industrial gas handbook - gas separation and purification

industrial gas handbook gas separation and purification ß 2006 by Taylor & Francis Group, LLC. ß 2006 by Taylor & Fran...

Author: Frank G. Kerry

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industrial gas handbook gas separation and purification

ß 2006 by Taylor & Francis Group, LLC.


industrial gas handbook gas separation and purification

frank g. Kerry


CRC Press Taylor & Francis Group 6000 Broken Sound Parkway NW, Suite 300 Boca Raton, FL 33487-2742 © 2007 by Taylor & Francis Group, LLC CRC Press is an imprint of Taylor & Francis Group, an Informa business No claim to original U.S. Government works Printed in the United States of America on acid-free paper 10 9 8 7 6 5 4 3 2 1 International Standard Book Number-10: 0-8493-9005-2 (Hardcover) International Standard Book Number-13: 978-0-8493-9005-0 (Hardcover) This book contains information obtained from authentic and highly regarded sources. Reprinted material is quoted with permission, and sources are indicated. A wide variety of references are listed. Reasonable efforts have been made to publish reliable data and information, but the author and the publisher cannot assume responsibility for the validity of all materials or for the consequences of their use. No part of this book may be reprinted, reproduced, transmitted, or utilized in any form by any electronic, mechanical, or other means, now known or hereafter invented, including photocopying, microfilming, and recording, or in any information storage or retrieval system, without written permission from the publishers. For permission to photocopy or use material electronically from this work, please access www. copyright.com (http://www.copyright.com/) or contact the Copyright Clearance Center, Inc. (CCC) 222 Rosewood Drive, Danvers, MA 01923, 978-750-8400. CCC is a not-for-profit organization that provides licenses and registration for a variety of users. For organizations that have been granted a photocopy license by the CCC, a separate system of payment has been arranged. Trademark Notice: Product or corporate names may be trademarks or registered trademarks, and are used only for identification and explanation without intent to infringe. Library of Congress Cataloging-in-Publication Data Kerry, Frank G. Industrial gas handbook : gas separation and purification / Frank G. Kerry. p. cm. Includes bibliographical references and index. ISBN 0-8493-9005-2 1. Gases--Separation--Handbooks, manuals, etc. 2. Gases--Purification--Handbooks, manuals, etc. I. Title. TP242.K47 2006 665.7--dc22 Visit the Taylor & Francis Web site at http://www.taylorandfrancis.com and the CRC Press Web site at http://www.crcpress.com


2006049163

Memoriam In Memory of my wife Mary Jane who patiently undertook the role of Penelope while I emulated the ever-roving Odysseus in search of new gases to recover and purify.



Preface As a practicing engineer in the field of industrial gases for the past 65 years, 40 with L’Air Liquide, and another 25 as a private consultant on an international basis, I have encountered only seven textbooks worldwide compiled for our industry, even though the industry is over 100 years old. Of the seven, five were edited by university professors or academic research scientists. The sixth was written in 1989 by a process manager of an industrial company in England which almost filled the bill, but not quite. Those were published in England, one in France, two in the United States, and one in Russia. With the possible exception of the one on process cycles published in England, the rest were heavy on the theoretical side and did not serve the majority of the engineers’ need for applied technology. The purpose of this manuscript, based on many years of field experience, is not so much for physicists or scientists who occasionally may require data relating to cryogenics and gas recovery, but for practicing engineers to help them master the fundamentals and advances in the field of industrial gas separation and purification. In planning the breadth of this manuscript the following colleagues were most helpful: Ted Pawulski. A former director of process design for separation units for Air Liquide in North America, now a private consultant, who reviewed the technical aspects of the text with diligence, patience, and made important suggestions. Nat Matlin. A graduate chemist and a polished copywriter by profession who went over the manuscript with a fine tooth comb. He was able to simplify the presentation and ease the flow of reading. To that end, he eliminated much of the technobabble so prevalent in current engineering reports. Joe Bernstein. A graduate chemical engineer with a stint at Air Products & Chemicals, now a private consultant, reviewed Chapter 3 on distillation and made useful recommendations. Homage must also be paid to the countless engineer-technologists whom I met in my professional life, who spend a lifetime in the field and away from their families, supervising the erection of machinery and equipment, the installation of instrumentation, the burdensome stroking of control valves, the startup of process units, and writing reports to head office, at the same time mumbling audibly that this is the last project they will participate in. And yet, the same faces show up in other projects, other countries, or other continents. They have amassed a precious lode of practical experience, which should be mined by enterprising young engineers at every opportunity. Many thanks should also go to my daughter, Nicolette Kerry, who functioned as a full secretarial staff correcting typos, maintaining my computer, running off copies, contacting outside sources for data and serving as both my most severe critic and most ardent supporter. This book is indebted to Elizabeth M. Mihaltse, whose professional graphic design expertise made a difference for the better, in many of the figures found in her grandfather’s book. Aside from individuals, the following organizations, listed alphabetically, who offered beneficial advice, as well as granting permission for use of data, diagrams, and photographs should be thanked: . . .

A.I.Ch.E Air Liquide Process & Construction American Institute of Physics


. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

ASTM International Atlas Copco BOC. Group BP Corp. Burckhardt Compression Carbo Tech Engineering GmbH Chart Energy & Chemicals, Inc. Chemical Engineering Chemical Engineering Progress (CEP) Compressed Gas Manufacturing Association Crane Co. Cryogenic Consulting Service, Inc. CryoGas International Consolidated Design Descon Engineering Co. Inc. ExxonMobil Upstream Affairs G.A.O. Hydrocarbon Processing (Gulf Publishing) Ingersol Rand Company Limited Kluwer Academic KOSO America Inc. (Hammel -Dahl) Leybold Vacuum LindeBOC Process Plants Los Angeles County Sanitation Districts MANTurbo AG Burckhardt Compressors McGraw Hill Education MV Engineering Chart Industries Metso Automation Oil & Gas Report Oxford University Press Perlite Institute Plenum Press Praxair, Inc. Southern Research Institute (University of Alabama) Bates & Hussman Springer Science & Business Media Taylor Wharton Thermax Thomson Learning Global Right Group Tyco Valves & Controls W.H. Freeman & Company

Nota Bene The reader is advised that because international industry uses SI units we have also used these with the possible exception of temperatures above 273 K where Celsius may be more convenient, especially for metallurgists; also US gallons for pumps and barrels for crude oil. To convert 8F into K, one may use K ¼ 5=9 (F þ 459.7) for all cases including if F is minus. Although kPa is used for small pressures up to and slightly above atmospheric, the term bar for higher pressures is preferable because the majority of international compressor


manufacturers employ it. In fact, if one deducts the pressure drop of an air filter of an air separation unit the process air enters the compressor at approximately 1 bar. 101.325 kPa ¼ 1 atm (most engineers use 101.3 kPa) 1 bar ¼ 100 kPa The term heat gain is used because heat enters the cryogenic equipment increasing the temperature and does not leak out. For vacuum applications the norm is bar, mbar, mbar, and Pascal.



Author Frank George Kerry is President of F.G. Kerry, Inc. with offices in Scarsdale, N.Y. and Sao Paulo, Brazil. He has been a consultant in the field of cryogenic systems and industrial gas technology for the past 30 years. He has directed major gas and air separation projects on six continents including the Dakota Gasification Company (USA), Companhia Siderurgica Paulista (Brazil), Siderurgica del Orinoco (Venezuela), and Barrick Goldstrike Mines (USA). Before establishing his own consulting organization, he spent 40 years with L’Air Liquide Group, first in Canada, then in New York where he established L’Air Liquide’s US operations. While being appointed vice-president, and serving as a member of the board he initiated international corporate development studies for the world headquarter in Paris, France. An honor graduate and medalist of McGill University (B. Eng Civil), Mr. Kerry continued on to do postgraduate studies in Paris, France (Ing. ESSA). During World War II, he was on loan to the Canadian Government to establish welding procedures for the fabrication of maritime tankers. He also designed and supervised the fabrication of piping networks to distribute oxygen and acetylene in shipyards, airplane factories, and steel mills. After the war, he concentrated on the application of oxygen in the open hearth furnace, and in 1946 he was granted a US patent (no. 2,446,511). He has presented numerous lectures in the USA and abroad, and authored many articles especially on gas purification and safety. He is a Member Emeritus of A.I.Ch.E, ASTM International, and ASM (Brazil).



Table of Contents Chapter 1 Gas Separation and Purification of Industrial Gases 1.1 Introduction 1.2 Thermodynamics 1.2.1 General Principles of Thermodynamics 1.2.2 Enthalpy (H) (J=kg) 1.2.3 The Second Law (The Availability of Energy) 1.2.4 Carnot Cycle 1.2.5 Entropy (S ) 1.2.5.1 Irreversible Systems 1.2.6 Third Law 1.2.7 Real Gases 1.2.8 Compression of Gases 1.2.8.1 Critical Temperatures and Pressures 1.2.9 Compressibility 1.2.10 Free Expansion through a Valve 1.2.11 Inversion 1.2.11.1 Deviation from Boyle’s Law 1.2.12 Adiabatic Expansion 1.2.13 Thermodynamic Charts and Tables 1.2.14 Cryogenic Properties of Air 1.2.15 Refrigeration and Liquefaction Systems (Ideal and Reversible) 1.2.16 Vapor Compression Systems 1.2.17 Liquefaction Systems 1.2.17.1 High-Pressure Free Expansion (Isenthalpic Linde–Hampson) System 1.2.17.2 Claude Isentropic System 1.2.17.3 Precooling Systems 1.2.17.4 Cascade Systems 1.2.18 Summary References Chapter 2 Industrial Applications 2.1 Early Development of Industrial Liquefaction Systems 2.2 Heat Exchangers 2.3 Expansion Machines 2.4 Contemporary Liquefaction Cycles 2.5 Linde Cycle (Free Expansion through a Valve) 2.5.1 Theoretical Analysis of the First Linde High Pressure Cycle 2.5.2 Theoretical Analysis of Linde Basic Cycle with Precooling 2.5.3 Theoretical Analysis of the Linde High-Pressure Dual Process


2.6

Theoretical Analysis of the Claude Cycle 2.6.1 Claude Cycle with Precooling 2.6.2 Claude Cycle with Dual Pressures 2.6.3 Claude Cycle with High Precooling to Liquefy Hydrogen or Neon 2.6.4 The Low-Temperature Refrigerator 2.7 Kapitza Cycle 2.8 Cascade Cycle References Further Reading Chapter 3 Air Separation Technology 3.1 Air Separation Overview 3.1.1 Linde’s First Fractionation Machine 3.1.2 Distillation and Fractionation 3.1.3 Fractionation 3.1.4 Stripping 3.1.5 Rectification 3.2 Theoretical Considerations of Fractionation 3.2.1 Evaporation and Condensation 3.2.2 Simple Separation by Condensation and Flashing (Separators) 3.2.2.1 Application 3.2.2.2 Procedure 3.2.2.3 Practical Example (It Requires Iteration) 3.2.3 Fractionation 3.2.4 Fractionation Methods 3.2.5 Fractionation Plates 3.2.6 Analysis of Flow in Equipment 3.2.6.1 Case I: Analysis of a Low Pressure Column 3.2.6.2 Case II: Consideration of Vapor Feed 3.2.6.3 Case III: Analysis of the High-Pressure Column 3.3 Practical Considerations 3.3.1 Bubble-Cap Trays 3.3.2 Sieve Trays 3.3.3 Structured Packings 3.3.4 Care in the Design of Structured Packing 3.3.5 Safety in the Use of Structured Packing 3.4 Operational Control 3.4.1 FCV-1 3.4.2 FCV-2 3.4.3 FCV-3 3.4.4 Refrigeration 3.5 Product Recovery 3.6 Optimum Reflux 3.7 Distillation Equipment 3.7.1 Upper (Low Pressure) Column 3.7.2 Lower (High Pressure) Column 3.7.3 Main Condenser 3.7.4 Liquid Subcoolers


3.7.5 Process Considerations 3.7.6 Crude-Argon Separation Column 3.8 Development of Low Oxygen-Purity Processes 3.8.1 The Lachmann Principle 3.8.2 The Oxyton Process 3.8.2.1 Thermodynamic Analysis of the Oxyton Cycle 3.8.2.2 Oxyton Development 3.8.3 Variable-Load Plants 3.8.3.1 Version A 3.8.3.2 Version B 3.9 Exergy References Chapter 4 Rare (Noble) Gases 4.1 Helium 4.1.1 Sources of Helium 4.1.2 General Principles of Recovery of Helium 4.1.3 Recovery Processes from Natural Gases 4.1.4 Applications of Helium 4.1.5 Conservation of Helium 4.2 Neon 4.2.1 General 4.2.2 Sources of Neon 4.2.3 Recovery of Neon 4.2.4 Industrial Recovery of Neon 4.2.5 Industrial Applications of Neon 4.3 Argon 4.3.1 General 4.3.2 Sources of Argon 4.3.3 Recovery of Argon 4.3.4 Recovery Procedure and Equipment 4.3.5 Secondary Rectification and Final Purification 4.3.6 Refining Operation and Equipment 4.3.7 Applications of Argon 4.4 Krypton and Xenon 4.4.1 General 4.4.2 Sources of Krypton and Xenon 4.4.3 Recovery of Krypton and Xenon 4.4.4 Refining of Krypton and Xenon 4.4.5 Recovery of Rare Gases from Ammonia Purge Gas 4.4.6 Applications of Krypton and Xenon References Further Reading Chapter 5 Front-End Purification Systems 5.1 Historical Background 5.1.1 Processes and Materials Used in Front-End Purification Systems


5.1.2 5.1.3

Original Prepurification Adsorbents 5.1.3.1 General 5.1.4 Introduction of Activated Alumina 5.1.4.1 Activated Alumina 5.1.4.2 Regeneration of Activated Alumina 5.1.5 Zeolites (Molecular Sieves) 5.1.5.1 Chemical Formula 5.1.5.2 Types of Molecular Sieves 5.1.6 Silica Gel 5.2 Design of Current Front-End Purification Systems 5.2.1 General Background 5.2.2 Equipment Used 5.2.2.1 Precooling Units Upstream of Adsorption 5.2.2.2 Direct Contact Aftercooler 5.2.2.3 Evaporative Water Chiller 5.2.2.4 Mechanical Refrigeration Unit 5.2.3 Adsorber Unit 5.2.3.1 Standard Design 5.2.3.2 Multiple Vertical Vessels 5.2.3.3 Horizontal Vessels 5.2.3.4 Radial or Concentric Design 5.3 Process Operation 5.3.1 Isolation Valve Downstream of FEP 5.3.2 Adsorption Kinetics 5.3.3 Regeneration Concerns 5.3.4 Warning against Excessive Heat during Regeneration 5.3.5 High-Pressure Vessel Regeneration 5.3.6 Regeneration Options for FEP Units 5.3.6.1 General 5.3.7 Summary 5.3.8 Operational Time Cycle 5.3.9 Prepurification Adsorbent Units and Operating Stability 5.3.9.1 Improving Operating Stability 5.4 Safety 5.4.1 Hydrocarbon Breakthrough 5.4.2 Safety Add-ons 5.4.3 Liquid Oxygen Purge 5.4.4 Analyzers 5.5 Activated Aluminas for Front-End Purification Systems 5.5.1 Background 5.5.2 Pressure Swing Adsorption 5.5.3 Industrial Applications in Air Separation Plants 5.5.4 Observations on PSA Prepurification 5.5.5 Field Observations References Further Reading on the Subject of Adsorption and Carbon Dioxide Build-up


Chapter 6 Product Liquefaction, Storage, and Transportation 6.1 Background 6.1.1 New Applications 6.2 Product Liquefaction 6.2.1 Enthalpy Balance 6.2.2 Direct Extraction 6.2.3 Basic Design Parameters for an Efficient Liquefaction System 6.2.4 Types of Liquefiers 6.2.4.1 Independent Liquefier 6.2.4.2 Integrated Liquefier 6.2.4.3 Very High-Pressure Liquefiers 6.2.4.4 General Summary 6.2.4.5 Energy Economics 6.3 Cryogenic Storage Facilities 6.3.1 General Considerations 6.3.2 Geographic Considerations 6.3.2.1 Ambient Temperature 6.3.2.2 Wind Loading 6.3.2.3 Seismic Loadings 6.3.2.4 Soil Conditions and Land Cost 6.3.2.5 Snow Loads 6.3.2.6 External Corrosion 6.3.2.7 Availability and Dependability of Utilities 6.3.2.8 Local Neighborhood Characteristics 6.3.3 Design Parameters 6.3.3.1 Low-Pressure Shop-Built Tanks 6.3.3.2 Storage Vessels with Internal Pressure 6.3.3.3 Low-Pressure Field-Built Aluminum Tanks 6.3.3.4 Flat Bottom Tanks 6.3.3.5 Spherical Containers 6.3.3.6 Cylindrical Vessels (Horizontal or Vertical) 6.3.4 Design Selection 6.3.5 Typical Designs of Cryogenic Storage Vessels 6.3.5.1 Vertical Cylindrical Tanks 6.3.5.2 Horizontal Cylindrical Storage Tanks 6.3.5.3 Spherical Tanks 6.3.5.4 Flat Bottom Tanks 6.3.6 Cryogenic Liquid Delivery Systems 6.3.6.1 General 6.3.6.2 Small Portable Containers 6.3.6.3 Customer Bulk Stations 6.3.6.4 LOX Distribution in a Shop 6.3.6.5 Liquid Deliveries by Truck 6.4 Cryogenic Pumps 6.4.1 Background 6.4.2 Variety of Applications 6.4.3 Materials 6.4.4 Present Designs 6.4.5 Net Positive Suction Head


6.4.6 6.4.7 6.4.8

Inlet Filter Screen Installation and Connections Typical Pump Piping Calculations 6.4.8.1 Velocity Constraints of Cryogenic Fluids 6.4.8.2 Pressure Drops due to Piping Components 6.4.9 Start-up of Pumps 6.4.10 High-Pressure Radial Pumps 6.4.11 Ultrahigh Pressure Pumps 6.4.12 Automation 6.5 Cryogenic Liquid Vaporizers 6.5.1 General Overview 6.5.2 Ambient Air Vaporizers 6.5.2.1 Modules Spread Apart 6.5.2.2 Modules in Alternate Operation 6.5.2.3 Modules with Pressurized Air 6.5.3 Direct Steam Vaporizers 6.5.3.1 Vaporization with Steam-Heated Water 6.5.4 Emergency Vaporization of Products References Chapter 7 Insulation 7.1 General 7.1.1 Theoretical Considerations 7.1.2 Insulations: General 7.1.3 Vacuum Insulation (Radiation) 7.1.4 Conductivity in Mass Insulations 7.1.5 Natural Convection in Mass Insulation 7.1.6 Vacuum Plus Powder or with Fibrous Insulations 7.1.7 Insulation (Multilayer, Super, or Simply MLI) 7.2 Industrial Practices 7.2.1 Industrial Applications of Insulation 7.2.2 Cryogenic Casings (Cold Boxes) for Process Equipment 7.2.3 Mineral Wool (Rock Wool) 7.2.4 Expanded Perlite 7.2.5 Glass Wool (Fiberglass) 7.2.6 Glass Blocks (Foam Glass) 7.2.7 Vermiculite 7.2.8 Silica Aerogel 7.2.9 Magnesium Carbonate 7.3 Cold Box Design for Insulation 7.3.1 Special Requirements for Liquid Hydrogen Processing Plants 7.4 Externally Located Process and Transfer Piping 7.4.1 Short Lines 7.4.2 Expanded Foams 7.4.3 Fiberglass Insulation 7.4.4 Prefabricated Vacuum-Insulated Piping 7.4.5 Multilayer Insulation 7.4.6 Cryogenic Liquid Piping Design


7.5

Insulation for Liquid Storage Tanks and Vessels 7.5.1 Large Storage Tanks (1000 t and over) 7.5.2 Smaller Storage Tanks (500 to 1000 t or 500 to 1000 kL) 7.5.3 Storage Tanks (50 to 500 t or 50 to 500 kL) 7.5.4 Storage Vessels (up to 50 t or 50 kL) 7.6 Vacuum Pumping Systems 7.6.1 General Overview 7.6.2 Vacuum Pumps 7.6.2.1 Roots Vacuum Pump 7.6.2.2 Rotary Vacuum Pump 7.6.2.3 Turbomolecular Pumps 7.6.2.4 Cryopumps 7.6.2.5 Adsorption Pumps 7.6.2.6 Getters 7.6.2.7 Small Laboratory Pumps 7.6.3 Periodic Purging and Deriming 7.6.4 Ancillary Equipment 7.6.4.1 Valves 7.6.4.2 Vacuum Measurement References For Further Study and Review Chapter 8 Special Gases 8.1 Hydrogen 8.1.1 Sources of Hydrogen 8.1.2 Recovery of Hydrogen 8.1.3 Hydrogen Use in Petroleum Refineries 8.1.4 Refinery In-House Recovery of Hydrogen 8.1.5 Recovery from Coke Oven Gas 8.1.6 Hydrogen Generation Plants 8.1.6.1 Electrolysis of Water 8.1.6.2 Thermal Cracking of Ammonia 8.1.6.3 Treatment of Hydrocarbon Feedstock for Hydrogen Recovery 8.1.6.4 Small Steam Reforming Plants (150–1000 Nm3=h) 8.1.6.5 Large Hydrogen Generation Plants (over 1000 Nm3=h) 8.1.7 Synthesis Gas, Partial Oxidation 8.1.7.1 History 8.1.7.2 Partial Oxidation Process 8.1.7.3 Ammonia Synthesis 8.1.7.4 Hydrogen Recovery from Ammonia Synthesis Plants 8.1.7.5 Other Uses for Synthesis Gas 8.1.7.6 Fuel Cells 8.2 Carbon Monoxide 8.2.1 Sources 8.2.2 Carbon Monoxide Recovery 8.2.2.1 General 8.2.3 General Process of Recovery 8.2.4 Basic Cryogenic Recovery Processes


8.3

8.4

8.5

8.6

8.2.4.1 Methane Wash Cryogenic Recovery 8.2.4.2 Simplified Carbon Monoxide Recovery 8.2.5 Compression and Conversion Machinery for Carbon Monoxide 8.2.5.1 Requirements for a Liquid CO Pump 8.2.5.2 Conversion from LCO to Gas 8.2.5.3 Safety of LCO Transport 8.2.6 Safety Equipment in General High-Purity Nitrogen 8.3.1 General Characteristics 8.3.2 Recovery 8.3.3 Applications for Inertness 8.3.3.1 Nitrogen as a Preservative 8.3.3.2 Nitrogen as an Emissions Controller 8.3.3.3 Nitrogen Use in Sparging 8.3.3.4 Nitrogen for Pressure Transferring 8.3.3.5 Liquid Nitrogen for Vapor Recovery 8.3.3.6 Liquid Nitrogen Makes Worn Rubber Tires Profitable 8.3.4 Process and Equipment Options 8.3.4.1 Cryogenic Process Cycle 8.3.4.2 Permeable Membrane Separation Process 8.3.4.3 Pressure Swing Adsorption 8.3.5 Ultrahigh-Purity Nitrogen 8.3.5.1 Removal of Outside Impurities 8.3.5.2 Process Cycle for Ultrahigh Purity Nitrogen 8.3.5.3 Outside Factors in Contamination 8.3.6 Other Atmospheric Nitrogen Compounds 8.3.6.1 General 8.3.6.2 Dinitrogen Monoxide (N2O) or Laughing Gas 8.3.6.3 Applications of Nitrous Oxide 8.3.6.4 Dangerous Side of Nitrous Oxide Carbon Dioxide 8.4.1 General Characteristics 8.4.2 Sources of Carbon Dioxide 8.4.3 Recovery Processes for Carbon Dioxide 8.4.3.1 Food Grade Recovery from Petroleum Off-Gases. 8.4.3.2 Food Grade Recovery from a Fermentation Source 8.4.3.3 Nonfood Grade Carbon Dioxide 8.4.4 Dry Ice: Food Grade 8.4.4.1 Production of Dry Ice 8.4.5 Applications of Carbon Dioxide Ozone 8.5.1 General 8.5.2 Properties of Ozone 8.5.3 Atmospheric Ozone Layer 8.5.4 Generation of Ozone 8.5.5 Applications Methane 8.6.1 Properties of Methane 8.6.2 High-Purity Methane for Chemicals 8.6.3 Natural Gas Peak Load Shaving


8.6.4

8.6.5 8.6.6

8.6.7

8.6.8

Base Load Natural Gas Plants (LNG) 8.6.4.1 Cascade Cycle with Mixed Refrigerants in General 8.6.4.2 ARC Process Cycle 8.6.4.3 Further Development of Mixed Refrigerant Cycles 8.6.4.4 Heat Exchangers 8.6.4.5 Propane Refrigeration System Pritko Process Cycle 8.6.5.1 General Process Cycle Final Product Purification 8.6.6.1 Nitrogen Rejection 8.6.6.2 Helium Recovery Natural Gas Prepurification 8.6.7.1 Acid Gases (CO2, H2S) 8.6.7.2 Water Removal (2H2O) 8.6.7.3 Mercury Contamination 8.6.7.4 Mercaptans 8.6.7.5 Butane 8.6.7.6 Propane and Ethane Economics 8.6.8.1 LNG Economics Safety

8.6.9 References Further Reading on LNG

Chapter 9 Noncryogenic Separations 9.1 Permeable Membrane Separation 9.1.1 General Principles 9.1.2 Mechanical Design of Membranes 9.1.3 General Applications 9.1.3.1 Nitrogen Separation 9.1.3.2 Disadvantages of Membrane Separation 9.1.3.3 Hydrogen Recovery 9.2 Gas Separation by Adsorption 9.2.1 General Overview 9.2.1.1 Adsorption Processes Studies 9.2.1.2 Regeneration of Adsorbent 9.2.1.3 Hydrogen Recovery from Coke Oven Gas 9.3 Nitrogen Recovery 9.3.1 Carbon Adsorbent (Carbon Molecular Sieve) (CMS) 9.3.2 High-Purity Hydrogen Recovery 9.3.3 Oxygen Separation and Vacuum Pressure Swing Adsorption 9.3.3.1 Process Description 9.3.4 Engineering Design 9.3.4.1 Basic Principles 9.3.4.2 Disadvantages of Adsorption 9.3.4.3 Economics References Additional Reading on Noncryogenic Separations


Chapter 10 10.1 Cryogenic Equipment, Materials, and Machinery 10.1.1 Heat Exchangers 10.1.1.1 General 10.1.2 Parameters of Design 10.1.3 Basic Principles 10.1.4 Typical Example for Designing Tubular Heat Exchangers 10.1.5 Brazed Aluminum Heat Exchangers 10.1.6 Effectiveness («) 10.1.7 Operability 10.1.8 Efficiency (h) 10.1.9 Industrial Applications 10.1.10 Development of Brazed Aluminum Heat Exchangers 10.1.10.1 Pressure Limitations 10.1.11 Vacuum Brazed Heat Exchangers 10.1.12 Mechanical Construction 10.1.13 Limitations 10.1.14 Operation and Maintenance References Further Reading 10.2 Expansion Machines 10.2.1 Expansion Machines 10.2.1.1 General 10.2.2 Reciprocating Expansion Engine 10.2.3 Radial Expansion Machines 10.2.4 Process Applications 10.2.5 Operational Factor (Air Separation Plants) 10.2.6 Refrigeration Availability 10.2.7 Process Technology 10.2.8 Expansion Turbine Efficiency 10.2.9 Expansion Turbine Losses 10.2.10 Measuring Efficiency 10.2.11 Various Expansion Turbine Systems 10.2.12 Mechanical Design Parameters 10.2.12.1 General 10.2.12.2 Operational Control 10.2.12.3 Shaft Speed (rpm) 10.2.12.4 Impeller Design 10.2.12.5 Materials of Construction 10.2.12.6 Bearings 10.2.12.6.1 Lubrication System 10.2.13 Instrumentation and Control 10.2.13.1 Process Control 10.2.13.2 Instruments Required 10.2.13.3 Computer Control (DCS) 10.2.14 Spares 10.2.15 General Applications for Expansion Machines References Supplementary Reading


10.3

Compressors 10.3.1 Compressors 10.3.1.1 General 10.3.2 Definitions 10.3.3 Centrifugal Compressors 10.3.3.1 General Parameters of Design (Per Stage) 10.3.3.2 Speed Variations 10.3.3.3 Surge Limitations and Pumping 10.3.3.4 Effect of Moisture 10.3.3.5 Effect of Altitude 10.3.3.6 Compressor Ratio Changes 10.3.3.7 Multistaging 10.3.3.8 Cooling Effect 10.3.3.9 Specific Speed 10.3.3.10 Stonewalling 10.3.3.11 Bearings 10.3.3.12 Seals 10.3.3.13 Lubrication System 10.3.3.14 Inlet Guide Vanes 10.3.3.15 Diffusers 10.3.4 Axial–Centrifugal Compressors 10.3.4.1 Bearings and Seals 10.3.4.2 Lubrication System 10.3.4.3 Inlet Guide Vanes 10.3.5 Axial Compressors 10.3.6 Integrally Geared Centrifugal Compressors (API-Standard-672) 10.3.6.1 General Overview 10.3.6.2 Functional Components and their Design 10.3.6.2.1 Gas Side 10.3.6.2.2 Mechanical Power Side 10.3.6.2.3 Bearings 10.3.6.2.4 Seals 10.3.6.3 Economics 10.3.7 Product Oxygen Compressors (RIO) 10.3.7.1 General Overview 10.3.7.2 General Safety Parameters 10.3.7.3 Safe Operation of Centrifugal Oxygen Compressors 10.3.7.4 Ultrahigh-Pressure Oxygen Compressors 10.3.7.4.1 Summary 10.3.8 High-Pressure Labyrinth Piston Compressor 10.3.8.1 Labyrinth Piston Compressor 10.3.8.1.1 General 10.3.8.2 Basic Design for Achieving Oil-Free Operation 10.3.8.3 Sealing Systems 10.3.8.4 Labyrinths 10.3.8.5 Internal Operating Elements 10.3.9 Compressor Drivers 10.3.9.1 Motor Torque 10.3.9.2 Enclosures 10.3.9.3 Power Factor


10.3.10

Operating Reliability versus Capital Costs of Compressors 10.3.10.1 Recommendations 10.3.11 Applicable Compressor Correction Factors Variables Summary of Factors Application of Factors References 10.4 Valves and Valve Stations for the Cryogenic Industry 10.4.1 General Design and Materials 10.4.2 General Design in Sizing 10.4.3 Sizing Parameters 10.4.4 Valve Categories 10.4.5 Nonmetallic Material 10.4.6 Manufacturers of Fluorinated Polymers 10.4.7 Warm End Switching Valves 10.4.7.1 Warm End Reversing Valves 10.4.7.2 Warm End Switching Valves for PSA Pre-Purification Systems 10.4.8 Flow Control Check Valves 10.4.9 Cryogenic Process Valves (General) 10.4.10 Hand-Operated Cryogenic Valves 10.4.11 Process Control Valves 10.4.12 Product Flow Control Valves 10.4.12.1 Gaseous Products 10.4.12.2 Liquid Products 10.4.13 Valve Connections 10.4.14 Insulation and Casing Designs for Cryogenic Valves 10.4.15 Liquid Purge Valves 10.4.16 Automatic Control of Cryogenic Valves 10.4.17 Cryogenic Liquid Storage Valves 10.4.18 Pressure Safety Relief Valves: Overview 10.4.18.1 Sizing for Pressure Safety Relief Valves (International Units) SI 10.4.18.2 Pilot-Operated Safety Valves 10.4.18.3 Pressure and Vacuum Relief Valves 10.4.18.4 Bursting Disks 10.4.18.5 Check Valves 10.4.19 Maintenance of Cryogenic Valves 10.4.20 Valve Stations: General 10.4.21 Valve Station Design 10.4.22 Destruction of a Pressure Reduction Station 10.4.22.1 Hypothetical Conclusions 10.4.23 Recommendations Applicable to Pressure-Reducing Stations Further Reading Chapter 11 Instrumentation and Controls 11.1 Overview 11.2 General Requirements


11.3 11.4

Controls and Control System Philosophy Minimum Instrumentation 11.4.1 General 11.4.2 Air Filter 11.4.3 Air Compressor 11.4.4 Direct Contact Aftercooler (If Applicable) 11.4.5 Front-End Purification 11.4.6 Air Separation Unit 11.4.7 Oxygen Product Compressor 11.4.8 Nitrogen Product Compressor 11.4.9 Liquid Oxygen Storage Tank 11.4.10 Liquid Nitrogen or Liquid Argon Storage Tank 11.4.11 Cooling Water System 11.4.12 Lube Oil System 11.4.13 Alarms, Shutdowns, and Interlocks 11.4.14 Analyzers 11.5 Possible Specific Requirements of Owner or Operator 11.5.1 Scope 11.5.2 Codes and Standards 11.5.3 Operational Philosophy 11.5.4 Distributed Control System 11.5.5 Field Instruments 11.5.5.1 Level Instruments 11.5.5.2 Temperature Instruments 11.5.5.3 Flow Instruments 11.5.5.4 Valves 11.5.5.5 Transmitters 11.5.5.6 Vibration Instruments 11.5.5.7 Local Controllers 11.5.5.8 Pressure Instruments 11.5.5.9 General 11.5.6 Interconnections 11.5.6.1 Pre-Packaging 11.5.6.2 Large Transformers Chapter 12 Safety 12.1 Safety Overview 12.2 Chemistry of Ignition, Combustion, and Explosion 12.2.1 Source of Combustibles 12.2.2 Ignition Energy 12.3 Critical Areas in an Air Separation Plant 12.3.1 General Description 12.4 Purification Systems 12.4.1 Adsorption Systems 12.4.2 Reversing Heat Exchangers: Revex 12.4.3 Nonreversing Heat Exchangers (Primary Heat Exchangers) 12.4.4 Distillation Column and Main Condenser 12.4.5 Auxiliary Vaporizers


12.4.6

12.5 12.6

12.7

12.8

12.9 12.10 12.11 12.12 12.13 12.14

12.15

Ancillary Equipment for Safety 12.4.6.1 Rich Liquid Filters 12.4.6.2 LOX Guard Filter 12.4.7 Liquid Oxygen Storage Tanks 12.4.8 Summary Parameters for the Safe Design of a Process Cycle General Design Procedures 12.6.1 Front End Prepurification 12.6.2 Reversing Heat Exchangers 12.6.3 Nonreversing Heat Exchangers 12.6.4 High-Pressure Column 12.6.5 Main Condenser Limits of Contaminants and Analysis 12.7.1 Argon as a Contaminant 12.7.2 Propane as a Contaminant Rotating Machines and Other Equipment 12.8.1 Expansion Machines 12.8.2 Liquid Oxygen Recirculating Pumps 12.8.3 Liquid Purge Lines 12.8.4 Liquid Oxygen Disposal Safe Practices 12.9.1 Analytical Equipment Summary Safety in the Design of Dynamic Oxygen Systems Causes of Combustion Test Procedures and Results as Explained by de Jessey The Following Recommendations Are in Order 12.14.1 Flow Velocities 12.14.2 A Very Careful Selection of Materials Consideration of Dynamic Oxygen Conditions 12.15.1 Tests 12.15.2 Example 12.15.3 Nonferrous Metals 12.15.4 Further Studies 12.15.5 Nickel and Its Alloys 12.15.6 Inconel Alloys 12.15.7 Stainless Steels 12.15.8 Copper and Its Alloys 12.15.9 Aluminum Bronze 12.15.10 Aluminum and Its Alloys 12.15.11 Compatibility of Aluminum and Its Alloys for Structured Packings (Structured Packing Consists of Corrugated Strips Coiled and Used as Distillation Trays) 12.15.12 Supplementary Tests on Aluminum 12.15.13 Replication of Aluminum Testing 12.15.14 Machines Used in the Fabrication of Structured (Corrugated) Packing 12.15.15 Summary 12.15.16 Iron Alloys


12.15.17 Nonmetallic Materials 12.15.18 Lubricants 12.15.19 Caution References For Further Reading Chapter 13 Cleaning for Oxygen Systems 13.1 Overview 13.2 General Considerations 13.3 Cleaning Requirements for Oxygen Systems 13.3.1 Inspection Standards for Fixed Surfaces 13.3.2 Inspection Standards for Movable Parts 13.3.3 Cleaning Procedures 13.4 Equipment other than Piping 13.4.1 Cleaning Procedures: General 13.5 Cleaning Procedures for Carbon Steel Piping 13.5.1 Definition and Recognition of Contaminants 13.6 Cleaning Procedures Available 13.6.1 Blast Cleaning: General 13.6.2 Sand Blasting in Place (Sandjet) 13.6.2.1 Equipment 13.6.2.2 Procedure 13.6.2.3 Inspection and Control 13.6.3 Secondary Cleaning Procedures 13.6.4 Pre-Cleaning before Erection (with Cleaning Reagents) 13.6.4.1 Precautions 13.6.5 Cleaning after Erection: General 13.6.6 Alternative A—with Solvents 13.6.7 Alternative B—Cleaning Agents 13.6.8 Alternative C—with Movable Pistons 13.6.9 Cleaning Stainless Steel and Nonferrous Metals Such as Copper, Associated Fittings, Parts, and Fabrications 13.6.10 Cleaning Aluminum Piping, Fittings, Parts, and Fabrications 13.6.11 Alternate Methods of Cleaning Stainless Steel Pipe, Aluminum Pipe, Copper Tubing, and Their Fittings 13.6.12 Oxygen Compressors 13.7 Cleaning Agents 13.8 Preparation of Cleaning Agents 13.8.1 Caustic or Alkaline Solutions 13.8.2 Acid Solutions 13.8.3 Agents for Stainless Steel, Copper, and Aluminum 13.8.4 Solvents 13.8.5 Aqueous or Semiaqueous Agents 13.9 Drying Gases 13.10 Testing and Inspection Procedures 13.10.1 Indirect Inspection 13.10.2 After Using Solvents or Chemicals


13.11

Ancillary Tools and Equipment 13.11.1 Wire Brushes 13.11.2 Immersion Tanks 13.11.3 Protective Clothing 13.12 Labor Force 13.12.1 Personnel 13.12.2 Supervision 13.12.3 Inspection 13.12.4 Cleaning Contractors 13.13 Protection and Storage 13.13.1 Piping 13.13.2 Small Equipment 13.13.3 Large Equipment References Chapter 14 Economics 14.1 General Overview 14.2 Historical Background 14.3 Post–World War II Development 14.4 Economic Overview 14.5 Energy Costs 14.5.1 Oxygen Purity 14.5.2 Liquid Oxygen Production 14.5.3 Pure Nitrogen Recovery (Purity at 99.9995%) 14.5.4 Argon and Rare Gas Recovery 14.5.5 Prepurification of Air 14.6 Investment Costs in General 14.6.1 Approximate Allocation of Investment Costs 14.6.2 Contingencies 14.7 Operating Costs 14.8 Maintenance 14.9 Marketing History of Industrial Gases 14.10 Challenging Market Conditions 14.11 Investing in a Project 14.11.1 Raising Investment Capital 14.11.2 Present Value of an Investment 14.11.3 Caveat on the Use of DCFROI 14.12 Envoi Appendix



1

Gas Separation and Purification of Industrial Gases

1.1 INTRODUCTION Although cryogenic science and engineering fully developed in the twentieth century, it will be fruitful to review and refresh our memory with the various people who dedicated their energies in uncovering the scientific principles and technology, which today form the basic foundations of low-temperature liquefaction and separation of industrial gases, now more elegantly known as cryogenics. One has to go as far back as the eighteenth century to uncover that in 1703 the work of a French physicist Guillaume Amontons in thermometry and mathematics led to the supposition of an absolute zero. Also in 1720, Gabriel Daniel Fahrenheit, born in Danzig, developed the idea of a temperature scale that bears his name, and wherein all the important numbers are divisible by four. In 1741, moreover, Anders Celsius formulated the centigrade scale, better known as Celsius, where zero represents the freezing of water, and 1008 represents the temperature at which water boils. At this point the discovery of dephlogisticated air or oxygen by Joseph Priestly in 1774 should be underlined. He reported his findings to Lavoisier (in France) who wasted no time in repeating Priestley’s experiments and went even further to recognize that it was the active ingredient of the Earth’s atmosphere. Priestley also discovered ammonia, sulfur dioxide, nitrogen, and a gas later identified as carbon monoxide. It was during the nineteenth century that serious attempts were made to liquefy industrial gases and to study their behavior as well as their characteristics. In 1823, Michael Faraday, then a young scientist at the Royal Institution of London, liquefied chlorine, and studied the liquefaction and characteristics of ammonia. He also discovered that ammonia changed from liquid to its gaseous state by absorbing heat from the surroundings, thus lowering their temperature. In 1824, a young French engineer, Sadi Carnot, published his thesis Reflexions, which became one of the pillars of thermodynamics. He demonstrated that a heat engine operating in his proposed ideal, reversible cycle between two heat reservoirs, one hot and the other cold, would be the most efficient engine possible. This ideal cycle is universally known as the Carnot cycle. In 1845, a little known Scot, Thomas Andrews of Queen’s College in Belfast, Ireland, experimented with the change of phase of carbon dioxide at various temperatures and pressures, and generated the first isotherms. These isotherms indicated that a gas cannot be liquefied, regardless of pressure, unless the gas is first cooled to or below a critical temperature. The origin of the absolute temperature scale was due to a British physicist, James Prescott Joule, who experimented with gases in the 1840s to confirm the efficiency of the Carnot cycle. In 1848, he wrote to William Thomson, a colleague, that his deduction was that the efficiency of the Carnot cycle is proportional to 1=T where T is from the equation pv ¼ RT, which represents the combined equations of Boyle and Charles, and T is based on t, the


actual temperature in degree Celsius plus 273. This sum t þ 273 was then known as the absolute temperature. Then in 1852 Thomson, later known as Lord Kelvin, formalized that the efficiency of the Carnot cycle for an ideal gas is W =Q ¼ (t1 t0 )=(t1 þ 273) ¼ (T1 T0 )=T1

(1:1)

After a long series of experiments with hydrogen, Joule and Thomson announced in 1862 that the difference between the above mentioned absolute scale and the air thermometer was insignificant. After experiments with helium, which was found closer to an ideal gas than hydrogen, 273 became 273.15 with slightly more additional corrections as the temperature dropped below 208C according to Keesom and Tuyen. During 1850, William Thompson (Lord Kelvin) followed by Rudolf Clausius in 1851, articulated the first and second laws of thermodynamics. Around 1865, a young German engineer, Carl Linde, designed a refrigeration machine based on the Carnot cycle and achieved low temperatures by vapor compression. Using the same principle and a spiralwound counterflow heat exchanger (designed by William Hampson of England), Linde built the first industrial air liquefaction machine in 1895. At this point it should be noted that in 1851, an American physician named John Gorrie patented the first ice-making machine using the compression and expansion of air to generate refrigeration1. The patent model is still in the National Museum of American History. In 1873, Johannes Diderick van der Waals developed an equation of state that offered an explanation of the critical phenomena in the change of state, and fitted nicely with the observations of isotherms of Thomas Andrews. Van der Waals used the values of pressure, temperature, and volume divided by their critical values. This led to the law of corresponding states, which enabled James Dewar and Kamerlingh Onnes to determine the necessary data for the liquefaction of gases. In 1877, a Frenchman, Louis-Paul Cailletet2 announced the liquefaction of oxygen. He used a glass capillary tube with thick walls wherein gaseous oxygen at 200–300 bar was cooled to 245 K by evaporating liquid sulfuric acid. After releasing the pressure almost adiabatically, the drop in temperature was sufficient to cause the formation of a mist of liquid oxygen. Almost simultaneously, Raoul-Pierre Pictet3 in Geneva, Switzerland, also liquefied oxygen by compressing oxygen to 200 bar, then cooling it to 173 K by the sublimation of dry ice, which in turn had been cooled by the evaporation of sulfuric acid. In both cases, however, a true continuous liquefaction of oxygen was impossible because the critical temperature of the gas is 155 K. Nevertheless, Pictet pointed the way to a new route to achieve low temperature, namely the cascade principle using a series of refrigerants, each lowering the temperature of the next one. This principle was fully developed by W. Keesom4 at Leyden in his work with helium using ammonia evaporating at 239.8 K, ethylene at 169.5 K, methane at 111.7 K, and nitrogen at 77.6 K. At this point, it is of interest to point out that during the latter part of the nineteenth century three experimental laboratories developed in Europe for the purpose of studying liquefaction, separation, and purification of the industrial gases at low temperatures. The men who worked in these laboratories should be given high credit, because they constantly faced risks from fire and explosion, given the limited availability of safety standards and process equipment at that time. It was only in the latter part of the nineteenth century, for example, that the Royal Institution in England decided to use reverse valve threads to differentiate between inflammable and nonflammable gases. The Royal Institution of London, England, founded in 1799, intensified its low-temperature studies especially under James Dewar, who developed the Dewar’s (vacuum) flask for conserving cryogenic liquids. He liquefied hydrogen as well in 1898, a high achievement at that time.


In 1882, the University of Leyden, Holland, hired Heike Kamerlingh Onnes4 as professor of experimental physics. During his tenure, Onnes, a close friend of van der Waals, founded the cryogenic laboratory in order to study equations of state, isotherms, and general thermodynamic properties of cryogenic liquids and gases. He became a master experimenter of physics. His laboratory included the development of measuring instruments and the fabrication of material for his experimental work. Postgraduate students worldwide were eager to intern at his laboratory. It was in 1908 that Onnes succeeded in liquefying helium at below 5 K. He also observed superconductivity at 4.19 K. He was unable to solidify helium, however. This latter phase was accomplished by his successor Wilhelmus Keesom in 1926. Keesom then refined the Onnes–Dana data on specific heat, and discovered that at 2.2 K another type of liquid was found, which was named helium II, and it exhibited a smaller density, a greater latent heat of vaporization, and a smaller surface tension—in other words, superfluidity5. Another smaller, but important cryogenic laboratory was founded at the Jagelonian University in Krakow, Poland, around 1876 under the team of Syzgmunt von Wroblewski, a very strong theoretician, and Karol Olszewski3, a capable technologist with a doctorate degree. In 1883, Wroblewski and Olszewski liquefied air, carbon monoxide, nitrogen, and finally oxygen. In fact, with the exception of hydrogen, they liquefied and studied the properties of all the known gases. They also improved the cascade principle that Pictet had pioneered. In January 1884, von Wroblewski produced a continuously changing liquid of hydrogen by cooling the gas with liquid oxygen and allowing it to expand rapidly, a most dangerous experiment by modern standards. In 1885, he also reported a change of conductivity in copper at low temperatures. Unfortunately, in 1914 an invading army closed down the promising laboratory.

1.2 THERMODYNAMICS It has been said many times that ‘‘engineering is the industrial application of science.’’ Unfortunately, many engineers get so involved with technology that they often forget the basic principles of science. In the case of liquefaction of gases and cryogenic systems, however, the two disciplines are integrally associated, and it is very difficult to do any worthwhile work involving cryogenic systems without a comprehensive understanding of the physical branch of science, thermodynamics.

1.2.1 GENERAL PRINCIPLES

OF

THERMODYNAMICS

The First Law or principle states very clearly that the quantity of energy in the universe is constant. It is the law of conservation of energy. Energy cannot be created or destroyed during a process, although it may be changed from one form to another. Indeed, thermodynamics deals with the relationships of various forms of energy. In essence, the First Law states that any change in internal energy Du is independent of the path and may be expressed as uf ui ¼ Q W

(1:2)

where uf is the final energy, ui the initial energy, W the work done, and Q the heat expelled. Energy may be classified into four basic categories: 1. Internal energy (u) due to the motion and configuration of a substance’s molecules, atoms, and subparticles 2. External energy (pv) due to a substance’s pressure and specific volume


3. Kinetic energy (V2=g) due to the motion of the substance, or system, where V is the velocity and g is the gravitational constant (9.80665 m=s2) 4. Potential energy (Z) due to the substance’s position in space There are other categories such as magnetic and surface energies, but they could be ignored for the time being, as this text deals only with gases. The external manifestations of energy are heat (Q) and work (W). The term heat is applied to the transfer of internal energy from one substance or system to another because of a temperature difference. The term work is applied to the transfer of internal energy to mechanical motion. Neither heat nor work can be considered as internal energy. If a system is subjected to a change, and the final state is identical in all respects to the initial state, and if only work and or heat has been exchanged between the system and its surroundings, then it can be said that . . . .

If the system has received heat, then the system has transferred work If the system has transferred heat, then the system has received work There is a constant relationship between work and heat The empirical relationship between work and heat is

1 cal ¼ 4.1868 J 1 kcal ¼ 4.1868 kJ One calorie is the heat required to raise 1 g of water from 14.58C to 15.58C. One kilocalorie is the heat required to raise 1 kg of water from 14.58C to 15.58C. One joule is the work done by the force of 1 N traveling 1 m in 1 s. It must be emphasized that heat Q and work W are not properties of a system, but rather the effects of the interaction of a system and its surroundings. When the effects of the interaction are balanced, the process is reversible. If they are not balanced, which is more realistic in practice, the process is termed irreversible.

1.2.2

ENTHALPY (H) (J=KG)

Enthalpy, or the total energy of a system, is the sum of its internal and external energies, or (u þ pv) where u is the internal energy and pv is the pressure–volume product. Moreover, the expression (u þ pv) should be considered as an entity and not separately. In the absence of any work other than that of expansion, any addition of heat to the system manifests itself in an increase in enthalpy. Conversely, any extraction of heat results in a decrease in enthalpy. Enthalpy, moreover, is a point property. It has no absolute value; and the engineer deals only with the difference in enthalpy between the initial and the final states of the system. When there is a temperature increase from T1 to T2 at constant pressure, there is an increase in internal and external energies, which can be written as follows: H ¼ H2 H1 ¼ (u2 u1 ) þ (p2 V2 p1 V1 )=J ¼ mcp (T2 T1 )

(1:3)

where m is the molecular weight of the gas and cp is the specific heat at constant pressure. The above equation is applicable to an ideal gas. If the specific heat is a variable and a function of the temperature, the change in enthalpy is an integrated quantity: H¼

ð Final mcp dt Initial


(1:4)

A change in the function H or (DH) is a measure of all heat effects, and is the basis of all total heat charts and tables. The latter tables are more widely used for the development of cryogenic process systems and operating procedures. To summarize the previous paragraphs, work and heat are manifestations of energy in transition. In a heat exchanger, energy is transferred from a warmer system A into a cooler system B. The energy in transition is heat (Q). On the other hand, if a gas, such as steam expands to drive a piston or turbine, the gas decreases in enthalpy, whereas the machine gains in work. In the latter case, the energy in transition is called work (W).

1.2.3 THE SECOND LAW (THE AVAILABILITY

OF

ENERGY)

The Second Law indicates the limitations of the First Law. The transfer of heat is an irreversible process. It is impossible to transfer heat from one system to another, which has a higher temperature without supplying external work. In general terms, during the transfer of energy, a part of the energy passes through a form that is no longer available for work. But to the engineer a knowledge of energy (enthalpy) is not enough. He must also quantify the availability or irreversibility of that energy for a process. A knowledge of the energy availability or irreversibility in a process, therefore, is mandatory. A short background of the development of the Second Law is as follows.

1.2.4 CARNOT CYCLE The Carnot cycle postulated in 1824 by Sadi Carnot assumes that an engine or a system operating as an ideal reversible cycle between two heat reservoirs (heat sinks) is the most efficient cycle possible. Furthermore, the cycle establishes the upper limit in the efficiencies of all possible cycles or engines. Because there is no change in enthalpy (heat exchange), the net work done by one cycle is equal to the net heat transferred into the system: W ¼ Q1 Q2

(1:5)

Efficiency is W=Q1 or (Q1 Q2)=Q1 or Efficiency ¼ (T1–T2)=T1, where T is in degrees (K). For a machine operating between two different temperatures, one higher at the inlet and the other lower at the exhaust, the maximum quantity of work that can be expected from the machine depends only on the difference between the inlet and exhaust temperatures and is independent of the nature of the fluid. The transferred heat is DQ ¼ (T1 T2 ) where Q is the heat absorbed by the system at T1 and the efficiency of the system is 1 – (T2=T1). From a practical point of view, since T2 cannot equal 0 K, a 100% efficiency is highly improbable for any machine or system. Conversely, a Carnot cycle in reverse, or transferring heat to a system, which has a higher temperature, is the basic process in achieving low temperatures for the liquefaction of gases, and that process demands the supply of external work. At the time, Carnot was completely unaware that the quantity of heat exhausted at T2 was not equal to the quantity of heat supplied at T1. It was about 25 years later, in 1851, that Rudolph Clausius calculated in a purely empirical manner that the quantity of heat exhausted at T2 was less than the heat supplied at T1. He stated, moreover, that this loss of heat (energy)


was characterized as entropy. Clausius also stated that the concept of entropy was a function of the absolute temperature. The principle of lost work or unavailability of energy can be a most serious challenge to the engineer for the optimization of any given process. According to Dodge6 Clausius’ original statement was: ‘‘It is impossible for a self-acting machine, unaided by an external agency, to convey heat from one body (system) to another which is at a higher temperature.’’

1.2.5

ENTROPY (S)

Entropy is not a property of any system or substance, but rather a mathematical function that quantifies the Second Law. It is a criterion of whether or not a process will take place in accordance with an engineer’s design. Furthermore, entropy is not an energy term by itself. It must be multiplied by the absolute temperature to be equivalent to energy. The product TS is then equal to that portion of energy in a process that is unavailable or unretrievable for further work. In a simpler form expressed by Dodge7: ‘‘It is the least amount of work that will be necessary to restore the system to the state in which it existed before the irreversibility occurred.’’ Dodge also articulated an analogy of entropy as follows: ‘‘If the individual components of the atmospheric air, such as nitrogen, oxygen, argon, etc., existed in individual cylinders, it would be stated that the system was highly organized, therefore, the entropy of that system can be considered as zero, or as a basic point of reference. If, however, the same individual gases were exhausted and mixed together as found in the atmosphere, the system is now highly disorganized, therefore, the entropy of the system has increased considerably, in fact, to the point that the probability of again separating the gases into their original organized state is highly unlikely without external work being applied.’’ Entropy is a useful mathematical function that quantifies the irreversibility of any process. With an understanding of its use, an enterprising engineer can attempt to achieve the maximum reversibility of a system or process. By studying the entropy of every subsystem (i.e., compressor, cooling, distillation, product compressor) of a process and by limiting any probable increase of the entropy of each operation, the engineer can increase the overall efficiency of the process. In other words, the engineer may extract the maximum energy available from the system. Hence the term exergy is now used by many practicing engineers. For example, over the years, engineers working on the low-temperature separation of air, have succeeded in reducing the energy required to produce one metric tonne of contained oxygen in a mixture of 99.5% purity from 350 kW h=h, to about 270 kW h=h and even lower, which represents a saving of 23%. For any reversible process cycle, the algebraic sum of all heat effects divided by the respective absolute temperatures at which the heat transfers occur is equal to zero: Q=T ¼

ð T2

(dQ=T) 0

(1:6)

T1

DS ¼ S2 S1 ¼

ð T2 (dQ=T)

(1:7)

T1

1.2.5.1

Irreversible Systems

Entropy depends only on the state of the system, therefore, the change in entropy depends on the difference of the initial and final states of the system. In an irreversible process, the change in entropy is


DS ¼

ð Tfinal (dQ=T) > 0

(1:8)

Tinitial

The above expression is an important one, because it indicates that the entropy differential is always positive. Since all separation processes are irreversible, the project engineer, as well as the process engineer have to pay extra special attention to anticipate and prevent any appreciable increase in entropy wherever it may occur, for example (a) Heat Transfer If an amount of heat Q at Th is transferred from a hot reservoir to a cold reservoir at Tc, the latter’s entropy increase is Q=Tc and simultaneously the hot reservoir has a final entropy of Q=Th. The difference in entropy is DS ¼ (Q=Tc ) (Q=Th ) > 0

(1:9)

because Th is larger than Tc, hence Q=Th is the smaller of the two terms. (b) Free Expansion of a Gas If an ideal gas is expanded within an insulated vessel with no work done, and no heat transfer carried out, and with no change in internal energy, then according to the first law Du (internal energy) ¼ Q W

(1:10)

Therefore W¼

ð Vf p dV

and

DS ¼ nRln Vf =Vi

(1:11)

Vi

Since Vf is larger than Vi, the change in entropy is larger. (c) Mixing Process If two gases or liquids for that matter, are mixed wherein the two masses are m1 and m2, the specific heats are c1 and c2, and the temperatures are T1 and T2 with a final temperature at Tf, then the change in entropy will be DS ¼ (m1 c1 ln Tf =T1 ) þ (m2 m2 ln T2 )

(1:12)

One of the terms on the right will always be positive, the other will always be negative, but the positive term will always be greater in value. In summary, therefore, in all reversible systems the change in entropy will always show an increase, and as indicated previously the engineer can only hope to limit the increase to a minimum in each and every item forming part of the overall system.

1.2.6 THIRD LAW Since the liquefaction and separation of gases involves very low temperatures, sometimes in the vicinity of absolute zero, it will be useful to keep in mind the third law of thermodynamics, which was articulated a long time ago by Lord Kelvin: ‘‘At absolute zero, the Carnot cycle cannot expel heat.’’ It was also stated by R. Nernst: ‘‘At absolute zero, entropy is constant.’’ It has been reported that a temperature as low as 0.000001 K has been reached.


1.2.7

REAL GASES

As stated previously, operations with real gases occur at low temperatures and high enough pressures so that a change of phase (liquefaction) does take place. This variation from ideal gases is due to the presence of molecules, which have volume and are the source of internal energy of the gas. Van der Waals8 succeeded experimentally to demonstrate that the former classic equation of state pV ¼ RT was no longer valid and that a new equation of state had to be developed. He proposed that: (p þ a=V 2 )(v b) ¼ RT

(1:13)

p ¼ RT=(v b) a=V 2

(1:14)

or

where v is the specific volume of the molecules in the gas, and a and b are constants for each gas. Since van der Waals formulated his new equation of state (about 1873), much scientific work has been carried out to modify or make the above equation more precise. Such attempts have proven encouraging, and have also helped the engineer to work efficiently in order to reduce the limitations of the Second Law. It is important to keep in mind that the word volume of a real gas is a measure of the number of molecules of the specific gas that are contained in a given volume or flow. The term normal cubic meter is always measured at 1 atm (101.325 kPa) and at a temperature of 08C (273.15 K). In these conditions, 22.412 normal cubic meters of any gas are equivalent to its molecular weight in kilograms. Since the molecular weight is constant for each component, a mass balance of any system is based on the number of molecules. It can be stated therefore that the mass flow, or weight of any gas depends on the total number of molecules contained in the gas. For example, 1000 g molecules (moles) of air can be separated into approximately 209 moles of oxygen, 782 moles of nitrogen, and 9 moles of argon. With these numbers in mind, it is easy to execute a mass balance of the system. ‘‘What goes in, must come out.’’

1.2.8 1.2.8.1

COMPRESSION

OF

GASES

Critical Temperatures and Pressures

In the nineteenth century, Faraday found that it was impossible to liquefy any gas by compression alone. He opined that the pressure for the liquefaction for any gas must be related to a certain temperature for each gas. Subsequently, Thomas Andrews published the results of his experimental work on the relationship between temperature and pressure for carbon dioxide. His charts demonstrated in graph form the critical temperature and pressure necessary to achieve any change of phase. Much above the critical value, a real gas follows the law for an ideal gas pv ¼ RT; but below the critical temperature, the gas can be liquefied by pressure. Furthermore, the lower the temperature below the critical value, the lower the pressure required for liquefaction (Figure 1.1).

1.2.9

COMPRESSIBILITY

At ordinary temperatures, near 300 K or below, and for small compression ratios, all the known gases with the exception of hydrogen are more compressible than what is indicated


T1 and higher indicates an ideal gas T2 and lower indicates a real gas T3 is critical temperaturature at critical pressure Pc T4 and lower, gas undergoes liquefaction Pc

C T1 T2 T3 T4 T5

B

A

C is critical point of CO2 To the right of CB fluid is all gas In the area ABC the fluid is mixed phase On the line BC the gas is at its dew point. On the line AC the fluid is all liquid To the left of line AC the liquid is supercooled.

FIGURE 1.1 Isotherms for CO2 at various temperatures developed by Andrews. (From Gomonet, E., Les Basses Temperatures, Production et Emplois, Librairie J.B. Bailliere et Fis, Paris, France, 1952. With permission.)

by Boyle’s law (pV ¼ constant). In other words, the product pV decreases when p increases. Depending on the temperature and the nature of the gas, this decrease reaches a minimum plateau and then begins to increase. At 100 K and below, hydrogen also behaves in a similar manner. (a) Constant pressure change, generally found in cyclic, or continuous flow processes: Q¼

ð T2 cp dT

(1:15)

T1

or p dv ¼

ð V2

p dv ¼ p(v2 v1 )

(1:16)

V1

(b) Isothermal compression (at constant temperature), and for ideal gases, p ¼ RT=V, and the work done is shown in the following equation: ð v2 ð v2 p dv ¼ RT dv=v (1:17) W¼ v1


v1

or ¼ RT

ð v2 dp=p v1

or W ¼ RT ln p2 =p1 or W ¼ 2:3026 log p2 =p1 per mole The above equations hold true for compression or for expansion. They also hold true for one only stage, and for one compression ratio, whether the latter is from 1 to 5 bar, or 5 to 25 bar. It is also seen that when an ideal gas is compressed isothermally, the quantity of heat equivalent to the work done must be removed from the gas. The amount of heat to be removed is Q(cal) ¼ W (J)=4:1868

(1:18)

If on the other hand, an ideal gas is expanded isentropically (Joule–Thomson effect across a throttle valve), the heat that is removed is also expressed as Q(cal) ¼ W (J)=4:1868

(1:19)

Dodge has stated that ‘‘isothermal work is the minimum amount of work that must be used to compress a gas over a given pressure range, or conversely, the maximum amount of work to be secured expanding the gas over the same range9.’’ (c) For an adiabatic compression of an ideal gas, the quantity of work required is W ¼ p1 v1 =(k 1)[1 (p2 =p1 )(k1)=k ]

(1:20)

W ¼ RT=(k 1)[1 (p2 =p1 )(k1)=k ]

(1:21)

or

The above equations are for theoretical adiabatic work per mole for either compression or expansion, where p1 is the initial pressure, p2 is the final pressure, and k stands for g the constant for the cp=cv ratio whose value changes for various gases, for example (Table 1.1). (d) In multistage compression, each stage must be treated and calculated as a series of single stage compressors adding the total work for the single machine. As stated previously, the above formulae are entirely theoretical and do not include any mechanical losses or driver inefficiencies.

1.2.10 FREE EXPANSION

THROUGH A

VALVE

At this point it may be profitable to consider the free expansion of a gas through a throttle valve, since its use was the first process to liquefy gases, and is still employed at the present


TABLE 1.1 Ratio of Specific Heats cp=cv or g for Various Gases Air dry Monatomic gases such as He, Ar Diatomic gases such as O2, N2, H2, CO, Cl2 Polyatomic gases such as CH4, CO2, H2O, SO2 Ammonia (superheated)

1.40 1.67 1.40 1.30 1.31

Source: From Dodge, B.F. in Chemical Engineering Series, McGrawHill, New York, 1944. With permission.

time. It has always been called the Joule–Thompson Expansion named after the early pioneers of cryogenics. Assuming an ideal gas and if only internal and external energies are considered, ignoring potential and kinetic, the effect of a free expansion through a throttle valve may be analyzed as follows: . .

No work is done, so W ¼ 0 No heat is exchanged, so Q ¼ 0

Therefore u 1 þ p1 v 1 ¼ u2 þ p2 v 2 H1 ¼ H2

(1:22)

or DH ¼ 0 As there is no change in enthalpy, an ideal gas cannot be cooled by a Joule–Thompson free expansion. Since there is no change in enthalpy, a Joule–Thompson expansion is called isenthalpic.

1.2.11 INVERSION In a real gas, however, consideration has to be given to van der Waals’ equation of state (p þ a=V2)(v b) ¼ RT, which includes the involvement of the molecules in a gas. The molecules in any gas have finite sizes and their volume is represented in the formula by the letter b. The molecules also display an intermolecular force or attraction, whose value is measured empirically in the formula by the letter a. In the Joule–Thomson effect10 it was found that not all gases showed a cooling effect when expanded through a throttle valve. Certain gases such as hydrogen and helium showed a heating effect. They calculated an inversion point for all gases, which is now known as mJT. The point at which mJT is zero is known as the Joule–Thomson inversion point. If van der Waals’ equation of state is applied to a free expansion of a real gas in order to determine whether or not a gas will be cooled or heated by free expansion, the following equation is useful: mJT ¼ (2a=RT )(1 b=v)2 ={cp [1 (2a=vRT)(1 b=v)2 ]}


(1:23a)

The above equation may be simplified for large values of specific volume to mJT ¼ 1=cp (2a=RT b)

(1:23b)

If T is less than 2a=bR, then mJT is positive and the gas will cool. If T is greater than 2a=bR, then mJT is negative and the gas will heat up. In general, those gases whose maximum inversion temperature is greater than ambient, say 290 K, will cool whereas others such as hydrogen, neon, and helium will heat up. In summary, when a real gas undergoes a free expansion through a throttle valve, there is no external work, and no heat is exchanged. It should normally behave as an ideal gas. There is a subtle change, however, in the internal energy of the molecules of the real gas as they disperse downstream of the valve. There is a minute increase in the potential energy of the molecule, which is more than offset by a decrease in its kinetic energy. This phenomenon results in an overall decrease in the temperature, however small, of the expanded gas. 1.2.11.1

Deviation from Boyle’s Law11

Figure 1.2 indicates a diagram for 1 kg of air at various isotherms and pressures. The dotted line reunites the various points of the minimum pv and the dotted diagram somewhat resembles a parabola. Within this dotted area, the air, or for that matter any gas, is subject

11000

0

20

40

60

80

100

120

140

T = 360⬚

7000

280⬚

280⬚

260⬚

260⬚

220⬚

5000

18

0⬚

16

0⬚

4000

200⬚

5000

180⬚

4000

14 0⬚

⬚ 160 ⬚ 140 ⬚ 0 2 1 ⬚ 100 80⬚

12 0⬚

0⬚

10

Va

po

2000

6000

220⬚

⬚

Pv

8000

240⬚

200

3000

9000

7000

240⬚

6000

11000 10000

300⬚

300⬚

8000

200

320⬚

320⬚ 9000

180

⬚ T = 340

340⬚

10000

160

r

3000 2000 1000

1000 id iqu

L

0

0

20

40

60

80

100

120

140

160

180

200

0

FIGURE 1.2 pV=p diagram for 1 kg of air. (From Gomonet, E., Les Basses Temperatures, Production et Emplois, Librairie J.B. Bailliere et Fis, Paris, France, 1952. With permission.)


TABLE 1.2 Maximum Inversion Temperature for Different Gases Gas

Maximum Inversion Temperture (K)

Helium-4 Hydrogen Neon Nitrogen Air Carbon monoxide Argon Oxygen Methane Carbon dioxide Ammonia

45 205 250 621 603 652 794 761 939 1500 1994

Source: From Barron, R.F. in Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.

to inversion and does not follow Boyle’s law. Along the inversion curve, the Joule–Thomson coefficient is zero because a point on the inversion outline is one at which the slope of the isotherm line is zero. This indicates that outside of the inversion area a gas follows Boyle’s law. For hydrogen or helium, the inversion point is below room temperature. Table 1.2 indicates the maximum inversion temperatures of various gases. It can be concluded from Table 1.2 that most gases can be liquefied by free expansion through a throttle valve, with the exception of helium-4, hydrogen, and neon. The latter gases have to be precooled before being processed with free expansion. Each gas, therefore, requires an individual study of its inversion behavior and the value of its Joule– Thompson coefficient (mJT). For hydrogen, it is necessary to precool the gas to 198 K before any refrigeration can be achieved by the use of a throttle valve expansion. In the case of helium, the precooling will have to reach 45 K.

1.2.12 ADIABATIC EXPANSION Even before Claude successfully applied a reciprocating expansion engine to air separation, the benefits of an adiabatic expansion with external work for refrigeration purposes were already known. It was only a question of how this could be accomplished from a practical point of view. In an adiabatic expansion, the energy in the gas is transformed into external work, and that loss of energy in the gas results in a substantial drop in temperature. It has been stated previously that the final temperature in an adiabatic expansion of an ideal gas is equivalent (in reverse) to the increase in temperature for the adiabatic compression of the same gas (Figure 1.3). The isentropic (adiabatic) expansion of an ideal gas may be stated by the following equation: WAD =m ¼ kRT=(k 1)[(p1 =p2 )(k1)=k 1]

(1:24)

This equation applies to both reciprocating and centrifugal expansion machines. The above expression can also be stated more simply as T2 ¼ T1 (p2 =p1 )(k1)=k


(1:25)

Y

T

X

t.

= t. ne co

p

Z

hs

ne

o =c

W

hs = h

con et.

Ri

sin

p

R

is

in

g

g

S

FIGURE 1.3 Isothermal compression, isenthalpic expansion, and isentropic expansion on (T1S) diagram. (From Springman, H. 1985. Cryogenics, principles and applications. Chemical Engineering. (May 13). pp. 58–67.) Note: The curved line from y to z indicates an isothermal expansion by a throttle valve, using the Linde high-pressure cycle. And the line y to w shows an isentropic (adiabatic) expansion with the Claude cycle. (From Springmann, H., Chem. Eng., May 13, 58, 1985. With permission.)

In the above equations, k ¼ g, the ratio of specific heats. In the case of adiabatic expansion of an ideal gas, the principles involved are the same as those for the adiabatic compression, except that the discharge temperature T2 is much lower that the inlet temperature T1. Furthermore, the expression for the work done, W, is negative since the expansion machine is actually the driver. The cooling effect of the adiabatic expansion of an ideal gas is considerable. For example, the expansion of air at 290 K and at 5 bar down to atmospheric can be calculated from the formula: T2 ¼ T1 ( p2 =p1 )0:29

(1:26)

and since T1 ¼ 290 K, and p2=p1 ¼ 1=5, therefore, T2 ¼ 182 K, or 273–182 K. The overall temperature drop is from 290–182 or 108 K. From the above calculations, it is obvious that for an increase in the cp=cv ratio, the lower will be the final temperature in an adiabatic expansion. For example, for a monatomic gas with a compression ratio of 1.67, the final temperature will be 152 K, i.e., a temperature drop from 290–152 or 138 K.

1.2.13 THERMODYNAMIC CHARTS AND TABLES Real gases do not follow any predetermined algebraic equation of state. For this reason, a series of charts and tables have been prepared over the years to represent almost all the thermodynamic properties of real gases in order to facilitate the engineer’s calculations over a wide range of temperatures and pressures. These charts involve the relationships of temperature–entropy, temperature–enthalpy, and enthalpy–entropy. These charts are also called


Mollier diagrams, which are familiar to most engineers. More recently, however, with the aid of computers many companies in the cryogenic field have developed their own tables indicating even small changes in enthalpy over a wide range of cryogenic temperatures as well as pressures. With the aid of such tables the process engineer can develop a more accurate process design than by dealing with charts and nomograms, which involve a certain amount of guess work. In dealing with charts, and more specifically with tables, it is necessary to keep in mind that there are no absolute values available for the thermophysical properties of real gases. The values of enthalpy or entropy found in the literature have been calculated above a definite reference state. For example in steam tables, the reference state is taken as the saturated liquid at 273.15 K (08C). This can be interpreted as follows: (a) The enthalpy and entropy for this state as both zero or (b) The properties recorded in the steam tables are really changes in the property between the state in question and the reference state Interpretation (b) is correct. In a flow process, where work effects and potential and kinetic energy changes can be ignored, the following equation is valid: Q ¼ hfinal state hinitial state

(1:27)

In cryogenic design, moreover, the reference point, or state may be anything depending on who prepared the chart, table, or software. For example (a) Saturated liquid at normal boiling point (NBT) or (b) Vapor at 273.15 K and 101.325 kPa. (1 bar) For tables relating to air at or near liquefaction temperatures, one must remember that air is a mixture of oxygen, nitrogen, and argon, all of which have different liquefaction temperatures. Some tables involving temperature–enthalpy figures may indicate negative quantities. Nevertheless, the engineer must use them in order to arrive at a proper heat balance in a system where the composition keeps changing. In the calculations of enthalpy and entropy around various points in an air separation system that may involve a composition slightly different than that of air, it will be prudent to calculate the properties of each component separately and to use the total sum as the final property. For example, if the fluid used in the expansion machine comes from the (lower) high-pressure column, it may have a composition different from that of air. When working with charts and tables, the engineer should take care to ascertain the system used for the values stated in the chart. At the present time, there is a certain amount of confusion in the selection of values from either the new International System (SI), or the American system involving ft=lbs. The gas industry standard cubic foot is measured at 608F, whereas the cryogenic industry maintains that a standard cubic foot is measured at 708F. Even in the metric system, a normal cubic meter is always measured at 273.15 K, or 08C, whereas some countries or companies refer to this as a standard cubic meter, and use a temperature of 15.58C, for a normal cubic meter.

1.2.14 CRYOGENIC PROPERTIES OF AIR12 The critical temperature of air is 132.6 K (140.558C) at 37.66 bar (3766 kPa) absolute. Above this temperature air cannot be liquefied by pressure alone. The following general formula may be used for the liquefaction of air.


For the beginning of liquefaction log p(bar) ¼ 4:640 0:458(1000=Tk ) þ 0:00733(1000=Tk )2

(1:28)

For the completion of liquefaction log p(bar) ¼ 4:764 0:476(1000=Tk ) þ 0:00733(1000=Tk )2

(1:29)

Between the temperatures of 81.15 K and 78.75 K (192.008C and 194.048C), the liquid produced is a variable mixture of oxygen, nitrogen and argon because each of these elements have a different liquefaction point, and great care must be taken in reading charts and tables.

1.2.15 REFRIGERATION AND LIQUEFACTION SYSTEMS (IDEAL AND REVERSIBLE) It is important to differentiate between refrigeration and liquefaction systems. A refrigeration system or cycle is essentially a closed system operating at a steady rate and where only work and heat are exchanged. In essence, a refrigeration system absorbs heat at a low-temperature level and rejects it at a higher temperature level. The system involves a compressor, a heat exchanger, a throttle (expansion) valve, and a liquid reservoir–evaporator. The system is normally employed to supply a predetermined quantity of refrigeration and at a fixed temperature. The measure of performance of a refrigeration system is called coefficient of performance (COP). The latter term is defined as the energy removed from the source divided by the work required to remove the energy in question. If the energy (heat) to be removed is Q, and work W is required, then the coefficient of performance, or COP is given by the formula: COP ¼ T1 =(T2 T1 ) where T1 is the lower temperature, and in such instances, the COP is greater than unity.

1.2.16 VAPOR COMPRESSION SYSTEMS13 In vapor compression systems using refrigerants such as ammonia, Freons, which are easily condensed when compressed, and are expanded in the dual phase, as in pressure–enthalpy (Mollier chart), which is more convenient to use than a T–s diagram (Figure 1.4). The energy balance is as follows: Heat absorbed Q ¼ HC HB Heat released Q ¼ HD HA Work required W ¼ Q1 Q2 ¼ HC HB HD þ HA ¼ HC HD

(1:30)

COP ¼ (HC HB )=(HD HC )

(1:31)

Since HB ¼ HA

where H is the enthalpy. As a rule, the COP of vapor compression systems is greater than 1. As noted, the simplified expression, COP ¼ T1=(T2T1) cannot be used for refrigerants using cold gases such as nitrogen.


Absolute pressure

Liquid region

Two-phase region A

Superheated vapor region E

Condensation

D

Throttle expansion

B

Compression

Evaporation C

Enthalpy

FIGURE 1.4 Mollier or pressure–enthalpy diagram used in vapor compression systems employing condensation by recompression followed by throttle expansion, evaporation. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

1.2.17 LIQUEFACTION SYSTEMS 1.2.17.1

High-Pressure Free Expansion (Isenthalpic Linde–Hampson) System14

In a liquefaction system, Figure 1.3 and Figure 2.3 the refrigeration duty Q is replaced by a continuous withdrawal of the finite liquid m from the reservoir. The uncondensed portion of the gas is returned to the system in countercurrent–heat exchange, and the liquid is removed from the system as product or for further treatment. In a liquefaction system identical to a Linde–Hampson free expansion cycle, the overall energy balance around the simple system of heat exchanger, throttle valve, and reservoir is as follows: mH2 ¼ (m mf )H1 þ mf Hf

(a) (b) (c) (d)

¼ ¼ ¼ ¼

(1:32)

(b) þ (c) enthalpy entering a heat exchanger enthalpy leaving heat exchanger enthalpy leaving with the liquid product

And if the portion liquefied is mf=m, then the liquid yield y is y ¼ (H1 H2 )=(H1 Hf )

(1:33)

Since the terms H1 and Hf are constant, it is obvious that the yield increases as the temperature H2 at the inlet of the primary heat exchanger decreases. That is to say, that by precooling the system, the overall efficiency of the cycle, and the liquid production increases. Incidentally, the same principle is applicable to a refrigeration system. Linde was well aware of this application, and used an ammonia precooling refrigeration cycle in a later development of his high-pressure system to improve the production of liquid.


1.2.17.2

Claude Isentropic System15

The Claude system (Figure 1.3 to Figure 2.6) replaces the throttle (expansion) valve with an expansion machine, either reciprocating, or centrifugal, which creates work thus eliminating the increase in entropy, at least theoretically. The throttle valve is not completely replaced in order to avoid the formation of liquid within the expansion unit. From Figure 2.6, it can be seen that the gas is compressed isothermally at ambient temperature, then it passes through a heat exchanger to be cooled to a predetermined temperature countercurrent to the uncondensed gas returning from the evaporator–condenser. A portion of the precooled gas is sent to the expansion machine where it reaches its dew point. Then it is returned to the liquefier– exchanger to condense the remaining portion of the process air. The remaining portion of the air is cooled by the returning air from the expander. It is then sent to the throttle valve for free expansion and condensation in the evaporator–condenser, sometimes called the liquefier. The expansion machine is loaded with an oil brake for very small plants, with a motor=generator set for larger plants, or a booster compressor for very large plants. The overall energy balance around the heat exchanger, expansion machine and the liquefier, applying the First Law, is as follows: mH2 ¼ (m mf )H1 þ me (He H3 ) þ mf Hf

(a) ¼ Enthalpy (b) ¼ Enthalpy (c) ¼ Enthalpy (d) ¼ Enthalpy

(1:34)

entering the first exchanger leaving the first exchanger and re-entering the compressor differential across the expander of liquid withdrawn from system

The yield is y ¼ (H1 H2 )=(H1 Hf ) þ r[(H3 He )=(H1 Hf )]

(1:35)

where r is the rate of flow of the refrigerant used in the expander.

1.2.17.3

Precooling Systems16

Auxiliary refrigeration cycles can and have been used with liquefaction systems in order to increase the liquid yield, and to liquefy gases whose inversion temperature is below ambient. Using Figure 2.6a, the overall energy balance is: mH2 þ mr Hd ¼ (m mf )H1 þ mr Ha þ mf Hf

(1:36)

and the total liquid yield y is y ¼ [(H1 H2 )=(H1 Hf ) þ r[(Ha Hf )=(H1 Hf )]

(1:37)

where r is the rate flow ratio of the refrigerant used for precooling. Moreover, since y ¼ mf=m, it can be shown that the maximum yield of liquid is limited by the temperature of the refrigerant.


1.2.17.4

Cascade Systems17

The cascade system for the refrigeration or the liquefaction of gases is merely an extension of the precooling system where each refrigeration cycle is precooled by an independent refrigeration cycle. It is the most efficient process in terms of energy requirements, the closest to a reversible process. The reasoning may be explained as follows. The minimum work W required for the liquefaction of any gas is expressed by Wreversible ¼ H T0 S

(1:38)

It is also established that the sum of all entropies leaving a system is always greater than that of all entropies entering the system and the difference between the two sums is a direct measure of the irreversibility of the process. It has also been established that by precooling a liquefaction system, the entropy differential is less as the initial temperature is lowered. In other words, if a gas is expanded in its saturated or liquid phase, it limits its increase in irreversibility considerably (Figure 1.5 to Figure 2.9). The cascade system, therefore, is an extension of the precooled liquefaction cycle wherein only liquids are expanded. The process employs a series of gases with progressively falling boiling points, each of which is liquefied under pressure through the refrigeration effect caused by the evaporation of the one next higher in boiling point. This was first carried out by Kamerlingh Onnes in Leyden, and treated theoretically by Keesom who succeeded Kamerlingh Onnes. In the liquefaction of air, one may begin the refrigeration process with either an ammonia or Freon cycle, then follow with ethylene or ethane, and finally with methane. There is a lower limit to the temperature attainable by this procedure, however, that is set by the fact that the critical temperature of the lowest usable gas must be higher than the triple point of the next higher one. For example, the evaporation of liquid nitrogen, with a

Specific energy consumption (kw h/kJ)

10⫺2

Hydrogen cycle

Nitrogen cycle 10⫺3

CH4 cycle C2H4 cycles NH3 cycles

10⫺4

Minimum energy expenditure (theoretical) Two stages One stage

10⫺5 280

240

200 160 120 Temperature (k)

80

40

FIGURE 1.5 As the temperature decreases, the specific energy consumption in kW h=kJ increases in order to a unit amount of refrigeration (cold). (Courtesy of Linde BOC Process Plants. With permission.)


triple point of 63.3 K, cannot be used for the liquefaction of neon with a triple point of 24.5 K, or for that matter, hydrogen with a triple point of 14.1 K. The evaporation of liquid nitrogen can and is being used, however, for the liquefaction of neon and hydrogen, but as an independent precooling agent in a separate neon or hydrogen process cycle. Although the cascade process cycle is the optimum selection for the liquefaction of gases in terms of energy savings, it also presents the disadvantages of mechanical complexity and high capital investment as well as operational problems due to leakages from the mechanical connections of the various refrigerants.

1.2.18 SUMMARY 1. The efficiency of any liquefaction system as defined by the ratio of the reversible work of liquefaction to the actual net energy consumed depends on the irreversible entropy increments occurring at various points in the system. 2. The principal sources of such increments are found in the isenthalpic expansions of the gas, and in the wide differences of temperatures at the cold end of the heat exchangers. 3. Precooling a compressed gas before isenthalpic expansion leads to a smaller irreversible entropy change, especially close to the saturation point of the gas. 4. Every effort should be made to expand a fluid in its liquid phase for a minimum change in entropy.

REFERENCES 1. Morse, M.S. 2002. Chilly reception, Smithsonian Magazine, June, pp. 30–32. 2. Gomonet, E. 1952. Les basses tempe´ratures, production et emplois, Librairie J.B. Bailliere et Fils, Paris, France, p. 66. 3. Gomonet, E. 1952. Les basses tempe´ratures, production et emplois, Librairie J.B. Bailliere et Fils, Paris, France, p. 67. 4. Gomonet, E. 1952. Les basses tempe´ratures, production et emplois, Librairie J.B. Bailliere et Fils, Paris, France, p. 172. 5. Timmerhaus, K.D., T. Flynn. 1989. Cryogenic process engineering, International cryogenic monograph series, New York: Plenum Press, pp. 26–31. 6. Dodge, B.F. 1944. Chemical engineering thermodynamics, New York: McGraw-Hill, p. 49. 7. Dodge, B.F. 1944. Chemical engineering thermodynamics, New York: McGraw-Hill, Chapters 2 and 3. 8. Gomonet, E. 1952. Les tre`s basses tempe´ratures, production et emplois, Libraire J.B. Bailliere et Fils, Paris, France, p. 31. 9. Dodge, B.F. 1944. Chemical engineering thermodynamics, New York: McGraw-Hill, pp. 265–267. 10. Baron, R.F. 1985. Cryogenic systems, Second edition, Oxford Science Publications, Monograph on Cryogenics #3, p. 65. 11. Baron, R.F. 1985. Cryogenic systems, Second edition, Monograph on Cryogenics #3, Cambridge: Oxford, p. 69. 12. Gomonet, E. 1952. Les tre`s basses tempe´ratures, production et emplois, Libraire J.B. Bailliere et Fils, Paris, France, pp. 70–71. 13. Dodge, B.F. 1944. Chemical engineering thermodynamics, New York: McGraw-Hill, pp. 430–435. 14. Barron, R.F. 1985. Cryogenic systems, Second edition, Monograph on Cryogenics, #3, Cambridge: Oxford, pp. 69–73. 15. Baron, R.F. 1985. Cryogenic systems, Second edition, Monograph on Cryogenics #3, Cambridge: Oxford, pp. 85–88. 16. Baron, R.F. 1985. Cryogenic systems, Second edition, Monograph on Cryogenics #3, Cambridge: Oxford, pp. 73–77. 17. Baron, R.F. 1985. Cryogenic systems, Second edition, Monograph on Cryogenics #3, Cambridge: Oxford, pp. 83–84.


2

Industrial Applications

2.1 EARLY DEVELOPMENT OF INDUSTRIAL LIQUEFACTION SYSTEMS1 Early experimentation on liquefaction of gases aroused the imagination of industrialists like Carl Linde who was already in the business of manufacturing refrigeration machines with ammonia as the refrigerant. His company, the Linde Eismachinen AG, was founded in 1871. With his background, Linde foresaw the possibility of recovering large quantities of oxygen from the atmosphere, and using it in the metallurgical industry as an enriching agent to increase operating efficiency in the recovery and production of metals. However, Linde’s initial trials in 1895 were not very successful, however. He tried to liquefy air by compressing it to 800 bar, and then expanding it through a throttle valve. He had hoped to cool the compressed air at a rate of 0.35 of a degree per 1 bar pressure drop. He soon found out that this procedure did not work. (The processed air was not at its critical temperature, and was outside of its inversion curve.) In 1899, having demonstrated the solubility of acetylene in acetone, Georges Claude also became interested in the possible use of oxygen in a low-cost production of calcium carbide.

2.2 HEAT EXCHANGERS In all the early experiments for the liquefaction of air and other gases, a crude apparatus resembling a countercurrent heat exchanger was used to help precool the gas to be processed. The lack of efficiency of this apparatus was a key obstacle to the successful development of a viable cryogenic industry as it is known today. At this point, therefore, it is only fair to give credit to William Hampson2 and to Heylandt for the ingenious design and development of the wound-coil countercurrent heat exchanger (Figure 2.1). This design offered a very efficient method of heat transfer in a highly compact configuration and gave the infant cryogenic industry a real impetus in its passage from the laboratory to a large-scale application in industry. This exchanger involved a large number of small copper tubings spirally wound within a cylindrical shell, also of copper. The incoming warm process air passed through the tubing in countercurrent heat exchange and at right angles to the colder outgoing products that flowed through the shell. With this design it was possible to enclose a large heat exchange surface within a relatively small external shell volume, as well as increase the efficiency of the heat transfer surface. It was the development of this concept that enabled Linde to launch the first industrial air liquefaction machine in 1896. In fact, some people call the process the Linde–Hampson cycle for air liquefaction. Although this heat exchanger design is still being used for many applications, especially for the liquefaction of large quantities of natural gas, it has been supplanted to a large extent by the development in 1950 of the brazed aluminum extended surface plate–fin exchanger, better known as BAHX. The latter exchanger offers an extremely large surface of heat transfer, which can be packed in a relatively small space. It also offers the economic advantage of being less costly to manufacture. This subject is discussed in more detail in a later chapter.


FIGURE 2.1 Coiling a subcooler. (ß Air Liquide, all rights reserved, 2006. With permission.)

2.3

EXPANSION MACHINES3

The possible use of an expansion machine for obtaining low temperatures isentropically was first suggested by William Siemens in 1857. In 1885, E. Solvay of Belgium attempted to use a reciprocating expander to operate at very low temperatures, but could not reach any temperature lower than 175 K (988C). The main problem was the lubrication of the machine at such low temperatures. Even Linde thought of using this concept but rejected the idea as impractical at low temperatures. In 1902, it was Georges Claude of France, who succeeded in applying the reciprocating expansion machine to the air liquefaction process. He used a burnt (degreased) leather packing as a piston seal and operated it without lubrication. With this adaptation, Claude was able to achieve an almost isentropic (adiabatic) expansion of the processed air, resulting in a lower temperature than was hitherto possible, with an air pressure of only 40 bar. In 1939, Pyotr Kapitza of Russia improved on the Claude cycle by suggesting the use of a centrifugal turbine as an expansion machine because of its higher operating efficiency. Since then the use of the centrifugal expander has taken over almost 100% of the liquefaction requirements of the cryogenic industry. During the past 20 years, moreover, there has been a tendency to couple the centrifugal expander directly to a blower compressor, and to boost a portion of the prepurified process air (normally 20% to 30%), to develop more energy to offset heat gain and any extraction of liquid product. This additional liquid product may be used either for storage and eventual transport, or for boosting to a higher pressure by means of a cryogenic pump. This eliminates a gaseous product compressor that has a high investment and operating cost.

2.4

CONTEMPORARY LIQUEFACTION CYCLES

The general parameters for the design of liquefaction cycles thus involved compression, prepurification, and precooling to reach the critical temperature for the gas in question, and the expansion through an expansion machine or a throttle (expansion) valve or both.


QR

Compressor m

2

1

1

2

We

con st

Expander

f

p=

T

We

f

f m

s (a)

Liquid

Liquid reservoir (b)

FIGURE 2.2 A thermodynamically ideal liquefaction system. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)

The variety of refrigeration process cycles that are available today to the process engineer are too numerous to be listed in detail. Final selection depends on the imagination and ingenuity of the process engineer, and on the constraints imposed by economics. Before outlining the principal cycles in use today, and on which all other cycles are based, it may be useful to review the original Linde, Claude, and cascade cycles in terms of efficiency and economic viability. These basic cycles are still in use along with an infinite variety of modifications and refinements to limit the increase in entropy to a minimum, increase the efficiency, and reduce power consumption and investment. Figure 2.2 shows an ideal cycle using a complete isothermal compression with a complete adiabatic compression.

2.5 LINDE CYCLE (FREE EXPANSION THROUGH A VALVE)4 The Linde cycle was the first successful industrial cycle for the liquefaction of air that involved compression, prepurification, heat exchange, and free expansion through a throttle valve. The air was first compressed at a pressure of 200 bar. In the prepurification phase, the air was first passed through a solution of 15% lump caustic soda in water for the removal of carbon dioxide, and then passed through a tower filled with dry lump caustic for the removal of entrained water and water vapor. As may be imagined, this method of prepurification was not very efficient, and plants had to be stopped almost every week for a complete deriming. Strangely enough, this method of prepurification continued for over 50 years. The purified process air then entered the primary heat exchanger system contained in an insulated ‘‘cold box’’ for the precooling of the process air in countercurrent heat exchange with the outgoing colder products. Downstream of the heat exchanger, the air passed through a throttle valve undergoing a Joule–Thompson free expansion with a slight decrease in temperature. The cooled air was then recycled through the primary heat exchanger system where it cooled down the incoming process air. Obviously, the cooling effect at the start of the process was minimal. After a long period of time, however, the process air coming out of the primary heat exchanger was sufficiently cool and near its dew point (around 100 K), so it became partially liquid after its expansion through a valve. The liquid portion could then be withdrawn, and


Makeup gas mf 1 m 1

Q 2

3

1 (m ⫺ mf) g

W

4

f

Liquid

mf

T = const 1

con

p=

p=

con

st

Temperature T

st

2

3

f

h = const

4

g

Entropy s

FIGURE 2.3 A simple Linde–Hampson high pressure and T–s diagram liquefaction system. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)

the uncondensed vapor portion could be recycled to maintain a temperature balance throughout the entire system. In the first Linde cycle, Figure 2.3, the refrigeration available was very small because it operated at almost constant enthalpy (isenthalpic). The small drop in temperature was provided by the kinetic energy of the molecules in the expanding gas, and not from any work done by the gas itself. A thermodynamic energy analysis of the original Linde–Hampson cycle indicates the following.

2.5.1

THEORETICAL ANALYSIS

OF THE

FIRST LINDE HIGH PRESSURE CYCLE5

Assuming a steady flow to the combined unit involving the heat exchanger, expansion throttle valve, and receiver, and applying the First Law, which states that 0 ¼ (m mf )h1 þ mf hf mh2

(2:1)

y ¼ mf=m ¼ (h1 h2 )=(h1 hf )

(2:2)

liquid yield is given by


Example In the application of the simple, original, and thermodynamically perfect Linde cycle, and assuming the following gas property values for nitrogen, if the latter is the refrigerant then h1 ¼ 462 J=g at 101.325 kPa and at 300 K h2 ¼ 432 J=g at 20265 kPa and at 300 K hf ¼ 29 J=g at 101.325 kPa and at saturated liquid phase s1 ¼ 4.42 J=g K at 101.325 kPa and at 300 K s2 ¼ 2.74 J=g K at 20265 kPa and at 300 K sf ¼ 0.42 J=g K at 101.325 kPa and at saturated liquid The liquid yield therefore from Equation 2.2 is y ¼ (462 432)=(462 29) ¼ 0:0693 and the work per unit mass compressed from Equation 2.2 is W1=mf ¼ T1 (s1 s2 ) (h1 h2 ) ¼ (300)(4:42 2:74)(462 432) or ¼ 504 30 ¼ 474 J=g

(2:3)

The work per unit mass liquefied is W1=mf ¼ 474=0:0693 ¼ 6840 J=g

2.5.2 THEORETICAL ANALYSIS

OF

LINDE BASIC CYCLE

WITH

PRECOOLING6

In a subsequent modification, Linde made use of a prior development of Joule–Thompson in their formula for expansion of air: Tk ¼ ( p1 p2 )(273=T)2

(2:4)

where the pressures are a few bars, and T is near ambient. If the temperature decreases, the expansion increases the temperature difference. Linde took advantage of this feature by placing an ammonia machine unit between the main air compressor and the primary process heat exchanger. This added unit lowered the temperature of the process air to 2408K (338C). Presently, ammonia chillers have been replaced by other more reliable refrigeration machines that are more environment-friendly. Example As before (see Figure 2.4), the air was compressed to 200 bar, but the primary heat exchanger was split into two sections. Between the two sections, process air entered a refrigeration machine lowering its temperature to 270 K, (338C). The property values from the T–s diagram for nitrogen used as the main refrigerant are the same as for the simple high pressure Linde cycle; ha ¼ 207.94 kJ=kg refrigerant at 101.325 kPa and at 300 K, hb ¼ 250.2 kJ=kg compressed refrigerant at 681.7 kPa and at 372.85 K, and hc ¼ 61.23 kJ=kg refrigerant at 300 K and saturated liquid. In this specific case involving an external refrigeration machine the thermodynamic balance is Yield, y ¼ ½(h1 h2 )=(h1 hf ) þ r½(ha þ hc )=(h1 hf ) ‘‘r’’ is the refrigerant mass-flow rate, or mf=m.


(2:5)

Cooling water Refrigerant compressor a mr

b

c

Refrigerant condenser

J−T valve

We2 a

QR

Makeup gas

d

2

4

3

m

1

6

1

J−T valve

Heat exchangers (m ⫺ m f)

We1

g

5

Main compressor

f

Liquid

mf

T = const

2

3

Refrigerant Boiling point

6

4

f

p=

p=

con

st

con

st

Temperature T

1

h = const

5

g Entropy s

FIGURE 2.4 A precooled Linde–Hampson high pressure liquefaction system, and T–s diagram. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)


The first term is the normal yield for a simple Linde cycle, the second term is the improvement by virtue of precooling. The mass flow rate of the refrigerant is 0.10. Therefore, the yield is y ¼ ½(462 432)=(462 29) þ 0:10½(207:94 61:23)=(462 29) ¼ 0:0693 þ 0:0339 ¼ 0:1032 which indicates an increase of 48.92% by precooling. The work requirement may be calculated from W=m ¼ T1 (s1 s2 ) (h1 h2 ) þ r(hb ha )

(2:6)

Hence W=m ¼ (300)(4:42 2:74) (462 432) þ (0:10)(250:20 207:94) ¼ 504 30 þ 4:2 ¼ 478:2 kJ And the work per unit mass liquefied is therefore W=mf ¼ 478:2=0:1032 ¼ 4634 J=g The improvement ¼ 6840 4634 ¼ 2206 J=g or 33%.

2.5.3 THEORETICAL ANALYSIS

OF THE

LINDE HIGH-PRESSURE DUAL PROCESS7

Linde also took the efficiency of the cycle another step further with the use of a dual compression cycle plus precooling. He employed the simple well-known fact that the compression energy varies as the log10 of the compression ratio. The compression ratio of the original cycle was 200=1 and its log10 was 2.310, whereas in the dual cycle the compression ratio of the main air compressor was 200=50, or 4 and its log10 was 0.6021. The process cycle, Figure 2.5, involved the initial compression of a smaller quantity of the process air, 20%–25% of the total, to a pressure of 50 bar and then recycling this quantity from 50 to a higher pressure of 200 bar. The higher-pressure stream passed through a threechannel heat exchanger and expanded to the basic pressure of 50 bar resulting in the liquefaction of a portion of the stream. The unliquefied portion is returned to the highpressure second compressor after passing through the heat exchanger cooling the incoming stream. Assuming that i is the intermediate-pressure stream flow-rate ratio (i ¼ mi=m) and m is the total mass flow rate through the high-pressure compressor, the yield of the liquid is y ¼ ½(h1 h3 )=(hi hf ) i½(h1 h2 )=(h1 hf )

(2:7)

And the work requirement=unit of mass of gas compressed in the high-pressure compressor is W=m ¼ ½Ti (s1 s3 ) (h1 h3 ) i½T1 (s1 s2 ) (h1 h2 )

(2:8)

Example Using nitrogen as the working fluid between 101.3 kPa (300 K) and 2030 kPa, to determine the liquid yield and work requirement per unit mass. The intermediate pressure is at 507 kPa and its flow rate is 0.80. In these conditions the properties are


N2 system Q

W W b Makeup gas

c

Q

m

2

a

mN

2

3

LN2 bath 4

1

7 (m ⫺ mf)

W

5 g

mf

6

Liquid

f

FIGURE 2.5 Linde dual-pressure process cycle. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)

h1 h2 h3 hf s1 s2 s3

¼ ¼ ¼ ¼ ¼ ¼ ¼

462 J=g at 101.3 kPa and 300 K 452 J=g at 507 kPa and 300 K 432 J=g at 2030 kPa and 300 K 29 J=g at 101.3 kPa and saturated liquid 4.42 J=g K at 101.3 kPa and 300 K 3.23 J=g K at 507 kPa and 300 K 2.74 J=g K at 2030 kPa and 300 K

From Equation 2.7 the liquid yield is y ¼ [(462 432)=(462 29)] 0:80 [(462 452)=(462 29)] ¼ 0:0693 0:0185 ¼ 0:0508 The work requirement is W=m ¼ [(300)(4:42 2:74) (462 432)] (0:80)[(300)(4:42 3:23) (462 452)] ¼ 474 277:6 ¼ 196:4 J=g


The work per unit of mass is W=mf ¼ 196:4=0:0508 ¼ 3866 J=g As one may see, the double cycle not only indicates a lower liquid yield than the simple highpressure cycle, but also a lower power consumption. Nevertheless, one may add a mechanical refrigerator operating at 240 K (338) and the operating efficiency is increased to 48.7%.

2.6 THEORETICAL ANALYSIS OF THE CLAUDE CYCLE8 The calculation of the Claude cycle may be a little burdensome, but it may prove useful to practicing engineers who may be interested in performing a heat and mass balance around a cryogenic system should an operating problem occur. Following Figure 2.6 and assuming that the expansion engine is reversible and adiabatic, the expansion process is isentropic. In this manner a much lower temperature is achieved than an isenthalpic expansion as practiced by Linde. In the Claude cycle, air or feed gas, is compressed to approximately 40 bar and passed through a primary heat exchanger. About 75% of the feed is diverted, expanded, and returned (reunited) with the cold vaporized gas from the liquid being produced. The combined streams are then recycled and passed through an ultimate expansion valve for total liquefaction. The use of an expansion valve is still necessary to avoid the formation of liquid in the cylinder of the reciprocating expansion machine. In reviewing the T–s diagram and applying the First Law for a steady-state flow through the heat exchangers, expansion valve, and liquid container as a single unit with no heat transfer, then 0 ¼ (m mf )h1 þ mf hf þ me he mh2 me h3

(2:9)

Assuming that x flow passes through the expander, or

x ¼ me=m

the yield is y ¼ mf=m ¼ [(h1 h2 )=(h1 hf )] þ x[(h3 he )=(h1 hf )]

(2:10)

Presuming that the last term is less than unity, the cycle is a definite improvement over the Linde high-pressure process. If the expander work aids in compression, then that work is expressed by We ¼ me (h3 he )

(2:11)

And the net work if the expander is used to aid compression is W=m ¼ [T1 (s1 s2 ) (h1 h2 )] x(h3 he )

(2:12)

It may be concluded that at the point ‘‘3’’ the inlet condition of the expander for temperature and pressure will yield the minimum work required. Example Using nitrogen as an example determine the liquid yield, total work of compression, and the work to liquefy a unit mass of gas for the Claude cycle. The system operates between 101.3 kPa, 300 K, and 5066 kPa. The expander inflow is at 60% and the conditions are at 270 K and 5066 kPa (50 atm).


Qr 1

3

2

9 m We

5

4

8

7

me g

We

6

Expander e Evaporator

Liquid

mf

Qa

T = const

2

1

3

p=c

s = const

p=

Temperature T

con

st

ons

t

9

4 8 5

e

7

h= t

ns co

f

6

g

Entropy s

FIGURE 2.6 A Claude low-pressure process cycle using an expansion machine and its T–s diagram. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)


The T–s diagram for nitrogen gives the following properties: h1 h2 h3 he hf s1 s2 s3

¼ ¼ ¼ ¼ ¼ ¼ ¼ ¼

462 J=g at 101.3 kPa and 300 K 452 J=g at 5066 kPa and 300 K 418 J=g at 5066 kPa and 270 K 238 J=g at 101.3 kPa and se ¼ s3 (Te ¼ 86.1 K) 29 J=g at 101.3 kPa and saturated liquid 4.42 J=g K at 101.3 kPa and 300 K 3.23 J=g K at 5066 kPa and 300 K se ¼ 3.11 J=g K at 5066 kPa and 270 K

From Equation 2.10 the yield is y ¼ [(462 452)=(462 29)] þ 0:60 [(418 238)=(462 29)] ¼ 0:0231 þ 0:2494 ¼ 0:2725 As 0.40% per unit kg was compressed and passed into the liquid vessel, one may assume that 70% was liquefied, and the total work is given by Equation 2.12, or W=m ¼ (300)(4:42 3:23) (462 452) (0:60)(418 238) ¼ 347 108 ¼ 239 J=g compressed The work required to liquefy a unit mass of nitrogen gas is W=mf ¼ 239=0:2725 ¼ 877 J=g liquefied: The above result is very high compared with other systems. Its theoretical operating efficiency has been calculated at 48.3%.

2.6.1 CLAUDE CYCLE WITH PRECOOLING (FIGURE 2.7) The Claude cycle can also be improved by adding a mechanical refrigeration unit immediately after the main air compressor. The calculations are the same as previously noted, but are lowered by dropping the process air temperature to say 278 K. In fact, all designers of present-day air-separation plants use precooling as a standard practice.

2.6.2 CLAUDE CYCLE WITH DUAL PRESSURES7 The basic Claude dual-pressure cycle is similar in principle to the dual-pressure Linde process; but in the Claude cycle, only the portion of the stream entering the expander is compressed to a higher pressure, which reduces the energy consumption considerably (Figure 2.8). Again using nitrogen as the refrigerant, the main process feed is compressed from 101.3 to 1013 kPa (10 bar). Approximately 60% is cooled to 160 K and sent to the expander. The remainder 40% is further compressed to 3530 kPa (35 bar), cooled, and undergoes a J=T expansion through a throttle valve to 101.3 kPa (atmospheric level). Following the T–s schematic, the following properties are derived: h1 ¼ 462 kJ=kg at 101.3 kPa and 300 K h2 ¼ 460 kJ=kg at 101.3 kPa and 300 K


Aftercooler Air Compressor

Purification Oxygen Nitrogen

Main exchanger

Expansion engine

Gaseous nitrogen Gaseous oxygen

Expansion valve

Rectification section

Air

Liquid oxygen Liquid nitrogen

FIGURE 2.7 Claude cycle with precooling. (Courtesy of F.G. Kerry, Inc., 1977. With permission.)

h3 ¼ 455 kJ=kg at 3530 kPa and 300 K h11 ¼ 309 kJ=kg at 1013 kPa and 160 K hf ¼ 29 kJ=kg at 101.3 kPa and saturated liquid hg ¼ 229 kJ=kg at101.3 kPa and saturated vapor s1 ¼ 4.421 kJ=kg K at 101.3 kPa and 300 K s2 ¼ 3.732 kJ=kg K at 1013 kPa and 300 K s3 ¼ 3.345 kJ=kg K at 3530 kPa and 300 K s11 ¼ 3.052 kJ=kg K at 1013 kPa and 160 K sf ¼ 0.418 kJ=kg K at 101.3 kPa and saturated liquid Excluding the compressors, the energy balance around the system is me h2 þ (m me )h3 ¼ (m mf )h3 þ me (h11 he ) þ mf hf Assuming that y ¼ mf=m and x ¼ me=m, then


(2:13)

Compressor Makeup gas

Heat exchanger QR2

QR1

2 4

m 1

2

3 10

1 11

Wc 1

Wc 2

me

We 9 5 Expander

e

8

6

g

mf

7

Liquid

f

FIGURE 2.8 Claude dual-process cycle. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)

y(h1 hf ) ¼ x(h3 h2 ) þ (h1 h3 ) þ x(h11 he )

(2:14)

The yield is y ¼ [(h1 h3 )=(h1 hf )] x[(h2 h3 )=(h3 hf )] þ x[(h11 he )=(h11 hf )]

(2:15)

The energy balance per unit mass can be computed by adding the energy balance per unit mass of the two compressors and then subtracting that of the expander work, or W=mnet ¼ T1 (s1 s2 ) (h1 h2 ) þ (1 x)T1 (s2 s3 ) (h2 h3 )] x(h11 he ) ¼ T1 (s1 s3 ) (h1 h3 ) x[T1 (s2 s3 ) (h2 h3 )] x(h11 he )


(2:16)

From a T–s diagram one finds that at 101.3 kPa, he ¼ 233 kJ=kg and the entropy ¼ 3.052 kJ=kg K. Therefore y ¼ [(462 455)=(462 29)] 0:6[(460 455)=(462 29)] þ 0:6[(309 233)=(462 29)] ¼ 0:1145 The net work is given by W=mnet ¼ 300(4:421 3:345) (462 455) 0:6[300(3:732 3:345) (460 455)] 0:6(309 233) ¼ 203:5 kJ=kg and the ideal value to liquefy nitrogen in these conditions is 768 kJ=kg.

2.6.3

CLAUDE CYCLE

WITH

HIGH PRECOOLING

TO

LIQUEFY HYDROGEN

OR

NEON9

The Claude cycle may be used to liquefy hydrogen or neon without modification by using liquid nitrogen precooling (see Figure 2.9), and assuming the following operating conditions. . .

. . . . .

Para-hydrogen compressed to 0.01 kg=s for liquefier and refrigerator Temperatures of all streams on the warm side of the three-channel heat exchanger are at 293 K Compressor operates between 1.01 and 10.1 bar Inlet temperature to expander is 60.8 K 70% of para-hydrogen diverted through the expander Assume ideality for expander, compressor, and heat exchangers The liquid nitrogen bath operates at a pressure of 1.01 bar (101 kPa)

From existing tables the following properties for para-hydrogen are assumed with reference to Figure 2.9. h1 (101 kPa, 293 K) ¼ 4100 kJ=kg s1 (101 kPa, 293 K) ¼ 64.5 kJ=kg K h2 (1010 kPa, 293 K) ¼ 4100 kJ=kg s2 (1010 kPa, 293 K) ¼ 55 kJ=kg K h4 (1010 kPa, 77.3 K) ¼ 780 kJ=kg h5 (1010 kPa, 60.8 K) ¼ 590 kJ=kg s5 (1010 kPa, 60.8 K) ¼ 82 kJ=kg K h12 (101 kPa, 77.3 K) ¼ 808 kJ=kg hf (101 kPa, 20.4 K) ¼ 256 kJ=kg sf (101 kPa, 20.4 K) ¼ 8 kJ=kg K ha (101 kPa, 77.3 K) ¼ 29 kJ=kg hc (101 kPa, 293 K) ¼ 455 kJ=kg (A) Make an energy balance including the three cold exchangers, expander, J-T valve, and the liquid reservoir to obtain the liquefaction rate of the para-hydrogen:


Compressor mN2

Makeup gas

LN2 bath

QR

LN2

m

(m−mf) We me

We Expander

J−T valve

Liquid mf

FIGURE 2.9 Claude precooled process cycle for liquefaction of neon and hydrogen. (From Timmerhaus, K.D., Cryogenic Process Engineering, Springer Science & Business Media, New York, 1989. With permission.)

mh4 ¼ (m mf )h12 þ mf hf þ me (h5 he )

(2:17)

If one assumes that me=m ¼ x, and that mf =m ¼ y, then (h4 h12 ) ¼ y(hf h12 ) þ x(h5 he ) (2:18) therefore, y ¼ [(h12 h4 )=(h12 hf )] þ x[(h5 he )=(h12 hf )]

(2:19)

For an ideal expander s5 ¼ se ¼ 32 kJ=kg K at 101 kPa, and he ¼ 235 kJ=kg and substituting numerical values y ¼ [(808 780)]=[808 (256)] þ 0:7[(590 235)=808 (256)] ¼ 0:260 For 0.01 kg=s compressed, the liquefaction rate ¼ 2.6 103 kg=s.


(B) To determine the liquid nitrogen consumption for this liquefaction rate one makes an energy balance around the three-channel heat exchanger and the precoolant liquid nitrogen reservoir. mh2 þ mN2 ha þ (m mf )h12 ¼ mh4 þ mN2 he þ (m mf )h1

(2:20)

m(h2 h4 ) þ (m mf )(h12 h1 ) ¼ mN2 (hc ha )

(2:21)

mN2 ¼ [(h2 h4 ) þ (m mf )(h12 h1 )]=(hc ha )

(2:22)

Substituting numerical values mN2 ¼ [(0:01)(4100 780) þ (0:01 0:0026)(808 4100)]=(455 29) ¼ 2:07 102 kg=s of liquid nitrogen consumed:

2.6.4

THE LOW-TEMPERATURE REFRIGERATOR

The Claude process cycle using nitrogen as a precoolant may also be used as a low-temperature refrigerator by using the vaporization of the liquid produced.

2.7

KAPITZA CYCLE10

In 1939, when Kapitza suggested the use of centrifugal expansion turbines for the Claude cycle, he also suggested the use of reversing regenerators packed with aluminum strips, the latter developed by Linde–Frankl in 1928. These reversing regenerators replaced not only the primary countercurrent heat exchangers but also served as prepurifiers for the process air. They consisted of diagonally corrugated aluminum strips about 5080 mm wide and approximately 3–5 mm thick. These strips were coiled to any diameter and stacked vertically to any height specified by the process engineer. These reversing regenerators replaced not only the primary countercurrent heat exchangers, but also served as prepurifiers for the process air. This combination became the basic process cycle during World War II. Reversing regenerators were widely used for large plants during the war and shortly afterwards, they have since been replaced by brazed aluminum heat exchangers developed in the United States in the late 1940s. Since the war, a wide variety of process cycles based on the basic Claude cycle and the throttle valve have developed. They are too numerous to be described individually but a few specific ones are described in Chapter 3.

2.8

CASCADE CYCLE11

The Cascade process cycle for the liquefaction of gases is certainly not new. For example, in 1890, Kamerlingh-Onnes in Leyden used a series of refrigerants for his experimental work on cryogenics. The refrigerants originally used were as follows: . .

Liquefied methyl chloride at ordinary temperatures between 4 and 5 bar. Liquefied ethylene in a bath of methyl chloride boiling at 183 K and a pressure of 10 mm of mercury. At this temperature ethylene liquefies at a pressure of 5–6 bar. Also ammonia vaporizes at 240 K. Ammonia has been replaced by propane.


.

Liquefied ethylene boiling at a pressure of 20 mm of mercury had a temperature of 131 K. At this temperature, oxygen can be liquefied at a pressure of 20 bar. Ethylene has been replaced by ethane.

By vaporizing liquid oxygen at 90 K, it was possible to reach a temperature down to 73 K; and by vaporizing it under a vacuum, one could attain much lower temperatures. Of course, the use of liquid nitrogen is safer and more effective, because it has a lower boiling point of 77.3 K. In an up-to-date cascade cycle, Figure 2.10, the refrigerant gases employed undergo compression, precooling, expansion through a throttle valve, liquefaction, vaporization, and subsequent recycling to their separate compressors. This cycle is an extension of the principles of precooling. The refrigerants are each liquefied and vaporized sequentially in conventional closed loops. Moreover, each refrigerant may be vaporized at several pressure levels in order to achieve lower temperatures, hence increasing efficiency (but at a higher cost of investment and operating complexity).

Cooling water

Ammonia NH3

Ethylene C2H4

Methane CH4

Makeup gas

Liquid nitrogen

FIGURE 2.10 Process Schematic of a classic Cascade cycle using external closed loops of refrigerants: ammonia, ethylene, methane, nitrogen. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)


In this process, each refrigerant has a progressively lower boiling point than the previous refrigerant used for precooling. For example, nitrogen may be liquefied by using four cycles in series: ammonia boiling at 240 K, ethylene boiling at 189 K, methane boiling at 112 K, and nitrogen boiling at 77 K. This process is thermodynamically efficient because it approximates an ideal reversible cycle. The expansion through a valve occurs at a lower pressure, and therefore results in a very low increase in entropy. In spite of its slightly higher process efficiency, this cycle is seldom if ever used in air separation because of its high refrigerant losses due to leaks, complexity in operation, poorer safety factor, and higher required investment. After World War II, with the tremendous development of the petroleum and petrochemical industries, some engineers did consider these four refrigerants for the separation of air for very large production of oxygen. Calculations confirmed that the cascade cycle was more efficient than the standard Claude cycle. In fact, an engineering company took a patent on it without knowing that it had been conceived and used many years before. In the 1960s, a merchant air-separation plant was built in Louisiana using a cascade cycle of several independent hydrocarbon refrigerants. When the plant went into operation, however, production costs were much higher than estimated due to excessive mechanical maintenance and continuous leaks of the refrigerants. Nevertheless, this cycle has been, and is still being, used for the liquefaction of hydrogen and helium and for the large-scale liquefaction of natural gas. Propane and ethane are abundant in natural gas. Therefore, they can be substituted in place of ammonia and ethylene with a minimal loss in efficiency. The cascade cycle is widely used in the treatment of natural gas for the recovery of pure methane, and is the standard for the large-scale production of liquefied natural gas (LNG). Further discussion can be found in Chapter 8. Chapter 2 has given the reader an overview of the energy required to recover industrial gases. An ideal quantitative summary is given in Table 2.1 at the following conditions of temperature at 300 K and a pressure of 101.325 kPa (1 atm).

TABLE 2.1 The Ideal Work Necessary for Liquefaction of Various Gases Gas Helium-3 Helium-4 Hydrogen, H2 Neon, Ne Nitrogen, N2 Air Carbon monoxide, CO Argon, A Oxygen, O2 Methane, CH4 Ethane, C2H6 Propane, C3H8 Ammonia, NH3

Normal Boiling Point (K)

Ideal Work of Liquefaction (Wl=mf) (kJ=kg)

3.19 4.21 20.27 27.09 77.36 78.8 81.6 87.28 90.18 111.7 184.5 231.1 239.8

8178 6819 12019 1335 768.1 738.9 768.6 478.6 635.6 1091 353.1 140.4 359.1

Source: From Barron, R.F. in Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.


REFERENCES 1. Gomonet, E. 1952. Les basses tempe´ratures, production et emplois, 71–91. Paris: Librairie J.B. Bailliere et Fils. 2. Gomonet, E. 1952. Les basses tempe´ratures, production et emplois, 54–79. Paris: Librairie J.B. Bailliere et Fils. 3. Gomonet, E. 1952. Les basses tempe´ratures, production et emplois, 80–87. Paris: Librairie J.B. Bailliere et Fils. 4. Barron, R.F. 1985. Cryogenic systems (Monogram on cryogenics 3), 2nd ed., 61–64. New York: Oxford University Press. 5. Barron, R.F. 1985. Cryogenic systems (Monogram on cryogenics 3), 2nd ed., 69–72. New York: Oxford University Press. 6. Barron, R.F. 1985. Cryogenic systems (Monogram on cryogenics 3), 2nd ed., 73–80. New York: Oxford University Press. 7. Barron, R.F. 1985. Cryogenic systems (Monogram on cryogenics 3), 2nd ed., 91. New York: Oxford University Press. 8. Barron, R.F. 1985. Cryogenic systems (Monogram on cryogenics 3), 2nd ed., 85–89. New York: Oxford University Press. 9. Timmerhaus, K.D. 1989. Cryogenic Process engineering, 132. New York: Plenum Press. 10. Gomonet, E. 1952. Les basses tempe´ratures, production et emplois, 91. Paris: Librairie J.B. Bailliere et Fils. 11. Gomonet, E. 1952. Les basses tempe´ratures, production et emplois, 67. Paris: Librairie J.B. Bailliere et Fils.

FURTHER READING Springman, H. 1985. Cryogenics, principles and applications. Chem Eng (May 13):58–67. Abbot, W.F. 1989. Thirteen ways of looking at the van der Waal equation. Chem Eng Prog (February):25–26. Kenney, W.F. 1989. Current practical appliations of the second law of thermodynamics. Chem Eng Prog (February):57–69.



3

Air Separation Technology

3.1 AIR SEPARATION OVERVIEW Oxygen is the most prevalent element on the Earth. It constitutes 46.6% by weight of the crust and 89% of the sea water. It also constitutes 23% by weight (20.9% by volume) of the Earth’s atmosphere (see Table 3.1). Oxygen was discovered by Scheele of Sweden (1773), followed by Priestly (1774), and investigated in depth by Lavoisier (1775) who named it oxygen (acid forming in Greek). Lavoisier studied its support of life by inspiration, as well as its acceleration of combustion. Oxygen is a diatomic gaseous element O2, colorless, odorless, and slightly soluble in water (48.9 mL in 1 L at 273.19 K). At 90.18 K and 1 barA, it is a pale blue liquid slightly heavier than water, and at 55.15 K it freezes into a crystalline solid. Both liquid and solid oxygen are paramagnetic, which is rare as other substances are diamagnetic. Though normal oxygen is diatomic, it also exists in the triatomic form as ozone and as monatomic O in the upper atmosphere (Table 3.2). As the industrial revolution accelerated in Europe, especially in the latter part of the nineteenth century, there was an increased demand for primary products such as steel and coal to satisfy production in factories. Engineers and scientists began to review the possibilities of employing oxygen, this newly discovered element that could accelerate combustion, and therefore increase the manufacture of steel products so much in demand. As noted, oxygen was available in the atmosphere but only at approximately 21%. To recover it economically required the design and construction of an industrial-size process unit. Moreover, the only process considered was liquefaction, separation, and revaporization. Gaseous oxygen could increase the smelting of iron ore. Oxygen, in liquid form, when mixed with carbon particles, could be used as a more powerful explosive to mine much needed coal in larger quantities. The liquefaction of air was only the first step to recover and purify large-scale oxygen production. Linde had proved that air could be liquefied on a commercial scale. The next step was to fractionate air and recover oxygen in as pure a state as required by the industry. Linde’s first venture in the separation of air was not very successful. He permitted liquid air to evaporate, assuming that nitrogen (the more volatile component) would evaporate completely leaving behind fairly pure liquid oxygen. Needless to say, the experiment did not work. Even though there was an important difference of 12.9 K in their respective boiling points at 101.325 kPa, a substantial portion of oxygen evaporated along with the nitrogen. In short, Linde found out that separating nitrogen from a mixture with oxygen was not an easy task. Even assuming that atmospheric air is a simple mixture of only two components, namely oxygen and nitrogen, as everyone believed at the time, the separation of these two gases required additional external work. Without wasting any time, Linde therefore turned his attention to distillation and fractionation.


TABLE 3.1 Constituents of the Earth’s Atmosphere in Mole Fraction Nitrogen Oxygen Argon Carbon dioxide Neon Krypton Xenon Hydrogen deuteride Deuterium Helium

0.78084 0.20946 0.00934 0.00033 1.818 105 1.14 106 8.6 108 3.12 104 1.56 104 1.0 107

Source: Courtesy of Lotepro Data Book, 1975. With permission.

3.1.1

LINDE’S FIRST FRACTIONATION MACHINE

Although Linde had taken out a patent in 1891 for an industrial machine that liquefied air, his first air fractionation column was actually designed only in 1902. It consisted of a single column with random packing. After the usual prepurification to remove CO2 and H2O using NaOH (see Chapter 5), the process air was compressed to 200 bar, and passed through a series of exchangers in countercurrent heat exchange with outflowing products. Air was liquefied using the Joule–Thompson effect. Part of this liquid product was sent to the top of the single tower as reflux. The procedure followed the basic distillation principles known at the time. Unfortunately, the gaseous product leaving the top of the column had an oxygen content of 7%. Calculating the oxygen recovery one finds (Figure 3.1a). 21 ¼ x þ (1 x) 0:07 x ¼ 0:1505 TABLE 3.2 Physical Properties of Oxygen Molecular weight Normal boiling point Vapor pressure Latent heat of vaporization Molar heat capacity of gas @constant pressure in the range of 0–1 bar and at 298 K Thermal conductivity of gas Thermal conductivity of liquid Gaseous real density Liquid density Critical temperature Critical pressure Flammability range Viscosity Dielectric constant

32.00 kg=kmol 90.18 K @ 101.325 kPa 1.52 mbar @ 54.35 K 6.83 MJ=kmol @ boiling point and 101.3 kPa (212.9 kJ=kg) 29.22 kJ=kmol K 0.024 W=m K at 273 K 151.4 mW=m K 1.429 kg=m3 @ 273.15 K and 101.325 kPa 1.141 kg=L at 90.15 K and 1.013 bar 154.576 K 50.4 bar None, but accelerates combustion 188 kg=m s 106 1.4837



N2 + 7% O2

B

Almost pure O2

A (a)

Pure N2

Pure O2 Pump for reduction of pressure

Liquid air

Liquid oxygen

Compressed air (b)

FIGURE 3.1 (a) Linde’s first attempt to separate air. (b) Linde’s second attempt to separate air. continued


Liquid 96% N2 and 4% O2

99% N2

C Liquid with about 40% O2

98% O2 D

E

B

Compressed air (c)

A

FIGURE 3.1 (continued) (c) Linde’s final and successful attempt. (Courtesy of Longmans, Green & Co. With permission.)

Recovery of oxygen ¼ 0.1505=21 ¼ 71.75%, which was commercially inadequate. Furthermore, the nitrogen product containing 7% oxygen was unusable. By trial and error, it was determined that a single column required pure liquid nitrogen reflux at the top. This was carried out by sending the liquefied air into the middle of the column. As the liquid entered the column, its vapor rose to the top and passed through a special apparatus containing pure liquid nitrogen boiling at a temperature of 75 K, even though its normal boiling point was 77.4 K. This lower temperature was achieved by boiling the liquid nitrogen at a very low pressure, around 93 mbar. The procedure stripped the rising vapor of all the oxygen, which after being recondensed descended to the bottom of the tower. The vapor leaving at the top was almost pure nitrogen. This process, however, required a vacuum pump with all its ancillary equipment, and also extra apparatus. Eventhough considered an important improvement, the process was not commercially acceptable and was rejected for air separation. But it did open the road for the development of the double column well known today. This development also emphasized the importance of the single column and promoted its refinement for the production of high-purity nitrogen and, ultimately, ultra-high-purity nitrogen so necessary today in the manufacture of electronic components (see Chapter 8). Finally, Claude’s adaptation of the expansion machine to operate at cryogenic temperatures in 1902 completed the general parameters of commercial air separation (Figure 3.1b and Figure 3.1c). At this point, the fundamentals of separation of air merit a review. ß 2006 by Taylor & Francis Group, LLC.

3.1.2 DISTILLATION

AND

FRACTIONATION

Distillation consists of vaporizing a more volatile component from a mixture using heat, and its subsequent recondensation by cooling and separation. The craft dates back to the ancient Alexandrians who produced oil of turpentine by distilling pine resins. In the early 1980s, industry had already developed large-scale stills for recovering alcoholic liquors equipped with bubble cap trays similar to those used today.

3.1.3 FRACTIONATION Fractionation, or fractional distillation, includes the separation of liquids or gases whose boiling points are very close to one another, for example, oxygen and argon. Its technology demands the closest possible contact between the rising vapor of the more volatile component, argon, and the descending liquid of the less volatile component, oxygen, so arranged as to provide the repeated vaporization and condensation needed to enhance separation.

3.1.4 STRIPPING Stripping means the maximum removal of the more volatile component from the liquid by heating, or reboiling, as it may be called. This occurs in the lower section of the upper (low pressure) column, where gaseous nitrogen is stripped from liquid oxygen.

3.1.5 RECTIFICATION Rectification means the enrichment or concentration of the more volatile component (gaseous nitrogen) in the upper section of the (low pressure) column at least in connection with air separation plants, and in the upper section of the (high pressure) column. In the latter case, it relates to the enrichment of liquid nitrogen.

3.2 THEORETICAL CONSIDERATIONS OF FRACTIONATION To separate even a simple binary system in order to recover its components, it is necessary to consider and understand a few basic facts of physics, with regard to vapor pressures and partial pressures.

3.2.1 EVAPORATION

AND

CONDENSATION1

Normally when one thinks of a boiling liquid one has in mind water boiling at 1008C (373.15 K), and its transformation into steam. It can also be easily understood that at elevated altitudes where the atmospheric pressure is lower, water boils at a lower temperature. Alternately, however, in a pressure cooker water boils at a higher temperature. It is a contest between the escaping molecules of the boiling liquid, and the pressure of the molecules of the ambient atmosphere. Figure 3.2a shows the vapor pressure of water as a function of temperature and pressure. If the analogy is carried over to all substances that undergo transformation from solid to liquid to gas and back again, it is clear that at higher temperatures the pressure exerted by the molecules of a vapor leaving a boiling liquid is greater as the temperature increases. The degree of increase in the ‘‘vapor pressure’’ is defined by the vapor pressure as observed in Figure 3.2b. All vapor pressure curves have an upper and a lower limit. The lower limit is known as the ‘‘triple point temperature’’ below which only a solid will exist. The upper limit is known as the ‘‘critical temperature’’ above which liquid cannot exist with vapor. For the latter reason, early pioneers could not liquefy air by compression alone, because they were


2.0

0.9

1.8 1.6

0.7

Absolute pressure in. Hg

Absolute pressure (psi)

0.8

0.6 0.5 0.4 0.3 0.2 Vapor pressure of water at 60⬚F is 0.52 in. Hg or 0.26 psia

0.1

20

30

40

50

60 70 80 Temperature (⬚F)

90

50 45 40

1.4

35

1.2

30

1.0

25

0.8

20

0.6

15

0.4 0.2

Absolute pressure (mm Hg)

1.0

10 5

100

FIGURE 3.2 (a) Vapor pressure of water vapor at varying temperatures. continued

unable to cool the air below the critical temperature as required by the laws of physics (see Figure 1.1). Figure 3.3 shows the vapor pressure curves between triple points and critical points of various industrial gases. Figure 3.4 also shows diagrammatically the equilibrium of liquid and vapor concentrations for an oxygen and nitrogen mixture. The lower curve is the liquid temperature curve, also known as the ‘‘bubble point curve.’’ Any one point or temperature in the lower curve is where the first bubble of vapor escapes from the liquid at a given temperature. The upper curve, also known as the ‘‘dew point temperature’’ curve is where the first drop of condensate forms at a given temperature. Figure 3.5 shows the vapor pressure curves of oxygen and nitrogen at various temperatures. It is important to note that at any given temperature, the vapor pressure of nitrogen is always higher than that of oxygen. In other words, nitrogen has a higher volatility than oxygen. It is this difference in vapor pressure between the two gases that provides the basis for the separation of oxygen and nitrogen. The same rule may be applied to all other gas mixtures. Figure 3.4 shows a vapor–liquid equilibrium diagram at a pressure of 1.0325 kPa. If a binary cryogenic mixture is placed in a vessel and allowed to evaporate, the liquid phase is at equilibrium at x, its bubble point at T1 kelvin, and the vapor phase is also at equilibrium at y at T1 kelvin, its dew point. The values of x and y are different and reflect the relative volatility of each component.


Critical

1000 800 600

Critical

400 f (r e

200

Pressure (psia)

nt

4) 74

ant

er frig

) 17

t7

id e io x nd o 13 rb nt Ca era g i r f Re

ran

502

Re

e frig

ia

n mo

(re

Am

So lid

100 80 60 40

r

ra ig e

20 10 8 6 4

2 50 nt 22 a t r ge an fri ger 2 e i R fr t1 an r Re ge fri Re

4

11

nt

g

fri

Re

a er

1

t1

an

r ige

fr

Re

13

t1

ge

fri

2

Re

n ra

fri

ge

ra

nt

71

8)

1 0.8 0.6

W

at

er

(re

0.4 0.2 0.1 –100

–80

–60

–40

–20

0 20 40 Temperature (⬚F)

60

80

100

120

FIGURE 3.2 (continued) (b) Vapor-pressure (boiling point) curves of some common refrigerant, to a logarithmic scale. (From King, G.R., Modern Refrigeration Practice, McGraw-Hill, New York, 1971. With permission.)

Figure 3.5 shows a series of phase equilibrium diagrams at various pressures. Note that as the pressure increases, the difference between the upper and lower curves also more or less decreases. It is of interest to note that at a pressure of 5 atm (506.6 kPa), the normal pressure in the highpressure column of an air separation unit, and at 100 K, the content of oxygen in the liquid mixture is approximately 40%, which is about the normal operating practice in the field. At this stage, it is essential to bear in mind several principles that apply to fractionation. As the ancient Alexandrians knew, the concentration of a component in either the liquid or gaseous phase can be made to increase or decrease by simply varying the temperature. The Gibbs phase rule may be interpreted to state that a separation will occur when a temperature gradient is placed across the distillation column. The more volatile components will rise to the top, and the less volatile components will move downward. According to Timmerhaus2 a distillation column is an apparatus that allows two phases to coexist in


150 140 130

2

ne

O

tha

110

n

go

Ar

TR

TR

120

Me

q)

(Li

90

TR

100

Air CO 2

–40 –60 1 10−1

10−2 2

2

3

4

3

4

5

1 9 8 7 6 5

9 8 100 7 6

60

TR TR 10−1 2

2

3

4

3

4

5

1 9 8 7 6 5

9 8 7 101 6

100 2

2

3

4

27

9 8

2

1 10 6 9 8 5 7 4 6 5 3

101

(VA)

Air

2

70

TR

80

TR

N

3

80 ⬚F

K

TR

NO

4

60 0

180 160 170

TR

KR

KR KR

3

O

KR

–80

210 200 190

TR

TR TR

TR n to yp Kr

KR

ne

yle

Eth

8 7

–340 –320 –300 –280 –260 –240 –220 –200 –180 –160 –140 –120 –100

230 220

2

CL

TR

KR KR KR KR

e

an

Eth

1 103 6 9 5 8 7 4 6 5 3

–20

250 240

2

SO

XE

Vapor pressure curves between triple points and critical points

1

LF

CC

bar

40

260

a ut Pr

S H2

ne yle op ne r P pa o

CO

2

CL 3 CH

ob Is

2

NH

F

1

L

CC

PSIA

20

300 290 270 280

e en tyl ne Bu uta B

F2 L2 CC

e

ne

e

an

nt

Pe

e Ac

tyl

TR

ne

xa

He

en

T

KR

Be n

KR

ne

ze

ne

ou e

FIGURE 3.3 Vapor pressure curves (between triple points and critical points). (Courtesy of Lotepro Data Book, 1975. With permission.)

equilibrium with one another. Since distillation deals with mixtures of vapor and liquid at the same conditions of temperature and pressure, two other principles are involved: (a) Raoult’s law, which states that p ¼ Po x

(3:1)

where p is the partial pressure of the component in the vapor phase, Po is the vapor pressure of the pure component at the temperature To of the mixture, and x is the mole fraction of the component in the liquid phase.


0

Condensation of a binary system Vapor component

100%

Vapor phase

A

y

Dew point

To

D

To

C

x B

Bubble point

Liquid phase

0

Liquid component

100%

FIGURE 3.4 Vapor–liquid equilibrium diagram. (Cryogenic mixture brought to To K and allowed to separate. Liquid phase is at its bubble point and at equilibrium at To K. Vapor phase is at equilibrium at y and at To K. The values of x and y are different and reflect their relative volatility. At the temperature To K, the equilibrium values are x (liquid phase) and y (vapor phase). The ratio of y=x ¼ k (which is the vaporization equilibrium constant). It is this constant at various values of x and y that gives the equilibrium line shown in Figure 3.10. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

(b) Dalton’s law states that the sum of all the partial pressures of all components in a gaseous mixture is equal to the total pressure of the mixture or p ¼ pa þ p b þ p c

(3:2)

where p is the total pressure. Moreover, assuming that the vapor phase in a fractionation column for air separation is an ideal gas, it is possible to equate the ‘‘Y’’ obtained from both Raoult’s Law and Dalton’s Law. Therefore, for all intents and purposes, pY ¼ Po x

(3:3)

where p is the total pressure. In summary, the above equations show that at a given temperature the vapor pressure of nitrogen is always higher than the vapor pressure of oxygen, and that in vapors coming from a boiling mixture of the two components the vapor will always be richer in nitrogen than the liquid mixture. In other words, the nitrogen component is more volatile than the oxygen component. This provides the basis for the separation and fractionation of air.


135 N2−O2 130 20 atm 125

120 15 atm 115

Temperature (K)

10 atm 110

105 5 atm 100

95 2 atm

90

85

1 atm

80

75

0

0.2

0.4

0.6

0.8

1.0

Mole fraction N2

FIGURE 3.5 Temperature–composition diagram for oxygen–nitrogen mixtures at various pressures. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)

3.2.2

SIMPLE SEPARATION

BY

CONDENSATION

AND

FLASHING (SEPARATORS)

Partial separation of a binary mixture can be obtained either by simple condensation or by evaporation. This can be accomplished quite easily if the two components have a large difference in their degree of volatility (helium–nitrogen), but with more difficulty when the difference in the degree of volatility is close (oxygen–argon). The cooling for separation is generally associated with the expansion through a throttle valve, and the yield in certain conditions is a gaseous and a liquid product with a higher concentration of the more volatile component in the vapor, and a higher concentration of the less volatile component in the


Vapor

Feed J−T valve

Liquid

FIGURE 3.6 Separator (flash condenser). (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

liquid product (Figure 3.6). This apparatus is called a separator or a flash condenser, and its throttle valve upstream of the device is adjusted to achieve a specific ratio between the gaseous and the liquid product. Though the design of the device is simple enough, its process calculation requires iteration and patience in the choice of the correct downstream temperature. The degree of separation between two phases can be achieved as soon as the equilibrium conditions exist and may be obtained when yi=xi ¼ Poi=pi ¼ Ki

(3:4)

which is a constant, and is known as the distribution coefficient or more practically, the equilibrium ratio. The above expression shows that the liquid obeys Raoult’s law, and the vapor obeys Dalton’s law, in other words, the relative volatility (ease of separation) between two components is equal to K1=K2. For real (or nonideal) gases, this expression is found through experimentation and is listed in tables as ln Kp=Kpo. In a binary system (oxygen–nitrogen), y1 ¼ K1 x1

(3:5)

y2 ¼ K2 x2 ¼ K2 (1 x1 )

(3:6)

y1 þ y2 ¼ 1 ¼ K1 x1 þ K2 (1 x1 )

(3:7)

and


or x1 ¼ (1 K2 )=(K1 K2 )

(3:8)

y1 ¼ K1 x1 ¼ K1 [(1 K2 )=(K1 K2 )]

(3:9)

or

3.2.2.1

Application

For flashing a binary mixture, F ¼ feed, V ¼ vapor, L ¼ liquid, H ¼ enthalpy of vapor, and h ¼ enthalpy of the liquid. The mole fraction of the j fraction in the vapor phase is yj, the mole fraction in the liquid phase is xj, and the equilibrium coefficient for the jth component is Kj ¼ yj=xj. A mass balance of component j around the separator Figure 3.6 will be xFj F ¼ yj V þ xj L ¼ yj (F L) þ xj L

(3:10)

The mole fraction of the jth component in the liquid phase xj ¼ yj=Kj, therefore, Equation 3.10 becomes yj ¼

xF j 1 þ L=F {(1=Kj ) 1}

(3:11)

and the energy (enthalpy) balance is hF F ¼ hL þ HV ¼ hL þ H(F L)

(3:12)

The liquid fraction of the stream leaving the system using Equation 3.12 is L=F ¼ (H hF )=(H h)

(3:13)

Since the sum of the vapor phase mole fractions must equal unity X J

3.2.2.2

yj ¼ 1 ¼

X

xFj =[1 þ L=F {(1=Kj ) 1}]

(3:14)

J

Procedure

To determine the expanded state of a binary mixture .

. .

.

.

Choose a reasonable temperature between the bubble point and the dew point of the mixture in the separator. Calculate y1 from K1=x1. Choose from Table 3.3 and Table 3.4 the K1 and K2 values of the two components at the chosen temperature. Select from Table 3.3 and 3.4 the enthalpies of the feed and of the expanded components at the selected temperature. Calculate the value of L=F from Equation 3.13.


TABLE 3.3 Distribution Coefficient Functions, In (Kp =po), for Nitrogen, Oxygen, and Argon (kPa). The Value of the Reference Pressure is po 5 101.3 kPa (1 atm) Nitrogen

Oxygen

Argon

Temperature (K)

101.3

202.6

505.5

101.3

202.6

506.6

101.3

202.6

506.6

78 80 82 84 86 88 90 92 94 96 98 100 102 104 106 108

0.0798 0.3040 0.5282 0.7582 0.9766 1.2008 1.4249 — — — — — — — — —

— — — 0.7048 0.9030 1.1012 1.2994 1.4976 1.6958 1.8939 — — — — — —

— — — — — — — — 1.5503 1.7017 1.8531 2.0045 2.1559 2.3073 2.4588 2.6102

1.3368 1.1164 0.8959 0.6755 0.4550 0.2346 0.0141 — — — — — — — — —

— — — 0.4573 0.3016 0.1459 þ0.0098 0.1655 0.3211 0.4768 — — — — — —

— — — — — — — — 0.6605 0.7877 0.9148 1.0420 1.1692 1.2963 1.4235 1.5506

0.9075 0.7157 0.5238 03319 0.1400 0.0519 þ0.1400 — — — — — — — — —

— — — 0.2516 0.0717 þ0.1082 0.2880 0.4679 0.6477 0.8276 — — — — — —

— — — — — —. — — 0.5518 0.7323 0.8629 1.0434 1.2240 1.4045 1.5851 1.7656

Source: From Timmerhaus, K.D. in Cryogenic Process Engineering, Springer Science & Business Media, New York, 1989. With permission.

TABLE 3.4 Enthalpy of Nitrogen–Oxygen Mixtures. hL 5 Enthalpy of the Saturated Liquid at the Bubble Point; hV 5 Enthalpy of the Saturated Vapor at the Dew Point. 1 J=mol 5 1 kJ=kg mol 101.3 kPa (1 atm) Mole Fraction N2 0.00 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.90 1.00

506 kPa (5 atm)

Temperature (K)

Temperature (K)

Liquid

Vapor

hL (J=mol)

hV (J=mol)

Liquid

Vapor

hL (J=mol)

hV (J=mol)

90.2 87.7 85.7 84.1 82.5 81.3 80.4 79.6 78.8 78.1 77.4

90.2 89.5 88.7 87.7 86.7 85.6 84.3 83.1 81.5 79.7 77.4

419 461 519 599 682 779 879 984 1084 1181 1273

7252 7231 7210 7185 7151 7118 7084 7042 6992 6933 6871

108.9 106.3 104.2 102.5 100.9 99.7 98.5 97.4 96.5 95.6 94.2

108.9 107.9 106.7 105.6 104.3 103.0 101.5 100.0 98.4 96.6 94.2

1315 1403 1499 1591 1675 1758 1851 1947 2052 2152 2248

7536 7507 7478 7448 7415 7377 7339 7298 7252 7201 7147

Source: From Timmerhaus, K.D. in Cryogenic Process Engineering, Springer Science & Business Media, New York, 1989. With permission.


.

Using summation Equation 3.14 calculate the sum of all components and determine if this adds up to unity. If not, then choose another temperature for the separator. Once the summation equation at the selected temperature adds up to unity or to within the fourth decimal point, find the condition of expanded stream using Equation 3.13 and x2 ¼ (1x1) for liquid components. Use Equation 3.5 and y2 ¼ (1y1) for vapor components.

3.2.2.3

Practical Example (It Requires Iteration)

Assume that a saturated liquid mixture of air of 79% nitrogen and 21% oxygen at 506.6 kPa passes through a throttle (expansion) valve and reaches a pressure of 101.3 kPa. The resulting mixture will be as follows: At 101.3 kPa, the mole fraction of nitrogen will have a dew point of 81.8 K, and a bubble point of 78.8 K. From Table 3.3, the liquid ¼ 932 J=mol and the vapor ¼ 6943 J=mol. The feed stream at 506.6 kPa has hF ¼ 2042 J=mol. Using Equation 3.11 L=F ¼ (6943 2042)=(6943 932) ¼ 0.8153 From Table 3.4 at 80 K and 101.3 kPa K ¼ ln K1 p=po ¼ 0:3040 (nitrogen) K ¼ ln K2 p=po ¼ 1:1164 (oxygen) The distribution coefficients may not be determined: K1 ¼ 1:3553 (nitrogen) and K2 ¼ 0:3275 The right hand side of Equation 3.14 may be determined: 0:790={[1 þ (0:8153)][1=1:3553] 1} þ 0:210={[1 þ (0:8153)][(1=0:3275) 1]}1:0048 þ 0:0785 ¼ 1:0833 > 1 is therefore too large One must repeat the calculations using a lower temperature, say T ¼ 79.07 K With T ¼ 79.07 K, one arrives at h ¼ 1005 J=mol and H ¼ 6916 J=mol; L=F ¼ 0.8309; K1 ¼ 1.2211, and K2 ¼ 0.2956. With these values the result of the right-hand side of the Equation 3.14 is 0:9299 þ 0:0705 ¼ 1:00041 which is near enough to 1. The condition of the exhaust stream is therefore L=F ¼ 0:8309 x1 ¼ (1 K2 )=(K1 K2 ) ¼ (1 0:2956)=(1:2211 0:2956) ¼ 0:761 y1 ¼ K1 x1 ¼ 1:2211 0:761 ¼ 0:929 The mole fraction of oxygen is x2 ¼ 1 x1 ¼ 1 0:761 ¼ 0:239 (liquid) y2 ¼ 1 y1 ¼ 1 0:929 ¼ 0:071 (vapor)


3.2.3 FRACTIONATION3 Fractionation of a binary system with two components with close boiling points can be effected by the sequential cascading of condensation and evaporation as explained in flashing, but in one single column. To carry out this concept efficiently, however, it is necessary to ensure that the vapor is in very close contact, and approaches equilibrium with the remaining liquid at every stage of the process. In air separation, the liquid feed is passed in countercurrent with the ascending gaseous mixture. When in equilibrium, the vapor just above the liquid mixture has a higher nitrogen content due to nitrogen’s higher volatility. The vapor passing through the perforations in the trays and the liquid flowing over them tend to remain in equilibrium by exchanging oxygen and nitrogen in contact with each other. This process involves heat exchange, resulting in the condensation of oxygen and evaporation of nitrogen. The ascending gaseous mixture will steadily become richer in nitrogen, and the descending liquid richer in oxygen. From Figure 3.4, it can be seen that a vapor and a liquid in equilibrium always have different compositions. Conversely, if a liquid and a vapor have the same compositions, they cannot be in equilibrium. (In certain azeotropic systems y ¼ x, but this does not apply to air separation.) If one considers the condensation of a binary system where a liquid phase and a vapor phase are brought together in a vessel and mixed thoroughly (Figure 3.4), the two phases will reach a common temperature To. At this temperature the resulting liquid phase will be at its bubble point x, whereas the vapor phase will be at its dew point y. The values of x and y will be different, reflecting the relative volatility of the two components. At temperature To, the equilibrium values are x for the liquid phase and y for the vapor phase. The ratio of y to x is a constant (y=x ¼ k). It is this constant (at relative values of y and x) that results in the slightly parabolic equilibrium line shown in Figure 3.10. Returning to the vapor–liquid equilibrium diagram in Figure 3.4, and drawing a vertical line at any section will show that the vapor at its dew point is always at a higher temperature T1 than T2, the bubble point of the liquid. It can be stated, therefore, that if a liquid and a vapor having the same composition are brought into contact, the compositions of the liquid and the gaseous phase will move apart. The liquid will contain more, and the vapor less of the less-volatile component. This is the basic principle of fractionation. In order to take full advantage of this characteristic behavior of gases at or near their boiling points, it is important to maximize the contact between the vapor and the liquid phases.

3.2.4 FRACTIONATION METHODS A standard fractionation tower consists of a vertical cylindrical vessel that encloses a number of bubble caps or perforated trays or the more recently designed structured packing (see Section 3.3.3 and Section 3.3.4). The liquid feed, containing the components to be separated, is brought in around the middle of the tower and flows horizontally across each tray and in a downward direction until it reaches the bottom sump of the tower. See Figure 3.7a. It is important to maintain the level of the liquid feed at a quasiuniform level across each tray. The vapor that has been stripped from the liquid at the bottom of the tower sump traverses the plates through perforations, and continues upward until it reaches the top of the tower. As the vapor passes up through the thin layer of liquid on each plate, it mixes with the liquid and almost reaches equilibrium before passing upward to the next plate. The plates are so designed as to permit the liquid to pass across and downward from plate to plate. The perforations in the plates are sized to permit a sufficient flow of vapor to be in equilibrium with the quantity of liquid flowing across them. The vapor is supplied with sufficient upward velocity to prevent the liquid from leaking through the perforations.


Condenser

Heat out

High volatility Low boiling Vapor up

Product out

High volatility Lower boiling constituent

Z

(N2) out

Z9 Y

Mixture in X9

Y9 Feed mixture (air) in

X W

Liquid down

W9

Heat in Boiler (a)

Product out Lower volatility

(O2) out Lower volatility

More volatile component evaporates and joins vapor Liquid layer on plate

Less volatile component condenses and joins liquid (b)

FIGURE 3.7 (a) Typical fractionation column. (Courtesy of Bernstein, J., Cryogenic Consulting Services, Inc., 2006. With permission.) (b) Transformation in the composition of both liquid and vapor during their intermingling on a fractionation plate. As the vapor of the more volatile component traverses the liquid on the plate, heat and mass balance transfer occurs thereby a small portion of the more volatile component in the liquid evaporates and enriches the vapor bubble whereas a small portion of the less volatile component bubble condenses and joins the liquid. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

As bubbles of vapor traverse the film of liquid (Figure 3.7b) heat is exchanged between the two phases. The less volatile component in the vapor condenses into the liquid film on the trays, whereas the more volatile component in the liquid vaporizes and is swept upward in the vapor. Within practical limitations of pressure drop and power consumption, it is important to produce very small bubbles of vapor passing through the liquid to attain optimum conditions of heat transfer and equilibrium balance. If the rate of flow of vapor increases beyond the design point, it will impede the downward flow of the liquid; frothing will occur and the equilibrium balance will be destroyed. If, on the


contrary, the vapor flow is below the design point, the liquid will drop through the perforations and dump to the sump. It is important, therefore, to operate the fractionation tower within the design parameters of liquid flow and vapor rate so that a proper mixture of gas and liquid will exist, thereby attaining the correct equilibrium conditions between the two phases. In the sump of the upper (low pressure) tower, the liquid, which is mostly oxygen (the less volatile component), is heated to release nitrogen (the more volatile component). In airseparation plants this heating, generally called reboiling, is accomplished with heat exchange from nitrogen supplied from the lower (high pressure) column. The reboiling oxygen is at a pressure of approximately 1.2 bar and 91.9 K, and the nitrogen is at a pressure of around 5.7 barA and 95.8 K. This difference of 3.9 K is enough to condense the nitrogen and vaporize the liquid oxygen. This condition was normal with the use of vertical copper-tube vaporizer–condensers. Presently, with the use of brazed-aluminum heat exchangers for vaporizer–condensers, a temperature differential of 2 K and even slightly lower has been achieved. The ascending nitrogen vapor comes into intimate contact with a reflux of nitrogen-rich liquid fed down from the top of the tower to complete the enrichment of the vapor. The top section is called the rectification or enrichment section of the tower. The bottom section is called the stripping section. Both terms here apply only to nitrogen, which is the more volatile component.

3.2.5 FRACTIONATION PLATES3 In the past, the number of fractionation plates to be used was usually a trial-and-error proposition. At present, however, designers have several options to calculate the number of theoretical plates to be used for air fractionation, or any other binary system. The options include the analytical method involving plate-to-plate equilibrium calculations, the McCabe– Thiele graphical method, and the Ponchon–Savarit graphical method. The analytical method is obviously the most accurate. Unfortunately, it is also very tedious and requires a lengthy procedure. In addition, there was originally a scarcity of data. Design engineers therefore abandoned the analytical method, and opted for a graphical method, which is simpler and quicker to calculate. The original graphical method was developed by Ponchon and Savarit of France (1922). This procedure was fairly accurate using enthalpy– concentration diagrams. But the lack of proper enthalpy diagrams placed the Ponchon– Savarit method at a disadvantage compared with the McCabe–Thiele method (1925). The latter method uses constant liquid and vapor rates or constant molal overflow throughout the tower. Constant liquid and vapor rates are based on the assumption that the components of the feed have about equal molal heats of vaporization. This assumption is acceptable when the components have comparable molecular structures and similar boiling points. In both methods heat losses are considered negligible. Although the Ponchon–Savarit method may have proved a little more accurate, the McCabe–Thiele method was found easier to use and was almost universally adopted for many applications such as air separation and fractionation. However, it is not accurate enough because enthalpy change or correction is important. Today it has been almost completely replaced by the original plate-to-plate analytical calculations, which, with the aid of fast computers and more precise liquid–vapor equilibrium data, has been rendered quick as well as accurate. It is now possible to calculate in less than 1 min the number of theoretical plates for an air separation column. Moreover, this permits the designer to start with a fixed plate feed or to place the feed at a given plate where the composition is more compatible with that of the feed. Once given the number of theoretical plates it is possible to recalculate very quickly the number of actual plates by varying the efficiency to a more optimum level. Finally, the number of plates may be altered with input from experience with similar process applications. Nonetheless, it may be of interest to have an overview of the older method in order to visualize the parameters involved.


Prior to a discussion of the number of plates required for the fractionation of air some definitions are in order. First of all, a fractionation tower is rated according to the number of theoretical plates contained, as well as according to plate efficiency. A theoretical plate may be defined as one that produces an improvement in composition as existing at equilibrium between a liquid mixture leaving a tray and the vapor rising from the same tray. Plate efficiency is defined as the ratio in percent of the number of theoretical plates to the actual number of plates required for fractionation. Since ideal conditions for equilibrium cannot be achieved in practice, one has to assume a certain plate efficiency. If the vapor has an initial composition of yo, and has a composition of y1 as it leaves the plate, compared with an ideal composition of y2, then the plate efficiency is Eff ¼ (y1 yo )=(y2 yo )

(3:15)

In practice the design engineer uses a specific efficiency for a given tray design. This efficiency may vary from one type of tray to another and from system to system, but an overall tray efficiency of 60%–65% or better is not uncommon for air separation systems. Finally, there is an experimentally proven relationship between the feed state (whether it is a subcooled liquid and a saturated liquid, a mixture of saturated vapor and saturated liquid, a mixture of saturated vapor and superheated vapor) and its composition, the two product compositions, the quantity and quality of reflux, and the number of theoretical plates. Any one of them can be calculated if the others are known.

3.2.6

ANALYSIS

OF

FLOW

IN

EQUIPMENT

Referring to Figure 3.8, Section A will apply to the upper or fractionation section of the tower, and B will apply to the lower or stripping section. V represents vapor, L the liquid, D the upper vapor product, B the bottom liquid product, and x and y are components. 3.2.6.1

Case I: Analysis of a Low Pressure Column

Section A (Rectification) Overall Material Balance Total balance is written as V ¼LþD

(3:16)

Vy ¼ Lx þ Dxd

(3:17)

y ¼ (L=V )x þ (D=V )xd

(3:18)

Component balance is written as

and Equation 3.17 can be written as

Equation 3.18 is the equation of the upper operating line. It defines the relationship of y and x between plates.


Case I: Fractionation (above feed plate) Case II: Stripping (below feed plate)

C

D

F —Feed mol/h D —Vapor product mol/h B —Liquid product mol/h

A section rectification

C —Condensing system y = Vapor composition more volatile component V

x = Less volatile component

L

liquid composition

V = Vapor L = Liquid

F V

L

V

L

Feed

B section stripping

Heat source reboiling

B

FIGURE 3.8 Analysis of low-pressure column. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

Section B (Stripping section) Overall Material Balance where F ¼ Liquid Feed, L ¼ Liquid descending from Section A. Feed is in liquid phase. Therefore, liquid feed below feed plate is given by LþF

(3:19)

Vy ¼ (L þ F )x (Bxb )=V

(3:20)

y ¼ [(L þ F )=V ]x (B=V )xb

(3:21)

and component balance is

or


Equation 3.21 is the equation of the lower operating line. To determine the point of intersection, y has the same value for Equation 3.18 and Equation 3.21 (equating the laws of Raoult and Dalton). Therefore, (L=V )xi þ (D=V )xd ¼ [(L þ F )=V ]xi (B=V )xb

(3:22)

where xi is the value of x at the intersection. Solution of Equation 3.22 gives the value of xi at the intersection when y is the same. (L=V )xi þ (D=V )xd ¼ (L=V )xi þ (F=V )xi (3xb )=V

(3:23)

(D=V )xd ¼ (F=V )xi (Bxb )=V

(3:24)

xi ¼ (Dxd þ Bxb )=F

(3:25)

Fxf ¼ Dxd þ Bxb

(3:26)

or

or

Overall balance states that

Substituting the value of F from Equation 3.25 into Equation 3.24 xi ¼ xf

(3:27)

Hence the operating lines intersect at x ¼ xf. 3.2.6.2

Case II: Consideration of Vapor Feed

As in Case I, the upper operating line is Vy ¼ Lx þ Dxd

(3:28)

But in Case II, the vapor line is below the feed plate and the operating line is below the feed plate. So (V F )y ¼ Lx Bxb

(3:29)

For the point of intersection what is the value of y when x has the same value in Equation 3.28 and Equation 3.29? Let yi ¼ y at intersection so that xi ¼ (Vyi Dxd )=L ¼ [(V F )yi þ Bxb ]=L

(3:30)

Vyi Dxd ¼ Vyi Fyi þ Bxb

(3:31)

or


Hence Fyi ¼ Dxd þ Bxb

(3:32)

yi ¼ (Dxd þ Bxb )=F

(3:33)

Fyf ¼ Dxd þ Bxb

(3:34)

or

But overall balance is given by

Substitute value of F from Equation 3.34 into Equation 3.33 yi ¼ yf

(3:35)

So the operating lines intersect at yf. Note: The slope of the lower operating line is larger for a vapor feed than for a liquid feed when the same condensation duty is used (see Figure 3.9). The height of a packed column divided by the number of theoretical plates is referred to as the height equivalent to a theoretical plate, or simply the HETP of a distillation tower.

MC

D Nitrogen 99.99% MC = Main condenser B = Rich liquid mol/h D = Liquid nitrogen mol/h F = Feed mol/h V = Vapor V

L

L = Liquid

F (Feed) yf B 40% Oxygen

FIGURE 3.9 Analysis of high-pressure column, Case III. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


3.2.6.3

Case III: Analysis of the High-Pressure Column

The case for the high-pressure (lower) column (see Figure 3.9) may appear a little easier than for the upper column. It should be remembered that the feed may not always be in vapor form: at times the feed into the bottom of the tower may be part vapor and part liquid depending on the process cycle involved. In this case, however, the feed is considered to be all liquid. Therefore, the component balance is Fy ¼ Dx þ Bxb

(3:36)

Fyf ¼ Dxd þ Bx

(3:37)

which is the inter-plate balance.

which is the overall balance. The bottom point of the operating line is given by y ¼ yf and x ¼ xb, since the value of yf from Equation 3.36 and Equation 3.37 gives xb and yf as points on the general operating line. Operating pressure in the tower also plays an important role in the calculation of the number of trays. A lower operating pressure increases the relative volatility and it becomes easier to separate the components, which means fewer theoretical plates or, for the same number, a higher purity for the more volatile component. To determine graphically the number of theoretical plates required, it is first necessary to find out what quantity of nitrogen to be retained in the liquid oxygen sump of the tower as specified by the designer. It is also required to know the purity of the nitrogen to be removed as product at the top of the distillation tower. This data is applied on Figure 3.10. The mole fraction of nitrogen in the liquid phase in the sump of the tower is on the x-axis. The mole fraction of nitrogen in vapor phase at the top of the tower is on the y-axis. A straight diagonal line is the theoretical operating line. The theoretical operating line represents operation of the tower at total reflux, or zero tower output (or feed). The curve joining the operating line at its ends is the equilibrium line, and is somewhat parabolic. At the left of the equilibrium line is the composition of the saturated vapor leaving the (nþ1)th plate, and to the right of the equilibrium line is the composition of the liquid at its bubble point also leaving the (nþ1)th plate at the same temperature as that of the vapor on the other side of the equilibrium line. The slopes of the operating lines are given by Equation 3.18 and Equation 3.19 as are also the intersections of the operating lines with the diagonal. Once the operating lines are drawn, the theoretical plates can be included in the diagram. Beginning at the value of x ¼ 0.95 (95% of vapor nitrogen in the product at the top), a vertical line is drawn from the equilibrium line to the upper operating line. This is the first plate. From this point, a horizontal line is drawn from the operating line to the equilibrium line, and a step formation follows until the steps are past the intersection of the two operating lines, the upper and the lower. With regard to the lower stripping section the same procedure is followed, but beginning at the x value of 0.5, which is the nitrogen content in liquid phase in the sump of the column, proceeding upward. At the end it is found that the stripping section requires five plates and the rectification section requires about 10 plates. It is also of interest to note that as the purity of the components gets higher, the number of graphical steps becomes greater, with each step producing smaller increments in rectification.

3.3

PRACTICAL CONSIDERATIONS

Obviously, the more the number of plates used in fractionation, the higher the purity of the vapor product. For example in air separation for very high-purity nitrogen, it is not unusual


B

Condenser

on

0.8

ti fica

line

cti

Re

Mole fraction of N2 in vapor

I Feed line C

lin

e

Eq

uil

ibr ium

lin

e

0.6

e

pin

g

0.4

g

rip St

ra

pe

r eo

Th

0.2

0

o al

ic

et

A

lin

tin

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9 0.95%

Mole fraction of N2 in liquid Air separation product nitrogen 95% Product oxygen = 95%, N2 = 5% Stripping section requires 5 theoretical plates Rectification section requires 10 theoretical plates

FIGURE 3.10 McCabe–Thiele graphical analysis for theoretical plates. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

to install close to 90 plates in the upper column. The use of a high number of plates also carries with it the disadvantage of a higher pressure drop, which however means a higher power consumption.

3.3.1 BUBBLE-CAP TRAYS The original bubble-cap tray used many years ago had certain advantages. It was cheap to construct, offered a very good liquid phase level, and had a good liquid retention whenever there was a wide variation or stoppage in vapor flow. Unfortunately it had low efficiency, vapor diffusion in the liquid was poor, and the cost became prohibitive with the increase in air-plant capacity (Figure 3.11).


Liquid flow

Vapor flow

Downcomer

Bubble-cap plate

FIGURE 3.11 Rectification column plates (or trays). (Courtesy of Bernstein, J., Cryogenic Consulting Services, Inc., 2006. With permission.)

3.3.2

SIEVE TRAYS

The bubble cap soon gave way to the more efficient perforated sieve tray fabricated with aluminum. It was more adaptable as well as more economical for high capacity plants—over 300 t=d of oxygen. The variety of designs for sieve trays are too numerous to describe. They have been used for oxygen plants with a capacity of over 2500 t=d. Even these trays, however, were rendered inefficient and costly for oxygen plant capacities over 1000 t=d. With a tray diameter of 4–5 m it became difficult and costly to construct a sieve tray that could hold a liquid phase absolutely level. In other words it was difficult to maintain a good equilibrium between the vapor and the liquid phase. The individual cost of such trays was anywhere from $2000 to $3000 per tray, and often higher (see Figure 3.12). Sieve trays with a diameter of up to 5 m have been installed in very large air separation plants.

3.3.3

STRUCTURED PACKINGS

In the recent past, however, designers have turned to the use of structured packing for large air-separation plants especially when high power cost was involved. The technology of packed towers is not new. (The use of random packing consisting of ceramic or plastic pall rings has been around for decades.) Structured packing has been used in refineries and chemical plants for a good number of years. It is only since 1980, however, that the industrial gas industry Liquid flow

Vapor flow

Downcomer

FIGURE 3.12 Perforated plate. (Courtesy of Bernstein, J., Cryogenic Consulting Services, Inc., 2006. With permission.)


has ventured to use structured packing for both lower (high pressure) and upper (low pressure) distillation columns in air-separation plants. The main objective has been to lower pressure drop across the columns to reduce overall energy consumption. At the same time, a well-designed, structured packing will also improve the mixing of the liquid and vapor phases for better rectification. The results from the field since 1990 have been encouraging. Pressure drops have been lowered dramatically and rectification has improved impressively. The technology of structured packing is quite simple and fairly well known. Thin strips of aluminum or copper of a predetermined width are perforated, and then corrugated. The strips may be assembled either in a parallel manner, or coiled in a circular configuration. In either case, their overall geometry resembles a tray. Finally, these trays are installed in the columns either in the shop or at the field. Designers have not yet agreed as to which type of assembly is better. The liquid reflux is fed from the top of the tower and distributed uniformly to the trays below by specially designed distributors. As packed towers are unusually tall, the designer has to consider the use of a number of distributors, as many as five in the lowpressure column to remix the liquid and to correct any tendency to develop poor uniformity of distribution. The functioning of the structured packing tray is similar to that of a reflux condenser, also known as a ‘‘dephlegmator’’ designed by Claude in the 1920s. The reflux condenser is still in use today especially for the recovery of crude argon. Refer to Figure 3.17 in Section 3.7.6. As a film of liquid reflux descends the vertical surface of the tray, it comes into intimate contact with the rising vapor and reaches equilibrium with it. As with a standard sieve tray, heat exchange occurs between the two phases. The more volatile component (nitrogen) in the liquid vaporizes and joins the ascending vapor, whereas the less volatile component (oxygen), entrained by the rising vapor, condenses and joins the descending liquid toward the sump. This structured packing design permits the production of a very thin film of descending liquid in intimate contact at all times with a continuous flow of minute vapor bubbles throughout the entire packing, with a minimal pressure drop. In fact, assuming an ideal design, if one were to take a vertical slice across the tray, one will find a fairly constant thickness of the film and a constant vapor-flow throughout the entire volume, which is the ultimate objective of the rectification process. The height of a packed tower divided by the number of theoretical plates is referred to as the height equivalent to a theoretical plate, or simply as the HETP of a distillation tower.

3.3.4 CARE

IN THE

DESIGN

OF

STRUCTURED PACKING

The use of structured packing, however, is not without certain disadvantages. The design imposes certain difficult demands: surfaces of the packing must be completely wet at all times, with a uniform liquid distribution throughout the entire ‘‘tray.’’ The packing has to be tightly fitted against the side wall of the tower to prevent leakage, and distributors have to be designed accurately and located at the right elevations in the tower. Moreover, it has to be kept in mind that the overall density of the packing per volume of tower is roughly 3.75 times that of an equivalent standard tray, which results in a heavier as well as a taller tower. In short, the use of structured packing, at least for the present, may be limited to air separation plants of a large capacity, 300 t=d and over, and in situations where power costs are much above average.

3.3.5 SAFETY IN

THE

USE OF STRUCTURED PACKING

Though it has been reported to the industry that the use of structured packing made from aluminum may lead to combustion, especially in the lower section of the upper column of an air separation plant; this has been examined in detail by the industry and has been found untrue. Details on safety can be found in Chapter 12.


OPERATIONAL CONTROL4

3.4

Normally, there are three operating control valves (see Figure 3.13) that are used in the operation of a standard air separation plant.

3.4.1

FCV-1

FCV-1 controls the flow of the liquid nitrogen reflux that is sent from the (lower) high-pressure column to the upper tray of the (upper) low-pressure column. By throttling the valve, the flow is reduced and the content of oxygen is also reduced. This action allows more liquid nitrogen reflux to remain in the high-pressure column reducing entrainment of oxygen. However, if the throttling is overdone, reducing the quantity of reflux available to the upper tray in the upper low-pressure column, it reduces the recovery of the oxygen product. If the valve is opened more than enough, the quality of reflux available to the upper column is excessive, and recovery is reduced because oxygen is entrained by the increase of waste nitrogen. Reflux 450 N2 @ 1% O2 FCV-1 Controls flow of reflux from HP to LP Maintains material balance

800 Waste N2

200 Pure O2

Lower LP column

Reboiler

FCV-3 Controls flow O2 Fine tunes balance

Upper LP column the condenser of the HP column also acts as the reboiler of the column

Condenser

1000 550 FCV-2 Maintains level in HP sump Prevents vapor transfer from HP to LP Lower temperature of RL

1000 mol air

Flow = 1000 – 450 = 550 mol Rich liquid @ 37.36% O2

O2 in RL = 0.21 x 1000 − 0.01 x 450 1000 – 450 = 37.36%

FIGURE 3.13 Diagrammatic view of the distillation of air (operational control). (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


3.4.2 FCV-2 FCV-2 maintains a constant level in the high-pressure column by throttling the rich liquid from the high-pressure column down to the operating pressure of the low-pressure column. This valve may also be of the ‘‘smart’’ type, with a programmed chip insert. This valve once set, seldom requires any further resetting.

3.4.3 FCV-3 FCV-3 controls the flow of the gaseous oxygen product leaving the upper column. This valve has to be monitored carefully so as not to prejudice the designed operation of the lowpressure column. An excess flow will lower the purity of the oxygen product. A decrease in the flow will increase purity but will also lower recovery of the oxygen product.

3.4.4 REFRIGERATION The quantity of refrigeration must also be monitored closely. Any excess cooling supplied to the plant will manifest itself in increasing the liquid level in the main condenser. It will then be necessary to take out some oxygen product as liquid, and reduce the flow of gaseous oxygen. Alternatively, one might extract part of the liquid nitrogen reflux out of the plant as an additional product.

3.5 PRODUCT RECOVERY The extraction efficiency of an air separation plant is based on its main product recovery. In terms of oxygen product, it may be described as h, the ratio of usable oxygen produced divided by the oxygen in the process air treated by the plant. The above definition can be verified very quickly by checking the quantity of oxygen entrained by the waste nitrogen leaving the low-pressure (upper) column. This observation will quickly verify if the plant is operating within the designed optimum parameters. In order to increase the oxygen-product recovery, one must increase the reflux at the top of the upper column, but this depends on the following: (a) The quantity and quality of liquid nitrogen reflux sent from the lower (high pressure) column (b) The quantity and quality of the vapor rising to the top tray of the upper (low pressure) column (c) The vapor rising to the top tray must always be close to equilibrium with the liquid at the top tray As can be seen from the examples shown diagrammatically in Figure 3.13, the recovery of oxygen product can be increased by throttling the reflux valve slightly and increasing the quality of the reflux liquid, but this procedure has a limit. Although it improves the quality of the reflux by lowering the oxygen content, it also lessens the quantity of reflux. If it is overdone it may decrease recovery. The operator has to be very careful in fine-tuning the recovery (Figure 3.14).

3.6 OPTIMUM REFLUX In reviewing a material balance of the lower (high pressure) column of an air-separation plant, it will be noted that the rich liquid accumulating at the bottom is about 550 mol per 1000 mol of the air feed. Since the total quantity of oxygen is to be recovered, the oxygen


Case I: Basic operative case 800 KX%

[email protected]% O2

[email protected]% O2 [email protected]% O2

In 400 • 0.015 800 • 0.019

Out 400 • X 800 • KX

6 + 15.2 = 800 X X = 0.0265 % O2 in WM2 = 0.0265 K = 0.01325 = 1.325% O2

Case II: Reduce reflux to improve quality and recovery of product Result: Recovery of O2 has improved 800 KX %

800@2% O2

350@1% O2

350@X% O2

In 350 • 0.01 800 • 0.02

Out 350 • X 800 • KX, K = 0.5

3.5 + 16 = 750 X X = 0.0260 % O2 in WM2 = 2.60 K, K = 0.5 = 1.30% O2

Case III: Reduce reflux still to improve quality Result: Recovery of O2 has decreased 800 KX %

300@5% O2

[email protected]% O2 300@X% O2

In 300 • 0.005 800 • 0.025

Out 300 • X 800 • KX, K = 0.5

1.5 + 20 = 700 X X = 3.07 % O2 in WM2 = K(3.07) K = 0.5 = 1.54%

FIGURE 3.14 Relationship of quantity and quality of reflux on product oxygen recovery. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

content is about 210 mol or 38.2%, the balance being 8.9 mol or about 1.62% of argon, and 331.1 mol or 60.2% of nitrogen. This quantity of 550 mol is withdrawn, subcooled, and sent to the middle of the upper column as feed. The remaining 465 mol are transformed into pure liquid nitrogen in heat exchange with the main condenser and sent to the top of the fractionation tower as liquid reflux. In the early 1920s, it was established that the 465 mol of reflux for every 1000 mol of process air was inefficient, because to produce oxygen even with a purity of 99.5%, only 200 mol of nitrogen reflux was required in theory. The balance 265 mol could be put to better use for the production of additional refrigeration. In actual practice, the nonessential nitrogen reflux was between 150 and 200 mol. There are several ways of exploiting the available excess nitrogen. One is to remove a portion of the uncondensed nitrogen from the high-pressure column, to recover part of its refrigeration capacity in a heat exchanger within the cold box, and then send it to the expansion machine for the generation of refrigeration. This refrigeration is recovered within the process system by sending the cold nitrogen through the primary exchangers in countercurrent with the incoming process air feed. The second alternative is to take a portion of the process air feed and send this stream directly to the upper (low pressure) column after first passing it through the expansion machine in order to reach the same operating temperature and pressure of the upper (low pressure) column. This latter procedure has been known as the Lachmann stream named after the developer.


TABLE 3.5 Argon Content in Oxygen Product Oxygen

Argon

Nitrogen

95% 98% 99.5%

3.65% 2.00% 0.50%

1.35% Less than 0.10% Traces

Source: From Kerry, F.G., 2005. With permission.

The production of liquid nitrogen is a costly process in terms of energy because of its boiling point of 77.3 K at 101.3 kPa. As the purity of the oxygen product increases, so does the requirement for liquid nitrogen reflux. This increase in liquid nitrogen reflux is to remove argon from the oxygen product. Argon has a boiling point of 87.3 K, which is only 2.9 K different from that of oxygen. Therefore, to increase the purity of oxygen from 95% to 99.5% only the removal of argon enters the process of fractionation. This requires a relatively important quantity of expensive refrigeration. Table 3.5 may be of interest to the reader. Because most pyrometallurgical and chemical applications can operate economically with an oxygen purity of only 95%, it is not feasible to use a standard Claude cycle for the production of high-purity oxygen. For low-purity oxygen it is essential to come up with low investment and low energy consumption. For argon removal as well, there are some rigid specifications for many commercial combustion applications concerning the quantity of nitrogen in the oxygen product in various states in the United States. For example, in California the quantity of nitrogen in the oxygen to be used for combustion must not exceed 0.10% in order to keep the amount of oxides of nitrogen to a strict minimum in the effluent gas. To meet this standard the oxygen product must be kept at a minimum purity of 98.0% (Figure 3.15). Also, there exist specifications for the production of certain steels that limit any introduction of nitrogen in the smelting process to a hard-to-obtain minimum.

3.7 DISTILLATION EQUIPMENT The basic elements of a double-distillation system as it is known today are as follows:

3.7.1 UPPER (LOW PRESSURE) COLUMN The technology and theory of this has previously been described in Section 3.3.3.

3.7.2 LOWER (HIGH PRESSURE) COLUMN The importance of the lower column is probably the most misunderstood item by many practicing engineers. In fact, this column is particularly important because of the multiple duties it has to perform. Its design encompasses the following parameters: a. It must be designed for the maximum recovery of both oxygen and argon. b. The liquid nitrogen product condensed for the top of the column must be of the highest purity so that it may be used as reflux for the upper (low pressure) column, or as a marketable product. c. The amount of gaseous nitrogen to be liquefied in heat exchange with liquid oxygen in the main condenser must be sufficient to condense the nitrogen used as reflux, and to return 100% of the oxygen as well as argon to the sump of the tower.


Oxygen nitrogen separation Theoretical calculation

5.0 4.0 3.0 2.0 1.0

Percentage of nitrogen in oxygen

Theoretical trays: Lower column 13; 19 practical Upper column 20; 28 practical

91%

92%

93%

94%

95%

96%

97%

98%

FIGURE 3.15 Oxygen–nitrogen separation—theoretical calculation. Theoratically the oxygen purity must exceed 96.5% to contain 0.10% nitrogen, the oxygen purity must reach close to 98% to attain less than 0.1% or less than 1000 ppm of nitrogen. The balance is argon. (ß Air Liquide, all rights reserved, 2006. With permission.)

d. It must be designed with an operating pressure low enough to save power, but high enough to drive the liquid nitrogen hydraulically to the top of the upper (low pressure) column for reflux. Actually, the pressure adjusts itself automatically. It keeps rising until the dew point is warm enough so that all the nitrogen gets liquefied. e. Sufficient distillation trays must be included in order to effect a 100% recovery of both the oxygen and argon products as stated above. f. The top of the high-pressure column should include a specially designed shelf to retain a portion of the high-purity liquid-nitrogen being condensed by the main condenser. It is this liquid nitrogen that is collected and used as reflux at the upper (low pressure) column.

3.7.3

MAIN CONDENSER

The main condenser, which has been called by many names such as ‘‘main vaporizer,’’ ‘‘vaporizer–condenser,’’ ‘‘reboiler,’’ etc., should rightfully be called simply ‘‘main condenser’’;


because the direct purpose of the unit is to condense the incoming process air feed. The vaporization or reboiling of the liquid oxygen within the unit is the indirect result of condensation. This element is nothing more than a large storage receptacle of liquid oxygen descending from the upper fractionation tower. Regardless of its simple function, this element imposes very important design criteria for heat exchange and safety. The design should create conditions that ensure all internal heat exchange surfaces be wet with liquid oxygen at all times. If any dry spots occur, deposits of dangerous hydrocarbons such as acetylene, propane, and ethane will accumulate and may cause a catastrophic detonation. To avoid such potential accidents, it is important that a liquid oxygen recirculating system be created, either by the use of a liquid-oxygen pump, or by a thermosyphon. In either case the recirculating rate should be in excess of 100% of the evaporation rate. Some designs may not require a high recirculating rate, especially in the production and the extraction of large quantities of liquid oxygen, but such designs should be examined carefully before acceptance. Another requirement for the design of the main condenser is a high rate of heat transfer. The heat exchange between the liquid oxygen product on one side, and the incoming process air feed on the other can be affected by (a) High process-air-pressure in the (lower) high-pressure column, which means a high power-consumption at the main air-compressor (b) High heat-exchange-surface at the main condenser, which means high fabrication-cost for the unit (c) Use of a material with high heat-conductivity in fabricating the main condenser, which again means high investment cost (d) An optimum combination of all the above options Various designs have been tried out over the past several decades, but most air-separation plant designers have fairly well standardized on the use of extended surface aluminum exchanger cores that are available on a worldwide basis. These cores are used in a variety of combinations and packaged modules in order to provide a maximum of heat transfer with minimum capital cost. A differential temperature of 1.8 K has been recorded and is in operation since 1990. A recent report indicated that during heat transfer between the liquid-oxygen product and the incoming process in a main condenser the temperature differential was lower than 1 K, but this has not been confirmed officially.

3.7.4 LIQUID SUBCOOLERS Plants producing high-purity oxygen on a small scale are generally not equipped with liquid subcoolers. The availability of liquid reflux is large enough to avoid their use. In large air-separation units, however, because a fairly high quantity of gaseous nitrogen is withdrawn from the lower (high pressure) column and expanded in a machine to provide the required refrigeration capacity; an efficient liquid subcooler is a must for saving energy. Subcooling liquids prior to expansion through a J–T throttle valve avoids vaporization (mainly of nitrogen) and provides an increase of liquid reflux by approximately 8%.

3.7.5 PROCESS CONSIDERATIONS When the process air enters the (lower) high-pressure column and ascends to the top of the column, the main condenser condenses it at about 99 K. A portion of the liquefied air rich in nitrogen boils off, rises to the top, comes into contact with the main condenser that is at a temperature of 90 K or higher, recondenses, and descends once again, serving as liquid reflux


for the ascending gaseous air flow. The oxygen and argon portions being less volatile are recondensed completely and accumulate in the sump of the tower. A portion of the liquefied nitrogen is used as reflux to make sure that all of the oxygen and argon are completely recovered. The liquid thus accumulated in the sump consists of nitrogen, oxygen, and argon, and is generally called ‘‘rich liquid’’ (liquid rich in oxygen). Care should be taken, however, that the oxygen content of this rich liquid be carefully calculated to arrive at an optimum balance within the designated operating pressure of the high-pressure column, and the equilibrium of the various components at that pressure. Assume that 1000 mol of air are divided into 550 mol of rich liquid with an oxygen content of approximately 40% and 450 mol of pure liquid nitrogen, which is retained on a shelf and serves as reflux after undergoing expansion to slightly above atmospheric pressure. This expansion across a throttle valve helps drop the temperature of the liquid to about 79 K (Figure 3.16). With regard to the heat-transfer capacity between the main condenser containing liquid oxygen and the ascending process air from the lower column, the following heat balance can be made: Assuming that 1 mol of air enters the lower (high pressure) column, 0.79 mol 6871 J/mol

0.21 mol 7252 J/mol

Unbalance differenial

5222.35 J/mol

0.45 2248 J/mol

1.0 mol 7252 J/mol

0.55 mol 1858 J/mol

FIGURE 3.16 Incoming process air ¼ 1 mol @ 7252 J=mol @ 506.67 kPa (5.06 bar) producing GOX þ GAN @ 6051 J=mol @ 101.3 kPa (1 bar) with an unbalance of 301 J=mol. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


Enthalpy of air at 506 kPa ¼ 7252 J=mol Enthalpy of rich liquid at 40% oxygen ¼ 1851 J=mol Enthalpy of liquid nitrogen ¼ 2248 J=mol 1 mol of air enters the lower column with 7252 J=mol and leaves with (1851 0.550) þ (2248 0.450) ¼ 2095.65 J=mol Therefore, the quantity of refrigeration being furnished by the liquid oxygen in the condenser to the lower column is equal to 72522095.65 ¼ 5222.35 J=mol, which is equal to the vaporization of liquid in the upper (low pressure) column. This energy is capable of vaporizing 5222.35=6830 (6830 J is the latent heat of vaporization per mol at 101.3 kPa) or 0.765 mol of gaseous oxygen of which 0.21 is sent to the primary- or front-end heat-exchangers as product, and the balance of 0.555% ascends the upper column for rectification purposes. The rich liquid at the sump of the lower (high pressure) column is withdrawn, subcooled, and sent to about the middle of the upper column for final rectification. As for the upper column, an enthalpy balance can be made for the 0.21 mol of oxygen and the 0.79 mol of nitrogen leaving the column: 0:21 7252 þ 0:79 6871 ¼ 6951 J=mol

(3:38)

This figure of 6951 J=mol is slightly below that of the incoming air feed, which has a total enthalpy of 7252 J. This imbalance of about 4% is made up by subcooling the rich liquid leaving the lower column, as well as cooling the liquid nitrogen reflux going to the top of the low pressure column, using outgoing gaseous waste nitrogen available from the top of the upper (low pressure) column as a refrigerant.

3.7.6 CRUDE-ARGON SEPARATION COLUMN Figure 3.17 shows the overall material balance treating 1000 mol of air including the separation of crude argon from the main distillation column.

3.8 DEVELOPMENT OF LOW OXYGEN-PURITY PROCESSES5 As noted, pyrometallurgical and many chemical operations did not require pure oxygen, and for many years air-separation designers strove to find an economically optimum solution for these industries. As plants grew larger in scale, however, so did the problem. Designers had to optimize the overall investment and operating costs for prepurification, cryogenic equipment, process air and product compression, as well as field costs for erection. Steel mills needed low-purity oxygen for smelting operations, but high-purity oxygen for oxy-cutting scrap and shape cutting for finished products. The product pressure for smelting had to be at 30 barG, and specified low nitrogen-content in the product oxygen to reduce nitriding of steel. Copper smelters needed only low-purity oxygen for smelting ore, but were willing to pay a premium for a product pressure of 3–4 barG. Syngas producers allowed a low purity for oxygen product, but at a pressure of 50 barG minimum. Since 1980, the increasing cost of oxygen compressors as well as the cost of field labor did not facilitate the problem. It is impossible to enumerate individually the variety of process cycles developed by imaginative process engineers, but a few merit some description. First of all, the question of oxygen purity has been quantified by industrial practice. For example, the steel industry smelts steel using the LD process with oxygen at 40 barG. Furthermore, although it can use oxygen with a low purity, the oxygen should be almost


Nitrogen products N2 —780.76 O2— 2.04 Ar—1.30 784.10 mol

Nitrogen reflux N2 — 432.4 O2 — ppm Ar — ppm 432 mol Crude argon product N2 — 0.24 O2 — 0.16 Ar — 7 . 5 0 ( 9 5 % ) 7.90 mol

Oxygen product ( 99.7%) O2 — 207.4 N2 — ppm Ar — 0.6 208.0 mol Expansion turbine N2 — 84.5 O2 — 35.0 Ar — 0.5 120.0 mol Rich liquid Process air N2 — 781.0 O2 — 209.6 Ar — 9 . 4

N2 — 264.5 O2 — 174.6 (3.9%) Ar — 8 . 9 448.0 mol

1000.0 mol Approximate material balance Conventional air separation plant Nitrogen Product(s) Recovery—9.96% Oxygen Product Recovery—98.95% Argon Product Recovery—79.78%

FIGURE 3.17 Approximate material balance—conventional air separation plant. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

nitrogen free, which means that the minimum purity should be 98%, the balance being argon, as per Table 3.1. The copper industry can use a purity as low as 95% but at 3–4 barG.

3.8.1

THE LACHMANN PRINCIPLE (FIGURE 3.18)

As far back as 1902, Lachmann of Germany indicated that since impure oxygen would be more acceptable and economical to major industrial users, there was no point in wasting costly refrigeration to liquefy the total gaseous nitrogen for reflux necessary to produce high-purity oxygen. This concept was put into use in 1920 for plants producing liquid products that require extra refrigeration. The analysis of the Lachmann principle is simple enough. If y is the rising vapor and x is the descending liquid, therefore y ¼ 0:14 and x ¼ 0:40 (rich liquid) y ¼ x þ 0:315 (the nitrogen contained in the rich liquid)


Waste nitrogen Turbine outage WP nitrogen HP air Gaseous oxygen MP air Main exchangers

Liquid oxygen to storage Liquid nitrogen to storage

LP column

FC

FC

Main vaporizer

Expansion turbines 2 x 100%

FC

HP column FC FC

LOX circulation pumps 2 x 100%

LOX filter LOX circulation pumps 2 x 100%

FIGURE 3.18 Use of nitrogen from the HP column to improve product recovery. (From Gomonet, E., Les Basses Temperatures et son Emplois, Librairie, J.B. Baillere et Fis, Paris, France, 1952. With permission.)

Hence the possible reflux in the upper (low pressure) column could be 0.17 instead of 0.47, which is required for high-purity oxygen. Therefore 30% of the nitrogen is available for liquefaction at 5 bar and could be used to supply extra refrigeration for other uses in the process cycle. For normal gaseous oxygen plants, this principle is carried out by taking 20%– 25% of the process air, passing it through the expansion machines at almost its dew point and inserting it into the upper (low pressure) column at a point just below the entrance of the rich liquid. This operation produces a low-purity product and saves energy.

3.8.2 THE OXYTON PROCESS5 In 1950, a paper was presented at the Physical Society in London introducing a new air separation process named Oxyton. It departed from the standard Claude cycle, because some effort was paid to the irreversibility of the rectification process of the lower (high pressure) column. The rectification was carried out at three pressure stages: 5.5, 2.5, and 1.3 barA. It included a third medium-pressure column (see Figure 3.19). This column received rich liquid from the lower (high pressure) column. After passing though a filter, the rich liquid was subcooled and expanded through a throttle valve before entering the middle tower. From there it flowed into an intermediate condenser–vaporizer built in the lower section of the high-pressure column providing refrigeration at 96 K, therefore offering a less irreversible separation to the rectification. The energy saved in the operation of the lower (high pressure) column was used at the middle-pressure tower to produce additional liquid nitrogen from the rich liquid at the top of the lower (high pressure) column. Thus, by improving the rectification of the latter tower, a greater quantity of nitrogen was withdrawn and utilized to provide additional refrigeration to the process. The power saved for the production of oxygen at 95% purity was significant for that era: 298 kwh=t. A plant with an oxygen-producing capacity of 300 t=d at 95% purity was designed and erected in northern Canada.


94 K

% Nitrogen in the vapor

100 95

Claude HP column

90 85 80 Feed 75 50

99 K

60

80 70 % Nitrogen in the liquid

90

94 K

100

% Nitrogen in the vapor

100

95

Oxyton HP column

90

96 K

85 80

Feed

75 50

99 K

60

70 80 % Nitrogen in the liquid

90

100

FIGURE 3.19 Comparison between the Claude and Oxyton high-pressure columns. (ß Air Liquide, all rights reserved, 2006. With permission.)

3.8.2.1

Thermodynamic Analysis of the Oxyton Cycle

In the stripping section of the rectifying tower, the reflux ratio l is the ratio of molal reflux R and molal vapor flow V, or l ¼ R=V

(3:39)

The above equation is based on the assumption that the heavier product must be stripped down in the liquid in equilibrium, or Rx1 ¼ V y1

(3:40)

or lm ¼ y1=x1 (minimum reflux) Assume that F is the molal flow of a pure more-volatile component produced at the top of the tower; it can be said that V R ¼ F, therefore


R(1=l 1) ¼ F

(3:41)

R ¼ [l=(1 l)] ¼ F

(3:42)

or

It can be said that the molal quantity of minimum reflux to obtain 1 mol of pure more-volatile product from a vapor containing y1% of the less volatile component may be expressed by the relation l=(1 l). If one assumes two separate stripping sections each operating at 5 barA with an infinite number of plates, and with a process air feed at its dew point of 99 K, the minimum reflux for the classical Claude and the Oxyton processes would be rm 99K ¼ lm=(1 lm )

(3:43)

where lm is the minimum reflux ratio ¼ y1=x1 ¼ 21=44 ¼ 0.447, resulting in rm ¼ 0:477=(1 0:477) ¼ 0:91 mol

(3:44)

In other words, for 1 mol of gaseous nitrogen extracted from the top of the fractionation column, one must obtain the above minimum quantity of liquid reflux at the feed. This quantity is produced by the main condenser–vaporizer at the nitrogen liquefaction temperature, which is 94 K at 5 barA. A less irreversible separation can be obtained, however, if an intermediate condenser operates at say 96 K. For example, lm 96 K ¼ y1=x1 ¼ 12:5=30 ¼ 0:417

(3:45)

rm 96 K ¼ lm 96 K=(1 lm 96 K) ¼ 0:715

(3:46)

which is lower than 0.91 mol of reflux necessary for the standard Claude cycle. The overall saving is 0.910.715 ¼ 0.195 mol, which is available for other purposes. The objective of this process was to narrow the area between the operating line and the equilibrium curve for the lower (high-pressure) column (see Figure 3.19). In this manner, the process was made more reversible. Although judged primitive when compared with modern systems, this cycle became the forerunner of other developments to reduce energy consumption and capital investment. Other more recent cycle developments using the same principle of a two-stage main condenser duty have evolved in the past few years. All of them, however, required extra cryogenic equipment and the inevitable increase in investment. With the expanding market for integrated coal gasification and nonferrous smelting applications, one cannot ignore or discount new process cycles that offer lower energy consumption combined with an attractive investment cost. 3.8.2.2 Oxyton Development In 1992, 52 years after the development of the original Oxyton, a large gas company in the United States sent out a request for a plant to generate 1989 t=d of contained oxygen in a 95% mixture, but at a pressure of 40 barA, and 85 t=d of pure nitrogen at 38 barA. The process engineers of Air Liquide remembered the original plant that had not been too successful in the market. They dug up the process drawings and went to work in improving the thermodynamic efficiency using contemporary ideas and equipment. The reversing regenerators for prepurification were replaced by molecular sieve adsorbers. Primary exchangers were


replaced by nonreversing aluminum exchangers, and the third column, with its separate cooler, was replaced by a lean subcooler and a mixing column. The results of these modifications were spectacular. The oxygen product came out of the mixing column at the exact quantity, analysis, and at 4–5 barA, which eliminated the first casing of the oxygen compressor, thus saving a great deal of energy. The oxygen product generated was produced at the outlet of the process cycle almost at the same pressure as that of the inlet air. The nitrogen product came out as liquid nitrogen, which was pumped to the required pressure and vaporized in the main heat exchangers. No nitrogen compressor was needed, saving both capital, and operation and maintenance costs. Extra refrigeration capacity was obtained by taking 7% of the main air-process stream and upgrading its pressure using the expansion turbine as the driver. No extra booster air-compressor was required. If one applies only the consumption of energy by the main air-compressor to the process separation, the result is approximately 225 kw=t of product oxygen. This is very low by any standard.

3.8.3

VARIABLE-LOAD PLANTS

When direct oxygen-injection process of steel making took effect in the early 1950s, it reduced the time of smelting a heat of 100–200 t down to less than 1 h. The oxygen plant designers had to formulate an oxygen supply system operating on an on–off basis. At first this was achieved by designing a large steel tank operating at 30 barG with a system of piping and valves controlled automatically to inject oxygen at the prescribed pressure and timing as the vessels came into line. This system was acceptable but costly. Finally, a new air-separation process was developed, which did away with the high-pressure storage tanks. This resulted in a welcome saving for steel mills and for pyrometallurgical industries operating on a batch system. For maximum operating efficiency, air-separation plants have traditionally been designed to supply products at a constant rate. To accommodate a variable demand, a new process was designed achieving the same result within the cryogenic system, but eliminating costly external high-pressure tanks and ancillary equipment. This process has been in successful operation since 1975 not only for batch production (as in metallurgy), but also for continuously variable demand over 24 h periods (as in waste-water treatment). These processes operate between 30% and 100% of peak load demand. To make the unit cost-efficient, the cold box equipment is designed with larger inflow capacity than the upper limits of the process air compressors. Several designs are available on the market to achieve this variation in demand. Two, however, have a basic principle in common. Because the process cycles must vaporize an above-average demand, they include an extra condenser–vaporizer capacity as an additional piece of equipment, and the cycles include buffer (accumulator) tanks operating at cryogenic temperatures to store and supply extra liquid-oxygen product when needed. Moreover, they also include buffer (accumulator) tanks to store and supply either liquid nitrogen or liquefied air to the process when needed to maintain the refrigeration balance of the overall separation system. Incidentally, regardless of their basic small differences in design, a variable load oxygen process cycle can operate as described below in version A, or as in version B. 3.8.3.1

Version A

When oxygen demand is near the upper limit in one process cycle, liquid oxygen is pumped from the Liquid Oxygen (LOX) buffer tank to the sump of the upper (low pressure) column and vaporized in the main condenser–vaporizer. Meanwhile, the expansion machine will operate with a lower flow of nitrogen. The increased liquid oxygen in the sump will be vaporized by the additional nitrogen gas entering the main condenser–vaporizer from the


lower (high pressure) column. This will increase Liquid Nitrogen (LIN) production, which will be sent to the LIN buffer tank for storage. When oxygen demand decreases LIN is pumped from the LIN buffer tank to the upper (low pressure) reflux plate, which will increase expansion machine refrigeration capacity, which in turn decreases vaporization capacity of the condenser– vaporizer. Excess LOX thus produced will be sent to the LOX buffer tank for storage. In other words, when the supplementary LOX demand is sent from the buffer tank to the condenser– vaporizer, the refrigeration from the condenser–vaporizer is used to liquefy extra gaseous nitrogen from the lower (high pressure) column, and when LOX demand is low, liquid nitrogen from the LIN buffer tank is sent to the cold box to maintain the designed refrigeration capacity of the process. During a 24 h period there is no gain or loss in liquid products. The only problem in this version is the fact that the buffer tanks are situated outside of the main cold box, thus requiring extra insulation, duplicate cryogenic pumps, and interconnecting piping for each buffer tank. This process cycle has been applied recently and very successfully in treating waste water at the west coast in the United States. In both Figure 3.20 and Figure 3.21 one can observe that the flow of oxygen follows the 24 h variation of the wastewater very closely and always at a slightly higher rate. Thus the reaction between the oxygen and wastewater is always complete. Four 24 h tests were carried out, and all tests performed as guaranteed. 3.8.3.2 Version B This alternative process cycle maintains the same basic principles as Version A, but has extra capacity for the main condenser–vaporizer, the LOX buffer tank and a liquid air (rich liquid) tank are supplied in a single but separate column. This extra tower is included in the main cold box, thus saving insulation and piping. The LIN buffer tank noted in version A has been replaced by a liquefied air (LAIR) buffer tank. Its contents are produced by taking a portion GOX production 325 305 285

Standard tons GOX

265

Curve 3 (Actual)

Curve 3 (Target)

245 225 205 185 165 145 125 14

15

16

17

18

19

20

21

22

23

24

1

2

3

4

5

6

7

8

9

10

11

12

13

14

Hours (02/28/02, 02:00 PM to 03/01/02, 02:00 PM)

FIGURE 3.20 GOX production. (Courtesy of Linde BOC Process Plants, 2006; CSDLAC. With permission.)


Gox production 440 420 400

Curve 4 (Target)

Curve 4 (Actual)

380 360 Standard gox

340 320 300 280 260 240 220 200 180 16 17 18 19 20 21 22 23 24 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 Hours (03/04/02, 04:00 PM to 03/05/02, 04:00 PM)

FIGURE 3.21 GOX production. (Courtesy of Linde BOC Process Plants, 2006; CSDLAC. With permission.)

of the main process air, boosting it to a higher pressure, precooling it, and then passing it through the auxiliary vaporizer. In this auxiliary condenser–vaporizer, liquid-oxygen product is vaporized at the prescribed operating pressure and exits the plant after passing through the main heat exchangers. In turn, process air is partially liquefied, subcooled, and sent to the liquefied air buffer tank. By the time it enters its buffer tank, one may call it rich in oxygen or rich liquid as it is known in normal air plants. It joins the rest of the precooled main process air in the bottom of the lower (high pressure) column. The LOX product is sent to the condenser by one of two parallel 100% capacity cryogenic transfer pumps (see Figure 3.21 and Figure 3.22). The former figure shows that during a single blow of oxygen the level of the oxygen tank drops quickly, whereas during a 24 h period, as shown in the latter figure, there is a series of oxygen blows and a subsequent series of level variations in the oxygen accumulator tank. The blows were carried out both in the reverbatory and convertor separately (see Figure 3.22 and Figure 3.23).

3.9

EXERGY

Although the term exergy was used by Trepp in 19626, and by Patela in 1964 in a few individual cases, the word was widely accepted by the industry after a general paper was published by Robert Evans of Georgia Institute of Technology in 1966. The second law shows the influence of irreversibilities within system components, and can provide methods to minimize the losses of energy and capital. Even in 1939, Kapitza concluded that conventional isentropic efficiencies were not a true measure of low-temperature applications when processes were at different temperature levels. In 1979 and thereafter, engineers began to use the exergy method for analyzing the heat exchanger and turboexpander performance. Exergy has been widely used in Europe and in


Tiempo real

Nallsis historic 11:17:37 1h

4h

Jul 24 11:17:37

Tiempo real

2 h 15 min

Zoom in Min Jul 24 11:51:22

Zoom out 30 min

Jul 24 12:25:07

Valores 13:32:37

Unavail

96.6

LocsaNivelEst95

10 min

Jul 24 12:58:52

AIT−2040

Jul 24 13:32:37

Unavail

56.5

FT-2040

100

110.0

Unavail

98 96 94

PT-2040

92

35.0

Unavail

90 88 86 84 82 80 11:17:37

Imprimir 13:32:32

FIGURE 3.22 Variable Oxygen used in stop start operation for converter and reverbatory furnaces covering 1 hour and 40 minutes. (ß Air Liquide, all rights reserved, 2006. With permission.)

Analisis

Nalisis historic 23:55:00

Zoom in 1h

4h

Jul 24 23:55:00

Min Jul 25 01:28:45

Valores

Tiempo real

6 h 15 min

Zoom out 30 min

Jul 25 03:02:30

Jul 25 04:36:15

06:10:00

AIT-2040 Unavail

96.6

LocsaNivelEst95

10 min Jul 25 06:10:00

79.6

64.9

FT-2040

180

138.4

162

168.5

144 126

PT-2040

108

34.2

32.4

90 72 54 36 18 0 23:55:00

Imprimir 06:09:48

FIGURE 3.23 Variable Oxygen used in stop start operation for converter and reverbatory furnaces covering 6 hours and 10 minutes. (ß Air Liquide, all rights reserved, 2006. With permission.)


Russia for maximum available work function. More recent work on the second-law application in cryogenics has been aimed at the analysis of component parts rather than in systems. A detailed second law–system analysis of energy availability improvement, the magnitude, location, and cause of each significant irreversible loss in the system must be determined. Losses may include mass and heat gains (external and internal), flow frictional losses, and entropy production from flow stream mixing. D.W. Townsend7 of ICI Petrochemicals, Process Division, clarified the term exergy lucidly in 1979. He stated that the application of exergy is the best single concept to carry out an energy-efficiency study. It is defined as the maximum work that can be derived from a reversible system, or Ex ¼ (H H0 ) T0 (S S0 )

(3:47)

A process depends specifically on its cooling-medium temperature T0 and by integrating it can be shown that DS ¼ DH{[ ln (T2=T1 )]=(T2 T1 )}

(3:48)

Linnhoff8 has shown that Equation 3.48 is an excellent approximation even when linear TDH profiles are involved. If gaseous pressure changes occur, the ideal gas assumption holds: DS ¼ [(Cp dT)=T] [(RDP)=P]

(3:49)

DS ¼ DH{[ ln (T2=T1 )]=T2=T1 } R ln (P2=P1 )

(3:50)

and it follows that

Equation 3.49 and Equation 3.50 have been accurate enough for most applications. An irreversible process is one in which the potential for doing work is lost and it can be shown that X Ex(in) ¼ Ex(out) þ Ex(irreversible) (3:51) This equation offers an objective basis for the evaluation of the second law. Exergy thus provides a useful tool for the investigation of efficiencies to reduce energy requirements in refrigeration systems, which include compressors, condensers, system exchangers, flash separators, and general process exchangers; to reduce energy concepts; and at minimum capital investments. Process and project engineers should strive to meet this target despite the poor Carnot efficiencies encountered in cryogenics. More recently exergy or the analysis of the second law has already been used to select optimal LNG process cycles employing front-end propane refrigeration, and to evaluate cryogenic versus noncryogenic processes for the production of pure nitrogen.

REFERENCES 1. King, G.R. 1971. Modern refrigeration practice, 48–61. New York: McGraw Hill. 2. Timmerhaus, K.D. and Flynn, D. 1989. Cryogenic process engineering, 291. New York: Plenum Press. 3. Barron, R.F. 1971. Cryogenic systems (Monogram on cryogenics), 2nd ed., 151–200. New York: Oxford University Press. 4. Garcia, L. 1995. Basic principles of air separation. In Practical operating manual. Paris: L’Air Liquide.


5. Grunberg, J.F. 1950. The oxyton cycle in the production of tonnage oxygen. In The Industrial Chemist. London: The Physical Society. 6. Trepp, C. 1962. Adv Cryog Eng 7. 7. Townsend, D.W.T. 1980. Second law analysis in practice. Chem Eng (October): 628–633.



4

Rare (Noble) Gases

The rare gases, sometimes referred to as noble gases, are situated in the zero group in the periodic table of elements between the most electronegative (the halogens) and the most electropositive (the alkali metals) elements. Because their electron shells are completely filled, they are inert, that is, do not normally form chemical compounds. Their ionization energy, the energy required to remove an electron from the valence shell, is greater than for any other element, and conversely, their electron affinity, the energy released in taking up another electron, is zero. Nevertheless, it has been found that for the heavier inert gases, whose outer electrons are farther removed from the nucleus and whose force of retention is weaker, and electrons can indeed sometimes be removed by strong electromagnetic fields. Fluorides of krypton, xenon, and radon have thus been prepared, and recently krypton has actually been inserted into acetylene mainly to form an ionic bond with carbon. (One may thus say that a more appropriate classification in the periodic table would be in group VIIIA.) (Table 4.1). During the early part of the twentieth century most interest in the rare gases was in the research laboratory. It was oxygen that took the center stage. In the latter part of the century, however, the industrial uses of rare gases grew by leaps and bounds in applications ranging from welding to deep-sea diving to medical imaging to industrial and commercial lighting.

4.1 HELIUM Helium was discovered almost simultaneously in 1868 by Pierre Janssen of France and Norman Lockyer of England when each conducted a spectrographic examination of the solar chromosphere during an eclipse in India in 1868. Their findings were presented at the same meeting of the Royal Society, but Lockyer has been granted precedence because he had invented the instrument that Janssen had also used. They had observed a line similar to but not quite the same as the D1 and D2 lines of sodium. It was therefore labeled D3. When it was later recognized that due to a new element the latter was named helium (Helios in Greek). In 1895, when Sir William Ramsay spectroscopically analyzed a gas given off by heating clevite, a uranium-bearing ore, he saw the identical D3 line. In 1903, both Ramsay and Frederic Soddy determined that helium was a product of the disintegration of radioactive minerals. When the concept of thermonuclear fusion was developed during the 1940s it was realized that this process, fusing hydrogen into helium, was responsible for the energy emitted by the sun and all other stars. Although helium accounts for 23% of the known mass of the universe it occurs in the Earth’s atmosphere only at 5 vppm since the Earth’s gravity is not strong enough to prevent its escape to outer space. Helium is a colorless, odorless, and tasteless gas. It can be liquified at 4.224 K and may be solidified, but only at a pressure of at least 26 barA. At any lower pressure it remains as a liquid, even at lower temperatures. As a liquid its index of refraction is very close to that of its gas so that it is very difficult to observe it in a flask. It has six known isotopes, but only two are stable, that is, helium 4 and helium 3. The other four isotopes are radioactive and decay


TABLE 4.1 Physical Properties of (Rare) Noble Gases Normal

Element

Atomic No.

Molecular Mass (kg/kmol)

BPt K (K at 1.103 bara)

Vapor Pressure (mbarb)

Latent Heat (MJ/kmolc)

Specific Heat (kJ/kmol Kd)

Helium Neon Argon Krypton Xenon Radon

2 10 18 36 54 86

4.0026 20.09 39.948 83.80 131.30 222.0

4.224 27.09 87.28 119.8 164.1 211.2

51 433 688 731 816

0.083 1.79 6.65 9.04 12.63

20.80 20.79 20.78 20.87 21.02

a

Normal point is at a pressure of 101.325 kPa. (Atmospheric pressure at sea level.) Vapor pressure is at the triple point. c Latent heat of vaporization is at normal boiling point 101.325 kPa. d Molar specific heat capacity at constant pressure in range from 0 to 1 bar at 298.15 K. Source: From Pauling, L. in General Chemistry, Periodic System of Elements, W.H. Freeman & Company, San Francisco, 1958. With permission. b

very quickly. Helium 4 is the most abundant of the isotopes. It is present in the Earth’s atmosphere about 700 times the concentration of helium 3. What helium 3 does exist in the air is due to the b-decay of tritium, hydrogen 3. Timmerhaus and Flynn1 have stated that when liquid helium 4 is subcooled below 2.17 K it splits up into two phases, that is, helium 1 and helium 2. At lower temperatures, the former has a viscosity of 3 106 Pa s; and the latter phase exhibits properties of superfluidity, inasmuch as its viscosity and resistance to flow almost vanishes. The viscosity of helium 2, at 1.3 K, is approximately 3 107 Pa s, or even lower (1012) in narrow channels, and its thermal conductivity reaches 86,500 W=(m K) that is 1000 times greater than that of copper.

4.1.1

SOURCES

OF

HELIUM

Originally, helium was recovered by heating monazite earth to a temperature of 1000 K. Shortly afterward, and with the industrial development of air separation plants, Georges Claude2 in 1908, suggested that it was possible to collect an uncondensed mixture of neon– helium from the top of the main condenser (MC) of the distillation section of an air separation plant. If this mixture was allowed to accumulate it would lower the efficiency of heat transfer of the MC, and therefore lower the productivity of the air separation plant. When syphoned off, this mixture could then be cooled to the temperature of liquid hydrogen, and when passed through a bed of activated charcoal, the neon could be adsorbed in a solid state, and the helium freed as a gas. This latter process was still used up to World War I. But although atmospheric air contains 5.3 vpm of helium its recovery is uneconomical, except for small laboratory work. Soon after World War I a natural gas well at Excell, Texas, in the Texas Panhandle was discovered to contain a fair quantity (0.70%) of helium. It was also shown that it was more economical to extract this helium for industrial uses by using a cryogenic process. Consequently, the U.S. Government told the Bureau of Mines to set up an office at Amarillo Texas, to take charge and develop the recovery and conservation of helium. From 1918 to 1961, the Bureau of Mines had fairly tight control over the use of helium, and its consumption was restricted to military operations, the inflation of weather balloons, and lighter-than-airships.


Natural gas processing

Helium processing

Helium distribution

Natural gas plant End-use customers

High-purity helium (gas)

Fuel gas

Gas distributors

Crude helium Gas tube trailer

(50% -70% He) Upgrader

Cylinders for end-use customers

Liquid storage

PSA

End-use customers Liquefier

Bulk liquid helium transport

Bulk distribution centers

Tube trailers, cylinders, dewars for end-use customers

Natural gas reservoir

FIGURE 4.1 Recovery of helium from natural gas reservoir to consumer in gaseous and liquid form. (Courtesy of Praxair, Inc., 2006. With permission.)

The Bureau’s other helium-bearing natural gases were located nearby at Keyes, Oklahoma; Shiprock, New Mexico; and Otis, Kansas. After 1960, private industry entered the helium recovery field by long-term contracts with the U.S. Government. In 1961, a privately owned plant started up in eastern Arizona with an annual production of 1.875 MM Nm3. Then, in 1966, another plant was built in Otis, Kansas, with an annual production of 4.019 MM Nm3. And in 1966, a third plant was erected at Elkhart, Kansas, with an annual production of 3.750 MM Nm3. These plants produced unrefined helium of 50%–85%, which was then purchased by the U.S. Government and stored underground in a natural gas deposit for future use (Figure 4.1). On a worldwide basis, the following sources are producing helium in the year 2000 (Table 4.2).

4.1.2 GENERAL PRINCIPLES

OF

RECOVERY

OF

HELIUM

Helium was early discovered in very limited quantities in radioactive mineral springs at Bath, England. However, not much work was carried out on its physical properties. It was at the turn of the twentieth century that helium was also discovered in monazite sands,

TABLE 4.2 Worldwide Sources of Helium Source Kansas Texas Panhandle Poland Holland Germany North Sea, Leman Bank France

He (%)

CH4 (%)

N2 (%)

C2 Plus

2.0 0.7 0.4 0.045 0.04 0.03 0.001

— 73.2 56.01 81.3 42.5 94.7 97.1

23.0 14.3 42.75 14.35 56.5 1.3 0.3

Balance Balance Balance Balance Balance Balance Balance

In Canada, Algeria at Arzew and Skikda, and Romania there are also small sources. Source: Courtesy of CryoGas International Publication, 2002. With permission.


TABLE 4.3 Typical Analysis of Uncondensed Stream from Main Condenser of Air Separation Unit Nitrogen Neon Helium

70.0% 23.0% 7.0%

Source: From Ruhemann, M. in The Separation of Gases, Oxford University Press, London, 1945. With permission.

phosphates-containing rare earth metals and thorium, at gravel pits in North Carolina, from which helium could be extracted, albeit with difficulty. After this important discovery, serious and extensive experiments were carried out on helium by Kamerlingh Onnes in Holland on the element’s physical properties, both in gaseous and liquid phases. When industrial air separation plants came into practical operation, Georges Claude suggested that helium could be recovered from the MC where both helium and neon accumulated in an uncondensed vapor. Unless removed, this created inefficiency in the heat transfer at the MC. The approximate analysis of the takeoff gas in the uncondensed stream was as follows3 (Table 4.3). The nitrogen was easily separated by compressing the mixture to 50 bar and by passing the gas through a reflux condenser that has been filled with liquid nitrogen. The remaining traces of nitrogen were removed by passing the effluent through activated charcoal, or over hot calcium, or magnesium in order to save the neon. To separate neon from helium, the use of nitrogen as a refrigerant was out of the question because its triple point is 63 K, and at this temperature its vapor pressure is 433 mbar, which is very difficult to maintain industrially due to the necessary use of a deep vacuum operation. Nevertheless, Meissner4 succeeded in separating the two elements with the use of liquid hydrogen as a refrigerant. This was the process used up to 1918, at the Physikalische—Technische Reichsanstalt at Charlottenburg in Germany. With the discovery of helium in natural gas at the Texas Panhandle the recovery of helium became a large-scale industrial process, and the Bureau of Mines soon found that for large flows of natural gas a helium content of less than 0.1% was not viable for production. In fact, the recovery of helium is of little interest when its content is less than 0.5%, at least in the United States. For a content of 0.5% and up to 2.0%, recovery is economically viable. Nevertheless, in former Soviet Union, the engineers were exploiting natural gas with only 0.1%–0.3% of helium, no doubt, for its military potential. The presence of significant nitrogen in natural gas more often than not indicates that helium is also present. Moreover, the presence of nitrogen is of definite advantage because it can be used as a refrigerant cycle after a methane cycle. The resulting liquefied nitrogen can be used to augment the recovery as well as the purity of the helium. In short, the combined use of both nitrogen and methane evolves into a classic cascade cycle, which is the most economical, especially if the natural gas also contains ethane and propane. Depending on the composition of the natural gas, a wide variety of process cycles have been designed, most of them operating on multistage refrigeration cascade process cycles. The first effect is a ternary mixture of methane, nitrogen, and helium. Methane is the first to be liquefied. As the solubility of helium in liquid methane is insignificant the helium remains in the gas. The binary mixture of nitrogen and helium remaining, however, is difficult to separate, irrespective of the nitrogen content. If one considers the binary system of helium and nitrogen, the latter being the main refrigerant, though the boiling points are far apart and the vapor–liquid equilibrium curves are well known, their separation is not easy, depending on the purity desired for the helium.


TABLE 4.4 Helium–Nitrogen Mixture P (bar) 10 20 40 60 100

% He in Vapor % at 80 K

At Equilibrium % at 70 K

79 91 95.5 97.2 98.5

93 98 99.2 99.6


For very high purities, of say 99.5%–99.85%, very high pressures are needed to attain high concentrations of helium vapors in equilibrium with liquids at 80 and 70 K. From Table 4.4 it is obvious that with nitrogen as a refrigerant, one can only hope to reach a refrigeration temperature of 77.3 K, the atmospheric boiling point of nitrogen. Therefore, one has to consider either the use of a higher pressure or the use of liquid nitrogen boiling at a lower pressure. Presently, it is common to use a reasonable pressure of around 40 bar for the closed nitrogen cycle and a much higher pressure for the helium cycle. The latter combination will result in an optimum consumption of energy for the final stripping of nitrogen. This concept is especially applicable for helium-bearing natural gases where the helium is present in a comparatively small quantity and nearly total recovery is mandatory (Figure 4.2).

N2 liquefier Crude He Crude He

Crude He separator He purifier Crude He

Heat exchanger

He heat exchanger

N2 gas

He heat exchanger We

Natural gas Compressed natural gas Natural gas

95% N2

Pure He separator N2

N2 + He in solution Separator

98.5% He

FIGURE 4.2 Helium separation by Bureau of Mines. (From Timmerhaus, K.D. and Flynn, T.M., Cryogenic Process Engineering, Springer Science & Business Media, 2005. With permission.)


4.1.3 RECOVERY PROCESSES FROM NATURAL GASES Before applying a cryogenic refrigeration cycle to any natural gas the entire gas stream has to be treated in order to remove all the cryogenic undesirables such as carbon dioxide, hydrogen sulfide, and water, which may be present in the process gas. If carbon dioxide content is above 2.9% by volume one must use either an amine scrubbing system, or refrigerated methanol scrubbing (Rectisol). A further treatment with a molecular sieve containing activated alumina upstream is then needed to make sure that the prepurification treatment operates at maximum efficiency. If carbon dioxide is less than 2.9%, the pretreatment can be, and has been carried out, by the exclusive use of molecular sieve adsorption. The refrigeration process may be considered as operating in two stages. In the first stage, the extraction involves the removal of both the methane and a major portion of the nitrogen by liquefaction. If too much of the nitrogen is removed, however, there may be a high loss of precious helium, because of the latter’s solubility in liquid nitrogen. If any small quantities of nitrogen and helium are left they will degrade the natural gas in terms of valuable calorific value. The remaining binary mixture of nitrogen and helium involves the denitrogenation of helium. This part of the overall process uses a standard Claude closed nitrogen cycle with a compressor and an expansion turbine. Depending on the final composition of the helium required it may also require the use of a high-pressure booster compressor for the helium. This high compression is also useful for transporting the helium product in cylinders elsewhere for further purification. Typical examples are from the Bureau of Mines (see Figure 4.2)6. The natural gas is compressed to a pressure of 42 bar, treated for the removal of carbon dioxide, hydrogen sulfide, and water, then sent through a primary heat exchanger (PHX), where it is cooled by the returning stream of cold natural gas containing some nitrogen, but free of helium. The main process stream continues through a throttle valve and is expanded into an MC at a pressure of 18 bar and a temperature of 125 K. In the MC all hydrocarbons are condensed and returned to the PHX, where they serve to cooldown the incoming natural gas. The MC is also refrigerated by an incoming stream of cold gaseous nitrogen at practically atmospheric pressure from the expansion turbine. Once this nitrogen gives up its refrigeration it is then returned to the inlet of the nitrogen-cycle compressor. The vapor at the top of the condensed natural gas in the MC consisting of approximately 60% helium, 40% nitrogen, and traces of methane, is extracted and sent to the inlet of the helium compressor after giving up its refrigeration in a helium heat exchanger (HHX). The principal refrigeration system is a standard Claude cycle employing a closed nitrogen cycle with a compressor boosting the gaseous nitrogen to a pressure of 42.6 bar. After passing through a nitrogen and nitrogen heat exchanger (NHX) where it is cooled by a returning stream of low-pressure nitrogen, it is split into two streams. One stream enters an expansion machine and is sent as a cooling gas to the top of the MC. The second stream is cooled by the expanded nitrogen stream, and is throttled as a liquid and is sent to the condenser–evaporator (CE), where it serves to strip the nitrogen from the helium mixture. It also upgrades the final product helium to a purity of 98.5% by maintaining a low temperature at an activated charcoal filter, as the gaseous product helium passes through it. At the bottom of the CE, the liquid nitrogen is removed, throttled to a lower temperature, and enters a separator (S). At the separator more of the nitrogen is condensed, whereas the uncondensed helium vapor containing progressively less nitrogen is removed and sent to the C-1 condenser to be recycled into the process for further stripping and upgrading. In order to achieve a higher purity, of say 99.995% or even up to 99.9995 þ %, it is necessary to treat the 98.5% helium separately using liquid nitrogen boiling at subatmospheric pressure (20 kPa), and removing the nitrogen with activated carbon. The vessels containing the activated carbon are also kept cold by encasing them in a liquid nitrogen cooled jacket.


4.1.4 APPLICATIONS

OF

HELIUM7

Because helium is chemically inert, nonflammable, has a high thermal conductivity, a low molecular weight, a small molecular size, and is nontoxic, it has many applications. It has a high arc temperature and a high heat transfer rate. This makes it ideal for its use in inert gas shielding, or plasma gas welding of steel, aluminum copper, nickel, stainless steel, titanium, and many other alloys. It is also beneficial for the smelting and refining of advanced materials such as niobium, tantalum, zirconium, and titanium. Having a very small molecule, helium’s high penetrating quality is very advantageous in leak detection, especially in aerospace equipment, as well as in industries dealing with semiconductors, and nuclear, cryogenic vacuum, refrigeration, and food canning equipment. Because of its low solubility and low boiling point (4.2 K), as well as its nonflammability even at very low temperatures, helium is employed as a purging and pressurizing agent for liquid hydrogen fuel systems of rockets and spacecraft. Helium is very useful in deep-sea diving and work in underwater caissons, replacing nitrogen in admixture with oxygen. Breathing it under high pressure does not result in the rapture of the deep, a dangerous narcosis caused by nitrogen, nor does it cause agonizing bends when a diver resurfaces rapidly. Helium–oxygen breathing mixtures have opened the door wide to undersea scientific exploration and made possible the development of offshore oil and gas resources even though the mixtures have the minor disadvantage of raising the pitch of divers’ voices. Helium is critical for producing optical fibers to be used in telecommunication cables. It also allows manufacturers to grow crystals for electronic semiconductors or crystals with optical properties needed to produce masers and lasers. Apart from applications of gaseous helium, liquid helium has developed as one of the most important new materials of science. It helps to cool certain compounds below a critical temperature, at which they become superconductors. Superconducting magnets are used in magnetic resonance imaging (MRI) that has become a powerful tool in diagnostic medicine. The application of a similar technology, nuclear magnetic resonance (NMR), is also invaluable in probing the structure and properties of chemical compounds. Liquid helium also finds application in power generation and transmission, fusion research, magnetohydrodynamic (MHD) propulsion of ships, superconducting magnetic energy storage (SMES), electronics, and magnetically levitated (Maglev) transportation systems. In order to give an idea of the year 2000 consumption of helium in the United States the following approximate figures in Table 4.5 are of particular interest8.

TABLE 4.5 Annual Consumption in the United States Industrial and scientific MRI Welding Lift gas Fiber optics Purge and pressure gas Leak detection Controlled atmospheres Breathing gases Other

18% 18% 16% 14% 9% 8% 6% 3% 3% 5%

Source: Courtesy of CryoGas International Publication, 2002. With permission.


FIGURE 4.3 Helium recovery after use. (Courtesy of Praxair, Inc., 2006. With permission.)

4.1.5 CONSERVATION

OF

HELIUM

In view of the simple fact that helium drifts off into outer space and is lost forever, every effort should be made by industry and all users to recover helium at every opportunity. The U.S. Government has already taken the lead in storing helium within natural gas storage caverns. This idea of conservation should be followed by ultimate users. This precious element should not be wasted for toy balloons and parade floats. Equipment has already been designed to recover helium at major users (Figure 4.3).

4.2 NEON 4.2.1 GENERAL Neon, which derives its name from the Greek (meaning new), has an atomic weight of 20.183, with three stable isotopes of 20, 21, 22, and is found in the atmosphere at 15.385 106 vppm. It is inert, but in certain conditions it can form a compound with fluorine. Liquid neon is clear and colorless, and has a boiling point of 27.09 K, a triple point of 24.56 K, and a vapor pressure of 433 mbar, at its triple point. The electrical discharge of gaseous neon at ordinary voltages and currents gives off an intense red.

4.2.2 SOURCES

OF

NEON

Neon has been extracted from the atmosphere as a by-product in air separation plants. As described in Section 4.1.2, neon can be recovered from the noncondensable stream removed from the MC of the air separation column. After catalytic oxidation to remove all traces of any hydrogen, and further treatment in an additional column and condenser assembly, the analysis is improved to the following9 (Table 4.6).


TABLE 4.6 Typical Analysis of Neon Stream after Complete Hydrogen Removal Nitrogen Neon Helium

70% 23% 7%


4.2.3 RECOVERY

OF

NEON

In the treatment of this last analysis it is important to remove the nitrogen completely, as any nitrogen left may spoil the color when neon is used for lighting. By passing the mixture at a pressure of around 50 bar through a reflux exchanger surrounded by a bath of liquid nitrogen, it is fairly easy to remove the bulk of nitrogen. The remaining traces can be adsorbed by a bed of activated charcoal. If the latter material also causes a high loss of neon, then one must consider the use of either hot calcium or magnesium. It should be noted, however, that the separation of neon from helium is very difficult as well as expensive. The use of nitrogen as a refrigerant is out of question because the lowest temperature attainable even as a liquid at its triple point is 63.2 K, and its vapor pressure at this temperature is only 126 mbar, which is difficult to maintain industrially due to the necessity of a continuous deep vacuum operation. On the other hand, the critical temperature of neon is 44.41 K and that of helium is 5.2 K. One, therefore, had to consider the use of liquid hydrogen. Alternatively, one may use Sterling cycle refrigerators to provide low-temperature cooling. Meissner10 succeeded in separating a neon–helium mixture in a fairly large scale by using liquid hydrogen in the early 1930s. This treatment, however, presents a problem because the boiling point of hydrogen is 20.4 K and that of neon is 27.2 K. The triple point of neon on the other hand is 24.6 K, which is above the boiling point of hydrogen. Care has to be exercised, therefore, to make sure that the neon does not solidify and block the tubes. To avoid this problem, one has to maintain the operating pressure above the liquid hydrogen at 3 bar, which is sufficient to raise its boiling point above the triple point of neon. Another method is to raise the temperature differential between the condensing neon and the evaporating hydrogen by reducing heat transfer surface in equipment.

4.2.4 INDUSTRIAL RECOVERY

OF

NEON

In recovering neon, one must consider the final purity desired, especially in terms of helium content. Helium is not really an undesirable impurity if the mixture is to be used in fluorescent lighting fixtures as a certain quantity of helium adds to the effect of the final color. If on the other hand, a high purity of neon is the objective (especially if the final product is liquid neon), then a two-phase system must be designed. The first phase (Table 4.6). The noncondensable mixture of helium, neon, nitrogen, and some traces of hydrogen is withdrawn from the top dome of the MC and led into the bottom of a separate column equipped with a reflux condenser, where the vapors are led through a bath of liquid nitrogen. The condensibles, mostly liquid nitrogen, fall to the bottom and are returned to the MC of the air separation unit. The remaining noncondensables, a mixture of gaseous neon–helium, are withdrawn from the top dome of the condensing unit and led into a second phase of purification involving liquid hydrogen.


In this secondary and final phase, the mixture from the auxiliary reflux condenser is mixed with a small amount of oxygen and fed into a catalytic burner to remove any traces of hydrogen. After the stream passes through an after-cooler and dryer, the final mixture of neon–helium enters an adsorption trap encased in, and cooled with liquid nitrogen. Finally, the resultant stream passes through a closed cycle hydrogen-refrigerated exchanger and enters a secondary reflux exchanger where it is liquefied in a bath of liquid hydrogen.

4.2.5

INDUSTRIAL APPLICATIONS

OF

NEON

Apart from its use in fluorescent light fixtures, neon has also found application in high-energy physics. According to Isalski11, the study of subatomic particles involves the use of liquid hydrogen-filled bubble chambers with a target residing at their center. When highly accelerated atoms strike the target, reactions may occur, releasing subatomic particles. Liquid hydrogen has the property of creating tracks that can be photographed when a subatomic particle is released. Liquid neon is more dense than hydrogen and has been used outside the hydrogen chamber in order to slow down the reaction, resulting in a slow-motion effect. Liquid neon also has an advantage over liquid helium as a low-level refrigerant. It is less expensive per unit heat of vaporization, and a precise temperature control can be easily achieved by varying the saturation pressure of liquid neon. Liquid neon is also used by the military in infrared detection equipment to improve light sensitivity. It actually reduces molecular motion within the equipment.

4.3 4.3.1

ARGON GENERAL

Argon, which derives its name from the Greek (meaning late comer), has an atomic weight of 39.948, and three isotopes 36, 38, 40, and is found in the atmosphere at 0.934%. The liquid phase is a clear colorless fluid with a normal atmospheric boiling point of 87.28 K at 101.325 kPa. Its triple point is 83.8 K, its critical pressure is 4860 kPa (48.6 bar), and its vapor pressure at its triple point is 688 mbar.

4.3.2

SOURCES

OF

ARGON

In 1785, H. Cavendish determined experimentally that about 5% of the so-called dephlogisticated air had different physical characteristics than ordinary air. In 1884, Ramsay and Rayleigh discovered that the atomic weight of atmospheric nitrogen was heavier than that of nitrogen produced by chemical reaction. The difference, they found, was due to a new inert gas in the atmosphere, which they have named argon. Further studies proved that this new element formed 0.934% by volume of the atmosphere, or 1.25% by actual weight. It also proved to be a very valuable factor in our modern industry.

4.3.3

RECOVERY

OF

ARGON

The atmospheric boiling point of argon (87.3 K) is closer to that of oxygen (90.2 K) than to that of nitrogen (77.3 K). As can be seen from its physical data, argon has a boiling point (87.3 K) between those of oxygen and nitrogen, namely, 90.2 and 77.3 K, respectively. In fact, it is closer to that of oxygen (90.2 K). In other words, a high recovery of argon implies a complete separation of oxygen and nitrogen, which requires extra energy and equipment. Forty to 50 years ago, a recovery of 40% of argon from a standard air separation plant was considered high for any definite guarantee. Today, however, designers are committing


themselves to recoveries as high as 95% in certain conditions of operation, especially in liquid plants. Optimum recovery of argon depends on three factors: (a) Increasing the number of trays in the upper (low pressure) column. This factor, however, increases the investment because of extra trays, longer piping, extra outer casing, and insulation. There is also an increase in overall pressure drop, which increases energy consumption. (b) Increasing efficiency of the distillation trays. This raises the recovery, but only marginally. (c) Increasing the total quantity of liquid nitrogen production, hence increasing the total nitrogen reflux both at the lower and upper columns. In this manner the argon will not escape in either the product or waste nitrogen streams. This option, however, requires extra refrigeration capacity with the use of an external nitrogen recycle compressor. This indicates a higher energy consumption. A combination of all the three factors can provide a good balance between energy and investment. One should note that there is a definite relationship between the quantity of nitrogen withdrawn from the lower column of an air separation unit, whether in gaseous or liquid phase, and the recovery of argon. The quantity of nitrogen available for reflux may be limited by (a) The amount of nitrogen, or nitrogen–oxygen mixture, expended for refrigeration in the expansion turbine (b) The amount of nitrogen withdrawn as a product either as a liquid or as a gas In other words, recovery of argon (RA) is proportional to N2 and processed air (PA). To overcome this above-mentioned limitation and to increase argon recovery to its maximum, it is mandatory to produce extra refrigeration by an external cycle using nitrogen as the refrigerant. This explains why some air separation units with a high production of liquids, and which use large nitrogen refrigeration recycles, also achieve close to 95% recovery of argon. In the past, the point of argon withdrawal from the upper (low pressure) column was not determined with any degree of accuracy. Rough calculations combined with past experience enabled engineers to guesstimate a few outlet nozzle locations, and then hope that one was the right one. Today, though, a reasonably precise location can be precalculated. With high recovery in mind, it is important to maximize the amount of liquid nitrogen reflux in the lower (high pressure) column to strip all argon and oxygen from the rising airstream and to accumulate them in the sump of the tower. This so-called rich liquid, rich in oxygen, also contains all argon that was in the process air. It is withdrawn, subcooled through an expansion valve, and sent to the upper (low pressure) column for rectification.

4.3.4 RECOVERY PROCEDURE

AND

EQUIPMENT

It should be noted that the upper or low-pressure column of an air separation plant is never in a steady state (Figure 4.4). The prepurification system upstream of the separation equipment—reversing exchangers, molecular sieve adsorption—plays an important role in the argon recovery. Each time there is a reversal of the exchangers or a change in the airflow due to the pressurizing and depressurizing of the adsorber vessels, there is also an upset in the upper column, which in turn affects the mixture of oxygen–argon–nitrogen at the point of takeoff. This loss of argon recovery will vary with the magnitude and frequency of the upsets caused by the prepurification system.


Crude argon

Crude argon column

Low pressure column Rich liquid filters

Lox Filter

Vaporizer condenser Liquid Nitrogen Lox Pumps

To and from heat exchangers

Oxygen

L o x

P u r g e

Nitrogen reflux cooler

High pressure column

Rich liquid cooler

Air Rich liquid Nitrogen

FIGURE 4.4 Crude argon recovery from standard air plant. (Courtesy of F.G. Kerry, Inc. With permission.)

From the upper (low pressure) column, a stream of argon is withdrawn and piped to a crude argon fractionation tower. The stream can contain 9% argon and 0.01% to a few vppm of nitrogen, and the remaining oxygen. To separate this large amount of oxygen, great number of trays are required; the higher the recovery of argon, the more trays are needed. The crude argon tower can thus be very tall and may require a separate cold box parallel to the main cold box. The tower is fitted with a reflux condenser developed by Georges Claude12 in the early 1920s. (Originally this piece of equipment was called a dephlegmator.) The shell side contains subcooled rich liquid. Oxygen in the argon-bearing stream is condensed and passes down to the sump whereas the slightly more volatile argon is condensed higher in the


column and collected in a specially designed shelf. At this shelf, the colder liquid argon acts as liquid nitrogen in the shelf of the air plant’s high-pressure column, but in this case to strip oxygen from the rising argon-rich stream.

4.3.5 SECONDARY RECTIFICATION

AND

FINAL PURIFICATION (SEE FIGURE 4.4)

The argon-rich liquid with an approximate analysis of argon 98%, with oxygen at 1% or considerably less, the balance nitrogen, and at a temperature of around 87 K, is withdrawn from the crude argon tower and sent outside to a special external heat exchanger. In this heat exchanger, the cold argon stream releases its refrigeration in countercurrent heat exchange to an entering refined warm argon stream. The term refined here means that all the oxygen has been completely eliminated. The most volatile components, namely nitrogen and some traces of argon, are removed from the top of the crude argon column, and may be routed to the low pressure of the main plant for reliquefaction and further rectification. The refined argon contains the original amount of nitrogen, sometimes less than 1 vppm oxygen, and some residual hydrogen left over from the deoxo catalytic unit. Now cold, it is routed to a trayed pure argon column, which is similar in design to the crude column. Its reflux condenser is filled with liquid nitrogen on the shell side at a subcooled liquid temperature. The refined and refrigerated argon enters the pure argon column near the sump in a closed circuit. The closed circuit coil that receives the refrigerated argon acts as a reboiler for the pure liquid argon product in the sump. As the refined argon exits from the tower, it enters a flash drum (separator) where the condensed refined argon is expanded and degassed. This side operation serves to get rid of any excess hydrogen and nitrogen, which now exits as vapor and is returned to the crude argon system for recovery. The balance is returned to the middle of the column as feed. The vapor from this feed rises to the top, enters the tube side of the overhead reflux condenser where it comes into countercurrent exchanger with subcooled liquid nitrogen in the shell side. This action serves to reliquefy the argon. Any gaseous nitrogen and residual hydrogen escape at the top of the condenser. Liquefied argon descends through the tower and acts as a reflux to the rising gaseous mixture. The small stream of nitrogen and residual hydrogen vapor passing through the condenser is either returned to the waste nitrogen exiting from the top of the main upper (low pressure) column of the air separation plant, or is vented from the top of the pure argon column through a special pressure indicator control (PIC) valve. Liquid nitrogen at the overhead condenser is fed from the pure liquid nitrogen found at the top of the lower (high pressure) column of the main air plant. It enters the sump of the pure argon tower in a closed circuit where it is maintained as a liquid, then as it exits, it is throttled slightly for subcooling and enters the reflux condenser at the top to replenish the liquid nitrogen bath. The purified argon product in the sump is now completely free of all impurities to a level of less than a few vppm for the total sum of oxygen, nitrogen, and hydrogen. The final liquid product is withdrawn and sent to a vacuum-insulated cryogenic storage tank. Vapor from the storage tank is recycled to the pure argon column for recondensation and recovery. It is also possible to recover some part of the hydrogen stream used in the refining operation by recycling the gas stream from the flash drum and separator into the crude argon system.

4.3.6 REFINING OPERATION

AND

EQUIPMENT

The recent trend of most cryogenic designers is to produce crude argon almost completely stripped of oxygen down to 1 vppm or less (Figure 4.5). To carry out this process, one must either design a very large number of theoretical trays, which is not practical, or better still, use structured packing. Otherwise the pressure drop would prove excessive. This procedure, if successful, completely eliminates the need for warm argon refining. Nevertheless, the following description is submitted, as it may prove useful as process technology.



F1

T1

Crude argon Nitrogen

P1

PIC

Crude argon compressor

Deoxo argon O2 AP

Separator

LSRE

PIC

dP1

Dust filter P1

P1

AIC

T1 T1

P1

T1

Dessication bottles Flame arrestor

Deoxo unit AIC

TLS

T1 T1

P1

T1

Vent Water Refrigeration cooler unit Separator Separator

T1

Atmospheric cooler

FIGURE 4.5 Oxygen removal from crude argon. (ß Air Liquide, all rights reserved, 2006. With permission.)

The refiner consists of a group of equipment to completely eliminate oxygen from the argon stream leaving the crude argon side arm column, and operating at normal temperature levels. This equipment consists of the following: an oil-free blower, a combustion tower containing a palladium catalyst, a refrigeration machine, a water separator, a dryer filled with activated alumina, and ancillary piping, valves, and controls. This entire assembly is prepackaged and skid mounted in a module for easy erection and connection to the cryogenic equipment. The argon-rich stream is compressed in the oil-free blower and injected with hydrogen in a proportion slightly higher than the stoichiometric ratio required for the removal of oxygen completely. In the combustion tower, the oxygen is consumed completely. Hot effluent from the combustion tower is first cooled by either air or water, and then chilled by a refrigeration machine. The water so formed is removed, first in a separator and then in a dual tower-activated alumina dryer. The final product is returned to the argon system where it is cooled in countercurrent heat exchange with the outgoing stream leaving the crude argon column. In catalytic combustion for oxygen removal, there is a marked increase in temperature, 808C, or higher, for every 1% oxygen content in the treated gas. For this reason, it is prudent to recirculate the effluent around the combustion chamber in order to dilute the oxygen content and to keep the temperature down to a permissible limit. Nevertheless, the temperature of the unit is sometimes maintained at 370 K to avoid any operational upsets. The necessity and inclusion of a refiner in the recovery of argon is an expensive, albeit a necessary concept in terms of capital, energy, and maintenance. It also involves the use of hydrogen, which is costly especially in large quantities. Hydrogen may be supplied in cylinders, as liquid hydrogen in bulk form, or from cracked ammonia. This last option is not recommended, however, because it increases the amount of nitrogen content to be removed in the pure argon column. The cryogenic industry had worked for years to develop a process for the cryogenic complete removal of oxygen in the crude argon column. Finally in 1989, Linde AG built an air separation plant in Germany with an argon production, which did away with the warm argon refining system. The crude argon column was filled with structured packing, and was capable of consistently removing oxygen down to less than 1 vppm. The entire air separation plant, including the argon purification, was operated automatically from a remote control center. The argon column was so tall that its casing would have overshadowed the main air fractionation tower. The project engineer, therefore, increased the elevation of the main tower in order to achieve the same top elevation for both towers. The space below the main tower had enough room to include the liquid recirculating pumps.

4.3.7 APPLICATIONS

OF

ARGON

In the 1950s, argon was sold in cylinders. Today, it is produced in liquid tonnage quantities and bulk-delivered in insulated tank trucks or even railroad cars. It is marketed in everincreasing quantities for manifold industrial applications. In the metallurgical industries, argon is used to inert molds in pressure die casting, in the fabrication of mild steel, to protect molten metal in the continuous casting of carbon steel, and it has been stated that about 90% of stainless steels produced in the United States use this argon–oxygen decarburization (AOD) process. A major application is also in tungsten-inert gas (TIG) and metal-inert gas (MIG) shielded arc welding, especially for aluminum alloys to prevent formation of nitrides. Whereas its original introduction for filling incandescent lamps is gradually encroached upon by krypton–xenon mixtures, it is extensively used as a carrier gas in the manufacture of semiconductors and other electronic components. It also finds use in the manufacture of plate glass and in fabricating hermetically sealed multipane insulated windows.


4.4 4.4.1

KRYPTON AND XENON GENERAL

Krypton derives its name from the Greek (meaning hidden). It has an atomic number of 36, a molar mass of 83.80 kg=kmol, no known isotopes, a vapor pressure of 731 mbar at its triple point of 116 K, a boiling point of 120.5 K, and a critical pressure of 54.82 barA. Xenon, which derives its name from the Greek (meaning stranger), has an atomic number of 54, a molar mass of 131.3 kg=kmol, no known isotopes, a vapor pressure of 816 mbar at its triple point of 161.3 K, a boiling point of 202.15 K, and a critical pressure of 58.99 barA.

4.4.2

SOURCES

OF

KRYPTON

AND

XENON

Both krypton and xenon are found in the atmosphere in extremely low concentrations. Krypton is at 1.14 106 vppm, whereas xenon is at 0.086 106 vppm. If 1 m3 of air is treated, one cm3 of krypton and one-tenth of a cubic centimeter of xenon will be recovered. Nevertheless, in 1918, Claude13 suggested that krypton be used in place of argon for incandescent lamps because it would retard the vaporization of the filament. This would make it last longer and increase the output of lumens. In 1928, Claude extracted krypton from an air separation plant. Apart from trying to recover these gases from the atmosphere, it is also possible to recover them from chemical processes, which use air, and periodically have to purge a substantial quantity of gas due to a buildup of inerts. These may also include krypton and xenon, for example, ammonia purge gas.

4.4.3

RECOVERY

OF

KRYPTON

AND

XENON

In view of the growing interest of using krypton for the lamp industry, I.G. Farben14 in 1933 decided to install a krypton recovery system in one of its three large air separation plants, each having a daily capacity of 90 t of impure oxygen, which were being constructed at Leuna for the gasification of lignite. In that first industrial adaptation, it was possible to enrich liquid oxygen in terms of 1 part krypton in 1000. It was impossible to achieve a further enrichment, however, because of the dangerous increase in hydrocarbons within the liquid oxygen sump in the MC. Following this preliminary pilot plant at Leuna, two subsequent plants were built around 1938, one unit at Boulogne Sur Mer in France treating 30,000 Nm3=h of air and another unit at Ajka in Hungary, treating 22,000 Nm3=h of air. In order to recover krypton at a time when oxygen plants were of a small capacity, the units had to be designed specifically for krypton as the main product15. It was at this time that the application of aluminum-packed recuperative exchangers, better known as reversing regenerators, were used for the removal of water, carbon dioxide, and some hydrocarbons. These reversing regenerators developed by Frankl16, opened the door to the design and fabrication of large air separation units. Unfortunately, while these units could eliminate water and carbon dioxide to a large extent, they were not successful in the complete elimination of the most dangerous hydrocarbons now in existence in the atmosphere. The recovery of a krypton and xenon mixture can be a very dangerous process as it involves the reboiling or vaporization of the more volatile component, namely oxygen, thus increasing the concentration of dangerous hydrocarbons in a bath of 100% liquid oxygen. Violent explosions have occurred in the past. In some places, the auxiliary cryogenic equipment carrying out the recovery of krypton and xenon has been separated and enclosed by a concrete barrier. In an early case, the air separation plant for the production of krypton and xenon was located at a high altitude in the European Alps in order to reduce contamination by dangerous hydrocarbons.


FIGURE 4.6 Oxygen generation of 2,000 MT=d producing approximately 12 M3=d of krypton and 1 M3=d xenon. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

Before World War II, an air separation process airflow of 30,000 Nm3=h would produce approximately 200 t=d of oxygen, which was then considered huge. Today, such a plant is considered as a small-prepackaged unit. In fact, nowadays a plant capacity of 1000 t=d may be considered as routine. For this reason, krypton and xenon are now considered as by-products of air separation plants with a capacity of at least 300 t=d. One plant built in the United States in the late 1980s consisting of two units, each with a daily capacity of 1550 t of oxygen, has been producing approximately 12.5 normal cubic meters of krypton and 1 normal cubic meter of xenon on a daily basis (Figure 4.6). The process cycle for the by-product recovery of a krypton and xenon mixture is somewhat similar to the recovery of argon. In this process cycle, as is described by Isalski17, liquid oxygen product is withdrawn from the MC of the air separation unit, and after passing through a well-designed hydrocarbon adsorber, is sent to an auxiliary rectification column somewhat similar in design to a crude argon column, but with a very wide diameter in order to enhance rectification. In the latter, a portion of the liquid nitrogen is sent from the topshelf of the high-pressure column of the main plant, to the top of the reflux condenser of the rectification column. This piece of equipment serves as a noncontact condenser for krypton and xenon as well as some of the oxygen all of which have higher boiling points. The remaining liquid oxygen, now enriched with krypton and xenon, is withdrawn and sent to the sump of a third vessel where the liquid oxygen is reboiled by a stream of warm process nitrogen that is sent to the sump of the main lower (high pressure) column. The remaining liquid in the sump of this third column contains a maximum enrichment of krypton and xenon, about 1000–2000 vppm, and is ready for the final refining process (Table 4.7).


TABLE 4.7 Boiling Points of Atmospheric Gases Element

Boiling Point (at 101.325 kPa)

Nitrogen Air Argon Oxygen Krypton Xenon

77.3 K 80.9 87.3 90.2 119.8 164.9

Source: From Pauling, L. in General Chemistry, Periodic System of Elements, W.H. Freeman & Company, San Francisco, 1958. With permission.

4.4.4

REFINING

OF

KRYPTON

AND

XENON

The quantity of recovered krypton and xenon mixture to be refined even from a very large air separation unit is not very impressive. The economics of the refining process, therefore, depend entirely on the quantity of the mixture to be separated. If the quantity is very small, refining does not involve any cryogenic distillation. After the usual means of vaporization, combustion, and adsorption to remove any residual hydrogen and hydrocarbons, the krypton and the xenon mixture can be extracted by the adsorption and desorption in a silica gel trap enclosed in a bath of liquid nitrogen. For the further separation of krypton from xenon, a small cryogenic column is necessary using liquid nitrogen at the top reflux condenser, and the conversion of air to liquid air as the reboiling medium at the bottom of the column. If, on the other hand, the quantity of the crude product is sufficiently large, a cryogenic distillation tower may be desirable. In the latter process, after the mixture goes through treatment for contaminant removal and cooling, it enters a side column with a reflux condenser at the top, which is filled with liquid air. This air is liquefied by using a high-pressure air, which is precooled in an exchanger in countercurrent heat exchanger with the vaporized air from the reflux condenser. This precooled air enters the bottom of the auxiliary column and acts as a reboiler stream to the liquefied mixture of krypton, xenon, and oxygen. As the vapor rises to the top and through the tubes of the reflux condenser krypton, xenon, and some of the oxygen are recondensed and descend to the bottom where they concentrate to a mixture containing 80% krypton and 20% xenon. This mixture, after heated to a proper temperature by countercurrent heat exchanger, is sent ultimately to a high-pressure storage tank for further separation either locally, or for transportation to another site with better facilities for the separation. Even if the latter method is chosen, it may be necessary to carry it out on the day shift, or perhaps when the total quantity is sufficient, to make it viable from an economic point of view.

4.4.5

RECOVERY

OF

RARE GASES FROM AMMONIA PURGE GAS

While several units have been built for the recovery of argon from ammonia purge gas, they were designed in a period when argon was in short supply; since then none has been built. As for the recovery of krypton and xenon from ammonia purge gas, while technically possible, it is doubtful if the units can be economically viable. The average quantity of purge gas from a 1000 t=d ammonia plant is 8558 Nm3=h. From this quantity, there is a possibility of extracting approximately 0.938 Nm3=d of krypton, assuming a 100% recovery. This is equivalent to a


production of krypton from an air separation plant having an oxygen production of 240 t=d, which is considered as a very small plant in present-day standards.

4.4.6 APPLICATIONS

OF

KRYPTON

AND

XENON

As krypton began to be recovered, it has been replacing, or has been mixed with argon for filling incandescent lamps. Krypton lowers the rate of degradation of the filament; therefore, the lamps last longer. Furthermore, because of its heavier molar mass, a smaller quantity of gas is required. In fact, even before World War II, almost all lamps in Europe were filled with krypton. With the increasing cost of energy in North America, so-called as energy saving lamps are filled with an argon–krypton mixture. Krypton is also used in high-energy physics and lasers. Another interesting use is for stimulating plant growth due to the fact that a fluorescent lamp filled with krypton can be designed to emit light of wavelength ideally suited to stimulate photosynthesis. As xenon gives off light similar in brightness and spectrum to sunlight, it is also used as an electric flash in photography, and since the year 2000, car manufacturers have been using xenon-filled lamps for the front headlights. Airport landing lights are also filled with xenon because it gives off a bright and sharp light. It also has applications in space programs.

REFERENCES 1. Timmerhaus, K.D., and T. Flynn. 1989. Cryogenic process engineering (The International Cryogenic Monographs), pp. 25–30. 2. Gomonet, E. 1952. Les Tre`s Basses Tempe´ratures, Production et Emplois, Librairie J.B. Bailliere et Fils. Rue Hauteville: Paris, 138–139. 3. Ruhemann, M. 1945. The separation of gases (The International Series of Monographs on Physics), p. 221. 4. Ruhemann, M. 1945. The separation of gases (The International Series of Monographs on Physics), p. 222. 5. Ruhemann, M. 1945. The separation of gases (The International Series of Monographs on Physics), p. 261. 6. Timmerhaus, K. and Flynn, T. 1989. Cryogenic Systems, 2nd edn. Science and Business Media, p. 361. 7. Praxair Company Publication, 2005. 8. CryoGas International. February Issue, 2002, pp. 24–31. 9. Ruhemann, M. 1945. The separation of gases (The International Series of Monographs on Physics), p. 221. 10. Ruhemann, M. 1945. The separation of gas (The International Series of Monographs on Physics), p. 222. 11. Isalski, W.H. 1989. Separation of gases (Monograph on Cryogenics, 5), 34–35. Oxford: Oxford Science Publications. 12. Ruhemann, M. 1945. The separation of gases (The International Series of Monographs on Physics), pp. 69–70. 13. Gomonet, E. 1952. Les Tre`s Basses Tempe´ratures, Production et Emplois, Librairie J.B. Bailliere et Fils. Rue Hauteville: Paris 19:140. 14. Gomonet, E. 1952. Les Tre`s Basses Tempe´ratures, Production et Emplois, Librairie J.B. Bailliere et Fils. Rue Hauteville: Paris 19:141. 15. Gomonet, E. 1952. Les Tre`s Basses Tempe´ratures Production et Emplois, Librairie J.B. Bailliere et Fils. Rue Hauteville: Paris 19:142. 16. Gomonet, E. 1952. Les Tre`s Basses Tempe´ratures, Production et Emplois, Librairie J.B. Bailliere et Fils. Rue Hauteville: Paris 19:60–65. 17. Isalski, W.H. 1989. Separation of gases (Monograph on Cryogenics, 5), 96–100. Oxford: Oxford Science Publications.


FURTHER READING Krepsche, R.H., B.W. Birmingham, and D.B. Mann. 1968. Technology of liquid helium, National Bureau of Standards Monograph 111. Springman, H. 1982. Methods for larger recovery to meet increased demand for the argon market. AIChE Sym Ser 79 (224):12. Springman, H. 1985. Cryogenics—Principles and applications. Chem Eng 92 (9):59.


5

Front-End Purification Systems

5.1 HISTORICAL BACKGROUND From the beginning of the industrial application of air separation by low temperatures, it was obvious that contaminants such as water and carbon dioxide had to be eliminated. Later on, it became painfully evident that from a safety point of view other dangerous hydrocarbons, especially acetylene, also had to be removed. Present day list of dangerous contaminants that may be found in the atmosphere of a highly industrial area is given in Table 5.1. The concentrations indicated, however, are to be considered as a yearly average and not as maximum. The project engineer should make a thorough study of the local atmospheric conditions. The dangerous contaminants that may be found in atmospheric air used for the generation of oxygen include acetylene, which is the most dangerous, ethane, ethylene, propane, propylene, hydrocarbons in general, oxides of nitrogen especially nitrous oxide, and even acetone.

5.1.1 PROCESSES AND MATERIALS USED

IN

FRONT-END PURIFICATION SYSTEMS

Before any discussion of processes and materials that are employed in prepurification systems, it may be prudent to define some terms that are commonly used in the industry. 1. Absorption is the assimilation of molecules into a solid or liquid substance, with the formation of a solution or a new compound. 2. Adsorption is the adhesion of molecules of a gas, liquid, dissolved substance, or of particles to the surface of a solid. 3. Adsorbent is the material, which carries out adsorption. 4. Adsorbate is the gas, liquid, dissolved substance, or particles that become adsorbed by the adsorbent. 5. Regeneration of an adsorbent involves the action (generally by heat or release of pressure or both) taken at the plant site in order to recover the adsorbent’s capacity to adsorb. 6. Reactivation of an adsorbent involves the action required by the original manufacturer of the adsorbent to recover the material’s usefulness for adsorption and regeneration.

5.1.2 ORIGINAL PREPURIFICATION The original process and material used in air separation was absorption with sodium hydroxide (NaOH), which is a white hygroscopic (water attracting) solid that readily attracts water and becomes deliquescent. It was used for the preliminary drying of air. In a 15% solution, it was used for the removal of carbon dioxide. The process involved is two towers and is in series. The first tower contained a 15% solution of NaOH for carbon dioxide and oil removal, and the second tower contained solid NaOH for drying process air. When the lump


TABLE 5.1 Assuming a Site Elevation at Sea Level at a Pressure of 101.325 kPa, the Contaminants Contained in the Industrial Quality of Air Expected in vppm Continuous Average vppm Hydrogen Carbon monoxide Carbon dioxide Methane Acetylene Ethane Ethylene Propylene Propane Butane and heavier Sulfur dioxide Hydrogen sulfide Mercaptans Ammonia Oxides of nitrogen (NO–NO2) Particulate matter

10 1 400 10 1 0.2–5.5 0.1 0.2 0–0.1 0.1 0.1 0.05 0.1 1.0 0.1 2.5 mg=m3

As mentioned, the above analysis is generally found in a highly industrial area. For a more accurate analysis, the project engineer must make his own study. Source: Courtesy of F.G. Kerry, Inc., 2006. With permission.

caustic became completely deliquescent (saturated with water), and rendered inadequate for removing anymore water it was taken out and the tower is refilled with new lump caustic. The removed solution was brought up to a 15% level and was used in the first tower for the removal of carbon dioxide. Needless to say again the process was primitive, inefficient, and messy. Yet, it withstood the test of time up to the early 1950s.

5.1.3 5.1.3.1

ADSORBENTS General

Because the use of adsorbents is so closely allied with the technology of air separation, it will be useful to review some of the adsorbents in general use. These adsorbents have been in use since 1950 in various forms and temperatures for the removal of such contaminants as water, carbon dioxide, acetylene, and many other dangerous hydrocarbons, both saturated and unsaturated. Commercial adsorbents, which exhibit ultraporosity, have been used for the selective separation of gases, and included activated carbons, charcoal, activated clays, silica gel, activated alumina, and crystalline aluminosilicate zeolites. They must be evaluated with care, however, before any selection is made. With the exception of the last, namely the zeolites, the other materials do not possess an ordered crystal structure. Consequently their ˚ ). On the other hand, pores may range anywhere from 20 to several thousand angstroms (A commercial synthetic zeolites, or molecular sieves, have pores of very uniform size from 3 to ˚ , which are uniquely determined by the unit structure of the crystal2. 10 A


5.1.4 INTRODUCTION

OF

ACTIVATED ALUMINA

With the development of activated alumina in the early 1950s, there was a sharp improvement in efficiency. The system still used a 15% caustic solution for the removal of carbon dioxide, but was then followed by a dual tower system of activated alumina for desiccation purposes. The dual-activated tower unit generally operated on an 8 h time cycle. One tower was in adsorption, whereas the other tower was regenerated with heat. 5.1.4.1 Activated Alumina3 Activated alumina is nothing more than a pure form of aluminum oxide, completely free of water. The product is supplied in small spheres in 3–6 mm in diameter and is an inexpensive as well as an efficient material for drying gases, especially process air. It can be regenerated at a fairly low temperature to elute water, assuming there is enough sweep gas. The regeneration gas must supply a driving force to remove water condensed in the pores and adsorbed on the high surface area of the activated alumina. Whereas the heat of condensation of water is not a function of the desiccant, the heat of wetting is. Activated alumina has a lower heat of wetting than molecular sieve. It therefore has an advantage in requiring less energy for regeneration. For this reason a dual bed including both alumina and molecular sieve is more advantageous in terms of energy saving and capital cost for the combined removal of water and carbon dioxide from process air. As termed by ALCOA, the dynamic desiccant capacity (DDC) of activated alumina is based partly on experience. A rule of thumb is that in a clean and dynamic drying system, it equals one-third of the static capacity. As the relative humidity of the stream decreases, it passes through the entire bed and is not used to maximum static capacity before it is regenerated. According to ALCOA, the following DDC maxima in weight percent may be used as guidelines (conservative design uses about 3=4 maximum DDC but this may be further reduced because of severe conditions)3 (Table 5.2). 5.1.4.2 Regeneration of Activated Alumina Retained liquid water may cause breakage or fracturing of the activated alumina spheres, especially in the regenerated condition rather than the saturated condition. A gradual decline in adsorption capacity is typical in process air dryers, but this condition plateaus after a while. The decline is caused by a loss of surface area and micropore volume by repeated wetting and heating of the adsorbent. Given good conditions for air and inert sweep gases, the operating life of the activated alumina should be in the neighborhood of 10 years. This estimate, however, is reduced by half if saturated hydrocarbons are present, a condition more likely to occur in the operation of air separation plants in highly industrial areas. The total quantity of heat required for the regeneration of activated alumina can be calculated and is given in Table 5.3.

TABLE 5.2 Following Dynamic Drying Capacity of Activated Alumina May Be Used as a Guidance Drying Applications Inert gas and air Saturated hydrocarbons Unsaturated

H-156

F-200

F-1

15 11 8

12 8 5

7 6 3

Source: Courtesy of F.G. Kerry, Inc., 2006. With permission.


TABLE 5.3 Distribution of Heat in the Regeneration of Activated Alumina a) Heat to alumina desiccant Ha ¼ ma (Tr þ Ta ) Specific heat (Cpa ) b) Heat for water desorption Sensible heat of liquid water þ heat of vaporization þ sensible heat of water vapor þ heat of wetting Hd ¼ mw (Tv Td )Cpw þ mw DHv þ mw (Tr Tv )Cps þ DHw c) Heat to vessel Hv ¼ mv (Tr Td )Cpv d) Heat to bed supports, screens, if present Ht ¼ mt (Tr Td )Cpt e) Heat to hydrocarbons, if present Hh ¼ 1860 kJ=kg of hydrocarbons f) Total quantity of heat estimated with 70% safety factor H total ¼ (Ht þ Hd þ Hv þ Ht þ Hh)1.7 wherein ma, mass of alumina (kg); mw, mass of water (kg); mv, mass of vessel (kg); mt, mass of bed support (kg); Tr, temperature of regeneration gas (K); Td, temperature of inlet drying stream (K); Tv, temperature of vaporization for water desorbing from activated alumina, assume 395 K; Cpa, heat capacity of activated alumina 91.7 kJ=kg=K; Cpw, heat capacity of liquid water 4.16 kJ=kg=K; Cps, heat capacity of water vapor 2.08 kJ=kg=K; Cpv, heat capacity of vessel 0.41 kJ=kg=K (carbon steel); Cpt, heat capacity of bed support 0.41 kJ=kg=K (if ALCOA T-162 is used) Hw, heat of wetting maximum F-1, 32.5 Btu=lb or 76 kJ=kg of alumina H-151, 44.0 Btu=lb or 102.1 kJ=kg of alumina F-200, 43.0 Btu=lb or 100 kJ=kg of alumina Source: Courtesy of ALCOA Bulletin, 2000. With permission.

5.1.5

ZEOLITES (MOLECULAR SIEVES)4

The study of zeolite crystals dates back to 1840 when Damour reversibly dehydrated them without any change in transparency or external form. Grandjean later found that dehydrated crystals of zeolite could reversibly adsorb inorganic vapors such as iodine, mercury, and ammonia. In 1925, Weirgel and Steinhoff found that the zeolite chabazite could adsorb water vapor, methyl and ethyl alcohol but could exclude acetone and benzene. In the early 1930s, x-ray diffraction studies revealed the zeolites as crystalline materials having within each crystal a system of precisely arrayed cavities and pores. In 1932, McBain was the first to use the term molecular sieve to explain the significance of these. Molecular sieves retain adsorbates by a strong physical force somewhat similar to magnetism. In other words, when the adsorbed molecule is removed by either heat or another material the remaining crystal of the sieve is in the same chemical state as before. Desorption of the adsorbate from the sieve produces no hysteresis effect. The adsorption and desorption phases are completely reversible, and with a few exceptions the respective isothermal curves coincide completely. 5.1.5.1

Chemical Formula5

Synthetic zeolites are better than natural types because they are produced with closely controlled chemical composition and a pore size suitable for the molecules to be passed (or excluded) in the specific purification involved. Commercially produced zeolites have the following basic formula of the unit cell:


TABLE 5.4 Chemical Formulae of Various Zeolites Used in Industry Type 4A Type 5A Type 13X

Na12[(AlO2)12(SiO2)12]27H2O Ca4.5Na3[(AlO2)12(SiO2)12]H2O Na86[(AlO2)86(SiO2)106]276H2O

˚ Pore size 4 A ˚ Pore size 5 A ˚ Pore size 10 A

Source: Courtesy of UOP Molecular Sieve Adsorbents, 2005. With permission.

M2=n O Al2 O3 xSiO2 yH2 O where M is a cation of n valence. Among the types that have been found to be most commercially useful are given in Table 5.4 with their unit cell formula. In all cases, the water of hydration is removed by heating before the product is ready for use. The commercial material may also include about 20% inert clay as a binder. Molecular sieves separate molecules not only by size and configuration, but they also adsorb preferentially based on polarity or degree of unsaturation. The less volatile, the more polar, or the more unsaturate a molecule, the more tightly it is held within the crystal of the molecular sieve. For example, molecules of water and methanol have positive and negative electrical poles. These polar molecules are strongly attracted to molecular sieves. On the other hand, nonpolar molecules such as methane or ethane are more weakly attracted. When a mixture of water and methane is passed over a molecular sieve, the water is readily adsorbed, whereas methane is excluded, even though both are small enough to pass through the pores easily. Water can be removed from a strongly attracted fluid such as methanol by using a molecular sieve with pore openings smaller than the methanol molecule but larger than the water molecule. When molecular sieves remove a water molecule from a process stream the reaction is exothermic. The increase in temperature depends on the affinity of the absorbate for the molecular sieve. Conversely, when water is desorbed from a molecular sieve, the reaction is endothermic, resulting in a drop in temperature, which is a hindrance to desorption. A molecular sieve can also be regenerated by a reduction in pressure. The lower the pressure, the better the regeneration. In fact, in some cases, a vacuum is employed. 5.1.5.2 Types of Molecular Sieves ˚ in diameter. At usual The Type 4A molecular sieve has a free aperture size of 4 A operating temperature, this size allows the passage of molecules with the same effective diameter. Moreover, the sodium ion can be exchanged to form other useful products. For example, replacement of sodium ions in Type A with calcium ions will produce Type 5A with ˚ . This latter sieve is highly recommended to remove dangerous a free aperture of 5 A hydrocarbons such as propane from the liquid oxygen (LOX) in the main condenser if it has been entrained beyond the prepurification system at the front end. Type 13X, which is the most widely used molecular sieve in air separation technology, has a nominal pore ˚ . It adsorbs all material with a smaller pore diameter and excludes those with diameter of 10 A a larger pore diameter. The rate at which a given material will be adsorbed on molecular sieve pellets in any operation is dependent on four variables: 1. The rate at which the adsorbates can diffuse to the activated crystals within the pellets, which means that an efficient distribution system is of the utmost importance


2. The relative size of the molecules in the vapor to be adsorbed, and of the pores in the molecular sieve 3. The strength of the adsorptive forces between the molecular sieve and the adsorbate 4. Temperature The life cycle of the adsorbent will depend on the continuous retention of water and heavy hydrocarbons, such as benzene. With the proper regeneration of the sieve freeing it from the adsorbates at each cycle, the life of the molecular sieve may well last up to 15 years. If water is allowed to accumulate, however, fracturing may ensue, and the sieve may last only 5 to 10 years. The same is true if hydrocarbons are permitted to coke the sieve. The life cycle of a molecular sieve, especially that of Type 13X, can be extended substantially by the use of a layer of activated alumina upstream of the sieve. Activated alumina has a very high affinity for water and for very heavy hydrocarbons. This procedure will shorten the life cycle of the alumina but will extend that of the molecular sieve, which has a higher capital cost.

5.1.6 SILICA GEL Silica gel is a colloid of silica and resembles very close in nature to the zeolite minerals. It has a structure, which contains corridors and chambers within its aluminosilicate framework, as well as alkali and alkaline earth ions. When a mineral such as chabazite CaAl2Si4O126H2O is heated, water molecules are driven off, but the crystal structure does not collapse. Therefore, the dehydrated mineral has a strong affinity for water, and for molecules of other vapors such as acetylene. Silica gel has long been in use as a drying agent for air conditioners, packing, etc. In 1950, it was also found to be very efficient as a filter trap in the removal of acetylene from liquid air and LOX at cryogenic temperatures. Tests on various drying agents are given in Table 5.5. As noted, using activated alumina as a base reference, the ordinary silica gel is slightly superior by 7% on an equal volume basis to the manufactured Mobil Sovabead, and very superior by 50% to Sovabead on an equal weight basis. The manufactured Sovabead has certain advantages, however, in that it can be handled easily, and does not dust, which is much appreciated by the operating departments. Results of thermal shock tests, that is, changing the temperature between 383 and 78 K in 10 cycles, indicated that there was very little loss of both, the Sovabead and the standard silica gel. The loss was around 0.003% for both. Results of mechanical vibration tests, however, indicated that the Sovabead was slightly superior to silica gel, and both were greatly superior to activated alumina (Table 5.6).

TABLE 5.5 Comparison of Various Drying Agents

On an equal weight basis Adsorption in LOX Desorption at 373–388 K On an equal volume basis Adsorption in LOX Desorption at 373–388 K

Alumina (%)

Sovabead (%)

Silica Gel (%)

100

103

154

100

117

125



TABLE 5.6 Screen Sizes of Typical Drying Agents

Retained on 50 mesh screen At warm temperature At cold temp.

Alumina (%)

Sovabead (%)

Silica Gel (%)

91.38 99.831

99.924 99.967

99.68 99.96


5.2 DESIGN OF CURRENT FRONT-END PURIFICATION SYSTEMS 5.2.1 GENERAL BACKGROUND6,7 For more than the past decade, the use of a front-end prepurification (FEP) system with adsorbents has played a dominant role in the design and construction of oxygen plants, both large and small. In fact, this option has pushed the previous use of reversing heat exchangers (Revex) into the background, the latter being mostly limited to small units less than 200 t=d, and to nitrogen-generating units. This is the case even though the present Revex cores are fabricated under vacuum and are more reliable. Overall, the FEP system has many advantages: it is more stable in operation; it is safer in providing a more thorough elimination of hazardous hydrocarbons; and it facilitates a faster turn around in periods of total defrosting, or for maintenance. On the other hand, one must accept a somewhat higher investment cost, and a marginally higher energy consumption. In contrast, Revex systems have not been able to recover as much of the air feed as pure products (only about 45%), but require 5%–10% less energy. During the 1950s and 1960s, Revex had pushed FEP systems into the background, especially for chemical and metallurgical plants that had limited use for pure nitrogen. In the early 1970s, however, a change of scale occurred in petrochemical plants: the amount of oxygen required increased substantially to a magnitude of 1000–2000 t and more per day. The increasing size of air plants required an arithmetic increase in the number of unit aluminum cores needed for Revex systems, and their application became expensive. Even more significantly, however, field performance of the large cores suffered. The old salt-bath brazing process, coupled with difficulties in maintaining quality control during the cores’ manufacture, made their use for large air plants suspect even though vacuum brazing of the aluminum cores had started to overcome this problem. Throughout this period, however, designers of prepurification units had exerted themselves to reduce the differential in capital costs and energy requirements. They had lowered pressure drop across the adsorbent bed, reduced regeneration temperature of the absorbents with the addition of activated alumina upstream, and eliminated the mechanical chiller whenever a large quantity of waste nitrogen or other inert gas was available. With the use of an evaporative water chiller (EWC) operating in conjunction with a direct contact aftercooler (DCA), it was possible to chill process air down to 283 K, which is the optimum. However, this required the use of about 50%–60% of the total process air as waste nitrogen. With regard to adsorbent vessels, pressure drop has been lowered to less than 6.9 kPa, by minimizing the bed depth and increasing the cross-sectional area in order to maintain bed volume. There is no easy solution to this problem, however, which has resulted in a variety of design configurations. It was also discovered that by using 15%–25% of the total air feed it was possible to regenerate the adsorbent bed at a reasonably lower temperature. Where high peak load demand was a problem using an electric heater for regeneration, it was also possible to use a heat accumulator to eliminate such peaks, or use the heat capacity of the processed air that is compressed (see Section 5.3.5).


5.2.2

EQUIPMENT USED

The equipment currently involved in the design of FEP systems may be summarized as follows. 5.2.2.1

Precooling Units Upstream of Adsorption

For small plants, 300 t=d of oxygen and under, a standard shell and tube aftercooler downstream of the main air compressor may prove more cost effective, depending on local conditions. 5.2.2.2

Direct Contact Aftercooler (Figure 5.1)

The design of the precooling unit may be modified depending on local conditions such as air temperature, relative humidity, feed air capacity, and cooling water temperature. For example, a noncontact heat exchanger aftercooler may be utilized, thereby eliminating the lower half of the DCA tower. On the other hand, if the availability of waste nitrogen is too low, then the use of a mechanical refrigeration unit may be unavoidable. Another point to consider at the outlet of the DCA unit is a well-designed and reliable mist eliminator in order to avoid serious entrainment of water into the adsorbent beds. The material of this mist eliminator should be stainless steel or plastic, preferably with overlapping layers. If horizontal sections are supplied they should be tightly fastened to avoid any open spacing between the sections to avoid any water entrainment. As shown in Figure 5.1, a DCA consists of two sections in a single vertical column. The lower (high pressure) section receives process air directly from the last stage of the main air compressor at a high temperature. It may be trayed or packed with rings. Polypropylene should not be used because of the high temperatures encountered at the bottom of the tower. Molecular sieve adsorber Process air 100,000 Nm3/h Temperature 8⬚C saturated CW—43 m3/h Temperature—6⬚C Waste N2—57,000 Nm3/h Temperature—30⬚C

CW—42.3 m3/h Temperature —11⬚C

LIC

Waste N2—57,000 Nm3/h

1.63 m3/h

Temperature—5⬚C dry

F1

Cooling water Process air Flow—100,000 Nm3/h Temperature—87.6⬚C

LIC

CW—111.5 m3/hr Temperature —51.26⬚C 6⬚C

5⬚C

CW—111.5 m3/h Temperature—51.26⬚C

C W—110.7 m3/hr Temperature—32⬚C DCA

LIC

Cooling water EWC

FIGURE 5.1 Typical cooling system for a front-end purification system. (Courtesy of F.G. Kerry, Inc. With permission.)


Field reports have indicated that such rings have fused when exposed to the heat. Corrosion resistance is important in this section. All internals should be stainless steel. The internal walls of the tower should also be coated to prevent or retard corrosion. Two separate liquid level indicators must be used, each with an alarm, and each with an independent connection with the tower. An abnormal rise in water level would permit water entrainment into the adsorbent vessel, thereby damaging the adsorbent. (Entrainment can be permanently damaging to the adsorbent if the cooling water is treated with a too high concentration of chemical additives.) The top section of the DCA tower is mechanically similar to the bottom section, but uses a closed water circuit in series with an EWC. Depending on makeup, cooling water enters at the top at about 279 K. The process air leaves the tower at approximately 281 K. If a mechanical chiller is used, any temperature lower than 299 K is a waste of energy. If the temperature is much over 283 K a larger quantity of molecular sieve and activated alumina will be needed to prevent a CO2, or even worse, a propane breakthrough. If conditions permit, one should use an open circuit in the upper section of the DCA. As shown in Figure 5.1, two cooling water injection pumps are required, one in operation, and the other a 100% hot-wired standby spare, with all required instrumentation and controls. Two independent pressure drop indicators should be used to further protect against any entrainment of water. The upper section of the DCA tower is piped in series with the EWC. 5.2.2.3 Evaporative Water Chiller The EWC is nothing more than an open vessel packed with rings, trays, grid, or other suitable packing. Dry waste nitrogen from the cold box enters at the bottom. This nitrogen can be a major portion of the air treated in the cold box. For instance, an oxygen plant processing 100,000 Nm3=h of air will require 50,000–60,000 Nm3=h of waste nitrogen for the evaporative chiller. Recirculated water from the DCA tower or from the water-cooling tower flows from the top countercurrent to dry, cool waste nitrogen from the cold box. Alternatively, if makeup water is at a low enough temperature, an open water circuit may be used. Care must be taken, however, to avoid reaching too low a temperature in the EWC, otherwise the recirculating pump and even the packing may freeze. This has already happened in the field (Figure 5.2). The design should include stainless sieve trays or adequate polypropylene ring packing. It should also include a storage section at the bottom of the vessel to serve as a surge tank for the chilled water system. The capacity of this storage should be enough to hold at least 3 min of water storage at the maximum chilled water flow. Two chilled water-circulating pumps should be supplied; one in operation, and the other piped and installed hot-wired as a complete 100% spare. The usual shutdown alarms should be considered for high and low levels. As indicated in Section 5.2.2.3, two independent water level indicators with separate attachments to the tower are recommended. 5.2.2.4 Mechanical Refrigeration Unit (Figure 5.3) If it is decided to use a mechanical refrigeration unit, its selection should be based on high reliability and proven field service. A centrifugal or a screw-type compressor is preferable. The refrigeration fluid shall also conform to EPA standards such as HCF 134a (1,1,1,2tetrafluoroethane) or equal.

5.2.3 ADSORBER UNIT6 During the past decade, there has evolved a variety of configuration designs, but aside from the standard two-vertical vessel design, the basic few options that have proven viable in the field are worth a brief mention.


AC/DC system (Airplant flow = 12,000 lb mol in.) 75 72.5 70 67.5 65

Water temparature (F)

62.5 60 57.5 55 52.5 50 47.5 45 42.5 40 37.5 1000 (b)

2000

3000 4000 5000 6000 Waste N2 flow (lb mol/h)

7000

8000

FIGURE 5.2 (a) Closed air-cooling water system. (ß Air Liquide, all rights reserved, 2006. With permission.) (b) Ideal cooling water temperature using waste nitrogen. (Courtesy of Bernstein, J., Cryogenic Consulting Services, Inc., 2006. With permission.)


Energy (kWh)

Conditions—air compressed to 6.5 bar Cooling water at 20⬚C

(C) (B)

(A) 30⬚C

25⬚

20⬚

15⬚

10⬚

5⬚

0⬚

(A) Shows the increase in energy for every degree in cooling of compressed air, using external refrigeration. (B) Shows the drop in energy requirements by using a DCA (C) Shows the new saving of energy between the two systems. After the compressed air reaches around 8⬚C there is no further saving of interest

FIGURE 5.3 Comparison of air cooling using mechanical refrigeration and a direct contact aftercooler. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

5.2.3.1 Standard Design The so-called standard design with dual vertical vessels is still used for oxygen plants up to a capacity of 300 t=d, or to a process airflow of 45,000–50,000 Nm3=h, assuming that the air pressure is at approximately 5–6 bar absolute, and that feed temperature is around 283 K. This basic design may still be used for larger capacities of oxygen production if the process air pressure is increased, for example, in the application of oxygen in integrated gas generation plants. In the latter case, however, one must take into account the critical velocity of the process air flowing through the adsorbent beds. A number of manufacturers still turn out prepackaged units of the basic design as illustrated in Figure 5.4a, which operate satisfactorily in fairly large capacities. Two vertical vessels are mounted on a skid module that contains all necessary switch valves and timers. These vessels are alternately on- and offstream: the off-stream was used to elute water, CO2, and various hydrocarbons including acetylene that have been removed from the process airstream during the adsorption phase, whereas the on-stream vessel is on adsorption. 5.2.3.2 Multiple Vertical Vessels As shown in Figure 5.4b, it has been possible to add two extra vessels operating in conjunction with a single set of switch valves. With this design, it has been possible to increase the capacity of an existing air plant from 300 to 600 t=d with no operating problems whatsoever.


Waste nitrogen vent

PIC

Waste nitrogen from cold box Reactivation heater

T1

Process air to cold box

Vent P1

(b)

T1

Process air from cooling system

FIGURE 5.4 (a) Standard twin vertical vessels with horizontal adsorbers for a 300 t=d oxygen plant. (ß Air Liquide, all rights reserved, 2006. With permission.) (b) Basic vertical-vessel adsorber design. (Courtesy of F.G. Kerry, Inc., 2006. With permission.) continued


Waste nitrogen vent

Waste nitrogen from cold box Reactivation heater Process air to cold box TI

Vent

(c)

Process air from cooling system

FIGURE 5.4 (continued) (c) Basic adsorber unit of vertical vessels with possible option for multiple vessels. (Courtesy of F.G. Kerry, Inc. With permission.)

Another variation of this four-vertical vessel design as shown in Figure 5.4c is to have the four vessels operating in a sequence. This design has been a successful option for many years, and has been used for units of airflows up to 125,000 Nm3=h operating at 7 bar, that is, a production of 700–800 t=d. Whereas this option involves a higher number of switch valves, the valves are of a much smaller diameter, resulting in a lower total purchase cost. The vertical vessels, mainly of a smaller diameter, also have a thinner wall thickness, therefore cheaper to roll and to fabricate. Moreover, equipment such as vessels, piping, switch valves, and timers, can be factory assembled into one prepackaged skid-mounted module complete with instrumentation, tested, and shipped to the site as a single unit. It also has an operating advantage because although it has more operating bumps in the upper (low pressure) distillation column than the two-vertical vessel adsorber design (due to the operation of the multiple switch valves), the bumps are much less severe, and the fractionation is therefore somewhat smoother. This is apparent in the resultant higher recovery of argon. Another option that has been used successfully in the field is a three-vertical vessel design, also operating sequentially. It has been applied in a plant producing 1500 t=d of oxygen. 5.2.3.3 Horizontal Vessels (Figure 5.5a) Several designers have opted for this configuration with satisfactory results, at least for very large plants (over 1000 t=d). A horizontal vessel orientation, however, increases the mechanical complexity of the grid platform supporting the adsorbents. Not only must the grid support the dead weight of the molecular sieve, but it also must withstand the cyclic thermal stresses resulting from temperature excursions from 423 to 523 K, down to 283 K, and


LBOUT

LEGRI

LGFIL

A

B

DRPIQ DRFIL

Air NR

DIBOU

B Steve

Molecular

Activated alumina

LGGRI

Air

NR

r su

Filtra t

A

ion

Cut B.B

(a)

FIGURE 5.5 (a) Molecular sieve adsorber horizontal vessel and horizontal beds. (b) Torn upper screen filter due to excessive fluidization. continued


FIGURE 5.5 (continued) (c) Disturbed surface of molsieve bed due to excessive fluidization. (Courtesy of F.G. Kerry, Inc. With permission.)

back again. In addition, the total assembly has to survive dynamic stresses caused by depressurization and repressurization from 6–7 bar to atmospheric pressure, and back again. Obviously, various members making up the grid will move by thermal expansion and contraction, but the movement must be such that small granules or beads of adsorbent will not fall through the grid or be ground to dust. The quantity falling through the grid must be kept extremely small—no more than a few grams per day. Another problem, but by no means less difficult, is how to distribute the process air uniformly along the large cross-sectional area of the bed. The solution chosen originally by almost all engineers was to let the process air enter from a bottom nozzle at full pressure, assuming that one simple wide baffle plate would distribute it equally across the area of the bed. They ignored a key fact, as memorably stated by an experienced designer, Layton Kitterman, ‘‘Fluids flow according to the laws of physics, and not according to the wishes of the designer.’’ Only in a few cases uniform flow has been obtained at a velocity that did not fluidize the bed, and then only when adjustable guide vanes or multiport inlets were employed. These solutions were only moderately successful, and also added significantly to the capital cost. Figure 5.5b and Figure 5.5c show results of fluidization. 5.2.3.4 Radial or Concentric Design7 Still another option that has been introduced successfully during the past decade involves two vertical vessels each with two concentric adsorbent beds (Figure 5.6). Each adsorbent bed is held together between two vertical cylindrical screens suspended from the top of the tower. Process air enters at the bottom and is directed toward the annular space at the wall of the vessel. It then travels horizontally from the wall to the center of the vessel. The process air collects in the central core of the vessel, and exits at the top. Waste nitrogen for regeneration enters centrally at the top of the tower, travels horizontally first through the molsieve, then through the activated alumina, and exits at the bottom. The advantages are obvious. The screens expand and contract freely during thermal cycles. The adsorbents can be installed in a thin bed, and tower height can be made as tall as necessary for the required


Air

N2

n Radial flow double bed Alumina Molecular sieve

N2

Air

FIGURE 5.6 Vertical radial flow vessel. (ß Air Liquide, all rights reserved, 2006. With permission.)

capacity. Only simple bottom and top baffles may be needed to make air and gas flow uniformly. No guide vanes are required. Yet all is not ideal. Because the design imposes some rigid and uncompromising procedures for fabrication, labor costs are high. Therefore, the concept may be expensive for small plants. Although units have indeed been built for plants as small as 600 t=d of oxygen, for those below 800 t=d, economics may become questionable. Each case has to be studied individually. In summary, one may say that up to 700 t=d, the market belongs to either the two-vertical vessel design or the prepackaged four-vertical vessel configuration, the latter operating in sequence. For oxygen units with a capacity of 700 up to 1500 t=d, the designer may choose between the three-vertical vessel option, the horizontal vessel configuration, or preferably, the radial design. Beyond 1500 t=d, the radial design has a very definite advantage in all respects.

5.3 PROCESS OPERATION 5.3.1 ISOLATION VALVE DOWNSTREAM

OF

FEP (FIGURE 5.7)

It is of the utmost importance to have an isolation valve between the adsorbent system and the cold box to make sure that the main process airflow is introduced slowly into the equipment, especially after a temporary shutdown when cryogenic liquids are still in the distillation columns. During initial start-up it would be foolhardy to subject the equipment



HIC 225

T1 240

ZIC 225

FIG 427

HV 225

HIC 223

From adsorber vessels

Legend DCS shared display in control room Discrete instrument in control room Shutdown signal

ZIC 223

1"

HV 223

TI

Temperature indicator

FIC

Flow indicator control

ZIC

Position indicator control

HIC

Hand indicator control

Modulating valve

HV

Hand valve

Check valve

Note:

Main process control valve does not open until bypass valve allows pressure down stream to build up to a permissible level

Permissive signal Solenoid

Venturi flow meter Piping insulation

FIGURE 5.7 Control system at discharge of adsorber vessels. (ß Air Liquide, all rights reserved, 2006. With permission.)

To cold box

to the full pressure and flow of process air instantaneously. Secondly, after a brief shutdown due to a compressor stoppage, the liquids on the trays in the distillation columns will dump to the bottom and create a large cold sink. If in this situation a large flow of air is permitted to enter the lower column, it will condense and create a partial vacuum. This in turn will increase the airflow velocity, which will create fluidization in the adsorbent beds, as well as upset the air compressors. This condition can be controlled by the use of the isolation valve with a bypass system between the FEP unit and the cold box. Alternatively, the isolation valve can be programed so that if there is an emergency shutdown, the valve will shutdown immediately, preventing cold air from backing into the warm section. During the restart period, warm airflow will ramp-up slowly in carefully controlled increments until the downstream pressure is almost equal to the design upstream pressure.

5.3.2 ADSORPTION KINETICS The adsorption kinetics of CO2, water, and hydrocarbons in process air on molecular sieves is not as simple as just letting the adsorbent in one vessel become loaded with contaminants, and then switching over to the next regenerated unit. Contaminants in the bed are in continuous movement, and the movement of each varies precisely according to the affinity of the adsorbent for it. Concentrations are continuously changing. As shown in Figure 5.8a, each component will establish a wave front and these will move as components and are displaced by more adsorbed material (Figure 5.8b and Figure 5.8c). Adsorption of hydrocarbons by molecular sieves generally increases with increasing molecular weight in a homologous series. Unsaturated compounds are more avidly adsorbed than alkanes. The affinity of a commonly used molecular sieve for various compounds of ˚ decreases in order: water, butanes, propylene, acetylene, molecular size less than 10 A ethylene, propane, CO2, ethane, and methane. Ammonia, ethanol, and mercaptans are also adsorbed more strongly than CO2. This does not imply that contaminants are totally removed. For example, this molecular sieve takes out only 89% of propane and 97% of ethylene. Methane is hardly removed at all, and exits with the process air. With the exception of the strict specification for aviation oxygen, methane poses no problem for industrial oxygen uses. The adsorption of water by molecular sieves is exothermic, and as the bed temperature increases, therefore, the adsorption of CO2 decreases. Conversely, the desorption of water is endothermic, so that more heat is required for desorption. Also note that if too many liquid water particles are entrained into the bed, adsorption will be impaired and fracturing of the beads may occur. The resulting dust may enter the primary exchanger system where it will be difficult if not impossible to remove. In one case, for an oxygen unit (1200 t=d), after 20 years of operation, the dust is still in the primary exchangers. During the adsorption phase, it is important, not only to make sure that the process air is diffused evenly over the entire surface of the bed, but also that its flow velocity through the bed is less than 20 cm=s, preferably closer to 10 cm=s. From empirical studies in the field, the minimum quantity of molecular sieve (Type 13X) is 50 kg per 1000 Nm3=h of process airflow per 1 h of adsorption. In circumstances where the atmosphere is heavily contaminated with propane it may be wise to increase the quantity of adsorbent. Unfortunately, this suggested increase is still in the realm of guesswork, but an increase of 10% is a reasonable minimum.

5.3.3

REGENERATION CONCERNS

The complexity of the complete regeneration of a molecular sieve adsorbent has not been fully appreciated by many designers. Prepurification units for smaller plants were often designed for 8 h duty, which fitted well with the 8 h labor shifts. Regeneration temperatures were


Position of hydrocarbons at the beginning

Position of hydrocarbons at the end

C3H8 C2H6

C2H2 C3H6 C4 H2O

C2H2 C3H6 C4

CO2

C2H6

Concentration

Length of the adsorber bed

CH4

CH6 C2H6

C3H8 C2H4

CO2

H2O

(a)

Flow direction

Vppm C2H6 in main condenser 200 180 160 140

Without LOX adsorber With

120 100 80 60 40 20

h

0 (b)

1.0

2.0

3.0

4.0

FIGURE 5.8 (a) Position of hydrocarbons in the adsorber bed. (b) Increased C2H6 concentration in the LOX of the main condenser at a content of 10 vppm C2H6 in the process air. continued

generally around 523 K. With the advent of large air separation plants, however, and with the introduction of automated control, such designs are no longer viable. Since large oxygen plant operators do not have a need for large volumes of pure nitrogen, this offers a designer the opportunity to use waste nitrogen to supply large quantities of heat for regeneration, without using high temperatures. It is now also possible to use waste steam at 3 bar and 407 K, or even heat from the intercoolers of the main air compressors at 353–363 K. The volume of waste nitrogen required for regeneration will be 15%–25% of the total process air, depending on regeneration temperature chosen.


6

5

C3H6

Molecules in each cavity (number)

C2H4 C2H2 4

C4H8⫺1

3

2

C3H8

1

C2H6 CH4

0 100 (c)

200

300

400

500

600

700

Pressure (mm)

FIGURE 5.8 (continued) (c) Adsorption of hydrocarbons by zeolites is much greater for unsaturated hydrocarbons whose molecules contain double or triple bonds. From top to bottom the curves show adsorption (at 1508C) of propylene, ethylene, acetylene and isobutylene (unsaturated), and propane, ethane, and methane (saturated). (Courtesy of Linde BOC Process Plants, 2005. With permission.)

The lower the regeneration temperature, however, the more difficult it may be to remove the adsorbed water. Residual water will build up and impair CO2 adsorption very quickly. It is prudent, therefore, to install an auxiliary heater, preferably electric even though it may be used only once a year or even less to prevent excessive accumulation of water. Since water is more strongly adsorbed than any other contaminant, it is a prime suspect when operational problems occur in prepurification systems. To minimize water’s deleterious effects, the following options are recommended: 1. Use a well-designed DCA with high contact efficiency and low entrainment. 2. Employ a good mist eliminator at the top of the aftercooler, or an efficient water separator, or both. 3. Eliminate all the water before the air feed reaches the adsorbent bed by using activated alumina or an equivalent material as a predryer immediately upstream of the molecular sieve bed. This recommendation is very important. Using an activated alumina predryer will also eliminate the harmful effects of sulfur dioxide, which can be entrained by water. Should entrainment of acids be a problem, a molecular sieve suitable for removing them should be installed upstream of the main adsorbent. Experience


with dual beds has shown that the life of the main adsorbent can be extended to well over 10 years. Furthermore, as water can be desorbed from alumina with a low heat input, a lower total quantity of heat will be required for a dual-bed system than for a single-bed unit. A final and not unimportant advantage of the dual-bed system is that less molecular sieve is needed in the main bed. With regard to the amount of activated alumina needed, one has to calculate the quantity of moisture contained in the main process airflow leaving the DCA tower, assuming a state of saturation at its exit temperature and pressure. In proportion to the mass weight of the molecular sieve, the quantity of alumina may be as low as 25%, or as high as 75%. These figures are entirely empirical, as no one has developed an algorithm to equate the necessary factors.

5.3.4 WARNING

AGAINST

EXCESSIVE HEAT

DURING

REGENERATION

The heater or the heating element should be switched off before the outlet temperature of the waste nitrogen reaches its maximum. The heat front will continue working its way through the bed (pulse heating). This will also save energy. If the heater is shutdown too late, however, there will not be sufficient time to cool the bed. After late switching, the process air will continue to flow through a still hot alumina bed, and it will not adsorb water effectively. This condition will load up the molecular sieve with water and reduce the efficiency of the molecular sieve for adsorption of CO2, C2H2, and other dangerous contaminants.

5.3.5 HIGH-PRESSURE VESSEL REGENERATION Another factor for regeneration not yet fully appreciated is the effect of pressure, especially in the range of 10–20 bar. At such high process air pressures, the thick walls (25–50 mm) of the steel vessels containing the adsorbents act as a heat sink, reducing the overall heat energy of the regeneration gas. This condition gets worse as the regeneration gas reaches the alumina at the lower level, and will impair elimination of adsorbed water. As noted, if water is allowed to accumulate in the adsorbent it will harm the overall adsorption phase. To remedy this condition, it may be necessary to incorporate an insulating material (possibly up to 50 mm) within the vessel to minimize the escape of heat through the vessel walls. To eliminate this condition completely it is advantageous to use the radial design so the regeneration gas first passes through both adsorbents beds before contacting the vessel walls.

5.3.6 REGENERATION OPTIONS

FOR

FEP UNITS

5.3.6.1 General If heat is to be used for regeneration of the adsorbent beds, any source of heat can be used: electricity, natural gas or propane, high- or low-pressure steam, hot water from the compressor coolers, or even heat from the compressed process air before the final aftercooler. What is important is constant energy availability and temperature stability. With regard to supply and degree of control, experience has shown that electricity is the most reliable source. Highpressure steam, if available, requires a fairly complex system of supply and a high degree of maintenance. Waste steam at 3 bar is more than hot enough to fit the need. In 1989, Isalski8 published an equation to include the heating and cooling times for a given regeneration gas flow: the equation is as follows: u ¼ th þ tc þ tx


(5:1)

and tc ¼ (WB CPB )=(GCPG ) ln{(T1 T3 )=(T2 T3 )}

(5:2)

where u is the total on-stream time for 1 bed (1=2 NEMA cycle time), th is the heating time, tc is the cooling time, tx is the changeover time (including depressurizing, equalization, pressurization), Wb is the weight of bed, CPB is the specific heat of bed, T1 is the initial bed temperature, T2 is the required bed temperature, and T3 is the coolant temperature. The above equation, however, does not take into consideration the amount of water or any other contaminants on the bed because it quantifies the cooling period after the water has been driven off. The quantity of the latter may be as high as 35%. Moreover, it does not include such factors as high peak demands if electricity is used, the cost of fuels such as propane if electricity is out of question, and the influence of the metal mass (heat sink) of the vessel itself, especially if the process air is delivered at a high pressure. If a high temperature above 425 K is specified, the use of a heat accumulator has been found useful to reduce energy costs. Because the regeneration phase requires only 25%–35% of the connected cycle time, average energy consumption should be about 35% of the connected energy rating. Therefore, the use of a heat accumulator has been especially economical when employing electricity for the heating phase. The accumulator is a vessel packed with either copper metal or quartzite stones. To avoid high peak demand penalties, designers have used heat accumulators in series with electric heaters. They have been used for air separation units with a capacity of up to 600 t=d of oxygen. This fixture has long been replaced by the use of a standard heat exchanger also called an economizer to provide a countercurrent heat exchange between the hot process air from the main air compressor and the cold waste nitrogen from the primary heat exchangers of the cryogenic unit (Figure 5.9). The other factor involved the use of higher pressures for the main process air, which increases the thickness of the vessels, if the standard design is used. This problem has been discussed in Section 5.3.4, and in these circumstances one may have to apply an insulation material between the inside surface of the vessel and the adsorbent material. Another system is to use multiple adsorbent vessels, as previously mentioned, and to operate each vessel in a well-defined sequence. Waste nitrogen being discharged from vessel A being regenerated flows through and warms up vessel B that has already been depressurized,

Preheated waste N2 to main heater

Cold waste N2 to heat economizer

Heat economizer

Warm process air from air compressor

Cooled process air to DCA lower section

FIGURE 5.9 Using heat from air compression to preheat waste nitrogen for the regeneration of adsorbers. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


FIGURE 5.10 Multivessel absorption system. (Courtesy of Chemical Design, Inc., 2005. With permission.)

and so on. Generally four vessels are used for this system, each one operating at a different cycle phase. In this manner no heat from waste nitrogen leaves the system (Figure 5.10 and Figure 5.11). Three vessels operating sequentially have also been used. A thermal pulse type

Operational sequence of FEP system at GPCG Adsorption--- 1.8 h Note: The sequence is staggered 0.52 hours per Depressure---0.1 h vessel in order to maintain a constant flow Heating--------0.52 h of heat and thus conserve energy. This factor Cooling--------0.95 h is an equal divider durring every four sequences. Depressure---0.15 h (The final timing has not yet been settled in the field.) 7/04/84 Total------------3.52 h 0:0

1:0

2:0

3:0

4:0

5:0

6:0

7:0

Hours

A H C A

DP PR

A H

C A

A H C

A

A H C A

FIGURE 5.11 Generation of multivessel absorption system. (Courtesy of Chemical Design, Inc., 2005. With permission.)


of regeneration has also been used to economize energy consumption. Regardless of the regeneration temperature selected, it is possible to adjust the heating element so that it is shut off before flowing hot gas reaches the end of the adsorbent bed. The heat wave, or pulse, will continue to move through the bed, and regeneration will be complete. An energy saving of approximately 25% may be realized with this procedure. Of course, the timing of the sequence has to be precalculated precisely. Another means of saving energy is to implant analytical probes into the adsorbent beds and monitor the CO2 already adsorbed. The regeneration phase can then be started only when the CO2 reaches a precalculated concentration, not on a time cycle. In all of those options, it is assumed that a dual adsorbent system is used with an activated alumina bed located upstream of the molsieve bed. One may conclude, therefore, that there is no single equation that can encompass all the variations to regenerate zeolites with an optimum economy.

5.3.7 SUMMARY In summary, two criteria must be satisfied in the design of successful prepurification units for regeneration purposes: 1. Regeneration gas must be diffused uniformly over the entire surface of the bed to prevent accumulation of residual water and the dusting of the sieve bed. This point demands great emphasis, as it has been a major problem for almost all designers. 2. Regeneration gas must have sufficient pressure to overcome the pressure drop of the absorbent bed, heater, piping, and valves. Otherwise, back pressure will develop in the upper, or low pressure, column of the air separation unit, thereby lowering the distillation efficiency. This problem can be overcome either by increasing process air pressure slightly, by using a small booster blower for waste nitrogen (which is generally anathema to plant operators), or by increasing the number of passages for the waste nitrogen stream in the front-end heat exchangers. In any case, there will be a slight increase in the total energy consumption or investment. 3. For large air separation plants, over 600 t=d of oxygen, consider the use of the radial design, because it has a low pressure drop.

5.3.8 OPERATIONAL TIME CYCLE (FIGURE 5.12) The total operational cycle, also known as the NEMA cycle, should cover both the adsorption and desorption phases, and should include . . . . . . . . . .

Equalizing pressure in both vessels Opening flow in the second vessel Isolating the vessel to be regenerated Depressurizing the vessel to be regenerated Heating waste nitrogen to the required temperature Passing hot waste nitrogen through the adsorbent bed Cooling the regenerated vessel to ambient temperature with cool, dry waste nitrogen Pressurizing the vessel Diverting process airflow through the regenerated vessel Using the prepared vessel for a complete adsorption cycle

All the steps are sequentially calculated, and then modified to suit field conditions. If they are programed and computerized, it is wise to install a manual override in the system. The total


Both vessels in operation

Vessel A Operating

Pressure

10 Bar

1 Bar

Heat

Cool

120⬚C

30⬚C

Regeneration time 240 min

480 min Time

FIGURE 5.12 Typical regeneration heating time 185 min minimum. Cooling time 118 min and depressurization 55 min. (ß Air Liquide, all rights reserved, 2006. With permission.)

NEMA cycle time may vary between 4 and 16 h. A realistic choice is 4 h, because it reduces the quantity of adsorbent needed yet does not increase maintenance cost unduly. In the 4 h cycle, 2 h of adsorption and 2 h of regeneration take place. The final choice, of course, depends on the supplier’s design, and experience with similar plants.

5.3.9 PREPURIFICATION ADSORBENT UNITS AND OPERATING STABILITY Stability of operation of the adsorption system is of the utmost importance. Because concentration of atmospheric contaminants continually fluctuates, the system imposes a need for constant vigilance. Its design involves the continuous and complete removal of not only carbon dioxide (for which it is basically designed) and water, but also of dangerous contaminants such as acetylene, ethane, ethylene, propane, and other hydrocarbons. Removal of CO2 is monitored either by an in-line gas chromatograph or a gas cell equipped with a dispersive or nondispersive infrared analyzer. These instruments have to be recalibrated at least once a week. Instrumentation should be capable of detecting CO2 down to at least 0.01 vppm. If it is observed that a serious quantity begins to breakthrough the adsorbent, action must be taken immediately. Otherwise dangerous hydrocarbons may be entrained. In fact, in a gaseous oxygen plant with minimal production of LOX, the mandatory concentration of carbon dioxide after adsorption should not exceed 0.01 vppm. An excessive concentration in the purified airstream can lead to an accumulation of solid CO2 on the main condenser–vaporizer surface and valves, especially in a gas-producing plant. This will block uniform flow of LOX, creating pockets of dry evaporation and dangerous concentrations of residual hydrocarbons. In LOX pumping, the entrained carbon dioxide will be carried away by LOX to be vaporized in the LOX pumps. In a LOX-producing unit, the CO2 will be carried out into the storage tanks (Table 5.7). 5.3.9.1 Improving Operating Stability To repeat, it cannot be overemphasized that it is of the utmost importance to operate any adsorbent purification system with high operating stability. If any instability occurs, one can improve the systems stability as follows: (a) Small breakthrough (1–3 vppm)


TABLE 5.7 Contaminant Breakthrough in Molecular Sieve Beds Methane Ethane Carbon dioxide, propane, ethylene Acetylene, propylene Butanes Water

Immediately Almost immediately Almost simultaneously, but in the reverse sequence Later Much later Very much later


This situation may be caused by the following: 1. An insufficiency of adsorbents (molecular sieve or alumina) to take care of impurities such as carbon dioxide and water 2. The adsorption phase is too long, and permits the carbon dioxide to breakthrough 3. The reactivation temperature is too low, its timing is too short or too long (overheating) Action to be taken: . . .

Reduce the process airflow slightly Shorten the adsorption time and observe the results after several complete NEMA cycles Increase the reactivation temperature

(b) Sudden serious breakthrough (5–10 vppm) Action to be taken: .

. . .

.

5.4

Strictly carry out the same action as mentioned for a small breakthrough, but more pronounced Reduce the process airflow Check the overall sequence and timing of each cycle Check the LOX product in the main condenser for any sign of milkiness. If it is pronounced, the product may require dumping Take a sample of the LOX product to the chemical laboratory and analyze it for any sign of acetylene, ethylene, propane, etc.

SAFETY

The low-temperature industry has a commendable safety record, mainly due to the extreme caution engineers and operators have used in designing and operating air separation plants. When the rare explosion had occurred, it was blamed on the presence of acetylene in the feed air or on acetylene-generating equipment in the vicinity. This was disproved in 1956; however, when an explosion in a large oxygen plant took place, but no traces of acetylene was found8. Investigation showed that the explosion was due to accumulation of nitrogen oxides, ethane, ethylene, and other paraffins and olefins. This incident and similar ones impelled designers of Revex units, which at the time were favored, to employ silica gel adsorbent to remove hydrocarbons from the oxygen-rich liquid downstream of the lower (high pressure) column in the cold box, and to add an extra LOX guard filter, also filled with silica gel,


to remove any remaining traces of dangerous hydrocarbons from LOX product in the main condenser. At the same time, it was recommended that a small stream of LOX (0.5%–1.0% of total product) be withdrawn from the main condenser. These design modifications reduced the number of accidents to almost zero, but they also increased investment and operating costs. With the resurgence of FEP systems, however, safety design has developed new parameters. Dangerous contaminants in the air are still present and in ever-increasing quantities. Their removal requires a change in the concepts as well as in procedures, which differ from those employed with Revex systems. Contaminants breakthrough a molecular sieve bed in reverse order of their adsorption.

5.4.1 HYDROCARBON BREAKTHROUGH9 Because acetylene, propane, propylene, and butanes breakthrough the bed more slowly than CO2, by carefully monitoring CO2 breakthrough it is possible to prevent the passage of dangerous hydrocarbons into the cold box. But the system is by no means foolproof: 3%–10% of propane and ethylene will not be adsorbed, methane breaks through immediately, and ethane soon after. Unfortunately, trying to prevent methane breakthrough is not economically feasible. Thus, residual traces of methane, ethane, propane, and ethylene will remain in the LOX located at the main condenser. Solubilities in LOX at process temperatures are about 1 mol % for propane, 2 mol % for ethylene, and 12 mol % for ethane and methane. The plant designer must also consider possible atmospheric inversions, which may take place especially in coastal areas, and which may increase pollutant concentration at the air plant intake equipment beyond design. Despite care in design and in operation, errors may allow CO2 to breakthrough in FEP systems, plugging up passages in the main condenser. Small, isolated, and quiescent pools of LOX may also form and if ambient air is heavily polluted, as is the case in refinery and petrochemical areas, these pools will develop a high concentration of ethane and propane. An enormous quantity of CO2 can accumulate in a cold box from even an insignificant trace passing through the FEP unit. For very large plants, the situation can become horrendous, as seen in Table 5.8, which is based on a plant producing 1000 t=d of oxygen with a process airflow of 150,000 Nm3=h, and an on-stream factor of 98% (8585 h). Ethane, on the other hand, has different properties; above 110 K liquid ethane is completely miscible in LOX. Between 110 and 90.18 K (the boiling point of LOX at atmospheric pressure), the two are partially miscible. Two phases can appear in the main condenser, one rich in ethane and the other rich in oxygen. Propane and propylene show behavior similar to that of ethane10.

TABLE 5.8 Annual Accumulation of CO2 Entrained in a Plant Producing 1000 mt=d of Oxygen CO2 vppm Concentration 0.10 0.25 0.50 1.00

Weight (kg=y) 253 632 1265 2530



Also, occlusion by a relatively soluble material such as ethane can occur in a less soluble compound that is precipitating from solution. Once precipitation has taken place, the ability of purging to remove deposits is questionable. It will depend on settling of solids in quiescent zones, adhesion to tube walls, and rates of resolubility. This means that less soluble materials can cause dangerous situations even through their concentrations are well below the recognized solubility and homogeneous flammability limits11,12.

5.4.2 SAFETY ADD-ONS The addition of an auxiliary (pressure) vaporizer can contribute to the safety of an FEP system (Figure 5.12). It will also cut down energy demands of the oxygen compressor. Conditions in this auxiliary vaporizer are more favorable for volatilization because of slightly higher vapor pressures and slightly greater turbulence. It is necessary to keep in mind, however, that hydrocarbons have a tendency to accumulate in the remaining LOX at the bottom of the auxiliary vaporizer. This location should be monitored very closely. The installation of a LOX guard filter is also strongly recommended. Many engineering firms do not offer this additional safety feature, believing that it is unnecessary. Its cost, however, is insignificant compared with the total investment of an air plant, so it may be considered as an inexpensive insurance policy. It should use a calcium silicate molecular sieve, similar to molsieve Type 5A, and not the one of silica gel. The latter substance is efficient only for removing acetylene, but has very low adsorption for ethane, particularly in the presence of CO2. Silica gel is only partially effective for C3–C5’s compounds10. For an unusual quantity of propane the only rule of thumb used by designers is to add a slightly greater quantity of molecular sieve 13X than calculated for normal amounts of contaminants.

5.4.3

LIQUID OXYGEN PURGE

The practice of maintaining a small, continuous LOX purge from the main condenser (as has been standard with Revex systems) is questionable. Purge rates of 0.5% are effective only when the contaminant concentration in the feed is low. If concentrations are high, the time needed to decrease their levels to acceptable values can be very long, because of the large volume of LOX in the main condenser compared with the volume of the purge. Using an auxiliary or pressure vaporizer and a LOX guard filter, however, enables LOX product to be drained from the main condenser only in a minimum continuous stream, or intermittently as needed. This will save substantial amounts of energy in plants of 1000 t=d or more of oxygen. In all cases, however, close visual monitoring by operators at every 4 h is strongly recommended. The increased reliability and safety of air separation plants using an FEP system should not lull operating departments into a false sense of security. Concentrations of CO2 and other hydrocarbons are continually rising in urban and industrial areas, so that designers are now using 400 vppm, and even 500 vppm, as the CO2 specification in heavily polluted regions. Some refineries and petrochemical plants now exhibit atmospheric concentrations of ethane and propane of 2 vppm and higher. Whenever this is the case a LOX guard filter is mandatory. In view of the dangerous, albeit low, concentrations of contaminants present, hydrocarbons should be monitored every 8 h, or at the very least, daily. If a total hydrocarbon analyzer is used, limits are sometimes set at 5–50 vppm in air and 40–150 vppm in LOX. It is far wiser and safer, however, to use an in-line gas chromatograph to check individual hydrocarbons, since a total hydrocarbon analyzer can promote a costly and unnecessary plant shutdown in the presence of high methane concentrations. Recently, a new methane–nonmethane hydrocarbon analyzer has arrived on the market. This apparatus employs a flame ionization detector (FID)


TABLE 5.9 Recommended Procedures for Analyzers Trace CO2 in purified air Hydrocarbons in LOX

Infrared (1–10; 10–100; 100–1000 ppm) In-line gas chromatograph With the following points to be analyzed every 4 or 8 h: rich liquid at lower (HP) column; main condenser=reboiler; upstream of LOX guard filter; downstream of LOX guard filter

The following analyzers are recommended to follow the progression of pollutants though the entire system. Source: Courtesy of F.G. Kerry, Inc., 2006. With permission.

as the sensing element and provides a digital readout of either the methane component or the nonmethane components, which are the chief worry of oxygen plant operators. The cost of this unit is somewhat similar in cost to a standard total hydrocarbon analyzer.

5.4.4 ANALYZERS The in-line gas chromatograph should include an alarm for any abnormal high content of the following hydrocarbons: methane, carbon dioxide, ethane, ethylene, acetylene, propane, propylene, butane, butadiene, and acetone. Caveat, unless the instrument is specified and tested for accuracy, any readings under 50 vppb should be debatable (Table 5.9).

5.5 ACTIVATED ALUMINAS FOR FRONT-END PURIFICATION SYSTEMS 5.5.1 BACKGROUND As far back as 1957, the research department of L’Air Liquide in Paris reported in a private in-house communication that during tests for the desiccation of process air with an activated alumina supplied by Pechiney SA (France), it was observed that the absorbent also had removed a fair quantity of carbon dioxide. At the time, however, the synthetic zeolites had come into vogue for contaminant removal, and any further testing for carbon dioxide adsorption with activated alumina was abandoned. Then in 1984, Goodboy and Fleming13 of the Aluminum Corporation of America published a paper on the possibilities of using modified aluminas for selective separation purposes by physical adsorption, in addition to their conventional role as desiccants. It should be understood that activated alumina has been defined as commercially available alumina, which may normally contain small quantities of combined ferric, sodium, and other oxides. The increased understanding of porosity control in the manufacture of activated aluminas has permitted producers not only to engineer pore size distribution, but also the degree of porosity in the adsorbent. Presently, manufacturers can supply various formulated aluminas with properties of selective adsorption, which can meet most specific process requirements of the industry. Moreover, certain commercially available aluminas are formulated to contain specific quantities of other materials to enhance their capacity for selective adsorption as well as other beneficial properties. There are various ways of treating aluminas to increase their capacity for adsorption. It can be done by calcining the adsorbent with an alkali metal oxide, or by impregnation with a decomposable salt at calcining temperatures. There are numerous patents on these processes.


The salts include hydroxides, bicarbonates, phosphates, and organic acid salts, with a preference for potassium carbonate. One may ask why there is such activity in formulating aluminas to enhance their capacity for adsorption, when zeolites, both natural and synthetic, are already available on the market and have had a long history of successful industrial application. The answer is very simple. Formulating aluminas to improve their affinity for carbon dioxide and other contaminants such as hydrocarbons offers a strong commercial challenge to synthetic zeolites in terms of energy savings in pressure swing adsorption (PSA) as it is known14–16.

5.5.2

PRESSURE SWING ADSORPTION

With the development of different types of formulated alumina adsorbents it is now possible to remove contaminants such as water, carbon dioxide, and dangerous hydrocarbons at close to ambient temperatures, say 300 K. This possibility is of high interest because it eliminates not only the requirements for a precooling system to reduce ambient air temperature down to 278 K, or 283 K, but also the need for extra energy to heat the regeneration gas (waste nitrogen) up to 363 K and even to 423 K, to regenerate the adsorbent bed when using a zeolite of the 13X variety. The process used for this type of adsorption is named PSA, in contrast to temperature swing adsorption (TSA) for the prepurification system using an adsorbent such as a zeolite 13X, which has a greater capacity for adsorption, but requires extra heat for its regeneration. Actually, when one uses a zeolite 13X adsorbent, operating pressure is also reduced to atmospheric before regeneration heat is applied. So that in reality, the latter system should really be called pressure–temperature swing adsorption (PTSA), according to D.M. Ruthven in his Principles of Adsorption and Adsorption Processes. Nevertheless, returning to the PSA system, the latter operates at normal process air pressure, of say, 700 kPa (7 bar). After process air passes through a standard compressor aftercooler, either a shell and tube or a DCA water wash tower, to return as closely as possible to an ambient temperature it is sent to the PSA unit. Depending on the specific design of the PSA system, two or three vessels are employed, either in vertical or horizontal configuration. The vessels are filled with a single bed or multiple beds of the appropriate selective-modified aluminas for the removal of the desired contaminants. Impurities entrained by the process air are retained by the absorbent at the same temperature and pressure as the inlet process air. In this regard, the operation of the PSA is similar to the TSA system, except that no extra chilling is required. A standard shell and tube aftercooler after the main air compressor, or a DCA will suffice. As soon as the first bed becomes saturated, the switch valves swing over, and process air is passed to a second vessel for treatment. The first vessel is then depressurized, and the impurities eluted by dry waste nitrogen. The complete regeneration involves the production of a very low partial pressure, always at ambient temperature, but at a lower pressure. The low partial pressure is obtained by dropping pressure in the vessel containing the adsorbent down to atmospheric, sometimes even below (using a vacuum pump), and by sweeping a waste gas (dry nitrogen) over the adsorbent. Precooling the air in a TSA system down to say 283 K is carried out to reduce the quantity of water entrainment from the DCA unit, to increase the efficiency of adsorption of the molecular sieve, and to reduce the quantity of water vapor. In the case of the PSA system operating at a higher temperature, more water is entrained increasing the quantity of activated alumina required in the vessel. As the latter adsorbent has a lower initial unit price, this saving may offset the total purchase cost of the adsorbent system.


Activated alumina has only about a 20% capacity to remove carbon dioxide, as compared with synthetic 13X zeolite, but it is said that the former has the advantage in that its effective capacity, that is the difference between adsorption and regeneration step loading, is just as high as 13X zeolite in PSA operations. Moreover, activated alumina adsorbs one-third less process air than the 13X zeolite, and that means a lower loss of energy in the desorption phase. Add the fact that the regeneration of water is an endothermic reaction, which lowers the temperature of regeneration. It should be borne in mind, however, that the use of a zeolite for water adsorption is not viable. The present practice in the design for TSA systems is to use a bed of activated alumina upstream for this purpose, and a 13X zeolite for the second bed downstream for the adsorption of carbon dioxide, and other contaminants such as dangerous hydrocarbons. In view of the temperature limitations of the regeneration cycle for the PSA system, the quantity of formulated alumina adsorbents is quite high. Therefore, in order to save on adsorbents, the total NEMA time cycle including adsorption and regeneration is very short, in fact at times, much less than 30 min. According to the proponents of the PSA system, there may be an advantage in this because the thermal effects due to the release of heat during adsorption and the consumption of heat during desorption are minimized. Moreover, short time cycles may lead to quicker start-ups. But these possible advantages are debatable. The very short time cycle of the switch valves in comparison with the TSA system of 4–8 h also means a higher cost in maintenance or replacement of the valves, whose switching action may now be compared with the Revex system. Another point to consider is that when formulated aluminas may adsorb one-third less feed air than the synthetic 13X zeolite, the high number of switching cycles compared with the TSA may also mean a higher loss of total feed air from blowdowns over a longer period of time. Furthermore, the thermal effects previously mentioned have been minimized enormously with the addition of activated alumina upstream for the almost complete removal of water, which is the prime culprit for the thermal phenomena. In fact, the quantity of activated alumina used for this purpose reached almost 90% of the molecular sieve by weight in one specific TSA application.

5.5.3 INDUSTRIAL APPLICATIONS

IN

AIR SEPARATION PLANTS

During the past few years, certain types of formulated aluminas developed commercially for selective adsorption have been used as prepurifying agents in air separation plants, for either nitrogen or oxygen production. The advantage of this application is that, when the affinity for physical adsorption of these aluminas may not be as strong as the comparable synthetic zeolites, the former can be used at normal ambient temperatures for both adsorption and desorption. No extra energy is required for regeneration. A fair number of air separation plants have already been in successful operation using the PSA system as a prepurifier16,17. According to recent publications, some of these new formulated aluminas, such as H-156 and Selexsorb CDX manufactured by ALCOA, outperform the molecular sieve 13X in carbon dioxide adsorption in PSA–VSA (vacuum swing adsorption) applications. Whereas the synthetic 13X has a higher equilibrium CO2 capacity, the carbon dioxide is more easily desorbed from Selexsorb CDX17,18. No heat is required for regeneration. Just a swing from operating to ambient pressure, and a simple sweep by a purge gas at low pressure, are sufficient to elute the undesirable pollutants. This results in an improved capacity of the system, which can then be made smaller or operated on a longer cycle. The operating sequence of a proposed system may be as follows: 1. On line 10–25 min 2. Depressurization, all the gaseous contents are vented


3. Purge phase, adsorbed pollutants are eluted by the sweep gas 4. Repressurization, a controlled stream of prepurified process gas is used to repressurize the desorbed vessel Under stated conditions for a two-vessel configuration overall valve switching loss is about 2%, which is rather high. It is considered prudent, therefore, to use a three-vertical vessel design followed by a pressure equalization procedure. In other words, the vessel to be purged is equalized in pressure with the vessel that has completed its purge phase. This latter step is then followed by a complete depressurization phase, and a complete purge. Finally, the sequence follows a pressure equalization, at one-half the pressure and a complete repressurization. In many air plants, the three-vertical design has more often been employed. It is also recommended, however, to install a hot-wired backup electric heater to be used in cases of malfunction or underperformance.

5.5.4

OBSERVATIONS

ON

PSA PREPURIFICATION

According to an article reviewing PSA prepurification for air separation plants published in Chemical Engineering of April 1999, and coauthored by Notaro, Ackley, and Smolack of Praxair, the authors noted the following comments19: ‘‘Prepurification unit adsorbents are often regenerated by reducing pressure alone, and without heating the purge gas stream, at air separation plants in which 30%–50% of the feed is available as purge gas. Examples include plants designed to deliver only oxygen or only nitrogen as products, or in total products plants where only part of the nitrogen is delivered as product.’’

In PSA regeneration the working CO2 capacity of the adsorbent is significantly reduced, particularly with zeolite adsorbents, from which desorption of CO2 is difficult due to their strong affinity. This reduced capacity would normally dictate a larger adsorber if regeneration took place at typical frequencies (2–7 h) for adsorption-step times in TSA prepurifiers. Adsorbent efficiency can instead be raised by the use of activated alumina as the predominant adsorbent, together with a six- or tenfold increase in the frequency of regeneration (20–40 min for adsorbent-step times). As a result, PSA prepurifiers may often be the more cost-effective solution for contaminant removal in small air separation plants. Removal of acetylene from feed streams, however, is a key issue in the design of PSA prepurifiers. In TSA prepurifiers the presence of a zeolite assures that CO2 breakthrough occurs ahead of acetylene breakthrough10. In PSA prepurifiers using only alumina the order is reversed20. Toward the end of the adsorption run, therefore, some acetylene may pass through the unadsorbed. These issues have been addressed by novel designs based on the use of adsorbent layering20,21 and mixtures22, especially with the use of aluminas formulated for this purpose. In addition to acetylene, one should also keep in mind the effects of ethane, ethylene, and especially propane, which are also dangerous contaminants and are not totally removed even by zeolites. The PSA system has always used the multiple bed (layer) design. The upstream layer uses activated alumina to remove water similar to the TSA system, but the PSA uses a much higher quantity of the material. In the downstream layer, the PSA system uses formulated aluminas for the removal of carbon dioxide and other dangerous hydrocarbons. This is also similar to the TSA system, but the latter uses a synthetic zeolite (13X), which has a higher affinity for the contaminants, and consequently requires a higher energy for its regeneration23,24.


In comparison to TSA prepurification, the faster time cycle of PSA processes results in a greater loss of compressed process air feed during depressurization (blowdown losses). The higher required purge flow also generates a greater pressure drop across the adsorbent. These additional losses may be offset (particularly for smaller plants) by eliminating purge heating equipment and related energy requirements, and by reducing the number of stages for aftercooling feed process air.

5.5.5 FIELD OBSERVATIONS In an actual study of an oxygen plant with a process air requirement of 40,000 Nm3=h the PSA system shows some definite disadvantages. It requires a larger quantity of adsorbent, and a high flow of sweep gas for regeneration (Table 5.10). The PSA unit proposed used only two vertical vessels. If the design involved three vertical vessels, as most larger plants require, the cost for equipment, adsorbents, and switching valves would be much higher, not to mention the additional operating cost for valve maintenance and replacement. The PSA system for the prepurification of process air has been in use with relatively good success in a number of plants ranging from small capacities up to a 1500 t=d oxygen plant. These units have been designed to treat process air even at a temperature of 323 K. The application of a newly developed selective adsorption aluminas has thus challenged molecular sieves in the field of prepurification of process air in some areas, especially those in which total product recovery is low, and the cost of energy to regenerate adsorbents is very high. In the long term, however, one has to examine very carefully both the advantages and disadvantages of each material for the prepurification of process air25–27. For example, in the choice of aluminas, the use of short switching time cycles of say, 12 min or 43,000 per year if three vessels are used, may lead to very high mechanical maintenance as well as high replacement cost for switching valves. Assuming an equipment design for an air temperature of 323 K to assure a good performance, thus may also lead to an overdesign in terms of material, as well as increased pressure drop. Also with synthetic zeolites such as 13X, its high affinity for contaminants may lead to a higher regeneration energy use. To summarize, both selective adsorption aluminas and synthetic zeolites have a definite place in the field of prepurification. Individual cases will determine specific applications.

TABLE 5.10 Comparison in Use of PSA and TSA (PTSA) in an Actual Operating Test Items NEMA cycle (min) Adsorb temperature (K) Regeneration temperature (K) Heater Evaporative cooler Pressure drop (kPa) Adsorbent(s) (kg) Kilogram of adsorb=K Nm3=h of process air= hour of adsorption Valve switching per year

PSA

TSA (PTSA)

28 303 303 N=R N=R 17.22 12,600 4,010

130 283 393 Required Required 6.89 1,900 113 (including alumina)

36,428

1,700



REFERENCES 1. Kerry, F.G. Inc., 2006. 2. Breck, D.W. 1974. Zeolite molecular sieves, ed. E. Robert, 3–16. Malabar, FL: Krieger Publishing Company. 3. ALCOA Bulletin F-35-14480. 4. Breck, D.W. 1974. Crystalline molecular sieves. J Chem Edu 4P:678–690. 5. UOP Pamphlet F.1979J-2M-1=90. 6. Kerry, F.G. 1991. Front-ends for air separation plants—The cold facts. Chem Eng Prog (August):48–54. 7. Ruthven, D.N. 1984. Principles of adsorption and adsorption processes. New York: John Wiley & Sons. 8. Grenier, M., J.Y. Lehman, P. Petit, and D. Eyre. 1984. Adsorption purification for air separation units. Cryogenic processes and equipment. Book No GOO283. New York: American Society for Mechanical Engineers. 9. Isalski, W.H. 1989. Separation of gases, 216. Oxford: Oxford Science Publications. 10. Sefton, V.B. 1962. Contaminants: Source, physical and chemical properties, basis of limits, presented at Compressed Gas Association Symposium on air Separation Plant Safety, October 5, New York. 11. Reyhing, J. 1983. Removing hydrocarbons from the process air of separation plants using molecular sieve adsorbers. Linde Reports on Science and Technology 36:14. 12. Ermenc, E.D. 1956. Wisconsin process system for recovery of dilute oxides of nitrogen. Chem Eng Prog 52:488. 13. Karwat, E. 1961. Hydrocarbon control in air separators. Chem Eng Prog 57:41. 14. Goodboy, K.P., and H.L. Fleming. 1984. Trends in adsorption with aluminas. Chem Eng Prog 80:63. 15. Keller, G.E. 1995. Adsorption building on a solid foundation. Chem Eng Prog 91:56–67. 16. Smith, D.L. 1996. Optimizing solid bed adsorption systems. Hydrocarb Process:129–132. 17. Kalbassi, M.A. 1996. Advanced process swing adsorption (PSA) air purification systems. International Institute of Refrigeration, Meeting on air Separation Technology (MUST’96), 10–12 October 1996, Munich, Germany, pp. 159–173. 18. Golden, T.C., and M. Kalbassi. 1998. Thermally enhanced pressure swing adsorption (TEPSA) air pre-purification system. AIChE Annual Meeting. Miami, FL, 15–20 November 1998. 19. Smith, D.L. 1999. Optimization of air plant pre-purifiers through use of combination adsorbent beds. Seminario Ass. Brasileira de Metais e Materiais, XIV Encontro de Produtores e Consumidores de Gases Industriais. 14–16 Setembro 17–23. Osasco, SP. Brasil. 20. Notaro, F., M.W. Ackley, and J. Smolack. 1999. Recover industrial gases via adsorption. Chem Eng:104–108. 21. Notaro, F. 1996. Advances in ambient temperature air separation. International Institute of Refrigeration, Meeting on Air Separation Technology, Munich, Germany, 10–11 October. 22. Jain, R. Carbon dioxide removal by pressure swing adsorption in an alumina adsorption bed. U.S. Patent 5,232,474, August 3, 1993. 23. Leavitt, F.W. A PSA pre-purifier for the removal of contaminants in a feed gas stream. U.S. Patent 5,769,928, June 23, 1998. 24. Golden, T.C., et al. A process for the adsorption of at least carbon dioxide, water, and oxides of nitrogen and preferably acetylene from a feed gas. U.S. Patent 5,779,767, July 14, 1998. 25. Industrial Gases 1997 Report MA28C (’97)-1. U.S. Department of Commerce, Bureau of Census, September 1998. 26. LaCava, A.I., A.I. Shirley, and R. Ramachandran. 1998. How to specify pressure swing adsorption units? Chem Eng 105(6):110–118. 27. Knaebel, K.S. 1999. The basics of adsorber design. Chem Eng Prog 106(4):92–101.


FURTHER READING ON THE SUBJECT OF ADSORPTION AND CARBON DIOXIDE BUILD-UP Norobilski, J.J., and J.S. Schneider. Particle loader. U.S. Patent 5,324,159, June 8, 1994. Kumar, R. 1996. Vacuum swing adsorption for oxygen production—A historical perspective. Separ Sci Technol 31:877–893. Ackley, M.W., et al. Multiple adsorbent loading method and apparatus for a radial flow adsorber. U.S. Patent 5,836,362, November 17, 1998. Houghton, J.T., et al. 1996. Climate change 1995. The science of climate change. Cambridge: Cambridge University Press. Case grows for climate change. Chem Eng News, August 9, 1999, p. 16.



6

Product Liquefaction, Storage, and Transportation

6.1 BACKGROUND Early air separation plants were designed primarily for the economic production of low-cost gaseous products to be supplied in high-pressure cylinders (150 bar) for industrial applications. The production of either liquid oxygen or liquid nitrogen was of a secondary consideration. Liquids were produced, however, but in small quantities, and essentially for use in research work at low temperatures in university laboratories. Small plants for total liquid production were also designed for the supply of liquid oxygen to coal mines in eastern France and in certain locations in the United States for the manufacture of liquid oxygen explosives. This application involved the use of activated carbon saturated with liquid oxygen enclosed in cartridges of nonflammable material. It has now been completely eliminated because of the relative high cost of cryogenic liquids versus the lower cost of standard chemical explosives. During World War II, however, when the availability of high-pressure industrial gas cylinders became very scarce because of the higher priority for forgings for the military, the concept of large-scale production and distribution of products in a liquid phase began to take form. Whereas the production cost of liquids was high in terms of energy consumption, this was more than offset by the lower transportation costs for the equivalent quantity of product. With the use of well-designed double wall vacuum-insulated storage tanks, liquefied gases were transported across the continent by trucks and railroad tank cars with commercially acceptable evaporation losses. The large-scale production and distribution of liquid products also initiated the development of more efficient liquefaction cycles, the design of larger liquid product storage tanks, a more extensive distribution of liquid products, and the newer applications of liquid cryogens. Oxygen, nitrogen, argon, hydrogen, and even helium have been liquefied and transported by railroad tank cars and large trucks, the latter hauling up to 20 t (liquid oxygen) at a time. Even relatively small individual customers are now supplied by the use of small double wall vacuum-insulated containers holding up to 265 L and more of cryogens. To provide these services, cryogenic producers are using large cryogenic storage facilities with individual capacities of up to 2500 t (2200 m3), and in one case even up to 5000 t.

6.1.1 NEW APPLICATIONS 1. Stored liquids are used to supply products to a process for peak load-saving applications, during an emergency shutdown of the main air separation plant, or during periods of planned maintenance. For example, in partial oxidation for the production of synthesis gas, if the oxyfuel burner is extinguished for some reason for an interval of 10 s or more, it may lead to a total plant shutdown for at least 2 or 3 d. This is economically unacceptable. For this reason, a small high-pressure storage tank of liquid


Bar graph showing relationship between power consumption per tonne of LOX at various production rates (the production is only from the cold box)

Power consumption, KWh/t of LOX

1000

800

600

400

200

20

40

60

80

100

Cold box production of LOX, t/d

FIGURE 6.1 Bar graph showing relationship between power consumption per tonne of LOX at various production rates (the production is only from the cold box). (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

oxygen is always connected directly to the main process oxygen pipeline. It is ready to supply oxygen instantaneously at the required operating pressure for at least 30 min while awaiting the start-up and operation of the stand-by liquid oxygen pumps, and vaporizers drawing product LOX from the main storage tank(s) (see Section 6.3.5.3, Figure 6.1). 2. To maintain the quality of food after preparation, cooling must be fast. Liquid nitrogen can be used to cool down perishable food at a temperature of 223 K very quickly, and to maintain this temperature at a steady level. 3. Liquid helium can be used for the refrigeration of processes that require a temperature of 13 K, such as MRI equipment in hospitals. 4. Liquid nitrogen can be used for cryosurgery. At liquid nitrogen temperatures of 77.3 K, the surrounding tissue is cauterized, thus preventing any abnormal cells from entering the bloodstream. Liquid nitrogen is also used for freezing human and animal sperm, which can be held alive for up to 20 years.

6.2

PRODUCT LIQUEFACTION

The production of pure liquid products from an air separation plant is possible either by direct extraction in small quantities or by the use of auxiliary refrigeration equipment for the liquefaction of large quantities. The process is fairly simple, as long as the designer obeys the second law of thermodynamics. As the extraction of pure products in a liquid phase increases, there is an equivalent loss of thermal energy, or enthalpy. In order to maintain a correct energy (enthalpy) balance in the process, therefore, an equivalent increase of external energy must enter the process. There is another factor (i.e., enthalpy balance), moreover, that


has been mentioned in a previous chapter 3, but not always appreciated or understood by engineers inexperienced in cryogenic technology.

6.2.1 ENTHALPY BALANCE In air separation plants that have been designed for low-cost gaseous products, process air is refrigerated, liquefied, fractionated, and the individual gaseous products are revaporized. The refrigeration capacity of the liquefied air is almost completely recovered in a countercurrent heat exchanger system within the cryogenic process equipment. Within 5 K, the gaseous products leaving the primary heat exchangers have approximately the same total enthalpy as the incoming air. The overall discrepancy in enthalpy is due to the heat gain from outside of the cold box, and by imperfect heat exchange. The enthalpy is returned into balance, however, by the generation of extra refrigeration derived from the expansion machine. The extraction of any liquid product, therefore, whether it be oxygen, nitrogen, or argon, in any quantity and combination requires additional external refrigeration equipment depending on the total quantity of liquid products extracted.

6.2.2 DIRECT EXTRACTION As noted, using a standard air separation plant for gaseous production, it is possible to extract a small quantity of liquid products because the addition and operation of the turboexpander may have the design capacity to generate enough additional refrigeration to offset the enthalpy loss from a small liquid extraction. This procedure involves extracting a small quantity of process air mixture from the high-pressure column, sending it, first through the primary heat exchangers in order to preheat it, and then to the expansion machine to generate additional refrigeration capacity. This procedure, however, is limited by the quantity of surplus heat exchange surface, if any, that may be available in the primary heat exchanger system. This surplus heat is necessary to preheat the extra air–nitrogen mixture extracted from the high-pressure column, and still maintain a correct heat balance within the primary heat exchanger system, thus avoiding opening up of the warm end temperature difference in the primary heat exchangers, and thus a loss of refrigeration. In a conventional process cycle, it is not possible to liquefy and extract more than 2% of the total process air feed without the addition of extra equipment. Increased turbine flow has another disadvantage in that it adversely affects the rectification of air in the upper (lowpressure) column, reducing the desired recovery of oxygen and argon. As explained earlier, there is no sufficient heat exchange surplus area in the primary heat exchangers in a standard air separation plant, to preheat the air–nitrogen fraction entering the turboexpander in order to compensate for the additional refrigeration required. Nevertheless, the additional liquefaction of 2% air feed represents an additional 10% of liquid oxygen production, which is more than enough in many cases. In the latter case, moreover, one may have to consider the use of two expansion turbines operating in parallel. Springman1 provided, in his classic article on liquefaction, the relationship between energy and air liquefaction for direct liquid oxygen extraction (Table 6.1). In actual practice and for a low-pressure plant with a capacity of 1000 t=d of oxygen, the total energy consumption per tonne of product has been found to be significantly higher than the one given in Table 6.1. For a liquid oxygen production of about 1% of the total, the net energy was 400 kWh=t, for a production of about 3% it was around 800 kWh=t, and for 5%–10% it was about 900 kWh=t. The net energy also includes all losses due to evaporation and pumping from the cold box to the storage tank (Table 6.2). According to Springman’s calculation, the energy cost for liquid oxygen in the aforementioned process mode is commercially


TABLE 6.1 Conclusion of Springman in Direct LOX Extraction from Air Separation Up to 0.6% of liquefied air Up to 1.0% of liquefied air Up to 2.02% of liquefied air

0.70 kWh=Nm3=h 0.85 kWh=Nm3=h 1.00 kWh=Nm3=h

490 kWh=t O2 595 kWh=t O2 700 kWh=t O2

Source: From Springmann, H., Chem. Eng., May 13, 58, 1985. With permission.

unacceptable when one reaches a production level of 5% of the total oxygen production capacity of the plant. During this liquid production, moreover, there is a simultaneous loss of gaseous oxygen production of approximately three times the liquid equivalent in tonnes. Although this high-energy cost may be viable in some rare cases, it is completely unacceptable

TABLE 6.2 Performance Tests 1200 ST=D Oxygen Plant Mode D 1 GOX 1 Maintenance LOX Level Duration of test Stability of operation Process adjustments Average airflow Average GOX production or Average LOX production (net) Total losses (LOX) 2500 t storage tank Small storage tanks Cooling and bumping cryopumps Total losses Total LOX produced Total O2 produced (c þ f) Net O2 produced (c þ d) Power consumed Energy=net tonne (i 24 h)=h Energy=gross tonne (i 24 h)=g Process efficiency (airflow) 6456.117 0.2095 O2 produced 1210.88 0.99526 Gas equivalent LOX ¼ 20.1 1.35 Total oxygen recovery Total oxygen loss (j k) Process efficiency (k=j) Waste nitrogen (a k) Calculated (l=n) Analyzed at plant

26 h Excellent Minor optimization 6456.117 KSCFH 1210.88 KSCFH 1202.96 ST=D 1.0 ST=D

(a) (b) (c) (d)

10.0 t=d 3.1 t=d 6.0 t=d 19.1 t=d 20.1 t=d 1223.06 ST=D 1203.96 ST=D 18,771.66 kWh=h 374.2 kW=t 368.355 kW=t

(e) (f) (g) (h) (i) (A) (B)

1352.556 KSCFH oxygen 1205.140 KSCFH oxygen 27.135 KSCFH oxygen 1232.275 KSCFH oxygen 120.281 KSCFH 91.11% 5223.842 KSCFH 2.3025% 2.31%

(j)

(k) (l) (m) (n) (o) (p)

N.B. Comparing (o) with (p) it is quite obvious that the plant is producing a fair tonnage of LOX (21.1) of which (20.1) unfortunately is used up to maintain the liquid level at all the storage tanks, also to keep the cryopumps chilled as well as in a ready and operable state. Source: Courtesy of F.G. Kerry, Inc., 2006. With permission.


Low-pressure nitrogen Warm exchanger

Feed compressor

Medium-pressure nitrogen

Cold exchanger

LIN to storage Separator Subcooler

Liquefier exchanger Recycle compressor

Turbo blower

Turbo expander

LAIN assist to ASU Pressurized nitrogen from ASU

FIGURE 6.2 Liquefaction of nitrogen from independent liquefier. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

for larger requirements of liquid products. It is obvious, therefore, that other more efficient liquefaction process systems have to be employed (see Table 6.2 and Figure 6.2).

6.2.3 BASIC DESIGN PARAMETERS

FOR AN

EFFICIENT LIQUEFACTION SYSTEM

Between the direct extraction of liquid product from a standard low-pressure process cycle and the classic high-pressure total liquid products plant, there is a multitude of process cycle selections depending on the imagination of the process designer and the market requirements. This chapter will limit itself to a few process cycles that are in wide use. The standard total liquid products and high-pressure cycle reached its peak during the 1950s and 1960s with a top overall capacity of around 100 t=d, using a reciprocating expansion engine with glass-impregnated Teflon rings (see Chapter 10.3). This system also employed an independent nitrogen recycle compressor in order to produce simultaneously high-purity product nitrogen. With the development of brazed aluminum heat exchangers that could withstand high pressures and the availability of highly efficient radial expansion turbines, however, the high-pressure system was quickly replaced with low-pressure auxiliary liquid-producing systems (so-called liquefiers), with very low-energy requirements and reasonable investment costs. An energy-efficient liquefier for the large-scale production of liquid products involves a separate system with its own closed refrigeration cycle. It includes a compression and an expansion system of its own. The refrigeration fluid used for compression and expansion is generally gaseous nitrogen, used because of its inertness, and therefore safety, even at high pressures (100 bar). The liquefier unit receives pure gaseous products from the nearby air separation plant, and converts them into a liquid phase. With the use of multipass, high efficiency heat exchangers, the system can liquefy independently any combination of products, within the preset refrigeration capacity of the liquefier. Standard process equipment for the liquefaction of gaseous products into the liquid phase is simple enough. The process involves a warm exchanger system wherein the temperature is lowered close to 176 K, a cold exchanger system that lowers the temperature still further close to 77.3 K, and a liquefier vessel. It is also prudent to add a separator and a liquid


product subcooler, in order to drop the temperature of the liquid produced by a few degrees, so as to minimize the evaporation and flash losses as the liquid is transferred to an outside storage tank.

6.2.4 TYPES OF LIQUEFIERS There are two general classes of liquefiers: separate or independent type, and the integrated type, which is tied to the process of an air separation plant. 6.2.4.1

Independent Liquefier (Figure 6.2)

Where the product liquefier is to be applied to an existing air separation plant, or where the owner and the operator wish the two process systems to be completely independent of each other, the liquefaction cycle involves an external refrigeration system, which will be completely independent of that of the air separation unit. The external refrigeration cycle proposed for a standard independent liquefaction system is as a rule of the closed recycle loop type. The medium used is nitrogen, which can be handled and compressed to high pressures safely. Nitrogen has a boiling point of 77.3 K at 101.325 kPa, and can be used as a refrigerant for liquefying almost all gases including hydrogen. Two compressors operating at two separate compression levels are used for the main refrigeration cycle. The first compression level is around 6.2 bar, which is very convenient, because it is somewhat close to the operating pressure of the lower (high-pressure) column. Gaseous nitrogen can be withdrawn directly from the high-pressure column and used in the second compression level of the liquefier. The second compression level is around 31 bar, which is also convenient, because it involves a standard radial compressor commonly found in the market. Further compression of the nitrogen refrigerant stream can also be achieved with the use of a booster compressor (turboblower) mounted on the same shaft as the expansion turbine. After a predetermined pressure level is reached, the high-pressure nitrogen is expanded and returned through the exchanger system providing refrigeration to the incoming gaseous product streams to be liquefied. After the expanded nitrogen refrigerant reaches the lower level of around 6.2 bar, and releases its refrigeration into the heat exchangers, it is recycled to the inlet of the highpressure level compressor. It may be useful to provide a small separate external mechanical refrigeration unit (MRU), using an environmentally acceptable refrigerant, and located between the warm and the cold exchangers, because it will assist in the overall conservation of energy. Such units operate between 233 and 216 K (408C) and can save up to 20% in the overall energy of the system. As can be seen from Figure 6.2, the process is fairly simple. The low-pressure nitrogen refrigerant from the low-pressure column LP and the warm exchanger WHX is fed into the medium-pressure recycle compressor FC, and boosted to 6.2 bar. This total stream is joined by a medium-pressure recycle stream also from the warm exchanger WHX, and both are fed into the high-pressure recycle compressor REC where they are boosted to a higher pressure of around 31 bar, as previously noted. This pressure level may also be raised to a higher level with the use of a turboblower mounted on the same shaft as the expansion turbine. Once the nitrogen refrigerant reaches its final pressure, it is passed first through the WHX, then through the cold exchanger CHX, then to the liquefier LX, and finally into a separator SEP and a subcooler SBC. In between the warm and cold exchangers, the designer may use an MRU. If the oxygen product has to be liquefied simultaneously, the same equipment is involved, but an additional pass is added to the exchanger and a separate oxygen liquefier–separator is included, as can be seen in Figure 6.3. In this figure, high-pressure gaseous oxygen HPO traverses the heat exchanger system of the liquefaction system, undergoing a progressive drop


High-pressure oxygen

Low-pressure nitrogen Refrigeration machine

Feed compressor

Warm exchanger

Cold exchanger

Medium-pressure nitrogen

LIN to storage

Separator Liquefier exchanger

Recycle compressor

Turbo blower

Pressurized nitrogen from ASU

Subcooler

Turbo expander LIN assist to ASU LOX separator and subcooler LOX to storage tanks

FIGURE 6.3 Simultaneous liquefaction of O2 and N2 from independent liquefier. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

in temperature without any serious change in pressure. Finally, this oxygen stream enters its separate liquefier by its own coil. It undergoes condensation as well as subcooling by heat exchange from the vaporization of liquid nitrogen, the latter having been fed from the liquid nitrogen separator. The liquid oxygen is then transferred to an outside storage tank, whereas the vaporized liquid nitrogen is returned through the heat exchanger system to recover its refrigeration, and recycled to the high-pressure compressor, REC. The process is designed for a preset refrigeration capacity of the liquefier unit for the production of liquid nitrogen. If oxygen is liquefied exclusively, the final quantity of the product will be slightly less than the calculated liquid nitrogen product, because the liquefaction of oxygen requires a higher refrigeration load. With multipass exchangers it is possible to liquefy oxygen and nitrogen simultaneously, but the total production of both products should remain within the refrigeration limits of the designed nitrogen cycle. Adding an environmentally acceptable refrigerant, it will be possible to upgrade the refrigeration capacity of the system by about 20% to 25%. 6.2.4.2 Integrated Liquefier (Figure 6.4) Integrated liquefiers are very advantageous for merchant plants because they offer a lower overall investment cost to users, as well as a wide variety of operational modes involving both liquid and gaseous products. These liquefiers also offer an operational simplification as the plant has a single central control system. The liquefaction section of the total plant is similar to a separate liquefier. It consists of a warm exchanger, a cold exchanger, a liquefier, and a subcooler. In the integrated cycle, however, nitrogen refrigerant is supplied directly from the fractionation column of the air separation plant within the same cold box. This nitrogen stream is first used in the warm exchanger of the liquefier to precool the high-pressure nitrogen cycle, then discharged from the cold box and fed to the recycle compressor. In turn, the liquid nitrogen produced is


Liquid oxygen Liquid nitrogen

Primary heat exchanger

Primary heat exchanger

Oxygen Waste nitrogen

Expansion turbine

Lowpressure column

Sub cooler

Main condenser

Air Nitrogen recycle Compressed nitrogen Highpressure column

Liquefier exchanger

Rich liquid External refrigeration

Expansion turbine

FIGURE 6.4 Gaseous oxygen plant with integrated liquefier. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

extracted and sent to the lower (high-pressure) column. There, a portion is withdrawn and passed through a subcooler before being sent to storage as liquid nitrogen product. A portion of the liquid nitrogen sent to the lower (high-pressure) column is used as reflux to allow withdrawal of oxygen in liquid form from the bottom of the main condenser and vaporizer. The liquefier section, therefore, must always be in operation in order to supply reflux to the low-pressure fractionation column even though no liquid nitrogen production may be required for outside storage. This latter problem may be eliminated, however, with the use of a separate turboexpander for the air fractionation unit. This solution is costly, however, and is really unnecessary in most cases. The other option during a low demand for liquid production is to supply refrigeration to the high-pressure column from the liquid nitrogen storage tank, which is known in the industry as a liquid-assist operation. All cryogenic equipment for both air fractionation and liquefaction is housed in a single cold box for a lower investment and convenience of maintenance. An earlier version of an integrated liquefier operating at comparatively low-pressure for the simultaneous liquefaction of oxygen and nitrogen is shown in Figure 6.4. In this particular flow sheet, nitrogen refrigeration cycle supplies liquid nitrogen to the top shelf of the lower (high-pressure) column, in exchange for pressurized gaseous nitrogen at 6.2 bar, also from the high-pressure column. With this procedure, one reduces the quantity of refrigeration available in the liquid oxygen at the main condenser, and is necessary for the condensation of gaseous nitrogen at the top shelf. Therefore, it is possible to withdraw liquid oxygen directly from the main condenser. In this process cycle, the recycled nitrogen is fed into the compressor at 6.2 bar, and the discharged nitrogen at 31 bar. As mentioned earlier for the independent liquefier, energy saving of about 20%–25% can be made with the use of an MRU operating with a difference of 40 K between the warm and the cold heat exchangers. With the


integrated process cycle, as described in Figure 6.4, the designer can avoid compressing the pure oxygen feed, and eliminate the use of special passages in the heat exchangers. Ancillary equipment, i.e., manifold piping and general instrumentation, can also be simplified. One may also consider eliminating the expansion turbine from the air separation unit, as the liquefier unit can be designed to supply the entire refrigeration requirements for both the air separation and the liquefier units. It should be kept in mind, however, that this scheme may lead to a slight loss in enthalpy from the extraction of pressurized nitrogen from the high-pressure column, extra losses from the heat gain due to insulation inefficiencies at the cold box, and temperature differences of the outgoing products at the outlet of the primary heat exchangers. Nevertheless, the designer can compensate for these losses by making adjustments in the overall capacity of the liquefier, and in the process airflow for the air separation unit. 6.2.4.3 Very High-Pressure Liquefiers Since 1995, designers have achieved higher nitrogen recycle pressures, 60 and 70 bar, for the liquefaction of oxygen and nitrogen, for greater economies of energy. In Figure 6.5 it can be seen that low-pressure nitrogen makeup feed is supplied from the air plant and sent to the feed gas compressor FC, where it is raised to 6.5 bar. This stream is joined by another

After cooler

After cooler

Recycle compressor

Feed from LPC

Recycle compressor Warm heat exchanger Mechanical refrigeration

After cooler

Warm expansion Warm turbo booster

GAN from PHX GAN from PHX Cold exchanger

After cooler

Cold expansion Cold turbo booster

Separator

LIN to top shelf HPC GAN from PHX

77.3 K LIN to top shelf HPC

FIGURE 6.5 High-pressure integrated liquefier. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


medium-pressure stream also at 6.5 bar from the prime heat exchangers of the air plant. Both streams are now sent to the nitrogen recycle unit REC and raised to 31 bar, which after compression is led to the first high-pressure cold booster CB, where it is compressed to about 41.3 bar. Then this stream, after passing through its own aftercooler, is sent to a second blower WB, coupled to the warm expander WXT, and where it is further compressed to about 60 bar. After this final compression, the main nitrogen refrigerant stream passes through another standard aftercooler, and at its highest pressure, is led to the warm exchanger WHX, where it is precooled to about 161 K. This precooling is brought about by the countercurrent effects of a medium-pressure nitrogen stream supplied by the primary heat exchangers of the air plant, and by the expansion of part of the recycle stream from 31 bar down to 6.5 bar at the ‘‘warm’’ expansion turbine WXT. Both of the precooling streams are returned to the outside nitrogen recycle compressor REC, with the warm heat exchanger WHX. The high-pressure stream still at 60 bar, and at 161 K, continues to the cold heat exchanger CHX. Before its entry to the exchanger (liquefier), a part of it is sent to the expansion turbine CXT, where it is expanded and reaches its final liquefaction temperature of around 77.3 K. This liquid enters the separator SEP. In the separator the portion remaining as liquid is now subcooled and sent outside to the storage tanks. The portion in the separator, which is in vapor phase, is returned through the exchanger system as a refrigerant, and finally fed to the recycle compressor. Another portion of liquid nitrogen from the liquefier is sent to the top shelf of the lower (high-pressure) column to be used as reflux in the fractionation columns. The final portion of the liquid nitrogen product may be transferred to the pure argon column, if the latter is used, to purify the argon. Calculated efficiency in each turboexpander is 85%–86%, whereas the turbobooster efficiencies are approximately 75%. As in the case of the independent liquefier, both oxygen and nitrogen can be liquefied either separately or simultaneously, as long as the total refrigeration requirement does not surpass the maximum refrigeration capability of the nitrogen recycle. It should also be understood that the nitrogen recycle must be continuously supplied with fresh gaseous nitrogen feed: (a) to make up for the liquid nitrogen extracted from the plant and supplied to storage; (b) to supply liquid nitrogen as reflux to the fractionation column; and (c) to make up for any losses within the process system. 6.2.4.4

General Summary

With the exception of a few cases involving metallurgical and petrochemical operations in isolated areas, the prime use of liquefiers is for merchant plants that have to modify their mode of operation from day to day depending on the ever-changing local market for liquids and gases. If local market conditions require a fairly large production of liquids, say above 300 t=d, it is important to design a liquefier with the highest possible energy efficiency, because energy is the most important factor in production costs. For large liquefiers, therefore, it pays to add any additional equipment that may save operating energy. On the other hand, having liquid oxygen, nitrogen, or liquid just sitting idle in large storage tanks means a loss of revenue. Above the liquefaction of 300 t=d of combined products, it is wise to consider the use of high-pressure ratios for compressing nitrogen, if the latter is the major and the final refrigerant. It is also recommended to employ a multistage MRU in the process, as a form of cascading principle. This will reduce the energy consumption considerably. If MRU is used, the refrigerant should be environmentally acceptable, such as HFC 134a (1,1,1,2 tetrafluoroethane) or equivalent. The use of an ammonia refrigeration unit, while acceptable from a process standpoint, can be a headache in mechanical reliability, continual maintenance, and toxic odor.


As for energy savings from high nitrogen compression, the use of pressures up to 70 bar and even higher is also recommended, but with the use of turboblowers directly coupled to the expansion turbines. For these very high pressures, one has to consider the choice between the availability of brazed aluminum and heli-wound coiled heat exchangers. The final choice for high pressures also depends on the availability and reliability of the turboblower and expansion turbine combination in one complete unit. Another possible thought is to use a liquid expansion turbine, which has shown some advantages in expansion from very high pressures, even up to 100 bar. 6.2.4.5 Energy Economics Some 30 years ago, energy consumption in production of liquid oxygen was approximately 600 kWh=t. Presently, the energy consumption has been reduced to 400 kWh=t or even less. If one considers the theoretical Carnot cycle, the power necessary to liquefy 1 t of nitrogen is around 80 kWh. Though unachievable, this gives the process designer a good target for the future.

6.3 CRYOGENIC STORAGE FACILITIES 6.3.1 GENERAL CONSIDERATIONS Liquid storage of products from an air plant is a viable procedure in order to coordinate the economic steady-state operation of the plant and the fluctuating demands of the market. These demands may range in capacity anywhere from a 1 L Dewar flask to a delivery requiring the use of a railroad tank car. Choosing a specific design for the cryogenic liquid storage tank or container, especially one of a very large size, requires a complete economic study. Inventory of cryogenic liquids represents a high fixed capital, as well as a high-energy operating cost. Therefore, a careful economic balance should be carried out on the desired inventory level, keeping in mind lost production due to a plant shutdown or loss of profit in sales to third parties. The design capacity of a cryogenic liquid storage system should be specified only slightly above the calculated inventory. In turn, inventory should be calculated to take care of: . . .

Peak demands of a process using the product Planned or emergency shutdowns Sale of product to third parties

6.3.2 GEOGRAPHIC CONSIDERATIONS 6.3.2.1 Ambient Temperature It is important to have a record of both the minimum and maximum temperatures, which occur in the area. This will affect the choice of materials, plate thickness, quantity, and quality of insulation. Daily variations in ambient temperature will affect the amount of thermal breathing, thus influencing the design of the vent recovery system, design pressure of the vessel, etc. The most economic storage temperature is one at which the vapor pressure of the contained liquid slightly exceeds the atmospheric pressure. 6.3.2.2 Wind Loading Considering the large surface of liquid storage tanks, it is necessary to specify the highest wind loads expected in the area. Under-specifying wind loads for reasons of economy could be potentially dangerous.


6.3.2.3

Seismic Loadings2

In areas of high seismic conditions, tanks must be designed accordingly. The Uniform Building Code (UBC) Division IV Earthquake Design is recommended. The calculations are complex and should be entrusted only to an experienced engineer, or a group specializing in this type of work. 6.3.2.4

Soil Conditions and Land Cost

It is necessary to determine the allowable footing loads, and also the need of piling. Cost and availability of land to be used for the storage tanks will determine the optimum geometry of the storage tank. 6.3.2.5

Snow Loads

If snow is a possibility, the load should be considered in the design of the roof structure if one is necessary. Subfreezing weather will also affect the choice and design of ancillary equipments such as pumps, instrumentation, steam tracing, and general maintenance. 6.3.2.6

External Corrosion

Because of the high cost of the product, the potential hazard from spillage or overpressurization of its vapor, it is of the utmost importance to maintain the integrity of the storage insulation system by protecting the outer shell from corrosion. This requires the proper choice of materials, the use of suitable preparation, and painting procedures. The inner tank is generally made of either stainless steel, aluminum, or a 9% nickel steel alloy, so that no special procedures need to be followed, aside from maintaining a high standard for descaling, degreasing, and welding. 6.3.2.7

Availability and Dependability of Utilities

This will affect the choice and design of pumps, lighting, refrigeration and vaporization systems, instruments, steam tracing, rainwater runoff, and communications. 6.3.2.8

Local Neighborhood Characteristics

The characteristics of the local area where the tanks are to be erected must be reviewed, as this will influence design features for safety and the environment. In populated and industrial areas, elaborate safety features are necessary: fire fighting equipments, remote control valves, concrete drainage systems, dykes, etc. Needless to state, an initial public relation program in the selected area is important.

6.3.3 6.3.3.1

DESIGN PARAMETERS Low-Pressure Shop-Built Tanks

These are restricted to a capacity of approximately 80 m3, and follow API Specifications 620 recommended for Design and Construction of Large Welded Low-Pressure Storage Tanks. These rules allow the pressure in the vapor space above the liquid to reach, but not exceed, 987 mbar G. Materials of construction depend on the temperature level.


6.3.3.2 Storage Vessels with Internal Pressure ASME Section VIII Division 1 or 2 gives design criteria for storage at pressures greater than 0.98 bar G. This code covers design, materials, fabrication, inspection, and testing of pressure vessels. The ASME Code is also an ANSI Standard, ANSI=ASME BPV—VIII-I. 6.3.3.3 Low-Pressure Field-Built Aluminum Tanks This design is covered by ANSI B96.1, Specifications for Aluminum Alloy Field-Erected Storage Tanks. The pressure limitation is close to atmospheric. 6.3.3.4 Flat Bottom Tanks These tanks have end enclosures designed for pressures near atmospheric. The shell must resist lifting forces arising from internal pressures. Hold down devices must permit the vessel to move readily in response to thermal displacements. This type of design provides cryogenic storage at a minimum capital cost, but special shapes may be dictated by service requirements: . . .

Pressure required for transfer of contents Evacuation (vacuum) of insulation space to reduce heat transfer Critical foundation conditions making other shapes or forms, such as spheres, more economical

The economic advantage of this type of design is that the inner container rests upon some form of load-bearing insulation, which carries the weight of the contents to the foundations. The bottom serves only as a seal, and is not subject to any significant stress. This permits it to be made of relatively thin material. 6.3.3.5 Spherical Containers A spherical inner tank enclosed in a spherical outer tank is geometrically the most economical for capacities up to 500 m3 as it offers a maximum internal volume for the minimum external surface. Nevertheless, it also demands a high cost of plate preparation and field erection. While it was a standard for the cryogenic industry over a long period of years, it has given way to the flat bottom shape, especially for very large storage volumes, for example, 1000 t and over for liquid oxygen. 6.3.3.6 Cylindrical Vessels (Horizontal or Vertical) This design is adaptable for internal pressures built up in cryogenic storage, originally designed for evacuated expanded perlite insulation, now redesigned for evacuated multilayer insulation. It is limited to small capacities because it is shop built as well as tested, and shipped as a complete unit, with all the ancillary equipment. They range in capacities from small liquid cylinders of 300 L to large 50,000 L (50 net tonnes) for customer on-site bulk stations. The cylindrical horizontal vessels with a capacity of approximately 20 t can also be mounted on trucks and transported to various clients.

6.3.4 DESIGN SELECTION For high-value products such as argon, hydrogen, and helium where the storage capacity is generally small, the obvious choice is the cylindrical vessel under pressure and with multilayer


insulation with a high vacuum, in order to reduce boil-off losses. These can be reduced to 0.1% per day or even less. For more common but still expensive liquid products such as liquid oxygen and nitrogen, where the storage pressure of the vapor is slightly above atmospheric (up to 9.867 kPa G), the economic choice, the economic choice is limited to pressurized bulk vessels. (Figures 6.6 and Figure 6.7). When soil conditions are very poor, such that a flat bottom tank may require a large number of foundation piles, a spherical design should be evaluated as it may prove less costly. A flat bottom tank has a lower boil-off loss than a spherical shape in similar conditions of vapor pressure and insulation, with storage capacities in excess of 1000 m3.

6.3.5 6.3.5.1

TYPICAL DESIGNS

OF

CRYOGENIC STORAGE VESSELS

Vertical Cylindrical Tanks (Figure 6.6)

6.3.5.1.1 General Design These vessels are fabricated with a capacity of up to 13,000 US Gallons (56 t of liquid oxygen). They have an outside diameter of 310 cm with an overall height of 11 m. Internal pressure is approximately 17.24 bar G. Also see the flow diagram (Figure 6.7). 6.3.5.1.2 Materials They may be vacuum insulated with either expanded perlite or superinsulation. The latter form of insulation results in a more compact configuration, and is somewhat lighter in weight. The inner tank containing the cryogen is made either from stainless steel Type 304 or a 9%

FIGURE 6.6 Vertical cryogenic bulk vessel. (Courtesy of Taylor Wharton, 2005. With permission.)


Safety valve ASME Safety relief selector SV-1

Rupture Rupture disc disc Dual safety device (optional) V-12 V-2 V-11

PBC-1

CN-1

SV-2

V-4 PCV-1

CV-2 V-14 SV-3 S-1

V-5 V-3

CV-1 CV-2 LI-1 PBC-1 PCV-2 PCV-1 PCV-2 PI-1 R-2 CV-1 VR-1 R-1 CN-4 R-2 S-1 SV-1 SV-2 V-8 SV-3 V-1 PI-1 V-2 V-3 V-4 LI-1 V-9 V-5 V-6 V-8 V-9 V-10 V-10 V-11 V-13 Vaporizer (optional) V-12 CN-5 V-13 V-14 V-6 CN-1 CN-2 CN-3 CN-3 CN-4 CN-5 R-1

VR-1

V-1 CN-2

Legend Economizer check valve Fill check valve Liquid level gauge Pressure-building coil Pressure-building regulator Economizer regulator Pressure gauge, inner vessel Vacuum gauge tube Rupture disc Outer vessel relief device Strainer Safety valve (ASME) Fill relief valve Pressure-building relief valve Bottom fill valve Top fill valve Pressure-building inlet valve Full trycock valve Vacuum gauge valve Evacuation valve Vapor phase isolation valve Equalization valve Liquid phase isolation valve Pressure-building isolation valve Vapor vent valve Vaporizer inlet valve Hose drain valve Fill connection Liquid withdrawal connection Liquid withdrawal connection Vapor connection Auxiliary product withdrawal connection

FIGURE 6.7 Flow diagram of vertical cryogenic storage vessel. (Courtesy of Taylor Wharton, 2005. With permission.)

nickel steel alloy. The inner container must also follow the most recent standards of the ASME Code, Section VIII with full radiographing inspection of welds. If necessary it must also follow the British Standards Institute, Standard 1500 or 1515. In general, these units are designed, fabricated, tested, inspected, and shipped to the site for erection on a prepared slab foundation. Ancillary equipment for these storage tanks includes such items as: vacuum-insulated pipe connections, ambient air vaporizers, and a gas-pressure control assembly connected to a customer distribution system. 6.3.5.2 Horizontal Cylindrical Storage Tanks (Figure 6.8 and Figure 6.9) 6.3.5.2.1 Design and Materials These units are generally an in-house storage system, close to the producing plant and are built up to a capacity of 222,000 L (250 t liquid oxygen). They may be made of either stainless steel Type 304 or a 9% nickel steel alloy. The use of either 5% nickel steel or aluminum is not generally recommended, because the tanks are usually designed for an internal pressure of 1.72 bar G and higher. They are usually vacuum insulated with expanded perlite, thus reducing the evaporated loss to less than 0.5% per day. The fabrication of the inner vessel must follow the most recent standards of the ASME Code, Section VIII, with complete radiographic inspection of welds, and thermal treatment for stress annealing. The outer tank is fabricated with carbon steel, usually 285 Grade C. The completed unit is designed, fabricated, fully tested, inspected, and shipped to the site for erection on the prepared concrete saddles. It is not unusual to find multiple horizontal tanks piped together by a manifold for reasons of economy. Flow diagram per vessel is shown in Figure 6.9.


FIGURE 6.8 Horizontal cryogenic storage vessel. (Courtesy of Industrias Brasileira de Gases, 2006. With permission.)

FIGURE 6.9 Instrumentation for a horizontal vessel. (Courtesy of Industrias Brasileira de Gases, 2006. With permission.)


6.3.5.2.2 Mechanical Stresses In this design, the engineer should keep in mind that the vessel must withstand the weight of the material of the inner and outer containers, as well as the weight of the contained cryogen. It must also support the bending stresses as the result of the beam-bending action when supported between the saddles of the foundation. Moreover, during the cooling down period when the cryogen is being introduced, there is a possibility of additional thermal stresses, depending on the rate of cryogenic liquid introduction. In regard to the outer tank, the designer must also take into consideration possible shell failure from elastic instability (collapsing or buckling), rather than any excessive stress. The ASME Code covers this type of failure from elastic instability. Charts are given for the design of cylinders and spheres, which may be subject to external pressure. Collapsing or buckling of the outer casing may be due to the great subcooling of the cryogen within the internal vessel. 6.3.5.3 Spherical Tanks 6.3.5.3.1 General Design In the capacity range of 450,000–900,000 L (500 to 1000 t liquid oxygen), the spherical design is possibly the most economical choice. Its design offers the maximum volume for a minimum external surface, in other words, the lowest cost for material for its volume. Each segment of the sphere can be cut, shaped, and the edges beveled for welding under controlled shop conditions, but subassemblies have to be field welded in whatever atmospheric conditions exist at the site during erection. 6.3.5.3.2 Materials and Codes Materials to be used are stainless steel Type 304, and possibly 9% nickel steel, which has a greater resistance to impact at low temperatures than 8% or lower nickel steel. Because of the higher pressures involved, it is not recommended to use a lower content of nickel in any steel. The same code requirements of ASME Section VIII including radiography of welds and stress annealing expressed for pressurized tanks also apply. With regard to stress annealing, the following procedure has been followed: A heavy layer of insulation is applied on the external surface of the inner spherical vessel, with the exception of a small opening. Through this orifice, a propane–butane burner is inserted to supply the necessary heat for stress annealing at the required temperature, in accordance with ASME conditions. Once the code requirements are fulfilled, the vessel is allowed to cool down, the external insulation is peeled off, and the external shell is constructed following the ASME requirements for elastic instability, as mentioned previously for cryogenic double wall tanks. The space between the inner and the outer shells is generally filled with expanded perlite, and is evacuated to a high vacuum. Whereas spherical tanks have proven popular in the past, they have been supplanted to a great extent by flat bottom double wall tanks that are more economical to erect. The latter also offer a smaller loss in boil-off per day. They are limited, however, to applications requiring low pressures. There is an exception to the latter premise as noted in Section 6.1.1 (1) whenever a high pressure and instantaneous supply of a liquid cryogen has to be injected into a pipeline to maintain a process in operation awaiting the start-up of liquid pumps. The normal capacity of such spheres is in the vicinity of 50 liquid tonnes or to maintain the process in operation for 30 min (see Section 6.4.9, Figure 6.10).


Inlet filter

Air compressor

D.C.A. Absorbers A

Waterwash tower

Air separation

B Expander Duct

HP LOX storage Tic Fic Ts Tell

LES

FC TC

TC FC FC

Consumer

B Steam fired waterbath LOX vaporizer

LP LOX storage

A HP LOX pumps

Fic

Oxygen compression system

FIGURE 6.10 Emergency spherical vessel system. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

6.3.5.4

Flat Bottom Tanks (Figure 6.11)

6.3.5.4.1 General For large cryogenic storage tanks with a capacity ranging from 800 to 5000 t (liquid oxygen), the advantages of the double wall flat bottom tank are self-evident. The metallic segments for the vertical walls are very easy to roll in the factory, and are easier to weld than spherical segments in the field. 6.3.5.4.2 Design Flat bottom tanks are designed for pressures near atmospheric in the range of 3–34 kPa G. The bottom is not subject to any important stress, except for compression from the contained cryogen and is made from relatively thin material. It consists of a welded metallic plate, which serves as a vapor barrier. On the bottom surface, underneath this plate, there may be applied a layer or two of load-bearing cellular-glass blocks acting as insulation. The inner vertical wall of the cryogenic vessel is, as a rule, anchored to the foundation to hold down the inner shell against any possible lifting forces due to rising internal pressure. Applied constraining devices, however, must allow any displacement of the vertical shell due to thermal variations. 6.3.5.4.3 Materials The material used for the inner shell of flat bottom tanks has been stainless steel, Type 304 originally, but because of the very low pressures involved, this has been replaced by aluminum, then by 9% nickel steel, which in turn has been displaced in some cases by 5% nickel steel. However, the project engineer should be warned. The use of 9% nickel steel has a greater resistance to impact than 8% nickel steel; and that if one considers anything lower than 5% nickel, the project engineer should be advised that around 3.5% nickel, the steel alloy may suffer from embrittlement due to very low temperatures involved, and its use should not be permitted. Flat bottom tanks up to 2500 t of liquid oxygen have been built with 5% nickel steel, but not without the permission of the ultimate user. Its use is one of the balancing


13.79 kPa pressure Liquid level

Volume:

2006 m3 530,000 US Gallons

17.20 m

Tank diameter: 12.19 m

Insulation: glass bricks

0.67 m

Pile cap

Piles

3.0 m

LOX pump

1.22 m (a)

Grade level

FIGURE 6.11 (a) Typical design for a large flat bottom Tank. (b) Construction of cryogenic LOX and LIN storage tanks. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


economics with long-term safety. Between the inner and outer shells the space is filled with expanded perlite as insulation.

6.3.6 6.3.6.1

CRYOGENIC LIQUID DELIVERY SYSTEMS General

Once large industrial liquid-producing plants became an economic reality around the end of World War II, the cryogenic industry attacked the next problem, namely, a viable delivery system of the liquid products from the producing plant to the point of use. The transportation of cryogens includes a variety of methods and vessels, ranging from small portable Dewar flasks to large trucks and railroad tank cars carrying liquid oxygen, nitrogen, argon, helium, and even hydrogen across the continent. 6.3.6.2

Small Portable Containers

The capacity of these units may be classified as follows: 1. Very small vessels with a capacity from 1 to 50 L are essentially used for laboratory work. These containers are double wall vacuum insulated, similar to Dewar flasks, which can be handled quite easily. The vessels are at a low pressure. The pressure relief valve is set at around 3.9 kPa G. 2. Heavier containers, which cannot be handled manually, are equipped with casters at the bottom of the outside tank (Figure 6.12). 3. Specially designed portable liquid cylinders are used for small industrial customers, whose consumption of gases is important enough to exceed those described in points (1) and (2). The capacity of these liquid cylinders is in the range of 165 L and operates at

FIGURE 6.12 Large portable liquid helium vessel on casters. (Courtesy of Praxair, Inc., 2006. With permission.)


a higher pressure. The relief valve is set at 152.75 kPa G, and the inner vessel safety disk is at 689.5 kPa. These portable, so-called, liquid cylinders can contain up to 148 L (liquid oxygen), at a DOT service pressure of 2014 kPa. Expected evaporation rate is around 1.5% per day for oxygen and 2.2% for nitrogen due to its higher vapor pressure. The inner vessel is generally made from Type 304 stainless steel and the outside container from carbon steel. As a rule they are maneuvered with a wheeled cart designed for that purpose. They can also be manifolded to a common pressure-reducing regulator to serve as an important factory area with high gas consumption. 6.3.6.3 Customer Bulk Stations (Figure 6.6 and Figure 6.7) Whereas these tanks may not belong, strictly speaking, to the category of transportable containers, they are part and parcel of a contractual delivery system. The tanks are rented from the supplier of gases in liquid form and can be removed by the supplier if the contract is terminated. These liquid containers serve hospitals and other large users of gases. Once these portable containers are installed at the customer’s site, they are supplied with liquid product from delivery trucks. The containers are large vertical cylindrical vessels, set on a predesigned concrete foundation pad. They have a capacity of up to 50,000 L, at a pressure of 1724 kPa G (17.24 bar), and an evaporation rate of about 0.22% per day for liquid oxygen. Their overall height is 11 m and outside diameter is 3.1 m. The weight is about 22,600 kg (empty) and 79,000 kg (filled) with liquid oxygen. The vertical tank is equipped with tripods at the bottom, which are anchored to the foundation pad. The top is equipped with shackles for hoisting the tank into place. The inner vessel is made from either Type 304 stainless steel or a 9% nickel alloy steel. The outer tank is from carbon steel, and the insulation is either expanded perlite or superinsulation under a very high vacuum. Fabrication of the inner vessel is by stateof-the-art welding procedures in shop-controlled conditions, including ASME Section VIII applications, stress annealing, and complete weld radiography. The outer carbon steel casing will also include ASME design for preventing elastic instability. Customer stations also include the following ancillary equipments: (a) a low-level alarm to signal a low liquid supply, (b) an ambient air vaporizer that may range from 6 to 135 Nm3=h, depending on the customer’s needs, and (c) a gas flow pressure regulator to control the flow from the supply tank to the client’s inlet pipeline. The system includes a bypass valve, a pressure gauge, and a safety valve (see Figure 6.7). Smaller capacity liquid stations can be completely prepackaged on a skid in the factory, along with an automatic pressure-building circuit and a liquid withdrawal unit. Connections may also be included for high pressure or centrifugal liquid transfer pumps, and a vacuuminsulated liquid pipeline ready to be connected to the client’s inlet pipeline. Smaller liquid stations can be completely prepackaged, instrumented, skid-mounted, as well as shop-tested, and delivered by truck to the client’s site, or by barge to an offshore oil rig. 6.3.6.4 LOX Distribution in a Shop The piping, valves, and accessories should be well-insulated or even vacuum insulated to reduce heat gain to a minimum. 6.3.6.5 Liquid Deliveries by Truck This popular delivery system involves the use of a horizontal cylindrical liquid storage tank, designed and fabricated as described in Section 6.3.5.2, but on a standard truck chassis. Of course, the design and structure of the truck will depend on the final dimensions and the weight of the filled storage tank. The design of the cargo will also involve certain alterations in


order to include a cryogenic liquid transfer pump, and any necessary ancillary equipment. Truck deliveries of cryogens generally supply customer stations, or function as a temporary on-site storage facility for special applications, such as a NASA launch site, an oil well site using liquid nitrogen for enhancing oil production, or even for an exchange of products between industrial gas producers. The fabrication of the inner vessel of the container incorporates the use of high-strength materials such as Type 304 stainless steel, or a 9% nickel steel alloy, and conforms to the existing codes of CGA, DOT, and ASME Section VIII. They also require high-technology welding procedures, and very efficient insulation with either expanded perlite or superinsulation, with a high degree of vacuum. The vessel capacity is generally around 16,000 L, or slightly higher, and working pressure is from 2.95 to 19 bar G, to keep product loss low during filling operations. The inner tank is also equipped with transverse baffles in order to minimize sloshing during a long journey on a highway. The tank should also include a high capacity pressure-building coil to facilitate the rapid fill of the receiving tank at the other end of the journey.

6.4 CRYOGENIC PUMPS 6.4.1 BACKGROUND One of the major contributions to the liquefaction of gases has been the development and application of the cryogenic pump that could operate at temperatures below 170 K. It was first applied to small high-pressure liquid-producing plants for the direct conversion of liquid oxygen to a high-pressure gaseous product for cylinder filling. This concept immediately eliminated the old and rather primitive water-lubricated reciprocating compressor, which required weekly maintenance, and was the bane of all operating departments in the compressed gas industry. The original cryogenic pump, introduced in the 1940s, was a positive displacement pump involving a horizontal stainless steel piston housed in a bronze casing, inserted into the cold box casing near the main condenser. The cold box insulation, either rock wool or glass wool, depending on the size of the plant, also insulated the pump. The pump was located close to the casing with a bolted aperture for easy maintenance. The electric motor drive, of course, was located outside of the casing. As good as the original concept was, stoppages due to freeze-ups from entrained solid carbon dioxide particles were not infrequent.

6.4.2

VARIETY

OF

APPLICATIONS

The original small positive displacement pump, whose technology is still used today, has been joined by a whole series of modern centrifugal pumps (Figure 6.13) ranging in application from the simple transfer of cryogenic liquids at low or medium pressures to larger highpressure positive displacement pumps operating at pressures in excess of 200 bar. These applications include: 1. The transfer of liquid products within the cryogenic process cycle itself, and is part of the cryogenic process cycle for process improvement 2. The transfer of liquid products to outside insulated storage tanks for back-up purposes 3. The transfer of liquid products from storage tanks to insulated cryogenic trucks or railroad tank cars for outside distribution 4. The transfer of liquid products from trucks or tank cars to back-up storage tanks 5. Compression pumps working in concert with specially designed vaporizers to convert cryogenic liquid products into high-pressure gas for distribution lines to various points of consumption


FIGURE 6.13 A cryogenic single-stage recirculating pump as the pressure increases, the number of stages also increases. (Courtesy of Lawrence Pumps, Inc., 2005. With permission.)

6. High-pressure positive displacement pumps operating over 200 bar that are used principally with liquid nitrogen for enhanced oil recovery, but may be considered for oxygen in very high-pressure services Each of these applications may require a specialized design depending on the flow and pressure requirements. It may be said, however, that cryogenic pumps are generally divided into (a) transfer pumps for high liquid flows, and low or intermediate pressures, and (b) compression pumps (including both centrifugal and positive displacement pumps) for the conversion of cryogenic liquids into a vapor phase for gaseous pipeline distribution and direct consumption.

6.4.3 MATERIALS Cryogenic pumps follow the same basic design as pumps manufactured for higher temperatures. Reliable operation below 169.5 K, however, strictly demands use of materials with the proper mechanical and thermal properties at low temperatures. Choice of materials is


controlled by the following parameters: operational reliability, safety, mechanical configuration, compatibility with oxygen at high pressures, prevention of contamination by gas entrainment or solid particles, and lastly, ease of cleaning and maintenance. Carter3 cited the experiments carried out in Germany by Baur et al.4 in which the following procedures were carried out: . . .

Pumps with aluminum impellers and aluminum casings Pumps with bronze impellers and bronze casings Pumps with stainless steel impellers and stainless steel casings

The operation of the pumps was programed with an axial displacement of the rotor, inducing rubbing between the rotor and the housing. Second, the injection of a foreign material into the pump. In the rubbing tests, two pump geometries were used, both single-stage and with similar tip speeds. The lowest tip speed was 65.85 m=s and the highest tip speed was 77 m=s. This corresponded to a pressure rise of 21.4 and 28.4 bar, respectively. Rubbing was initiated by a special mounting at the thrust bearing, which permitted the rotor to be moved against the housing during the operation. The tests terminated when the torque became high enough to stall the motor. The same tests were repeated with a mixture of liquid and vapor. In the latter tests, both the aluminum combination and stainless steel combination started internal fires, which burned through the housings. There was no ignition or fire with the all-bronze unit. During a test involving an aluminum impeller with a bronze-casing ignition occurred even with only liquid as the medium, but the test with a bronze impeller and an aluminum casing there was no ignition. The use of aluminum, therefore, is not the sole culprit. In the test involving the introduction of a foreign object, the all-aluminum pump caught fire, but the allbronze combination did not indicate any ignition. The preceding tests were carried out with peripheral speeds of about 76 m=s, which include most, if not all, medium pressure and pipeline back-up pumps. In regard to transfer pumps, which operate around 5 bar, a series of tests were also carried out with peripheral tip speeds of 29 to 30 m=s. In these tests, the material combinations used were as follows: An aluminum impeller in a stainless steel volute and an all-aluminum pump. Ignition was attempted by thrusting the impeller toward the housing, and by the introduction of foreign particles into the process. No ignition or fires were observed during the latter tests. One may conclude that up to a discharge pressure of 5 bar, all the above-mentioned material combinations are safe to use against gross axial rubbing conditions. As for medium pressure applications, up to 28 bar, an all-bronze pump will resist gross axial rubs and entrained foreign objects in both liquid- and two-phase operations. However, a pump with an aluminum impeller with a bronze casing is the most sensitive to ignition than all other combinations. There is no doubt that choice of materials and peripheral tip speed are the critical items. A high peripheral tip speed may induce rubbing between the impeller and the casing. The combination of bronze impellers with a bronze casing is the least sensitive to ignition. Carter also reported that radial failures can occur from the rubbing between a rotating shaft and the metal bellows seal, which may be too closely spaced around the shaft. This critical situation may be diminished, however, with the use of a nonrotating bronze sleeve between the rotating shaft and the stationary bellows. In the choice of materials for pumps, one has to bear in mind the application of the pump. For example, if the pump is to handle liquid nitrogen or any other inert fluid, it may be more economical to choose an aluminum impeller with a chrome–nickel steel casing; but if the same


pump were to have a dual purpose within a few years in the future and were to pump liquid oxygen, then an all-bronze pump is mandatory for safety reasons.

6.4.4 PRESENT DESIGNS As the preceding tests were carried out, all manufacturers of cryogenic pumps have exercised great care in the choice of materials to ensure mechanical reliability at low temperatures and compatibility with oxygen under pressure. Nevertheless, ultimate responsibility rests with the designer, so one must be extremely careful in the preparation of technical specifications for pumps, and in the final acceptance tests at the manufacturer’s shop. It is also important to obtain and study the Compressed Gas Association Pamphlet G-4.15. The design of a cryogenic pump should be as simple as possible to make maintenance easy. Proper selection of seals and bearings is critical. Labyrinth seals with nitrogen injection at one end are commonly specified. Sleeve bearings are used and thrust bearings are lubricated with halocarbon, a synthetic lubricant compatible with oxygen. Impellers are generally made of bronze, and shafts of stainless steel or Monel. Maximum safety casings are made of a tin–bronze alloy. Bronze stops are recommended to prevent the impeller from wandering axially. The impeller should be keyed to the shaft. Electric motor drives, direct gear or belt driven, are normal. For pumping low viscosity fluids such as liquid hydrogen special pumps are required. Their design has to accommodate high differential thermal shrinkages, not only industrially acceptable manufacturing tolerances.

6.4.5 NET POSITIVE SUCTION HEAD6 (FIGURE 6.14) Aside from the importance of a good basic design, the net positive suction head (NPSH) for a cryogenic pump is of greater significance than that for a standard pump operating at a higher temperature. It has to be borne in mind that a cryogenic fluid is always very close to its boiling point. Unless the liquid is kept subcooled, it will vaporize and enter the pump as a two-phase fluid. The expression NPSH may be defined as the difference between the vapor pressure of the liquid and the static head at the suction of the pump. More problems have ensued from this factor than any other single cause. A cryogenic pump cannot be operated on a saturated liquid because in order to establish flow into the pump, there has to be a lower pressure in the pump. If there is a pressure drop in the liquid, the latter will start to boil, and the resultant vapor will enter the pump along with the liquid as a two-phase feed. This condition will create cavitation, and will cause seal or bearing failure. In order to prevent cavitation there must be provided a minimum NPSH at the suction of the pump. According to Martin7, the NPSH available is dissimilar to the NPSH required because the former defines the excess of pressure over the liquid’s vapor pressure at the pump suction flange. With some rare exceptions, centrifugal pumps specify that the NPSH available be greater than the NPSH required for avoiding cavitation. This minimum NPSH depends on the design of the pump, and is usually stamped on the nameplate by the manufacturer. A good rule of thumb for transfer pumps is to specify a minimum NPSH of 1.22 m. Another good rule to adopt is to make sure that all incoming piping beginning from the source of the cryogenic liquid is of ample diameter, minimum length, straight flowing as much as possible, with a minimum of bends and valves. Of course, all the piping should be very well insulated for continuous service. Moreover, the piping on the discharge side should not be undersized. One should allow a maximum velocity of 0.61 m=s, and have a low-pressure drop. In order to stabilize the flow at


Inducer

FIGURE 6.14 Schematic of an inducer in a pump. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

the entrance port of the pump it is prudent to specify the use of a bronze inducer, which will serve to streamline the fluid flow at the entrance port. This piece of equipment is a small axial type of impeller similar to a conveyor screw, which is placed close to, and on the same shaft as the main pump impeller. The inducer has a very low NPSH, and generates just enough head to provide a suction pressure for the principal impeller. When vaporization occurs at the inlet, the bubbles so formed are small and collapse over the entire length of the inducer. Therefore, when a shock occurs, the inducer keeps operating with no vibration. If no inducer is used the vaporized bubbles implode along a very small portion of the impeller vane because the pressure rise is sudden and steep. The efficiency of the inducer is lower than that of the main impeller, but the effect on the overall operation in terms of energy is minimal. Problems at the pump may also be due to a long and complex piping between the storage tank and the cold box. The liquid cryogen leaving the cold box is generally subcooled at about 1 K or more below its boiling point so as to arrive in storage as a completely saturated liquid. If it arrives slightly above its boiling point one may expect problems at the pump. When a cryogenic pump especially in high-pressure service exhibits a malfunction, the local supervisor may be little too quick to point at cavitation resulting from a low NPSH. The mechanical operation should also be checked for any part malfunction.


Example In a recent project involving a large LOX storage tank, the high-pressure cryogenic pumps were ordered with a minimum NPSH of 8 ft. (2.44 m). As shown in Figure 6.17 the available NPSH is more than adequate including the pressure drops including the liquid level at the tank, vapor pressure at the top of the tank, and the various valves and piping. NSPH evaluation for LOX back-up pumps Calculate maximum=minimum suction pressure

1. Determine tank’s maximum liquid height Tank diameter 12.19 m; tank capacity 2006 m3 (530,000 US gallons) Cross-sectional area [(12.19)2 3.1416=4]H ¼ 2006 m3, so H ¼ 17.2 m From Figure 6.15 maximum liquid above grade 17.2 þ 0.66 þ 3.0 ¼ 20.86 m Deduct pump elevation 1.22 m Net height 19.64 m Max. suction pressure 19.64 m 9.7927 1.141 ¼ 219.447 kPa Add tank vapor pressure 13.790 233.237 kPa Min. liquid above grade 3.00 m þ 0.66 m ¼ 3.66 m1.22 m ¼ 2.44 m Min. suction pressure 2.44 9.7927 1.141 ¼ 27.26 kPa Add tank vapor pressure 13.79 41.05 kPa LOX pump capacity to be specified on a daily plant production rate: Plant daily production 1088 t=d 1 t LOX 231.5 US gal or 251,880 gal=d Pumping rate 251880=24=60 175 gpm Minimum NPSH required ¼ 2.44 m or 8 ft. or (8=2.31) 1.141 3.95 psi System pressure drop: 1–4 in. tank shutoff manual valve ¼ 2 ft. equivalent drop 1–4 in. tank shutoff automatic valve ¼ 4 ft. 1–4 in. pump inlet manual valve ¼ 2 ft. 4–4 in. elbows 74 ells ¼ 28 ft. 1–4 24 in. long flex hose ¼ 10 ft. 30–4 in. pipe ¼ 1 ft. Total pressure drop ¼ 47 ft., say 50 ft. Pressure drop for 4 in. pipe at 180 gpm ¼ 3.53 ft.=100 ft. length Pressure drop for system ¼ (3.5350)=100 ¼ 1.765 ft. or 1.765=2.31 ¼ 0.764 psi 1.141 ¼ 0.871 psi or 6.005 kPa Actual NPSH required ¼ 8 ft. þ 1.765 ft. ¼ 9.765 ft. say 10 ft. or 3 m Note: Industry does not recommend draining a vessel or tank to the bottom. It generally refills vessel when liquid level descends to approximately 10% of its total volume. Conversions: 2.31 ft. of water ¼ 27.72 in. ¼ 1 psi ¼ 6.8948 kPa 1 m of water ¼ 9.7927 kPa Density of liquid oxygen ¼ 1.141 at 90.18 K 1 US Gallon LOX ¼ 9.522 lbs ¼ 4.319 kg 1 cu. ft. LOX ¼ 71.23 lbs ¼ 32.31 kg All piping calculations have employed US sizing.


Another problem may exist, however, and this involves the temperature of the liquid oxygen in the tank itself, and the temperature of the product when it arrives at the inlet port of the pump. To take care of this problem the product LOX leaving the process is normally subcooled by 1 or 2 K. In the previous problem it was found that the LOX had been subcooled to 88 K, or two degrees below its normal boiling point of 90.18 K, which ruled out the possibility of cavitation. Finally, when the pump arrived at the factory, it was found that one of the parts had been made from a wrong material. As soon as the part was changed to Monel as originally planned, the pump was returned, installed, and operated perfectly.

6.4.6 INLET FILTER SCREEN The use of a strainer at the pump suction is of prime importance. This screen will trap entrained solids such as frozen carbon dioxide, weld chips, small pebbles, etc. These strainers should be examined and cleaned at regularly scheduled periods. Clogged screens may create sufficient pressure drop to start cavitation even in well-installed pumps. Some of the suggested standards of the Commission Permanente Internationale of Paris are as follows: A strainer should be included in all pumps and should be located as close to the inlet flange as possible. The screen size of the filter should be smaller than the design gap between the impeller and the pump casing.

In general the screen openings should be no more than 0.5–1.0 mm maximum. In fact, some air plant suppliers recommend a screen opening less than 0.3 mm. Moreover, the crosssectional flow area should be 1.5 times the cross-sectional area of the installed piping at the entrance port.

6.4.7 INSTALLATION

AND

CONNECTIONS

A common problem with the installation of all pumps, and especially those operating at cryogenic temperatures, involves the ancillary piping around the pump. Poorly designed piping both at the inlet and outlet can create enough strain to unbalance the pump and lead to malfunction, breakdown, and even a fire. In order to overcome the problems of strains due to piping, one should consider the use of flexible metallic connections at both the suction and the discharge sides of the pump. Of critical importance is the sizing of the piping at the suction side. Too large a cross-sectional area will lead to thermal losses due to a low flow velocity, but inadequate diameter will cause a greater friction loss. It is also recommended that pumps have antivibration mountings to prevent excessive shocks from nearby rolling stock and reciprocating compressors. With cryogenic pumps handling high-pressure liquid oxygen, i.e., 30 bar and over, it is recommended that the units including ancillary equipment such as valve bodies be housed behind a concrete barrier for protection of personnel and equipment (see Chapter 12 on safety).

6.4.8 6.4.8.1

TYPICAL PUMP PIPING CALCULATIONS Velocity Constraints of Cryogenic Fluids

Piping calculations for cryogenic fluids involve an appropriate selection of pipe diameter and fluid velocity, and the project engineer must balance piping costs and constraints of the cryogenic fluid in motion. There are very few published articles on the subject, and advisory


bodies such as Compressed Gas Association offer a graph on permitted velocities based on safety, but not on efficiency. In practice most designers accept a velocity for liquid oxygen in the range of anywhere from 1.8 to 2.4 m=s, but in a straight horizontal length of pipe. In actual application, however, this velocity drops to 0.61 m=s to allow for a fairly high friction factor due to type of valves, elbows, fittings, and changes in direction. It has been stated that at the exit of a pump a cryogenic fluid such as liquid oxygen becomes subcooled, therefore higher velocities may be used, but then how high? Whereas the previous premise is valid, experience has also shown that once the cryogenic fluid has gone well past the pump, energy generated by the friction factor of the piping, valves, and fittings can well cause volatility, which should be avoided at all costs. Example As piping usually involves US Standards for most applications, the following calculations may be easier to understand: Let us assume that the piping run is 300 ft. (91 m) The vertical head is estimated at 100 ft. (30.3 m) The static head ¼ (12 in. 100 ft.)=24.27 in. LOX LOX density ¼ 71.21 lbs=cu. ft. If LOX flow is 45 gpm And 1 US Gallon of LOX ¼ 9.52 lbs Then weight ¼ 45 gpm 9.52 lbs 60 min ¼ 25,704 lbs=h (3.24 kg=s) Volume of liquid ¼ (45 gpm 9.52 lbs)=71.21 lbs=cu. ft. ¼ 6.016 cu. ft.=min (2.84 L=s) Case A If a 1 in. Schedule 1.0 stainless pipe is selected The internal diameter ¼ 1.097 in. Area ¼ 0.945 in.2 ¼ 0.00656 ft.2 Velocity ¼ (6.016 cu. ft.=min)=(0.00656 60) ¼ 15.28 ft.=s ¼ 4.685 m=s The above velocity is much too high. Case B Select a 2 in. Schedule 10 pipe Internal diameter ¼ 2.157 in. Area ¼ 3.65 in.2 ¼ 0.25347 ft.2 Velocity ¼ 3.95 ft.=s or 1.20 m=s Above velocity is still too high. Case C Select a 3 in. Schedule 10 pipe Internal diameter ¼ 3.26 in. Area ¼ 8.35 in.2 ¼ 0.057986 ft.2 Velocity ¼ 1.73 ft.=s or 0.524 m=s The above velocity is acceptable.


Case D The same calculation for a 4 in. Schedule 10 pipe will indicate a velocity of 1.01 ft.=s or 0.306 m=s. In terms of economy, select a 3 in. Type 304 Schedule 10 stainless steel pipe. In a 3 in. line with a velocity of 1.73 ft.=s, the following pressure drop calculation is in the following order: Reynolds number (NRe) ¼ (50.6 Q r)=(d m) where r is the density and m is the viscosity. Therefore 50.6 45 gpm 71.21 lbs=cu. ft. 3.26 in 0.12 (cent. poise) ¼ 414480 And the friction factor from graph8 is 0.021 and p=100 ¼ (f r Q2)=D5 or [0.0216 0.021 71.21 (45)2]=(3.26)5 or 0.18 psi=100 ft. 3 ¼ 0.54 psi Executing the same calculation with a 2 in. line it is found that the pressure drop will be 4.2 psi (too high). Assume a 3 in. pipe, and a pump discharge pressure to be 50 psig or 344.75 kPa. In specifying a pump for the above conditions, the following may be used: Suction line ¼ 4 in., Schedule 10, SS 304 Discharge line ¼ 3 in., Schedule 10, SS 304 Fluid is liquid oxygen Inlet temperature is 90.15 K Fluid flow is 50 gpm Head is 50 psig Minimum suction head is 0.2 psig NPSH available is 4 ft. (1.22 m) LOX composition is 99.5% oxygen 0.5% argon 6.4.8.2

Pressure Drops due to Piping Components

There is enough published empirical data available from suppliers such as Crane8 relating to the flow resistance of various components, i.e., valves, tees, elbows, bends, etc. It is known as the K factor, the flow factor is known as C factor, and the friction factor relates to the internal roughness of the pipe7. The pressure drop (Dp) in any piping system depends on: 1. The pipe friction ( f ) or ( ft for complete turbulence), which is a function of the inside diameter of the pipe, its fluid velocity, its density, its viscosity, and the internal roughness of the pipe 2. The change in direction of flow path involving bends, elbows, and change in elevation 3. Any obstructions in the flow path such as control valves, orifice plates, and check valves 4. Changes in cross-section and shape of the flow path whether sudden or gradual Once the resistance of the various components is known, they can be summed and converted into total pressure in terms of psi, or kPa or additional lengths of feet or meters in the


pipeline. Because most of the data is empirical a certain safety margin should be factored into the calculations. Example In Figure 6.15 it has been assumed that the piping follows the previously chosen 3 in. nominal stainless steel Schedule 10 Type 304 pipe in terms of fluid, flow, and physical properties. In addition, the pipe also includes the following fittings: a 3 in. gate valve, a 3 in. elbow with a 6 in. angle valve. All valves are wide open. Length Internal diameter Fluid and flow Vertical head, fT Viscosity, m Density, r Friction factor Reynolds, NRe Mean velocity

300 ft. (91 m) 3.26 in. (8.28 cm) Liquid oxygen at 90.18 K and 45 gpm (2.84 cm=s) 100 ft. (30.3 m) 0.19 71.23 lbs=cu. ft. (1141 kg=m3) 0.018 261,778 1.73 ft.=s (53 cm=s)

Pumping system of 2500 tonne LOX tank total liquid drop 6.706 m Inner tank size 12.19 m diameter 17 m high 2.54 m

Center line of suction

43.2 mm

5.309 m

1.83 m

1.46 m to top of grating

1.6 m

1.27 m

1.37 m

Ground level

1.96 m to flange

Pump suction flange 48.25 mm Ground level

0.915 mm Pump suction

FIGURE 6.15 Example calculations for NPSH. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


Case A The piping system as described but following a horizontal plane in space9. Gate valve K factor 8 fT or 80.018 ¼ 0.144 Elbow K factor 8fT or 120.018 ¼ 0.216 Angle valve K factor 150fT or 1500.018 ¼ 2.85 Total K factor for fittings ¼ 3.21 Pressure drop for all fittings ¼ [181063.21 71.23(45)2]=(3.26)4 ¼ 0.074 psi (0.51 kPa) K factor for piping ¼ (fT L=D) or (0.01830012 in.)=3.26 or 19.88 Pressure drop for piping ¼ [18106 K r Q2]=(3.26)4 ¼ 0.457 psi (3.151 kPa) Total pressure drop for piping and fittings is 0.074 þ 0.457 ¼ 0.531 psi (3.661 kPa) Case B If the elbow at 175 ft. (53.8 m) is raised at 908 with a 6 ft. (1.83 m) radius so that the pipe section from elbow to angle valve 50 ft. (15.25 m) is vertical. The last 75 ft. (23 m) is once again on the horizontal. All the K factors for the valves and fittings remain as before. The K factor for the piping remains the same, but the change in elevation has to be factored in. Pressure drop for 50 ft. rise ¼ (50 ft. r)=144 in.2 ¼ (50 71.23)=144 ¼ 24.73 psi (170.5 kPa), which should be added to that in Case A or 0.531 þ 24.73 ¼ 25.26 psi (174.17 kPa). Case C The same case as in Case A but the 3 in. gate valve is replaced with a 3 in. globe valve with a K factor of 340 0.018 or 6.12. In this case the resulting K factor becomes a net total (340 – 8) (0.018) or 5.97, and means an extra pressure drop 0.137 psi (0.95 kPa) will be added to Case A. In summary, the project engineer should make an effort to select valves and fittings with the lowest possible K factor, piping with a very smooth internal surface, and avoid as much as possible any rise in elevation. During execution of the piping system strict supervision should also be the key word to avoid additional resistance to fluid flow from poor workmanship.

6.4.9

START-UP

OF

PUMPS

Before starting up any cryogenic pump, it should be cooled slowly and completely down to the temperature of the liquid. One has to make sure that the impeller can operate freely at all

Angle valve

B Elevation 50 ft. 15.25 m

A Get valve

Radius 6 ft. 1.8 m Elevation zero

175 ft. to bend 53.4 m

FIGURE 6.16 Pressure drops due to piping components (i.e., valves, tees, elbows, etc.). (Courtesy of Crane Company, 1986. With permission.)


times. In some operations a transfer pump is completely immersed in the liquid to be pumped to make sure that, when the start-up signal is activated, the pump will start-up immediately. In this type of operation, the vertical type of transfer pump may have an advantage over the horizontal configuration. In some field applications the cryogenic pump will not start immediately, and may require anywhere from 5 to 30 min to start pumping regardless of what the pump supplier may claim. (Inadequate priming of the pump is the principal reason for unsuccessful starts.) All cryogenic pump installations must be equipped with a valved vent line and a drain at the suction flange of the pump. If the process involving a cryogenic pump imposes a maximum cutoff time of less than 10 s, as may well be the case in the partial oxidation of hydrocarbon feedstocks for synthetic fuels, it is prudent to connect the oxygen distribution line directly to a small pressurized liquid storage tank (see Figure 6.10). In case the pump does not activate immediately, the storage tank will feed a gaseous product into the pipeline for 30 min at least, so that the main process does not stop. During start-up, the pump will begin to rotate on its own accord due to the gas flow through the pump. This phenomenon is caused by the initial and possibly appreciable vaporization of the boiling liquid. As long as this rotation is well below the operating speed of the pump, the overall effect is negligible. It is imperative, however, to make sure that this rotation does not accelerate close to the design speed. Cooldown may be judged as complete when liquid begins to emerge for the vent line. The vent line must be kept open all through the start-up, and shutoff only when liquid begins to appear. The start-up is then complete. If the pump has been idle for a longer period of time, say over a month, it may be prudent to turn on the sealing gas (dry oil-free nitrogen) for several hours before beginning the cooldown operation, so that if any moisture that enters the machine will be removed, thereby avoiding any icing conditions.

6.4.10 HIGH-PRESSURE RADIAL PUMPS (FIGURE 6.17a,b) High-pressure centrifugal cryogenic pumps have had a long history of successful operation since the early days of partial oxidation, in the 1950s, when reliable and safe oxygen compressors were not yet available. These pumps are vertical multistage pumps involving anywhere from 3 to 21 stages depending on the flow and pressure required. They have been used in various services including oxygen, nitrogen, argon, and liquefied natural gas. They can be used in two ways and are as follows: 1. First, as a column pump within the process extracting liquid oxygen from the main condenser, pumping at a high pressure through the primary heat exchangers where the liquid is vaporized in countercurrent heat exchanger with incoming warm air. In this process cycle the primary heat exchangers must be designed to withstand the high product pressure. For a pressure of over 30 bar indicates a higher plant investment. Moreover, the same process demands a higher energy consumption, around 8%, because of the loss of the latent heat of vaporization during the change of phase from liquid to vapor. This concept must be evaluated against the use of a standard low-pressure Claude cycle equipped with an oxygen compressor, which is always more efficient compared to pumping cycle. The Claude cycle, however, may not be found viable for plants where the high cost of an oxygen compressor, with all its ancillary piping and controls, makes its application uneconomical. This has been found to be the case whenever the oxygen plant is tied in with a liquid nitrogen-scrubbing unit for the production of ammonia. In the latter process, the pumping cycle is economically competitive.


Pump can purge out Insulation fill Connector Nitrogen purge inlet Gasket Fiberglass batting

Barrel drain

Pump can purge in Discharge Perlite fill connector Nitrogen purge exhaust filter 10.2 mm

Grade

Pump mount barrel and insulation 81.3 mm diameter Perlite insulation Insulation Outer barrel 51 mm Pump can 20.3 mm

(a)

Nitrogen sparger End plate (wrap for protection) Concrete base

Per pump length

15.25 mm

FIGURE 6.17 (a) Pump mount barrel. continued

2. Another application of the high-pressure pump is to serve as a back-up, supplying large quantities of product to keep the overall process in operation during a temporary emergency shutdown of the air separation plant. In this case, the pump is tied in with a large liquid product storage vessel at the intake, and linked with a high volume steam vaporizer at the discharge side. The storage tank supplying the pump must be designed in such a manner to deliver enough products for at least 2 d depending on the longest possible shutdown foreseen for the oxygen plant. A 3 d supply is more prudent, especially for a metallurgical operation wherein any shutdown is very costly. Capital cost of liquid storage may be minimized by supplying liquid product from large tanker trucks from nearby merchant plants. The cost of this option is very high, however, both in terms of capital and anxiety.


LOX from tank

Break out spool Vent, etc. Connector

Add spool for additional connections if required

Vacuum jacketed valves Support

Support

(b)

FIGURE 6.17 (continued) (b) Cryogenic pump connections. (From Kerry, F.G., Inc.)

6.4.11 ULTRAHIGH PRESSURE PUMPS In recent studies for the production of synthetic fuels, the compression of large quantities, up to 2000 t=d of oxygen to pressures of 100 bar, and even to 140 bar, became a problem. Oxygen compressors are already operating at pressures of up to 60 bar, but above this pressure seal losses become very significant, and increase geometrically as the pressure gets higher. In these conditions efficiencies drop markedly. In fact, the projected efficiencies for centrifugal compressors as a function of pressure are given in Table 6.3. An experimental radial machine has been developed to compress oxygen from 60 to 125 bar with a separate casing housing four extra stages. Polytropic efficiency was stated as 66%, but the average was closer to 60%, as projected in Table 6.3. The cost of such an industrial

TABLE 6.3 Projected Efficiencies at Various Pressures Pressure (bar) 40 70 80 100 125

Efficiency (%) 63–65 62–64 61–63 60–62 59–61



machine may prove excessive, as it used Monel metal for all piping and valves. A review of a pumping cycle with reciprocating pumps may well worth the effort. Such pumps have already proved their worth in the pumping of nitrogen up to 145 bar for enhanced oil recovery, but with relatively small flows (2–4 gpm). They are rugged, easy to maintain, and can be operated in tandem and at a relatively low capital cost. The design of a multistage centrifugal cryogenic pump may not be described here, because of economic and safety considerations. According to Karassik et al.9, the design parameters for a high-pressure pump have the following relationship: Q2 ¼ Q1 (D2 =D1 ) and

DH2 ¼ DH2 [(D2 )2 =D1 ]

(6:1)

It may involve a pump flow of 300 gpm (19.14 L=s of liquid oxygen), an impeller diameter exceeding 254 mm (10 in.), and a tip speed of 47 m=s. Moreover, the shaft can have a length of around 3 m, and support approximately 24 stages, which poses quite a design as well as an operating problem. The design for shaft bearings as well as for the motor will be a major problem in terms of safety according to Baur et al.4. Added to this is the question of seals. The choice of materials for high-pressure cryogenic pumps follows the same parameters as those for low- and medium-pressure transfer pumps in terms of operational reliability and safety. In the case of vertical multistage pumps, however, extra care must be taken in the design of the shaft and impeller to avoid rubbing against the casing. The choice of metals is most critical. For example, a greater use of softer and more ductile metals such as Monel and cuprous alloys with a minimum content of 40% nickel when in contact with stainless steel is highly recommended. In summary, before concluding that liquid oxygen pumping is the way to go one must keep in mind that liquid oxygen pumping extracts from the cryogenic system the ability to recover the equivalent of the product’s latent heat at its operating conditions, resulting in a fairly large cold end heat load. This loss will be a power penalty of 3%–14% depending on the final product pressure.

6.4.12 AUTOMATION As with pumps for higher temperature service, cryogenic pumps can be computer-controlled from a central distributor system, or put under fully programable logic controls (PLC) and distributed control systems (DCS) for cooldown operation. Proximity probes will indicate any axial movement of the rotor assembly, forestalling radial-bearing failure, and imbalance of rotating members. A resistance temperature detector (RTD) may be used as an automatic shutdown device. Monitoring bearing temperature is necessary on the fan side and especially on the pump side. Pump side-bearing temperature must not be allowed to drop below 24.8 K. Specified minimum and maximum bearing temperatures should be relayed to the central control system and to an alarm and trip.

6.5 6.5.1

CRYOGENIC LIQUID VAPORIZERS10 GENERAL OVERVIEW

When oxygen, nitrogen, or argon are liquefied for back-up storage or transport to a distant point the cryogens must then usually be vaporized for use. A wide variety of equipment is available for vaporization. Selection will depend on the locale, utilities available, and cost. Such equipment falls into two categories:


. .

Low duty ambient air vaporizers High duty steam and hot water vaporizers

6.5.2 AMBIENT AIR VAPORIZERS (FIGURE 6.18) These vaporizer heat exchange elements are generally made of extruded aluminum fins called Thermofins. They have a tapered geometry, with a cross-section surface similar to a star. Each of these Thermofins has a central core, which serves as a passage for the cryogen to be vaporized. Several of these fins are attached to form an individual module. The vaporization capacity of each module depends on the number of Thermofins that are in the module. The cryogen is distributed through the central core of the Thermofins in the module, in heat exchange with the ambient air passing around the combined external surface of the module. Depending on the specific design of the various groupings, modules can be rated to vaporize anywhere from 170 to 1340 Nm3=h of gaseous product. Using stainless steel for the internal cores, the cryogen can be pumped and vaporized at a higher gaseous product pressure (up to 98 bars), but of course, at a slightly lower rate of vaporization. 6.5.2.1 Modules Spread Apart A more recent design of the ambient air vaporizers has involved the use of a wide gap between two groupings of the Thermofins in the same module, extending the time necessary to shut down the operation for purposes of defrosting. In this manner, the new design, called Super Gap Thermofins, extends the operating time to be at least 60%. 6.5.2.2 Modules in Alternate Operation In a further development of the same design of the ambient air vaporizer, two modules can be piped in a parallel formation and operated alternately, so that one module can be in operation whereas the parallel module is undergoing defrosting. This design can be operated automatically in a manner similar to reversing exchangers in a standard air separation plant. 6.5.2.3 Modules with Pressurized Air Other variations of the same vaporizer involve the use of a cryoduct fan-boosted ambient air vaporizer with a unit capacity of 2700 Nm3=h, and the use of a defrostable fan ambient vaporizer with two units and an automatic electric defroster for continuous operation.

6.5.3 DIRECT STEAM VAPORIZERS (FIGURE 6.19) Also available are high-performance automatic vaporizers using steam ranging from 0.7 to 10 bar (70–1000 kPa). These vaporizers are of special interest to industries such as steel mills, which have a fair quantity of waste steam available at 3 bar (300 kPa). These units have a vaporization capacity ranging from 700 to 34,000 Nm3=h (oxygen). The main components of these high-performance vaporizers involve: . . .

.

A horizontal steam chamber A removable U-bend tube bundle of 18-BWG-stainless steel tubes A gas discharge temperature controller connected to a pneumatic steam valve and a condensate trap assembly A computerized full ASME VIII Division 1 Code design


Useful range days before excessive ice clog 6 21 Cold/dry 2 7 (winter) Warm/humid (summer)

10

O2/N2 flow (SCFH)

7

70,000

5 3

30,000

Superg

ap (S-G

2

)

Sta nd

ard

1 10,000 7 5

5,000

0.1

0.2

60% Gain

0.3 0.5 0.7 1.0 2 3 5 7 10 14 21 Days continuous operation

28

FIGURE 6.18 Super gap vaporizers. (Courtesy of Thermax Corporation, 2000. With permission.)

Operation is fairly simple. Liquefied gas enters the tube and is vaporized and superheated by the steam condensing in the steam chamber. The warm gas leaving the unit activates the controller, which in turn throttles the steam valve. The final condensate trap drains the spent steam from the chest when holding the steam under pressure within the chest. At lowtemperature settings, or low throughputs, a vacuum breaker on the chest admits air in it, cutting heat to the liquefied gas, and preventing condensate buildup in the chest. The attached diagram illustrates the equipment involved. 6.5.3.1

Vaporization with Steam-Heated Water

Another variation involves the use of a steam-heated hot water bath (Figure 6.19). This application is useful if more than one cryogen needs to be vaporized at different capacities


Steam vaporizer supplied by thermax

Steam in

Pneumatic steam control valve

Steam strainer

Steam shell relief set Valve bypass (optional)

SI H

Options

Gas discharge N2 N4

Vacuum breaker

RE Gas discharge

Temp. control instr. air 1.7 Barg. trim FPT

Steam valve a/o trap block and bypass valves prop and reset temperature controller low temperature shutdown

N3

Weatherproof insulating jacket

Drip trap connector 13mm FFT

RE Liquid in FE

N1 Main liquid in

CI

Other ratings Main trap Ethane: same as connector propane

L

Bolt hole 4 PLCs

Leave (5 + 50 cm) clear to pull bundle

Pressure drop S style 1.4 BARD at 13.8 BARG/SH style 2 BARG at 13.8 BARG For 75% flow use 1/2 of these values Other values call thermaz for bulletin PC-13

FIGURE 6.19 Steam vaporizer. (Courtesy of Thermax Corporation, 2000. With permission.)

simultaneously. Two or more separate coils, each carrying a different cryogen, are submerged in the same bath heated by steam or electrically, depending on the local cost of utilities.

6.5.4 EMERGENCY VAPORIZATION

OF

PRODUCTS

Another application is also very useful. For example, in an emergency shutdown of the local air separation plant, a product flow from a small high-pressure storage tank can be released and vaporized immediately in order to maintain product flow, whereas the emergency highpressure product pumps are activated to supply product from the larger storage tank. In such a situation, the small auxiliary high-pressure storage tank should have a capacity of around 30 min supply of product.

REFERENCES 1. Springman, H. 1978. The liquefaction of oxygen, nitrogen and argon. Linde reports on science and technology, No. 28, pp. 23–30. 2. UBC Uniform Building Code Division IV Earthquake Design. 3. Carter, T.A. 1971. Oxygen Compressors and Pumps Symposium, November 9–11, 1971. Compressed Gas Association Inc. pp. 57–60. 4. Baur, H. et al. Fire tests on centrifugal pumps for liquid oxygen. Kaltechnik, Klimatisienung No. 4, 1970, and No. 4, 1971. 5. Compressed Gas Association, Installation guide for horizontal. Stationary electric-motor driven liquid oxygen pumps, 1st ed., CGA G-4.7, 2006.


6. Grohmann, M. 1979. Extend pump applications with inducers. Hydrocarb Process (December) 121–124. 7. Martin, G.R. 1996. Pumps and NPSH: Avoid problems and improve reliability. Hydrocarb Process (May) 63–65. 8. Flow of fluids through valves, fittings, and pipe, 25th ed., Technical Paper No. 410 and 410 M, Crane Publication, 1991. 9. Karassik, I.J. et al. 1976. Pump handbook, 2–32. New York: McGraw-Hill. 10. Bernert, R. Thermax Incorporated, 3 Pleasant Street, South Dartmouth, MA 02748.


7

Insulation

7.1 GENERAL In cryogenic systems for the liquefaction, separation, and purification of gases, one has to deal with temperatures of around 112 K for liquid methane, and as low as 4.2 K for liquid helium. The design engineer, therefore, has to provide adequate insulation for the processing and ancillary equipment subjected to these very low temperatures, in order to reduce the heat gain to a practical minimum. Any heat gain from the outside ambient atmosphere means a corresponding loss in refrigeration applied to the process cycle. Since the cost of externally applied refrigeration, energy for compressors, mechanical refrigeration units, far outweighs the cost of insulation, it is therefore prudent to spend the necessary time to study the requirements for insulation, instead of leaving it in the hands of others or of the cheapest insulation supplier. The lower the processing temperatures, the greater the need for properly specified insulation, in terms of type, quantity, quality, and density. For industrial purposes any temperature lower than 120 K may be considered as in the field of cryogenic processing. Cryogenic process systems may involve any or all of the following materials and equipment operating at very low temperatures: 1. Processing equipment, such as heat exchangers, distillation columns, and piping, enclosed inside steel casings commonly referred to as the cold box 2. Externally located process or transfer piping 3. Storage vessels for cryogens (cryogenic liquids) All these elements require properly specified insulation in order to reduce heat gain (heat flow from the outside ambient conditions into cold equipment), and thus maintain a low-energy consumption for the process. Moreover, each element may require a separate study because of the size, configuration, and duty of the specific elements.

7.1.1 THEORETICAL CONSIDERATIONS Although the term heat gain (W=m2) is synonymous to heat flux and heat transfer, the former should be used for cryogenic systems because it implies that heat is entering the system. The basic principles still involve the same concepts as heat transfer. Therefore, a quick review of these principles and their terminology may be useful. This chapter is not intended for design purposes, but for understanding and evaluating various insulations available in the present market, and their proper selection.

7.1.2 INSULATIONS: GENERAL The thermal conductivity of any insulation is based on interaction of any or all of the following parameters: 1. Thermal conductivity of the solid components forming the insulation material or structure 2. Thermal conductivity of air, or of any residual gas within the insulation


3. Heat transfer by natural convection within the cells or pores of the material 4. Heat transfer by radiation within the structure of the material

7.1.3 VACUUM INSULATION (RADIATION) Heat transfer with no material carrier Vacuum insulation is almost synonymous with the term Dewar flask named after James Dewar, who in 1892 not only conceived the idea of vacuum insulation, but also built the first double wall vacuum-insulated flask. However, the theory behind the concept is a little more involved than the original basic product. This is especially true in the use of multilayer insulation (MLI), which involves a large number of reflecting surfaces. The Stefan–Boltzmann law deals with the quantity of heat transferred by radiation, Q, from a blackbody that absorbs all radiant energy, which strikes it. It is proportional to the fourth power of the body’s absolute temperature. Since a blackbody absorbs and emits the maximum amount of energy at a given temperature, it is necessary to introduce an emissivity factor Fe to estimate the radiation of an actual body. The emissivity is the ratio of the actual radiation to the radiation of an analogous blackbody. To calculate heat gain from radiation, the modified Stefan–Boltzmann equation may be stated as follows1: Qr ¼ sFe F12 A1 (T24 T14 )

(7:1)

where Qr is the radiant heat transfer rate per unit area, s is the Stefan–Boltzmann constant (56.69 nW=m2=K4 (n ¼ nano ¼ 109)), Fe is the emissivity factor, F1–2 is the configuration factor, A1 is the area of colder surface (m2), and T is the absolute temperature (K). For standard cryogenic double wall vacuum-insulated vessels F1–2 ¼ 1, where subscript 1 refers to the inner vessel, and subscript 2 refers to the outer vessel. The emissivity factor for diffuse radiation for concentric spheres or cylinders is given by 1=Fe ¼ [1=e1 þ (A1=A2 )(1=e2 1)]

(7:2)

where e1 is the emissivity of the colder surface, A1 is the area of the colder surface, e2 is the emissivity of the warmer surface, and A2 is the area of the warmer surface. If a thermally isolated floating highly reflective shield is placed between the warm and cold surfaces, then the emissivity factor becomes as follows: 1=Fe ¼ (1=e1 þ 1=es 1) þ (N 1)(2=es 1) þ (1=e1 þ 1=es 1)

(7:3)

This equation shows that heat gain can be reduced tremendously. For example, in the basic equation with no radiation and no shields, if e1 ¼ e2 ¼ 0.80, then Fe ¼ 0.6667, but if 10, highly reflective shields each with an emissivity of 0.05 were interposed, then the overall emissivity becomes 0.00255. In other words, radiation heat gain is reduced by a factor of 0.00383. The advantage of using reflecting shields with any insulation becomes obvious as will be shown in MLIs (Figure 7.1). In a vacuum insulation the principal heat gain by radiation is also supplemented from a smaller heat gain by free molecular conduction. The residual molecules remaining in the evacuated space carry energy from the outer warmer surface to the colder inner surface without transferring energy to other molecules. They never remain on either surface long


T2

! T1

T1

Warm surface

Cold surface

T1

FIGURE 7.1 Energy transportation from free molecular movement. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

enough, however, to establish a thermal equilibrium. This inefficiency, therefore, is expressed by the accommodation coefficient a, which is a ¼ actual energy transferred=maximum possible energy transfer The accommodation coefficient a is based on a specific gas-surface combination, and the surface temperature at the warm end and at that on the cold side. For example, a1 ¼ (T21 T11 )=(T21 T1 ), at the cold surface a2 ¼ (T21 T11 )=(T2 T11 ), at the warm surface

(7:4)

T2 T1 ¼ (1=a1 þ 1=a2 1)(T21 T11 ) ¼ (T21 T11 )=Fa

(7:5)

and

Fa is the accommodation factor that has the same form as the emissivity factor, and can be calculated in the same manner 1=Fa ¼ 1=a1 þ A1 =A2 (1=a2 1)

(7:6)

Approximate typical accommodation coefficients are given in Table 7.1.

TABLE 7.1 Accommodation Coefficients for Air, Hydrogen, and Helium Temperature (K) 300 77.7 20

Air

Hydrogen

Helium

0.8–0.9 1.0 1.0

0.29 0.53 0.97

0.29 0.42 0.59

Source: From Timmerhaus, K. and Flynn, T. in Cryogenic Engineering, Springer Science & Business Media, New York, 1989. With permission.


To determine the total heat gain of a typical double wall vacuum-insulated vessel, the following factors must be considered: 1. The emissivity factor, 1=Fe ¼ [1=e1 þ (A1 =A2 )(1=e2 1)]

(7:2)

2. The radiant heat transfer, Qr ¼ sFe F12 A1 (T24 T14 )

(7:1)

3. The accommodation coefficient factor, 1=Fa ¼ [1=a1 þ A1=A2 (1=a2 1)]

(7:6)

4. The heat gain from free molecular conductivity heat gain, or Qc (for concentric cylinders and spheres). The energy supplied by the free residual molecules is given by De ¼ (cv þ (1=2)R)(T21 T11 )

(7:7)

where R is the specific gas constant (R ¼ Ru=M), Ru is the universal gas constant (8.31441 kJ=kmol K), M is the molecular weight of the gas, cv is the specific heat of gas [R=(g1) for an ideal gas], and g is the specific heat ratio, cp=cv (g ¼ k at times). Eliminating effective molecule temperatures, and from the kinetic theory of gases, the mass flow rate of molecules per unit of surface is m=A ¼ p(T=2pgc R)1=2

(7:8)

where T is the gauge temperature of gas (K), and gc is the conversion factor in Newton’s second law of motion, which states that acceleration of an object is directly proportional to the resultant forces acting on it and inversely proportional to its mass, or SF ¼ ma. This is a vector expression, hence proportional to three component equations. The equation is valid only when the speed of the particle is much less than the speed of light. In SI units, mass ¼ kg, acceleration ¼ m=s2, force ¼ N ¼ kg m=s2, and N ¼ 105 dyn (9.80665 m=s2). The energy transfer rate by molecular conduction is therefore Q=A ¼ (m=A)De

(7:9)

or Q ¼ GpA1 (T2 T1 ) and G ¼ [(g þ 1)=(g 1)][(gc R=8pT)1=2 Fa


(7:10)

Free molecular conductivity to occur, however, the mean free path (l) of the gas molecules must be larger than the spacing between the two surfaces. To determine this l ¼ (m=p)(pRT=2gc )1=2

(7:11)

where m is the gas viscosity at the temperature (K), p is the absolute pressure of the gas, gc is the correction factor in Newton’s second law of motion, R ¼ Ru=M or (universal constant=molecular weight). The total heat transfer, therefore, can be calculated by Qt ¼ Qr (radiant) þ Qg (molecular energy)

7.1.4 CONDUCTIVITY

IN

(7:12)

MASS INSULATIONS

Conductivity follows Fourier’s law for gases as well as for solids and can be expressed by Q ¼ km Am (Th Tc )x

(7:13)

where Q is the heat transferred (W=m K) across the mean heat transfer surface of A (m2), Th is the temperature (K) of the warm surface, Tc is the temperature (K) of the cold surface, Dx is the thickness of the insulating material (m), and km is the coefficient of thermal conductivity of the mass insulation. The negative sign indicates that the temperature decreases as the thickness from the warmer side increases. Heat transfer by conduction can be reduced by using an insulating material having a low thermal conductivity coefficient.

7.1.5 NATURAL CONVECTION IN MASS INSULATION Convection can be reduced by packing the mass insulation so the dead airspace between the outside warm casing and the cold cryogenic equipment within can be broken up into an unlimited number of very small cells or spaces. This, however, will also increase the effects of conduction from the material used. The best possible mass insulation, therefore, will be one that will reduce convection as much as possible, yet have a very low overall coefficient of conductivity. Nusselt and Bayer2 developed the following expression for the apparent thermal conductivity of gas-filled powders and fibrous insulation, but not for evacuated insulation materials km ¼ [r=ks þ 1={kg =(1 r) þ 4sT 3 d=r}]1

(7:14)

where km is the thermal conductivity of the material, r is the ratio of solid volume to total volume, ks is the thermal conductivity of the solid insulation material, kg is the thermal conductivity of the residual gas within the material, s is the Stefan–Boltzmann constant (56.69 nW=m2=K4) (n ¼ nano ¼ 109), T is the mean temperature of insulation (K), and d is the mean diameter of powder or fiber. Taking into consideration that at very low temperatures the term involving T3 is much smaller than kg, the above expression may be reduced to km kg =(1 r)


(7:15)

The thermal conductivity of gas-filled powders or fibers approximates that of the gas within the insulation. Except for very fine powders, where the distance between the powder particles is smaller than the mean free path l of the gas molecules. In such cases the effective gas thermal conductivity is reduced because the gas thermal conductivity is proportional to the mean distance traveled by the molecules. In fine powder insulation, the gas used should not react with the insulating material. For safety purposes an inert gas such as either nitrogen or carbon dioxide is used.

7.1.6

VACUUM PLUS POWDER

OR WITH

FIBROUS INSULATIONS3

As shown, the thermal conductivity of powder and fibrous insulations approaches the thermal conductivity of the residual gas within the insulation at cryogenic temperatures: km kg =(1 r)

(7:15)

Q ¼ km Am (Th Tc )Dx

(7:13)

and

It is also possible to further reduce the heat gain by evacuating the gas from within the insulation. As shown in Figure 7.2, from atmospheric pressure to about 13.332 mbar (1.33 kPa), there is a very little change in thermal conductivity. But from 1.33 kPa to 1.33 Pa, the conductivity 105

s

Note: Nitrogen and helium identify the interstitial gas in the powder.

ere

k

lac

ph

b mp

cs

La

Ka (mW/m K)

Ph

en

oli

104

Silica

us eo c a m th to r ia ea

aerogel

D

103

Perlite

T2 = 300 K T1 = 76 K Nitrogen

el

rog

e aa

ic

Sil

Perlite

102 10−3

10−2

T2 = 76 K T1 = 20 K Helium 10−1

10

101

102

103

104

Pressure (Pa)

FIGURE 7.2 Apparent mean thermal conductivities of several powder insulations as a function of interstitial gas pressure. (From Perry, R.H. and Green, D., Chemical Engineer’s Handbook, 6th ed., McGraw-Hill, New York, 1984. With permission.)


is reduced in a linear manner. This is the region of free molecular conduction for the gas within the insulation wherein kg is nearly proportional to the gas pressure. For any further reduction in gas pressure, the gaseous heat transfer becomes smaller than the heat transfer by radiation and solid conduction. In such applications, the heat transfer rate can be determined from the Fourier rate expression for conduction heat transfer using the apparent thermal conductivity of insulation. Q=Am ¼ [k(Th Tc )]Dx

(7:16)

where Dx is the thickness of insulation, Th is the temperature of warm surface, Tc is the temperature of cold surface, Am is the mean heat transfer area given by Am ¼ (A2 A1 )=ln (A2 =A1 ) for concentric cylinders

(7:17)

Am ¼ (A1 =A2 )1=2 for concentric spheres and elliptical heads of tanks

(7:18)

where A1 is the surface area of the enclosed surface and A2 is the surface area of the enclosure. For torispherical and elliptical heads on ends of cylinders the area may also be calculated from (A1=A2). The surface area of a standard ASME torispherical head is given by A ¼ 0:264pD2

(7:19)

where D is the diameter of the straight flange. For elliptical heads the surface area is given by A ¼ (1=4)pD2 {[1 þ e2 =2«] ln (1 þ «)=(1 «)}

(7:20)

where D is the major diameter of head, D1 is the minor diameter of ellipse, and « ¼ [1(D1=D2)]1=2. For standard 2:1 elliptical heads D1=D ¼ 0.5 and « ¼ 0.865, therefore A ¼ 0.345D2. The Fourier rate expression (Equation 7.20) on heat transfer must be modified for rods supporting cryogenic storage vessels and which may have the following variable thermal conductivity: Q ¼ [km A(Th Tc )]=L ¼ (Kh Kc )A=L

(7:21)

where km is the mean thermal conductivity of rod, or (KhKc)=(ThTc), Th is the temperature of the warm end of rod, Tc is the temperature of cold end of rod, A is the cross-sectional area of rod, and L is the length of rod. K¼

ðT

kt dT ¼ thermal conductivity integral

4K

Values for thermal conductivity integrals are given in Table 7.2.4 Example Heat gain through evacuated insulation, vessel, and ancillary equipment


(7:22)

TABLE 7.2 Representative Apparent Thermal Conductivity Values for Various Insulations Type of Insulation

Apparent Thermal Conductivity ka (Heat Gain) between 77 K (mW=m K)

Pure gas at 1013 mbar (180 K) H2 N2 Pure vacuum 106 mbar

104 17 5

Foam insulation Polystyrene foam (46 kg=m3) Polyurethane foam (11 kg=m3) Glass foam (140 kg=m3)

26 33 35

Nonevacuated powders Perlite (50 kg=m3) Perlite (210 kg=m3) Silica aerogel (80 kg=m3) Fiberglass (110 kg=m3)

26 44 19 25

Evacuated powders and fibers (1.3103 mbar) Perlite (60–180 kg=m3) Silica aerogel (80 kg=m3) Fiberglass (50 kg=m3)

1–2 1.7–2.1 1.7

Opacified powder insulations (1.3103 mbar) 50=30 wt% Al=Santocel (160 kg=m3) 30=50 wt% Cu=Santocel (180 kg=m3)

3.5 101 3.3 101

Multilayer insulations (1.3106 mbar) Al foil and fiberglass 12–27 layers=cm 30–60 layers=cm

3.5–7 102 1.7 102

Al foil and nylon net 31 layers=cm

3.5 102

Al crinkled, Mylar film 35 layers=cm

4.2 102


Assume a horizontal vessel with a water capacity of 26,000 US gallons (98.4 m3); evacuated perlite insulation with apparent thermal conductivity of 1.20 mW=m K; an inner shell 13 m long by 3.070 m OD and hemispherical heads 3.3062 OD; outer shell 13 m long by 3.636 m ID and hemispherical heads 3.648 m OD; 52 ss Type 304 support rods with diameter of 20 mm; 20 vertical, 1.15 m long; 20 transversal, 1.10 m long; 12 longitudinal, 1.905 km long; fill drain pipe 80 mm nominal Schedule 5, stainless steel 7 m long; vent pipe 100 mm nominal Schedule 5, 304 stainless steel 8 m long. Liquid oxygen temperature in vessel at 90.2 K; external ambient temperature 301.5 K (Figure 7.3).


Safety valve Relief valve Rupture disc Vent valve

Fill/ drain Outer shell ID = 3.636 m Inner shell ID = 3.070 m Insulation Δx Δx = 0.283 m

Inner vessel Length = 13 m for cylinder Hemispherical ends = 3.3062 m OD External vessel Length of cylinder = 13 m Hemispherical ends = 3.64 m OD

FIGURE 7.3 Heat gain through evacuated insulation and ancillary equipment. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

Heat transfer through evacuated perlite insulation Inner surface area ¼ pDiL ¼ 3.14159 3.07 13 m ¼ 125.381 m2 Outer surface area ¼ pDoL ¼ 3.14159 3.636 m ¼ 148.497 m2 Mean area from Equation 7.21 Am ¼ (148:497 125:381)=0:1692 ¼ 136:613 m2 Thickness of insulation around shell: (3.636 3.070)=2 ¼ 0.283 m Heat transfer rate from Equation 7.20 Q1 ¼ (1:2 103 )(147:1)(301:5 90:2)=0:283 ¼ 131:797 W Heat transfer through insulation around spherical heads Inner surface both heads ¼ pD21 ¼ 3.14159 (3.062)2 ¼ 29.455 m2 Outer surface both heads ¼ pD21 ¼ 3.14159 (3.648)2 ¼ 41.808 m2 Mean area from Equation 7.22 Am ¼ (29:455 41:808)1=2 ¼ 35:092 m2 Thickness of insulation around heads: (3.648 3.062)=2 ¼ 0.293 m Heat transfer rate from Equation 7.20 Q2 ¼ (1:2 103 )(35:09)(301:5 90:2)=0:293 ¼ 30:367 W Heat transfer through ancillary equipment Consider one support member. Cross-sectional area ¼ (p=4)D2 ¼ (3.1416=4) (0.02)2 ¼ 3.14 104 m2 Total of all three types of support members: A=L ¼ (20)(3:14 104 )=(1:15) þ (20)(3:14 104 )=(1:10) þ (12)(3:14 104 )=1:905 ¼ 0:01315 m Heat transfer rate from Table 7.2 and Equation 7.21 Q3 ¼ (3060 438)(0:01315) ¼ 34:49 W


Heat transfer through fill=drain, and vent piping Fill=drain cross-sectional area: p(Do t)t ¼ p(0.1016 0.0021) 0.0021 ¼ 6.564 104 m2 Vent pipe cross-sectional area: p(0.1143 0.0021) 0.0021 ¼ 7.402 104 A=L ratio for both pipes: (6.564 104 (7.0)) þ (7.402 104)=(8.0) ¼ 1.863 104 m Heat transfer rate: Q4 ¼ (3060 438) 1:863 104 ¼ 0:49 W Total QT ¼ 131.797 þ 30.15 þ 34.49 þ 0.49 ¼ 196.897 or 196.9 W Total heat gain from insulation ¼ (131.797 þ 30.15)=196.9 ¼ 82.25% Calculation of normal evaporation rate (NER) per day Energy required to evaporate liquid oxygen: Et ¼ rf (fluid) hhg (latent heat at boiling point) V (volume) ¼ 1141:0 213 98:4 ¼ 2:39 107 kJ Energy transfer in 1 d: Ed ¼ (196:9 103 ) 60 s 60 m 24 h ¼ 1:7012 104 kJ Therefore, NER ¼ Ed =Et ¼ 1:7012 104 =2:39 107 ¼ 0:0712% per day

7.1.7

INSULATION (MULTILAYER, SUPER, OR SIMPLY MLI)5,6

Radiant heat transfer rate as described in Section 7.3 can also be reduced considerably by interposing floating (thermally isolated) low-emittance radiation shields, i.e., aluminum foil, copper foil, or aluminized Mylar, between the outside warm surface and the inside cold surface. For example, a modified Stefan–Boltzmann expression can be used for a variable number of concentric enclosures (N). Q ¼ sFeo A1 (TN4 T14 )

(7:23)

where TN is the absolute temperature of the outermost surface, T1 is the absolute temperature of the innermost surface, and Feo is the overall emissivity factor given by s, where s is Stefan– Boltzmann constant ¼ 56.69 nW=m2=K4. Considering the emissivity factor for N shields N1 1=Feo ¼ Si¼1 A1 =Ai [1=e1 þ Ai =(Aiþ1 )(1=eiþ1 1)]

(7:24)

If the emissivities of the outermost and innermost surfaces as well as the interposed shields have equal emissivities, then e1 ¼ eo ¼ eN, and for N shields or (N þ 2) surfaces and using a shield emissivity of e1 ¼ es (i ¼ 2, 3, . . . , N 1) the emissivity factor becomes Fe (Ns shields) ¼ [2(1=eo þ 1=es 1) þ (Ns 1)(2 es )=es ] Assuming that eo ¼ 0.9 and es ¼ 0.05, then Fe (Ns shields) ¼ (1:2 þ 39Ns )1 and Fe (no shields) ¼ 0:889 and for 10 shields Fe=Feo (N) ¼ 348.


(7:25)

If 10 reflecting shields are used the emissivity factor Fe (radiation heat transfer rate) will be reduced by 1=348, or 0.0029, which is very appreciable indeed. For a well-evacuated spacer insulation, heat is transferred primarily by radiation and solid conduction, the latter through the spacer material placed between the shields. Using the basic principle of multiradiation shields, it is possible to reduce the apparent thermal conductivity to a very low point employing alternating layers of highly reflecting material such as aluminum foil, copper foil, or aluminized Mylar. These may be separated by low-conductivity spacers, such as Mylar, glass-fiber material Dexiglas, or Tissueglas having a nominal thickness of 0.7 and 1.5 mm, respectively. The entire system is then highly evacuated as low as 1.33 106 mbar. In these conditions, the mean apparent thermal conductivity can be calculated as follows: ka ¼ 1=N Dx{hs þ (seTh3 =2 e)[1 þ (Tc =Th )2 ][1 þ (Tc =Th )]}

(7:26)

where hs is the solid conductance of spacer material per unit of thickness, s is the Stefan– Boltzmann constant, e is the effective emissivity of reflecting shields, Tc is the cold end boundary surface, Th is the warm end boundary surface, and N is the number of layers. Example What is the mean apparent thermal conductivity of an MLI insulation between 300 and 20 K, if the insulation is made of 24 layers of aluminum foil with e ¼ 0.05, and fiberglass paper spacers whose solid conductance is 0.0851 W=m2 K: ka ¼ 1=(24 100){[0:0851 þ (5:669 108 )(0:05)(300)3 ]=(2 0:05)} [1 þ (20=300)2 ][1 þ (20=300)] ka ¼ 55:5 mW=m K

Equation 7.25 shows that the difference in the warm and cold temperature and number of shields are major factors in calculating the apparent coefficient of thermal conductivity. Also as shown by the work carried out by Black and Glaser,7 heat flux is influenced by compressive loads on the insulation. In fact, it is proportional to 1=2 to 2=3 power of the externally applied pressure. With the use of an MLI, it is possible to reduce all modes of heat transfer: radiation, solid conduction, and gaseous conduction, simultaneously to a practical minimum. To summarize, radiant heat transfer is inversely proportional to the number of interspaced shields, and directly proportional to the emissivity of the shields. Radiation can be minimized by using many layers of a low-emissivity material such as copper or aluminum. Convection is eliminated by lowering the pressure so that the mean free path of the molecules is much larger than the spacing between the layers of insulation. Even free molecular conduction is minimized by employing a high vacuum. This factor is proportional to the residual gas pressure. Heat transfer through the spacer material is proportional to the thermal conductivity of the spacer material, and inversely proportional to the resistance to heat flow at the areas of contact between the particles of the spacer material. Such areas of contact involve size, geometry, and discontinuity of fibers in the spacer material (Table 7.2 and Table 7.3).

7.2 INDUSTRIAL PRACTICES 7.2.1 INDUSTRIAL APPLICATIONS

OF INSULATION

The use of a double wall vacuum-insulated flask for storing cryogenic liquids to serve cryogenic laboratories was cost efficient, but very costly and impractical for industrial scale


TABLE 7.3 Emissivities of Various Metals Metal Aluminum(annealed)

Aluminum vaporized on 12.7 mm Mylar (both sides) Aluminum foil Kaiser annealed (25.4 mm) Aluminum foil (household) Aluminum (Alcoa No. 2, 510 mm)

Brass (rolled plate) Brass Brass (73.2% Cu, 26.72% Zn) Brass shim stock (65% Cu, 35% Zn) Brass, shim stock (65% Cu, 35% Zn) Copper Copper foil (127 mm) 301 Stainless 316 Stainless 347 Stainless Electroplate silver Silver polish Copper

Surface Preparation Electropolished Electropolished Electropolished — — — Hot acid cleaned Alcoa process Alkali cleaned Natural surface Clean, some scratches Highly polished Highly polished Highly polished Commercial emery polish Dilute chromic acid Cleaned with toluene and methanol Cleaned with toluene and methanol Cleaned with toluene and methanol Commercially supplied Commercial emery polish

Surface Temperature (K)

Emissivity Total Normala

300 300 300 300

0.03 0.018 (77 K) 0.011 (42 K) 0.04 (77 K)

300

0.018 (77 K)

273

0.043 (77 K)

300 300 295 273 520 295 295 292 300 297 297 297 300 300 292

0.029 (77 K) 0.035 (77 K) 0.06 0.10 (77 K) 0.028 0.029 (77 K) 0.018 (42 K) 0.03 0.017 (77 K) 0.021 0.028 0.039 0.017 0.0083 (77 K) 0.03

a

Denotes absorptivity at the temperature listed for blackbody radiation from a source at the temperature listed in the preceding column. Source: From Timmerhaus, K. and Flynn, T. in Cryogenic Engineering, Springer Science & Business Media, New York, 1989. With permission.

applications. Now, vacuum insulation is primarily used for small-scale liquid storage, or in combination with some other form of mass insulation. The latter combination of vacuum and porous insulation is also used for the efficient transfer of liquid product in pipelines. Over the years it has been found that mass insulation such as fibers and porous powders is more economical and easier to use for carbon steel casings housing cryogenic processing equipment. Materials that are used include mineral wool, fibrous asbestos, glass wool, vegetal cork, Santocel, calcium silicate, and expanded perlite. The choice depends on factors such as price, shipping costs, local availability, and ease of handling in the field. Presently, the popular choice for mass insulating material, especially for cryogenic boxes, has been narrowed down with some minor exceptions to mineral wool and expanded perlite. Vacuum insulation in combination with either porous or MLI is applied to small vessels, bulk-liquid storage vessels (up to 10,000 L, or 10 t), for liquid transfer lines of significant length (over 20 m) or for liquid hydrogen and liquid helium lines regardless of length. Small cryostats are also included in this field of application.


7.2.2 CRYOGENIC CASINGS (COLD BOXES)

FOR

PROCESS EQUIPMENT

Normal industrial practice is to arrange processing equipment such as heat exchangers, distillation columns, and process piping as tightly as possible within limits of casing (cold box). The casing is then filled with the specified insulation. Sometimes several cold boxes of different dimensions may be involved. These may have one common interface or be completely separate, in which case they will be connected to external process piping also housed in its own insulated casings. While the foregoing may sound simple enough, the design engineer must consider the following factors: 1. Temperature differential between the process equipment and the outside ambient conditions throughout the year 2. Heat transfer rates of the insulating material 3. Resistance to fracturing (dusting) of the insulating material 4. Moisture permeation and its effect on the insulating properties of the material 5. Settling and its effect on process piping and instrumentation 6. Optimum compactness or density in terms of weight per unit volume 7. Fire hazards from the insulating material in the presence of oxygen 8. Facility of filling or emptying the cold box to execute repairs within the box 9. Ease of storage or disposal of the insulating material once removed With the exception of liquid hydrogen and liquid helium plants, cryogenic process equipment can be satisfactorily insulated with either fibrous or porous material. As noted with a few exceptions, the majority of cryogenic plants have been insulated either with mineral wool or with expanded perlite. The latter material is getting the larger share in the market because of its ease of handling, and overall lower cost on an in situ basis.

7.2.3 MINERAL WOOL (ROCK WOOL, FIGURE 7.4A) When properly packed to a density of 250 kg=m3 the thermal conductivity of a granulated mineral wool is very low, around 0.035 W=m K. Mineral wool is all fibrous. Its sulfide content should be less than 0.30% by weight, and the combustible content should be less than 0.20% by weight. The material should be relatively dust-free and capable of tamped down from shipping density to an average of 250 kg=m3. It should be used for filling all void spaces in the cold boxes between the inside equipment and the panel structure. During tamping care must be taken to ensure no damage to equipment and especially instrumentation lines. Ignore vessels and piping when calculating the volume. If more than one section is involved for the cold box, then calculate the volume of each section and add these volumes together to arrive at the total volume (in cubic meters). The final volume is (m3) multiplied by 250 kg, which will give the total weight of the required rock wool. The rock wool should be well compacted in order to arrive at the smallest possible airspaces (cells) within the insulation. It has a high resistance to fracturing and dusting, and when properly compacted mineral wool does not settle. It also has the advantage that minor maintenance or inspection can be carried out by tunneling through insulation. There is no need to remove the material totally. If moisture enters, however, it will penetrate the mineral wool and freeze. The mass of insulation plus the accumulated ice has been known to settle because of its enormous weight and damage piping and instrumentation. In one incident the combined weight was heavy enough to settle at about 3 m, and in its downward path collapsed on a stainless steel process


λ 0.12

64 kg/m3 96 kg/m3 128 kg/m3

K cal/h m C

0.10 0.08 0.06 0.04 0.02

100 (a)

200 300 400 Average temperature (C)

FIGURE 7.4 (a) Thermal conductivity of rock wool. (b) Crushing of stainless steel pipe with diameter of 273 mm by weight of frozen mineral wool. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

pipe, which had a diameter of 273 mm and a wall thickness of 3.4 mm in terms of volumetric space (Figure 7.4b). Unfortunately, mineral wool is very difficult to handle manually and compact evenly within a cold box. Its use requires a considerable number of man-hours and strict supervision to avoid unevenness in compacting. It is also very difficult to remove from a cold box, especially if it has been hardened by internal icing. Recently, in a cold box that had been operating continuously for 5 y, jackhammers had to be used to remove when attempting to make repairs to leaking reversing exchangers. After repairs were finally made, the mineral wool was replaced with expanded perlite. Once removed, mineral wool is hard to store and


TABLE 7.4 Specifications for Granulated Mineral Wool (Cold Box and Ducts) Service Rating Material Thermal conductivity Sulfide content Combustibles General

Ordering

Cold box, valves, and duct insulation 77 to 311 K (1968C to 388C) Granulate mineral wool At 08C (273 K) and packed density of 240–250 kg=m3 should be a maximum at 0.03 kcal, m=h, m2, 8C Maximum content 0.30% by weight Maximum 0.10% weight Material should be relatively dust-free and capable of being tamped down from shipping density to an average of 240 kg=m3. It should be used for filling all void spaces in the cold boxes between the inside equipment and panel structure. During tamping care must be taken to ensure no damage to equipment and especially instrumentation (a) Calculate the volume of the cold box ignoring vessels and piping. Volume ¼ length (m)width (m)height (m). If the cold box is made of different sections of various heights or widths, calculate the volume of each section and add these volumes to find the total volume (b) To find the total quantity of wool (kg) needed, multiply the total volume by 250 and this will equal the total quantity of wool (kg)


dispose of because of its tremendous uncompacted bulk. Because of its bulk, new mineral wool is very costly to transport especially overseas (Table 7.4). An 8’’. diameter stainless pipe deformed by the sheer weight of frozen rock wool insulation is shown in Figure 7.5.

7.2.4 EXPANDED PERLITE8 Perlite is a hydrated volcanic rock which, when crushed and expanded by heat to remove the moisture, has been used successfully as insulation in cryogenic systems. It has a low thermal conductivity at 100 K and low density. It is nonhygroscopic, chemically inert, and free of organics. When packed to a density of 50 to 100 kg=m3, its coefficient of thermal conductivity is about 0.02 W=m K. Expanded perlite is very easy to handle and can be poured into the cold box directly from the expansion equipment. In a recent study for a 1000 t=d oxygen plant, it was estimated to require 30 d to pack the cold box with mineral wool, compared with 7 d using expanded perlite. Moreover, if the cold box is provided with proper drainage valving, the perlite can be drained directly into plastic sacks, stored, and reused. It is readily available in industrial countries and can be shipped overseas reasonably at low cost. It does not cake in use and is used in vacuum insulation. For purchasing, the particle size should be 2%–25% on a 16 mesh and 88%–95% on a 100 mesh, respectively (Figure 7.5). Expanded perlite, however, also has a few disadvantages, and some are quite serious. Two or three months after installation it will have settled 4%–5% in volume and must be replenished at the top of the cold box. If moisture seeps into the cold box, the perlite has a tendency to fracture and cause dusting. This, too, results in more settling that increases the coefficient of heat transfer. Because it is dusty and abrasive it cannot be used near moving parts. Thus, valve bodies within the casing must be isolated in carbon steel shells called valve boxes. These compartments, packed with mineral wool, facilitate maintenance from outside the casing (Figure 7.6a). Presently, this design is outmoded by the use of bayonet-type control valves which can be withdrawn, repaired, and reinserted without


Thermal conductivity density (kg/m3) 25

50

75

100

125

150

200

0.50 0.45

−0.00

0.40 −0.05

C)

−0.04

F 75

0.30

F(4 40

0.25

2C)

8 F(− −115

−0.03

) 29C

0.20

F(−1 200

−

0.15

) 84C

F(−1

−300

0.10

W/m K

Btu· in. /h·ft2·F

) 1C C) (24

F(4

105

0.35

−0.02

−0.01

0.05 0.00 0

1

2

3

4

5 6 7 8 Density (Ibs/ft3)

9

10

11

12

FIGURE 7.5 Perlite loose fill insulation. (From Technical Data Sheet, Perlite Institute, New York, 2005. With permission.)

contacting the insulation (see Chapter 10.4). When maintenance or inspection has to be carried out inside the cold box, moreover, the perlite must be drained down to the level of the proposed work. Presently, control valves are available with a body fully protected within the cold box, but its internals can be fully withdrawn externally for maintenance. Figure 7.6b shows workers saving expanded perlite in plastic sacks during a maintenance period. If any leak develops in the gaseous process piping or valves, in its path the escaping gas will entrain particles of expanded perlite which, because of their jagged edges and hardness, will cut through the thin aluminum sheets of the brazed aluminum heat exchangers. For this reason, many designers specify the use of mineral wool in the cold box containing heat exchangers. An alternative solution to this problem is to wrap the exchangers with glass wool batting (25–50 mm) to prevent contact with the expanded perlite. It is also prudent to use heavy gauge tubing for instrument and sample lines to avoid any damage from entrained expanded perlite. The use of stainless steel thin wall tubing for this purpose gives adequate protection because of its structural strength, and is lower in final cost than other materials, including copper. Care must also be taken in the geographic choice of the source of the perlitic ore. The best quality ore is found in the state of Arizona. Other ores, somewhat similar to those from Arizona, may be inferior in quality for insulating purposes (Table 7.5).

7.2.5 GLASS WOOL (FIBERGLASS) Glass wool or fiberglass is a very good insulation material as it has a low thermal conductivity, but is seldom used except in very small portable cryogenic systems, because of its high


FIGURE 7.6 (a) Typical compartment packed with mineral wool to facilitate maintenance of valve. (b) Saving expanded perlite during maintenance of cold box for reuse. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


TABLE 7.5 Specifications for Expanded Perlite Insulation Service Rating Material Density Thermal conductivity Water content Organic material content

Particle size Sampling

Cold box and duct insulation 77 to 311 K (1968C to 388C) Permalite Industrial Perlite-Grade No. 1 or Silbrico Corp., Grade 38 or equal Compacted density of about 64 kg=m3 At a mean temperature of 828C (191 K) and at a compacted density of 64 kg=m3 should be a maximum at 0.031 kcal, m=h, 8C, (0.036 W=m) Less than 1% by weight loss on heating at 1058C (378 K) for 2 h or more Substantially free of organic material and must not contain more than 0.1% by weight of matter soluble in carbon tetrachloride. Material should not spark or burn when in contact with embedded glowing platinum wire in an oxygen atmosphere 90% by weight should pass a 16 mesh sieve (890 mm), and at least 60% should be retained on a 100 mesh sieve (140 mm) For material expanded before shipment, a representative sample is to be obtained for each lot. Lot should be one truck trailer or carload or fraction thereof For field expanded material, a representative sample is to be obtained every 2 h In both cases, sample should be tested for particle size, compacted density, moisture or organic content to ensure that quality meets requirements. Copies of test certificates are to be sent to owner of cold box and duct work Material which fails to conform to this specification should be rejected. Rejection should be made on many quantities defined above


initial cost, and its difficulty in handling at the field. It is dimensionally stable, readily available, but somewhat difficult to install in a vacuum-insulated vessel.

7.2.6

GLASS BLOCKS (FOAM GLASS)

This is used principally for flat bottom cryogenic liquid storage tanks as a support for the weight of the cryogen. It is strong, readily available in industrial countries, nonflammable, and nonhygroscopic, but requires skilled labor for installation.

7.2.7

VERMICULITE

This material is a hydrous silicate related to the micaceous minerals. This insulation is used primarily in geographic areas where neither mineral wool nor expanded perlite is available or may be suspect in quality. Vermiculite has a low initial unit cost, but its coefficient of thermal conductivity is slightly higher than other insulating materials. A greater quantity, therefore, has to be provided. It is also highly water adsorbent. It has fairly good structural characteristics, but it also has a fairly high convective heat transfer.

7.2.8

SILICA AEROGEL

The insulating advantages of this material are many, even more than expanded perlite, but its price is very high (close to 10 times that of expanded perlite), and its insulating properties degenerate greatly after any contact with water.


7.2.9 MAGNESIUM CARBONATE Magnesium carbonate has been greatly used in the past because of its low cost, ease of availability, and fairly good insulating properties. It is also effective in vacuum insulation, but its properties degenerate after exposure to moisture. It cannot be regenerated. Presently, its use is minimal.

7.3 COLD BOX DESIGN FOR INSULATION8 To maintain the specified minimum heat gain for the insulating material selected, it is necessary to design the casing (cold box) with the following mechanical engineering prerequisites. The cold box should be completely sealed by welding with the exception of a few bolted panels to facilitate maintenance of internal equipment. Manhole covers should be tightly bolted with weatherproof gasketing. Apertures for extended bonnets of control valves should be covered with weatherproof neoprene, and attached to the casing with bolted rings. All efforts should be taken to seal the apertures against moisture. Moisture permeation is by far the major problem. The entrance of moisture into the insulating material will greatly increase the conductivity, especially if the moisture enters as vapor and condenses as it approaches the cold process equipment. The vapor pressure of water is lower at the colder part of the insulation than at the warmer part near the casing. This difference tends to push the moisture farther into the insulation, to an extent proportional to the difference in temperatures between the cryogenic equipment and the outside atmosphere. This difference can be as much as 300 K, an important consideration when designing the cold box. The yearly average relative humidity of the geographic area for the plant should also be taken into account. The cold box design should provide a constant sweep of a dry inert gas (preferably waste nitrogen) throughout the mass of insulation to maintain it dry at all times. The sweeping gas should be at a slightly positive pressure of 0.245 to 0.49 kPa (25 to 50 mm of water) and flow evenly throughout the entire cold box. Channeling should be avoided at all costs. A minimum of two pressure per vacuum valves should be located on the top of the casing, and set to actuate at a pressure of 0.49 to 0.735 kPa (50–75 mm of water). These valves should undergo preventive maintenance on a strict schedule. The foregoing procedure has been supplanted in some instances by the use of a flat metal plate lying horizontally on the top of the cold box surface. This plate is raised about 76 mm above the top surface by a pipe section about 10 mm thick by 76 mm leg completely welded to the top of the cold box surface. The plate lies unbolted across a standard manhole neoprene-gasketed flange located on the roof of the casing (Figure 7.7). The weight of the plate maintains a slight pressure within the casing, but being unbolted still allows a small outward seepage through the gasket. If there is an accidental rupture in the process equipment within the casing with a sudden large flow of gas, the unbolted flat plate at the top of the casing will act as a rupture disk. The main casing will not be damaged. The plate itself should be attached to the casing top by a strong metallic link chain so that it will not act as a projectile when blown off. If expanded perlite insulation is used, drainage valves should be provided, located near the bottom of the casing around 2 m above ground level, to facilitate drainage of the expanded perlite, if required. Two to four drainage valves should be provided depending on the size and configuration of the cold box. Separate silos connected with piping and vacuum pumps have also been used to store the expanded perlite, but this is very high in investment, therefore not very popular. It is used primarily in areas with strict environmental ordinances. In either case, the extracted perlite insulation is stored, saved, and reused. On the top of the cryogenic cold box, a combination of safety relief and vacuum break valves should be located in duplicate (Figure 7.8).


Note: Cover to be held down securely during shipping by straps tackwelded to cover material at shop, but removed after erection. Plan

1

2-1 1/8 dia. 1-10 5/8 dia.

3

1/8

1/8

1/4

(See note above)

Weld to cover 2 3

Weld to pipe

1

4 2-0 dia.

Cold box pressure safety device 24 (610 mm) diameter. Item

Quant

1

1

2

1

3

1

4

1

Material 24 Pipe 3/8 wall 3 leg 610 mm 10 mm wall 76 mm leg Cover plate 2-1 1/8 diameter 318/ thick 638 10 thick mm Gasket 22–5/8 ID 24 5/8 OD 1/8 thick 57 mm ID 627 OD 3 mm thick 14 D chain 1-6 (457 mm Leg 6 mm D)

Type Commerical A106 grade B Commercial Carbon steel neoprene shore hardness 5C Carbon steel

FIGURE 7.7 Cold box pressure safety device 2400 . (610 mm) diameter. (Courtesy of F.G. Kerry, Inc. With permission.)

If expanded perlite is used in the main casing, with mineral wool used for the casing containing brazed aluminum heat exchangers, and the two casings have a common interface, the casings should be separated by a common carbon steel plate to prevent the seepage of the expanded perlite into the mineral wool. If expanded perlite insulation is used in the casing housing brazed aluminum heat exchangers, it is prudent to wrap all the exchangers, ancillary piping and valves with glass wool batting, and at a minimum thickness of 50 mm. The wrapping should be tightly bound so there will not be any exposure of aluminum surfaces to perlite insulation. Another factor, though rarely considered, is the material used for the fabrication of the cryogenic equipment itself. If the material used has a high thermal conductivity, the heat gain will also be higher and are given in Table 7.6. There are definite advantages in the use of a metal with a low thermal conductivity. Table 7.6 shows a comparison of thermal conductivities (W=m K) at various temperatures for materials most commonly used


FIGURE 7.8 Relief and vacuum break valves in duplicate on top of flat bottom tank. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

in the fabrication of cryogenic systems. Values are approximate and should be used only for comparison.

7.3.1 SPECIAL REQUIREMENTS

FOR

LIQUID HYDROGEN PROCESSING PLANTS

Casings for liquid hydrogen plants are filled with porous insulations such as expanded perlite. Though they generally follow the basic rules for air separation plants, special requirements must be added for greater care and safety. The designer must keep in mind the difficulty of fabrication and erection of equipment in order to assure an absolutely leakproof plant. To minimize heat gain, process components should be arranged within the casing, that equipment operating at the lowest temperatures is innermost, surrounded by other

TABLE 7.6 Thermal Conductivity (W=m K) Cryogen Temperature (K) 77.30a 20.40b 4.23c

6073-T5 Aluminum

Beryllium and Copper

K Monel

304 Stainless

Low Carbon Steel

234.86 172.4 35.19

36.81 10.82 1.97

13.69 4.40 0.46

8.08 2.00 0.28

59.33 24.32 3.19

a

Liquid nitrogen. Liquid hydrogen. c Liquid helium. b



components operating at a higher temperature. Process piping and valves require special consideration to withstand the high temperature gradients imposed. Nitrogen is used as an inert gas purge before start-up, and during operation. At process temperatures below 80 K, the saturation temperature of nitrogen, problems may occur, however. Purging with hydrogen gas, therefore, may be considered, but only during operation. For nitrogen purging, an internal pressure of 0.49 kPa G (50 mm of water) is adequate to prevent seepage of air and moisture. A hydrogen-blanketed casing requires special design concepts to maintain it leak tight, including where the casing sits on the foundations. The designer should also take into account that hydrogen has a higher coefficient of thermal conductivity than nitrogen. Therefore, more insulation should be used. A greater distance should be allowed between the cold equipment and the casings as well as supports. The casing should be protected with the use of two relief valves at the roof of the casing plus a dead weight relief membrane (carbon steel plate), as previously explained for air separation plants. This latter relief component will act as a rupture disk, and should be set to actuate at a pressure higher than the primary relief valves. Normal operating procedure is to begin with a nitrogen purge. Then, as internal temperature approaches 90 K, start introducing a stream of hydrogen, and simultaneously cut back on the nitrogen stream until the total flow is almost completely hydrogen. This can be done by using the same purge lines, but having two separate feeds, one for hydrogen, and the other for nitrogen, each with both block valves and flow control valves on the separate feed pipes. Before shutdown operations, the reverse procedure can be followed. Reduce the hydrogen flow and increase the nitrogen flow until the total is completely nitrogen just before shutdown. Another cold box design is to have two concentric casings. The internal casing housing the liquid hydrogen equipment has a separate hydrogen purge, whereas the annular space between the two casings will be its own nitrogen purge. In practice, however, there is always a leakage between the two casings so that the end result is a mixture of hydrogen and nitrogen in both casings. But as already explained a single casing with a changeover of purge gases is more advantageous as well as less costly.

7.4

EXTERNALLY LOCATED PROCESS AND TRANSFER PIPING9,10

For this type of application the design engineer has a variety of choices. The materials used are the following: free expanded perlite, expanded perlite under vacuum, vermiculite, expanded foams such as polystyrene, foam glass, fiberglass, MLI (superinsulation), and any combination of the above. The design factors to be considered are as follows: maintaining the process design flow conditions, the prevention of vaporization in liquid transfer lines, the temperature differential between the cryogen and the outside ambient conditions, simplicity, ruggedness, safety, and of course cost.

7.4.1

SHORT LINES

For relatively short lines (20 m or less) and for cryogens such as liquid oxygen, liquid nitrogen, and liquid argon, it is possible to encase the individual lines in a steel duct and fill the duct with enough free expanded perlite to maintain a low heat gain. The duct can then be wrapped with special sheeting to prevent further heat gain. In fact, it is quite possible to design a common duct for two or even three lines, provided they are kept separated from one another by material with low heat conductivity. If this concept is adopted, it is advisable to insulate the liquid argon piping separately and at a distance from the liquid nitrogen line to keep the liquid argon from freezing should the latter’s flow be interrupted.


7.4.2 EXPANDED FOAMS Expanded foams are also used for liquid nitrogen and liquid argon lines, but not recommended for liquid oxygen unless a few layers of fiberglass are first placed to act as a safety barrier. Though expanded foams have the advantage of low cost and ease of installation, they also have the following disadvantages: poor safety, higher thermal conductivity, and thermal degradation over time. One of the most serious disadvantages is that they have a very much higher coefficient of thermal expansion and contraction compared with that of either carbon or stainless steel. During the cooldown period, cracks may occur in the expanded foam permitting air and moisture to seep in, degrading the insulating qualities of the material. If foam insulation is used, it should be protected with some form of outside cover. When using foam glass, the thickness should be determined based on the particular conditions for the vessel or piping system being insulated. The heat gain should be limited to a maximum of 25.25 W=m2.

7.4.3 FIBERGLASS INSULATION The use of fiberglass is satisfactory for cryogenic piping. This material is readily available in most places and relatively easy to use, but is very expensive. In some instances, a combination of fiberglass for the inner layers, with polystyrene foams for the outside has been found to be a worthwhile compromise. The final layer of insulation for straight lengths of piping should be applied with all joints sealed with a joint sealer, which shall cover the joint surface through the entire thickness of the insulation. The outer layer of fiberglass should be secured with thick stainless steel bands at the rate of two bands per section of equal spaced insulation. Over the completed insulation, installation applies a coat of weatherproof-coating paint at the rate of 8 to 12 L per 10 m2. It is important to keep in mind that when porous insulations are used, a vapor barrier should be applied to the external surface of the insulating material. Otherwise moisture will seep into the insulation degrading its quality. Porous insulations are not recommended for liquid hydrogen transfer lines. If the insulation is degraded, air seepage plus moisture condensation will occur. This will result in a fire and possibly an explosion, if the insulation is combustible.

7.4.4 PREFABRICATED VACUUM-INSULATED PIPING Prefabricated vacuum-insulated piping for long cryogenic lines is also available in the market. This piping involves two concentric pipes. The inner pipe carries the cryogen, and the annular space between the two pipes is filled with either evacuated perlite or MLI under vacuum. With evacuated expanded perlite, per foot heat gain may be as low as 0.07 to 1.58 W=m2 depending on the diameter of the process or transfer piping. The two concentric pipes are kept apart by carefully designed spacers (triangles, squares, or rollers) with minimum contact between the two pipes. These spacers are made up of material with a very low coefficient of conductivity. Whereas the use of vacuum-insulated piping is very easily field erected for erection purposes, it is expensive because it has to be shop fabricated under very close supervision and quality control (Figure 7.9).

7.4.5 MULTILAYER INSULATION If prefabricated piping is to be considered, then the best choice is an MLI better known as superinsulation, with which a heat gain of 0.025 to 0.0073 W=m2 can be easily attained. MLI is by far the best insulation for long liquid lines carrying cryogens, and especially for liquid


Vent Vent Phase separator

Vacuumjacketed valve Operating pressure regulator

Electronic keep cold and relief valve

Vent

Pipe system relief valve

Bayonet connection or field joint

Gas trap and extended stem valve

Tank hook-up

(a)

Primary relief valve

FIGURE 7.9 (a) Design of vacuum insulated cryogenic system. (Courtesy of Consolidated Design, Inc., 2005. With permission.) (b) Typical cryogenic liquid tank area in a large air separation unit. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


hydrogen and liquid helium. Such insulation involves 15 or 60 layers per inch of a material, which is made up of alternate layers of treated aluminum reflection shields, and specially made fine spacer material such as nylon net, ultrafine grains of glass, and fiberglass. The insulated piping is inserted into a concentric outer pipe and a very high vacuum, 1.33 106 mbar, is applied within the annular space. Here, as in evacuated powder-insulated lines, the piping and accessories must be shop fabricated under very close supervision, with highquality control. The insulating materials such as aluminized Mylar–Dexiglas, aluminized polyamide-tissueglass, and others are readily available in the market. Having a low heat gain (W=m2) than evacuated expanded perlite, MLI is also light in weight, resulting in a lower energy loss during the cooldown period. MLI, however, is not easy to apply on irregular shapes. Therefore, the finished product has a high cost, because fabrication is a high degree of automation and precision to control such variables as tension, overlap, density, compactness, as well as the number of layers. Fabricators, however, have been working hard to reduce the cost. One final word of caution on the use of a multilayerinsulated pipeline for oxygen, several incidents have occurred where there was a sudden break in the line and high pressure oxygen ignited the aluminum reflection shields. Always keep in mind that aluminum, especially in small thicknesses, can ignite very rapidly in the presence of high pressure oxygen given even a small input of energy.

7.4.6 CRYOGENIC LIQUID PIPING DESIGN Parameters for Liquid Piping Design In the design or cryogenic pipeline distribution systems overall cost is the dominant factor but does not necessarily lead to an optimal solution. Other important factors include extent of equipment maintenance, heat gain reduction of the selected insulation system, and cost value of the cryogen that is transferred. These factors vary case by case. For example, nature and temperature of the transferred cryogen, volume and rate of usage, continuous or intermittent operation, and operating efficiency in terms of minimal liquid loss will significantly influence final design. To maintain the cryogen in liquid phase throughout the distribution system up to the point of use, the engineer has a variety of insulations and designs to choose from. Insulated copper piping with outer plastic coating has very low initial cost but is very expensive in the long run because of high heat gain that results in excessive loss of valuable cryogen. At the other end of the scale there is MLI piping plus high vacuum. This offers extremely low heat gain and therefore minimal loss of cryogen. In between there is a variety of insulated equipment to satisfy a clients’ needs. There is much latitude, for example, in choice of external coating (foams, fibers, or powders with or without evacuation), or one may choose pure vacuum (Table 7.7). As noted previously, when a cryogen sits idle in a storage or distribution system, evaporation results in loss of product. Vapor within the system will seek a high point and can produce slugs of gas when flow is restarted. Such two-phase flow can cause operational instability, hinder or even arrest the entire process, a costly outcome. Slug flow can be prevented, however, by using a vent valve actuated mechanically or electronically by temperature differential. An alternative solution is to use a phase separator which receives the liquid phase from the main piping system, discharging it at lower pressure at the bottom and allowing the vapor phase to escape at top. The selected design should specify the vacuum-jacketed valves, bayoneted-connected vacuum piping, relief valves, and rupture disks in the piping system, and an operating pressure regulator for the pressurized liquid storage tank.


TABLE 7.7 Stainless Steel IPS Pipe Sizes in Metric Nominal Size (mm) 15 20 25 32 40 50 65 80 100 125 150 200 250 300 350 400 450 500 550 600

OD (mm)

Wall Thickness Schedule 5S

Schedule 10S

Schedule 40S

Schedule 80S

21.3 26.7 33.4 42.2 48.3 60.3 73.0 88.9 114.3 141.3 168.3 219.1 273.0 323.9 355.6 406.4 457.2 508.0 558.8 609.6

1.65 1.65 1.65 1.65 1.65 1.65 2.11 2.11 2.11 2.77 2.77 2.77 3.40 3.96 4.00 4.20 4.20 4.80 4.80 5.50

2.11 2.11 2.77 2.77 2.77 2.77 3.05 3.05 3.05 3.40 3.40 3.76 4.19 4.57 4.80 4.80 4.80 5.50 5.50 6.40

2.77 2.87 3.38 3.56 3.68 3.91 5.16 5.49 6.02 6.55 7.11 8.18 9.27 9.52 9.50 9.50 9.50 9.50 9.50 9.50

3.73 3.91 4.55 4.85 5.08 5.54 7.01 7.62 8.56 9.52 10.97 12.70 12.70 12.70 12.70 12.70 12.70 12.70 12.70 12.70


7.5 INSULATION FOR LIQUID STORAGE TANKS AND VESSELS In this area of application the choice of insulation depends on the volumetric capacity of the storage, the cost of the cryogen, the required boil-off rate during storage or during transportation, and finally on the overall costs. For the most part, the clear-cut choice is nonevacuated expanded perlite for storage tanks with a capacity of 500 t and over.

7.5.1 LARGE STORAGE TANKS (1000 T AND

OVER)

For very large tanks, the most economic design is the use of a double wall flat bottom tank with nonevacuated expanded perlite insulation. The design and quantity of insulation to be used should limit the total boil-off rate to less than 0.25 of 1% per day, depending on the tank capacity. A fairly substantial layer of foam glass blocks should be used at the bottom of the tank to reduce the heat gain to a practical minimum. A small low-pressure stream of dry nitrogen gas at a slight positive pressure should be kept flowing through the expanded perlite insulation to prevent any moisture from seeping into it. The external shell should include dual relief and vacuum break valves (Figure 7.9).

7.5.2

SMALLER STORAGE TANKS (500 TO 1000 T

OR

500 TO 1000 KL)

The most economic design for this capacity range is a double wall sphere with expanded perlite insulation between the inner and the outer walls. Depending on size, the boil-off rate


should range from 0.05 to 0.20 of 1% per day, or lower. A small low-pressure stream of dry nitrogen should be kept flowing through the expanded perlite to prevent any moisture from seeping in. The external wall should include dual relief and vacuum break valves.

7.5.3 STORAGE TANKS (50

TO

500 T

OR

50 TO 500 KL)

Such capacities for cryogens are designed in jumbo horizontal cylinders up to 300 t (oxygen), smaller interconnected cylinders, or in spheres. Preferred industrial design is the use of a jacketed cylinder with evacuated expanded perlite between the inner cold wall and the outer warm wall. Depending on size, the boil-off rate can be reduced to less than 0.05 of 1% per day. The vacuum applied should be held to 3 102 mbar, if possible. To maintain this vacuum, it may be necessary to provide a small vacuum pump, working intermittently or even continuously. It is also very important to provide vacuum-proof materials for valving. If nonmetallic diaphragms are used, they shall be inspected meticulously for porosity at the factory or at the site. Quality control is paramount. Improperly selected diaphragms are a major source of vacuum degradation. If a continuous flow of the cryogen is critical and if the location is far from any developed industrial area, it may be necessary to install a second vacuum pump fully connected and hot-wired.

7.5.4 STORAGE VESSELS (UP

TO

50 T

OR

50 KL)

Up to 50 t these storage tanks can be classified in the following manner: small transportable liquid cylinders, sometimes called cryocylinders with a normal net usable capacity of up to 240 L, but also available up to 800 L for exceptional users. These cylinders are double wall stainless steel using MLI and evacuated at 1.33 101 mbar. This has replaced expanded perlite almost completely for this size. Customer bulk stations, either vertical or horizontal, with a net capacity of up to 20–25 t also involve double wall stainless steel using MLI, evacuated down to a pressure of 1.33 101 mbar. This classification also includes transport trucks whose load is limited by the state highway regulations. Finally, large vertical customer bulk stations which may be either vertical or horizontal. With a construction similar to the previous class these have a specially designed insulation called XCELL (Table 7.8), which has a daily evaporation loss of almost one-third of that for evacuated expanded perlite. In addition, some designers employ a vapor shield to keep evaporation losses even lower. This is carried out by coiling the vent gas line around the inner vessel before taking it outside. In this manner, the outflowing vent gas intercepts and absorbs some of the heat gain, which would otherwise reach the stored liquid. The

TABLE 7.8 Relationship of Latent Heat of Vaporization to Cargo Weight and Capacity Liquid Oxygen Argon Nitrogen Hydrogen Helium-4

Boiling Point (K)

Density (lbs=G)

LHV (Btu=G)

LHV (kJ=kg)

LHV (kJ=G)

LHV (kJ=L)

90.18 87.28 77.35 20.27 4.22

9.52 11.71 6.75 0.59 1.03

875 826 574 116 9

212.5 161.6 198.3 445.6 20.73

920 858 607 120 10

243 227 160 32 3.6

Note: G ¼ US Gallon ¼ 3.785 L. Source: Courtesy of Chart Industries, 2005. With permission.


TABLE 7.9 Advantages and Disadvantages of Evacuated Perlite versus Superinsulation Evacuated Perlite Advantages Lower initial cost Easier to install Permits complex designs

Disadvantages Compacting and settling Adsorbs and retains water Needs more insulation space Vacuum filters required Harder to find leaks Evaluation of various insulations in relation to storage capacity

Superinsulation

Lower heat gain Less insulation space required Lighter in weight Easier leak detection Faster evacuation Higher initial cost Requires adsorbents to maintain vacuum Requires heat for evacuation

Source: Courtesy of Chart Industries, 2005. With permission.

effectiveness of this design depends on the ratio of the sensible heat of the escaping vapor to the latent heat of the stored liquid. The more common distribution of cryogens has been by semitrailers carrying up to 20 t or more (liquid oxygen), and in general, evacuated expanded perlite has been the preferred choice. In similar conditions, the chosen insulation should possess a light weight, good durability, a long vacuum, and a low thermal conductivity. The choice of perlite, however, has given problems from uneven roads resulting in vibrations, perlite breakup, packing, and sometimes to reinsulate the vessel completely. In some severe cases, the entire vessel had to be replaced. This problem has been eliminated, however, by the use of superinsulation, which though initially more costly, has more than paid off in terms of durability, thermal efficiency, and vacuum life. The norm for choosing an overall efficient insulation system for a container to be used in long range transportation must include: (a) a very low heat gain to minimize any possible increase in pressure; (b) a long vacuum life; and (c) a certain amount of structural flexibility to maintain its original design. The pros and cons for this choice are given in Table 7.9. Because of any uncertainty in scheduling that may be caused by any unforeseen delays, vessels to be transported by railcar must be specified with a very low heat gain performance to minimize venting important quantities of valuable products. Generally, railcars have been insulated by evacuated expanded perlite, but the application of superinsulation has been increasing especially for the transportation of liquid hydrogen because of its low heat of vaporization (Table 7.10).

7.6 VACUUM PUMPING SYSTEMS 7.6.1 GENERAL OVERVIEW When using liquefaction for the separation, purification, and liquid storage of industrial gases, the importance of vacuum systems cannot be swept aside. Vacuum systems are widely employed for increasing the efficiency of insulation used for low-temperature vessels; evacuating cold traps normally refrigerated by liquid nitrogen used to filter out impurities


TABLE 7.10 Normal Evaporation Rate (NER) per Day US Gals 500 900 1,500 3,000 6,000 9,000 11,000 13,000

Approx. (L)

Approx. (t)

Evacuated Perlilte

Evacuated XCELL

Evacuated Superinsulation

2,000 3,400 5,000 11,600 23,250 35,000 42,600 50,250

2.15 3.90 6.45 13.00 26.00 39.00 48.00 56.00

0.55 0.40 0.40 0.50 0.30 0.26 0.25 0.23

0.45 0.35 0.25 0.14 0.12 0.08 0.08 0.08

— — — — 0.22 0.10 0.10 0.10

Source: Courtesy of Chart Industries, 2005. With permission.

such as water vapor, carbon dioxide, and other frozen contaminants; cryopumping in large environmental simulation chambers; and He3–He4 dilution refrigeration. Vacuum Technology10 ‘‘We are grateful to Leybold Vacuum, Inc., USA for its kind and generous permission to use material and figures from its Product and Reference Book of 2003 to 2004 as our own experience has been limited.’’ Typical pressure ranges of industrial vacuum processes are given in Table 7.11. See Appendix for relationship of vacuum pressures in different measurements. In this specialized field, it will be useful to become familiar with the following terms: p pt

¼ Pressure measured in millibar, or mbar ¼ 1 103 bar ¼ 0.7547 Torr. The latter is no longer used. ¼ Total pressure, or the sum of all partial pressures of gases or vapors. TABLE 7.11 Typical Pressure Ranges of Industrial Vacuum Processes Annealing of metals Melting of metals Degassing of molten metals Degassing of steel Electron beam melting Electron beam welding Molecular distillation Degassing of liquids Sublimation Drying of insulating papers Freeze drying of mass materials Production of incandescent lamps Production of electron tubes Production of gas-discharge tubes

103–104 103–105 100–103 101–101 1.5 10 –1010 103–105 100–104 101–103 100–104 102–103 100.5–101 101–104 104–107 102–106

Source: From Leybold Product & Reference Book, 2003 to 2004. Leybold Vacuum, Inc., USA. With permission.


pi ps pn pend pamb pe pw n

¼ ¼ ¼ ¼ ¼ ¼ ¼ ¼

Partial pressure of a certain gas or vapor exerting its own pressure alone. Saturation vapor pressure and is a function of temperature of any substance. Normal pressure at 1 atm, or 1013.25 mbar. Lowest pressure that can be attained in a vacuum vessel. (absolute) atmospheric pressure. Over or excess pressure. pe ¼ pabspamb. Negative in a vacuum operation. Working pressure. Particle number density or number of gas molecules (cm3), dependent on pressure and temperature. p ¼ nkT where n is the particle number density and k is the Boltzmann’s constant. r ¼ Gas density (kg=m3 or g=cm3) r ¼ nmT

(7:27)

where mT is the particle mass. Ideal gas law states that the relationship between mass mT of a gas molecule and molar mass M of this gas is M ¼ NAmT, and Avogadro’s constant NA indicates the number of gas particles contained per mole. NA is also the proportionality factor between the gas constant (R) and Boltzmann’s constant (k) or, R ¼ NAk. Therefore, one can derive that p ¼ r(RT )=M In practice, one considers a certain enclosed volume V in which the gas present is at a pressure p, and if m is the mass of the gas, then r ¼ m=V, therefore, it follows from the ideal gas law that pV ¼ m=MRT ¼ vRT, and for 1 mol or m=M ¼ 1 the following is the equation: pV ¼ RT

(7:28)

With vacuum conditions, one must keep in mind that a specific mass of a light gas (helium) exerts a greater pressure than an equal mass of a heavier gas (nitrogen), because assuming an equal temperature, there are more particles of the lighter gas (large n, small m) than those of the heavier gas. The principal task of vacuum technology is to reduce the particle number density (n) inside a given volume (V). However, as shown in Equation 7.28, this requires not only lowering the pressure p, but also lowering the temperature T. v ¼ Volume (L, m3, cm3). qv ¼ Volumetric flow (L=s, m3=h, cm3=s). The flow indicates the gas volume flowing through a piping element per unit of time at the prevailing pressure and temperature and at a specific moment. S ¼ Pumping speed (L=s, m3=h, cm3=s). It is the volumetric flow through the pumps intake port. S ¼ dV=dt

(7:29)

pV value ¼ Quantity of gas in mbar L. Calculated from pV ¼ (m=M) RT or m ¼ (pVM)=RT


(7:30)

Outgassing: This term refers to the liberation of gases and vapors from the walls of a vacuum chamber or other components on the inside of a vacuum system. The outgassing rate is in mbar L=s (mbar) ¼ 1 s1

(7:31)

For this measurement one must calculate the size of the interior surface area, its material, and the surface characteristics. When referenced to the surface area the rate is expressed as (mbar L=s=cm2) rate in mbar ¼ s1 cm2

(7:32)

Example of using a vacuum plus insulation For storage of cryogens, when using expanded perlite under vacuum at a pressure of 1.33 103 mbar, it is possible to achieve an apparent thermal conductivity (heat gain) of ka ¼ 1–2 mW=m K. With MLI (MLI of 31 layers=cm of Al foil þ nylon net) at a vacuum of 1.33 106 mbar, the apparent thermal conductivity (heat gain) is kt ¼ 3.5 102 mW=m K. With pure vacuum and the same vacuum pressure as with MLI, the apparent thermal conductivity is kt ¼ 5 mW=m K, which means that MLI with vacuum plus reflective shields is about 20 times more efficient than pure vacuum. Cold traps are employed by process designers to upgrade the purity of main product. These involve the use of adsorbent material such as silica gel, charcoal, Cr2O3 and molecular sieve 13X operating at temperatures of liquid nitrogen, 77.3 K. These adsorb water vapor, carbon dioxide, carbon monoxide, and even oxygen and nitrogen for the ultrapurification of hydrogen or helium. The degree of vacuum desired may be classified as follows: . . . .

Rough vacuum (atmospheric—1 mbar) Medium vacuum (1103 mbar) High vacuum (103107 mbar) Ultrahigh vacuum (1071011 mbar and higher)

7.6.2 VACUUM PUMPS To achieve a vacuum there is a variety of vacuum pumps operating at different principles. They range from the reciprocating positive-displacement type, such as the diaphragm pump and its rotary cousins, to the modern entrapment types such as the adsorption pump and cryopump. A few examples are described below. 7.6.2.1 Roots Vacuum Pump This pump has been well established in vacuum technology for a long time. It is a rotary positive-displacement pump where two symmetrically shaped impellers rotate inside the casing are in close proximity. This operating principle allows the assembly of units having very high pumping speeds resulting in a vacuum pressure of 105 mbar, which makes the additional use of diffusion or turbomolecular pumps unnecessary in many cases. Because of the noncontact rotation of their impellers these pumps are able to operate at high speeds, in excess of 1000 m3=h. Through an integrated pressure equalization line it is possible to use the pumping speed of the roots vacuum pump at high pressures and large quantities of gas at an early stage. Canned motors are hermetically sealed preventing contact with the atmosphere,


1 2

2

3

5 4

(1) Intake flange (2) Rotors (3) Chamber

(4) Exhaust flange (5) Casing

FIGURE 7.10 Schematic cross section of a roots pump. (Courtesy of Leybold Vacuum Products, Inc. With permission.)

so a service life of over 20,000 h without maintenance is quite common. Finally, if the medium being pumped is not contaminated with lubricants or sealants, exhaust filters or separators are not necessary (Figure 7.10). 7.6.2.2

Rotary Vacuum Pump

The annular space between the inner vessel containing the cryogen and the outer casing can be evacuated employing a vacuum pump of the rotary piston type. These pumps have two or even three pistons rotating on a common shaft providing an internal compression of a 2:1, or even a 4:1. They achieve a vacuum of 1.33 104 mbar, which is more than enough for evacuated expanded perlite, and satisfactory for most MLIs. While producing such low pressures may be possible in shop conditions with proper supervision, they may not be achieved in field operations because of careless welding of small instrument tubing connected to the outer shell, poor connections between the outer shell and the analytical instrument, outgassing from the annular surfaces of the vacuum space or poor calibration of the instrument. In one specific case concerning a large vertical bulk vessel containing high purity liquid nitrogen with a specified internal vacuum pressure of less than 5 mbar, the reading fluctuated frequently up to 30 mbar, and the vacuum pump had to be adjusted several times per day. In another case where three horizontal vessels were interconnected, it was necessary to install two vacuum pumps. One was on operation continuously, and the other was a 100% standby. Pumping speeds of 1000 m3=h can be attained. In this type lubrication is necessary. The oil must be of a high quality, its vapor pressure must be low at high temperatures, its water content and water uptake must be minimal. The use of synthetic oils may be considered if they meet the rigid standards of the pump manufacturer (Figure 7.11 through Figure 7.13).


Valve stop

Leaf spring of the valve

(I) High vacuum stage (II) Second forevacuum stage

FIGURE 7.11 Schematic cross section of a two-stage rotary vane pump. (Courtesy of Leybold Vacuum Products, Inc. With permission.)

7.6.2.3 Turbomolecular Pumps This pump is a high speed turbine with blades which when operating downstream of a suitable backing pump can generate an ultrahigh vacuum of 1010 mbar and lower. By the transfer of momentum from the rapidly rotating rotor blades to the gas molecules, the initially nondirected thermal motion of the latter is changed to a directed motion. At pressures below 103 mbar the mean free path (l) of the molecules is larger than the spacing between the rotor and the stator blades; therefore, the molecules collide with the rotor blades increasing the efficiency of pumping. At higher pressures over 101, the effect of the rotor is impaired by the frequent collision between the molecules, thus lowering the efficiency. The rotor speeds range from 36,000 rpm for the larger unit of 20 cm diameter, and to 72,000 rpm for the smaller unit of 6 cm diameter. Pressure (Torr) 104

103

102

101

100

101

750 500

DK 200 102

100 50

101 8

10 5

6 4

Pumping speed (cfm)

Pumping speed (m3/h)

103

2

100

1 2 4 68

104

103

102

101

100

101

102

103

Pressure (mbar) Without gas ballast With gas ballast

FIGURE 7.12 Pumping speed characteristic of the DK rotary piston vacuum pump (60 Hz curves at the end of the section). (Courtesy of Leybold Vacuum Products, Inc. With permission.)


750

103

101

101

100

100

101

101 DK 200

102

102 8 6 4 2

103

Pressure (Torr)

Pressure (mbar)

102

103 0

1

2

3

4

5

6

7

8

9

10

Time (min) Without gas ballast With gas ballast

FIGURE 7.13 Pumpdown characteristics of a 1000 L vessel for the DK type rotary piston vacuum pump (60 Hz curves at the end of the section). (Courtesy of Leybold Vacuum Products, Inc. With permission.)

7.6.2.4

Cryopumps (Figure 7.14 through Figure 7.16)11,12,13

Cryopumping is another simpler way of affecting a vacuum. The annular space to be evacuated is first sealed hermitically, and then filled with a gas, preferably carbon dioxide. As soon as the inner vessel is filled with a cryogen having a much lower boiling point, the gas

FIGURE 7.14 An assortment of cryopumps for vacuum generation. (Courtesy of Leybold Vacuum Products, Inc. With permission.)


6 1

7

2

5

3

4

10

8

11

9

(1) High vacuum flange (2) Pump body (3) Foreline flange (4) Safety valve with flange connection for connection of an exhaust line (5) Thermal radiation shield

(6) Baffle (7) Second stage of the cold head (8) Cryopanels (9) First stage of the cold head (10) Helium gas connections (11) Cold head motor with housing and electrical connections

FIGURE 7.15 VAC refrigerator cryopump. (Courtesy of Leybold Vacuum Products, Inc. With permission.)

Pressure (Torr) 107 106 105 104 103 102 101 100 101 102750103

Adsorbed quantity of gas per quantity of adsorbent (mba·L/g)

103

N2 (195C)

102 101

N2 (20C)

100 101 102 103

Ne (195C)

104 105

He (195C)

106 7 6 5 4 3 2 1 10 10 10 10 10 10 10 100 101 102 103 Pressure (mbar)

FIGURE 7.16 Adsorption isotherms of zeolite 13X for nitrogen at –1958C and 208C, as well as for helium and neon at –1958C. (Courtesy of Leybold Vacuum Products, Inc. With permission.)


TABLE 7.12 Measured Values of the Adhesion Coefficient f Gas Pumped Ammonia Argon CO2 CO2 CO2 CO2 CO2 CO2 CO2 Nitrogen Nitrogen Nitrogen Water

Cryosurface Temperature (K)

Warm Surface Temperature (K)

Adhesion Coefficient f

References

77.3 20 77.3 20 77.3 20.2 77.3 21.0 77.3 20 20 20 77.3

300 77.3 300 195 195 300 300 400 400 — 85 77.3 300

0.45 0.79 0.62 0.90 0.80 0.65 0.63 0.44 0.49 0.55 0.73 0.87 0.92

Brown et al. (1965) Dawson et al. (1964) Brown et al. (1965) Dawson et al. (1964) Dawson et al. (1964) Dawson et al. (1964) Dawson et al. (1964) Dawson et al. (1964) Dawson et al. (1964) Moore (1962) Hurlbut and Mansfield Dawson et al. (1964) Brown et al. (1965)

molecules within the annular space freeze out, attaching themselves on the inner cold wall surface, thus creating a vacuum. With an inner storage vessel containing either liquid hydrogen or liquid helium a vacuum pressure of 2.7 105 mbar has been obtained. This alternative to mechanical vacuum pumps has a fair following with cryogenic engineers who deal with all types of gases, and for companies that fabricate cryogenic vacuum vessels. Engineers who specify the level of vacuum to be expected in these types of vessels must be familiar with the range of gases used for cryopumping, the cryogenic surface temperature necessary for adhesion to succeed, and the measured values of the attachment coefficient f (Table 7.12). Cryopumping can be carried out either by bath cryostats, continuous flow cryopumps, or refrigerator cryopumps. The latter are used almost exclusively today and operate somewhat similar to the common household refrigerator (cold on demand). Its operation is based on the Gifford-McMahon process developed by A.D. Little using helium as the refrigerant. It involves a compressor connected to the cryopump by a flexible pressure line to be free from any vibration. The cryopump unit consists of a pump casing and the cold head inside. The refrigerant is helium circulated in a closed cycle by the compressor. Inside the cold head there is a cylinder, which is divided into two working spaces V1 and V2 by a displacer. During operation the right space V1 remains warm, whereas the left space V2 remains cold. The Leybold Vacuum Co., uses a two-stage cold head.10 At a displacer frequency (f) the refrigerating power (W) of the refrigerator is W (J=s) ¼ (V2, max V2, min )(pH pN )f

(7:33)

where pH is the inlet pressure and pN is the outlet pressure. The bonding process is carried out in three steps: the mixture of gases and vapors meets the baffle at 80 K, where H2O and CO2 are condensed. The remaining gases penetrate and impinge in the outside of the cryopanel of the second stage at 10 K where N2, O2, and Ar will condense. Any remaining gases such as H2, He, and Ne cannot be pumped by the cryopanels, and after several impacts with the thermal radiation shield to the inside of these


panels which are coated with an adsorbent (cryosorption panels) they are finally bonded by cryosorption. All the aforementioned gases may be divided into three groups depending on at which temperatures within the cryopump their partial pressure ps drops below 109 mbar. First group ps less than 109 mbar at T ¼ 77 K (LN2): H2O, CO2. Second group ps less than 109 mbar at T ¼ 20 K: N2, O2, Ar. Third group ps less than 109 mbar at T less than 4.2 K: H2, He, Ne. A difference is made between the different bonding mechanisms and are as follows: Cryocondensation is the physical and reversible bonding of gas molecules through van der Waals forces on sufficiently cold surfaces of the same material. The bond energy is equal to the energy of vaporization of the solid gas bonded to the surface and decreases as the thickness of the condensate increases as does the vapor pressure. Cryosorption is the physical and reversible bonding of gas molecules through van der Waals forces on sufficiently cold surfaces of other materials. The bond energy is equal to the heat of adsorption which is greater than the heat of vaporization. As soon as a monolayer has been formed, the following molecules impinge on a surface of the same kind (sorbent) and the process transforms into cryocondensation. The higher bond energy for cryocondensation prevents further growth of the condensate layer thereby restricting capacity for the adsorbed gases. However, the adsorbents used such as activated charcoal, silica gel, alumina gel, and molecular sieve have a porous structure with very large surface areas, say about 106 m2=kg. Cryotrapping is the inclusion of a low boiling point gas difficult to pump as hydrogen, in the matrix of a gas having a higher boiling point and which can be pumped easily such as Ar, CH4, or CO2. At the same temperature, however, the condensate mixture has a saturation vapor pressure, which is by several orders of magnitude lower than the pure condensate of the gas with the lower boiling point. Partial regeneration of cryopumps The service life of a cryopump depends on the capacity limit for the adsorption of the low boiling gases such as nitrogen, argon, and hydrogen, which are pumped by the second stage. This means that only the second stage needs to be regenerated. To carry out this procedure, however, one must maintain the first stage at or below 140 K, or the partial pressure of water vapor will become high enough allowing the water molecules to contaminate the adsorbent on the second stage. In 1992, Leybold Vacuum developed a fast regeneration using a microprocessor controlled process reducing the regeneration time to 40 min, instead of the usual 6 h. Nevertheless, one may have to carry out one complete regeneration after 24 partial regenerations. 7.6.2.5 Adsorption Pumps (Figure 7.17) Adsorption pumps operate on the principle of physical adsorption of gas molecules on molecular sieves (zeolites) or other materials as explained in detail in Chapter 5. Using a molecular sieve 13X, it is possible to adsorb gases such as nitrogen, carbon dioxide, water vapor, and hydrocarbons, but not helium or neon because of the latter’s smaller diameter compared with 13X. The capacity for adsorption, however, depends not only on temperature


2 1

3

4

5

(1) Inlet port (2) Degassing port (3) Support (4) Pump body

(5) Thermal conducting vanes (6) Adsorption material (e.g., zeolite)

6

FIGURE 7.17 Cross section of an adsorption pump. (Courtesy of Leybold Vacuum Products, Inc. With permission.)

but also on pressure. If vacuum pressures below 103 mbar are desired, one must avoid the presence of neon or helium. Furthermore, after the pumping operation is completed the pump must be regenerated to room temperature or even higher to 473 K if the gas has contained a large quantity of water vapor. For pumping very large vessels one must use several adsorption pumps in series or in parallel. 7.6.2.6

Getters

Employing getters is a low-cost procedure to enhance and maintain a high vacuum. It involves the use of material that has a strong affinity for gases at very low pressures. Charcoal, silica gel, molecular sieves, barium, zirconium, and titanium have been employed for this purpose. These products can remove nitrogen, oxygen, and hydrogen by solid solution. Outgassing is a fact of life in the industry and the use of getters along with the insulating material is small cost insurance for maintaining a high level of vacuum for a long period of time. As metallurgists know, when metals or alloys are poured and rolled, there is a fair quantity of air and other gases trapped on the surface of the finished product. When a high vacuum is applied the trapped gases have a tendency to escape. This outgassing spoils not only the applied vacuum, but also any analytical instrument that may require a vacuum in its operation. As shown in Chapter 5, adsorption can also be used to enhance a high vacuum. Molecular sieve 13X has a large surface area per mass, approximately 1000 m2=g, therefore its capability ˚ well within the size of water vapor, oil of adsorption is considerable. Its pore diameter is 13 A ˚ , i.e., nitrogen, oxygen, and carbon dioxide. Adsorpvapor, and large gas molecules of 10 A tion of gases at surfaces is dependent not only on temperature, but more importantly on the pressure above the adsorption surface. Adsorption pumps are connected by a valve to the vessel to be evacuated. By immersing the body of the pump in liquid nitrogen the adsorption effect is enhanced. Because of different adsorption properties, the pumping speed and ultimate pressure of an adsorption pump are different for various gas molecules. Optimum values are obtained for nitrogen, carbon dioxide, water vapor, and hydrocarbon vapors. Light noble gases, however, are not used because of their small molecular diameter. As the


adsorption of particles increases the pumping speed decreases, which means that the speed of the adsorption pump is not constant, but dependent on the quantity of gas already adsorbed. Vacuum pressures of less than 1.33 102 mbar can be obtained. 7.6.2.7 Small Laboratory Pumps Water injector vacuum pumps soon gave way to diaphragm pumps for environmental reasons and greater convenience. During compression the pumping chamber leads to the exhaust chamber, and the diaphragm provides a hermetic seal between the gear chamber and the pumping chamber, so in essence it becomes a dry compression pump. The diaphragm and valves are the only components and are the only contact with the medium. This pump is very well suited for a chemistry laboratory (Figure 7.18 and Figure 7.19).

7.6.3 PERIODIC PURGING

AND

DERIMING

During a planned plant shutdown the insulation in a vessel and especially in transfer piping should be examined for insulating efficiency. This can be carried out by purging with warm nitrogen, but not air. Any condensates or solids from atmospheric air will reduce the

1

2 5 4

3

(1) Steam inlet (2) Jet nozzle (3) Diffuser (4) Mixing region (5) Connection to the vacuum chamber

FIGURE 7.18 Schematic representation of the operation of a steam ejector pump. (Courtesy of Leybold Vacuum Products, Inc. With permission.)


1 2 3 4 5 6 7 8

(1) Casing lid (2) Valves (3) Lid (4) Diaphragm disk

(5) Diaphragm (6) Diaphragm support disk (7) Connecting rod (8) Eccentric disk

FIGURE 7.19 Schematic on the design of a diaphragm pump stage (Vacuubrand). (Courtesy of Leybold Vacuum Products, Inc. With permission.)

effectiveness of the insulation and even damage it if the purging is done at an abnormally elevated pressure during the warming phase. For liquid hydrogen storage, it is wise to use helium for purging as it will not condense. A nitrogen purge should be used for liquid oxygen storage.

7.6.4 7.6.4.1

ANCILLARY EQUIPMENT Valves

High quality valves, so necessary for vacuum systems, are designed to meet exacting standards so that gas molecules adhering to the surface of the valve shaft are not transferred from the outer atmosphere into the vacuum during operation. These valves are equipped with metal bellows isolating the valve shaft from the atmosphere or totally encased. The valves can be sealed with oil or grease for strict specifications. Their leakage rate is approximately 109 mbar L=s. Aside from these isolation valves, there are others such as motor-driven variable leak valves suitable for remote control, and when these are connected to a pressure gauge the process pressure can be set and maintained. Safety valves are also available to quickly and automatically cutoff diffusion pumps or vacuum systems incase of a power failure. 7.6.4.2

Vacuum Measurement

Present-day vacuum technology permits measurement in a range from 1013 to 1012 mbar, depending on the order of accuracy required. As it is impossible to design a single vacuum gauge to cover the entire vacuum range, the market offers a variety of vacuum gauges each with a desired accuracy for a specific range of vacuum. It is impossible to cover the variety of measuring instruments available in this chapter, and it is recommended that the reader contacts a company specializing in vacuum technology for more details.


REFERENCES 1. Barron, R. 1966. Cryogenic systems (McGraw-Hill Series in Mechanical Engineering), pp. 386–391. 2. Nusselt, W., and Z. Bayer. 1913. Revision-ver. November 13–14. 3. Perry and Green. 1985. Chemical engineers handbook, 6th ed., Psychometry, Evaporative Cooling, Refrigeration and Cryogenic Processes, pp. 12-56–12-58. 4. Stewart, R.B., and V.J. Johnson. 1961. A compendium of the properties of materials at low temperatures, Phase II, Part IV, WADD Technical Report, pp. 60–56. 5. Timmerhaus, K.D., and T. Flynn. 1989. Cryogenic process engineering (The International Cryogenics Monograph Series), pp. 382–392. New York: Plenum Press. 6. Caren, R.P., and G.R. Cunnegton. 1967. Properties and applications of multilayer insulation systems, 61st National Meeting, A.I.Ch.E., pp. 19–23. 7. Black, I.A., and P.E. Glaser, 1966. Advances in Cryogenic Engineering, Vol. II, Plenum Press, New York, p. 26. 8. Kropschot, R.H., and R.W. Burgess. 1963. Perlite for cryogenic insulation (Advances in Cryogenic Engineering), vol. 8. New York: Plenum Press. 9. Lutgen, H.M. 1984. Cryogenic insulations for the distribution of liquefied gases, IOMA Broadcaster, May–June, pp. 13–16. 10. Umrath, W. 2003–2004. Fundamentals of Vacuum Technology, Leybold Vacuum, Bonner Strabe, Germany. pp. D007-182. 11. Brown R.F., et al. 1965. Capture coefficients of gases at 77 K. In Advances in Cryogenic Engineering, vol. 10, 283–291. New York: Plenum Press. 12. Dawson, J.P., et al. 1964. Temperature effects on the capture coefficients of carbon dioxide, nitrogen, and argon. In Advances in Cryogenic Engineering, vol. 9, pp. 443–450. New York: Plenum Press. 13. Hurlbut, F.C., and R.J. Mansfield. 1963. Calculated and observed pumping speeds of a shielded cryogenic pumping surface. In Advances in Cryogenic Engineering, Vol. 8, pp. 46–56. New York: Plenum Press.

FOR FURTHER STUDY AND REVIEW Tien, C.L., and A.J. Stretton. 1987. Heat and mass transfer in refrigeration and cryogenics, p. 3. New York: Hemisphere Publishing. Steward, W.G., 1986. Cryogenics 26:97. DiPirro, M., and M. Castles. 1986. Cryogenics 26:84.



8

Special Gases

Excluding the generally well-known atmospheric gases such as outlined previously, there is a group of other gases, which are constantly encountered in the study of air separation. These gases are hydrogen, carbon monoxide, ultrahigh-purity nitrogen, ultrahigh-purity oxygen, nitrous oxide, carbon dioxide, ozone, and methane. It is important, therefore, to review these gases, as they have become important components of the industrial gas industry.

8.1 HYDROGEN Hydrogen is the most abundant element in the universe. The sun is nothing more than a ball of hydrogen undergoing a nuclear fusion reaction to produce helium. On the Earth’s crust the hydrogen content was about 0.14 wt%. Wherever water exists in liquid or vapor form it is present, as it is in all organic compounds and living organisms. Hydrogen was first studied by Henry Cavendish of England in 1766. In 1781, he discovered that when hydrogen was burned, it produced water. Lavoisier of France who simultaneously had been studying this substance’s physical properties, named the gas hydrogene, or hydrogen, which in Greek means producer of water. Hydrogen occurs as a diatomic molecule. Its atomic weight is 1.008, and the physical properties of hydrogen are given in Table 8.1. Molecular hydrogen has weak intermolecular forces, and its temperature rises when the gas expands at normal room temperatures. This may be explained by the fact that at these temperatures the repulsive forces are stronger than the attractive forces between molecules. At a temperature of 204.55 K and lower, however, the attractive forces become stronger, and there is a pronounced cooling. In fact, at the temperature of liquid nitrogen (77.3 K), hydrogen cools by expansion and liquefies. It has two stable isotopes namely the normal hydrogen and the so-called heavy hydrogen, or deuterium, which has a neutron, as well as a proton in its nucleus. The neutron has the same weight as the proton, but has no electric charge. The ratio of deuterium to normal hydrogen protium is about 1 to 3200 or less. Hydrogen also has a third radioactive isotope, called tritium, but this is very scarce. It has a half-life of 12 years. Molecular hydrogen has two additional forms, namely ortho- and para-hydrogen. In orthohydrogen the two protons of the molecule spin in the same clockwise direction, whereas in para-hydrogen the two protons of the same molecule spin in opposite directions. At normal temperatures, around 300 K, hydrogen is formed (75% ortho-hydrogen and 25% parahydrogen). As the temperature decreases, however, the ratio of para-hydrogen increases to the point where at liquid hydrogen level, 20.27 K, the percentage of para-hydrogen reaches almost 100. This conversion from ortho to para is exothermic and takes place without any outside influence. In the liquefaction of hydrogen, therefore, it is important to convert the ortho to para artificially whereas the hydrogen is in gaseous form before the liquefaction process proceeds, so the heat generated during the conversion can be extracted by external cooling.


TABLE 8.1 Physical Properties of Hydrogen Molecular weight Normal boiling point Vapor pressure Latent heat of vaporization Molar heat capacity of gas at constant pressure in the range of 0–1 bar and at 298 K Thermal conductivity of gas Gaseous real density Liquid density Critical temperature Critical pressure Flammability range

2.02 kg=kmol 20.268 K at 101.325 kPa 70 mbar at 13.9 K 0.92 MJ=kmol at boiling point and 101.321 kPa 28.97 kJ=(kmol K) 0.171 W=(m K) at 273 K 0.089 kg=m3 at 273.15 K and 101.325 kPa 0.071 kg=L at 20.268 K and 1.013 bar 32.98 K 12.93 bar 4.1%–95% in air, 5%–95% in oxygen


8.1.1

SOURCES

OF

HYDROGEN

It is found as water in rivers, oceans, ice packs, and in the atmosphere as clouds. It is also found in all organic compounds in animal, vegetable, petroleum, and carbon products.

8.1.2

RECOVERY

OF

HYDROGEN

Before considering the recovery of hydrogen and the process to be used, one must take into account the purity of the hydrogen specified by the user. For example, certain industries demand a high purity, namely 99.95% and even 99.9999%. These industries include: . . . . . .

Hydrogenation of oil for the food industry Bright annealing of tin plate in the steel industry Production of float glass in the glass industry Manufacture of transistors in the electronic industry Hydrogenation of propylene and polymerization of ethylene for the chemical industry Production of liquid hydrogen as a propellant for space flights or to be transported to a point of use for other applications

For applications requiring very low contents of impurities, it may be wise to consider purification by adsorption in place of cryogenic distillation, or a combination of both, with adsorption being used for final purification. Adsorption has been proven successful for the recovery of hydrogen from a wide variety of gases, at pressures up to 40 bar. It has been used with partial oxidation (POX), gas reforming, ammonia purge gas, cryogenic process gas, and other applications. Hydrogen adsorption units have been designed with capacities up to 47,000 Nm3=h at operating pressure of 23 bar. In such units, however, more than two vessels have to be used depending on the contaminants to be adsorbed. Depressurization may also have to involve the use of a slight vacuum. Because the total switching cycle can be very short, demands on switching valves may be severe and their control must be extremely accurate. The use of an electronic program timer and a comprehensive preventive maintenance program is a must.


TABLE 8.2 Choice of Adsorbents for Purification of Hydrogen Molecular sieve Activated carbon Carbon prefilter Activated alumina

Contaminant removal Contaminant removal Contaminant removal Contaminant removal

N2, CO Ar, O2, CH4, CO2 H2S, BTX, CnHm H2O, NH3


In both pressure swing adsorption (PSA) and vacuum pressure swing adsorption (VPSA), the removal of impurities, which may vary in chemistry, can also require the use of a specific adsorbent for each impurity (e.g., see Chapter 9, Table 8.2).

8.1.3 HYDROGEN USE IN PETROLEUM REFINERIES The demand for hydrogen in petroleum refineries has been mounting exponentially for use in 1. Hydrotreating to reduce sulfur content, which is removed as H2S, and converted to elemental sulfur in Claus units. Unfortunately sulfur is currently a glut on the world market, as there is an excess over demand of 2 million tonnes annually. 2. Hydrotreating to remove oxygen as H2O. This usually proceeds simultaneously with sulfur removal. 3. Hydrotreating to remove nitrogen as ammonia. 4. Saturation of olefins by the simple addition of hydrogen. This is an important factor in stabilizing some types of edible oils. 5. Hydrogenation of aromatics to naphthenes, and dehydrogenation of naphthenes to aromatics. (This is a reversible reaction, as occurring in catalytic reforming, an important source of refinery hydrogen.) 6. Hydrodealkylation of alkyl-substituted ring compounds, i.e., manufacture of naphthalene by dealkylation of suitable kerosene cuts. 7. Hydrocracking (complements catalytic reforming) for gasoline and other light fractions producing a higher yield of light ends than catalytic reforming without producing carbon. 8. Hydrogenation and hydrocracking for lubricating oil production. Hydrogenation improves the color of solvent-extracted oils replacing traditional clay finishing. Hydrocracking also improves the viscosity index and removes sulfur.

8.1.4 REFINERY IN-HOUSE RECOVERY

OF

HYDROGEN1

US refiners have preferred catalytic steam reforming (SR) over POX because of an attractive price for natural gas, lower energy demands, and relatively low quantity of heavy residuals. Presently, however, with the increased cost for natural gas, more serious evaluations may be on the table for the selection of POX. Refineries have also been looking into the recovery of hydrogen from various off-gases and heavy residuals (petroleum coke), which in the past have been flared or used as fuel. Processes used for hydrogen recovery from refinery off-gases are cryogenic separation, PSA, and membrane separation. Cryogenics has the advantage of a high product recovery (90%–98%), a product purity of 95%–99%, and usable by-product recovery, but the disadvantage of high investment and a variable product pressure.


PSA has the advantages of low investment, high product purity 99.9% plus, and the ability to handle the separation with a product pressure almost equal to process feed pressure. Disadvantages are low product recovery (75%–90%) because of product loss during depressurization. Maximum feed pressure is also limited to 55 bar, and there is no profitable by-product recovery. Membrane separation has the advantage of low cost and can handle a feed pressure of up to 125 bar, but it has a low recovery (less than 90%), a low product purity, less than 95%, and delivers a product at a much lower pressure. An important problem to consider is that feed gas impurities such as hydrogen sulfide or aromatics can damage the membrane, which in certain circumstances may require a complete replacement.

8.1.5 RECOVERY

FROM

COKE OVEN GAS

Coke oven gas has a hydrogen content of at least 50%, but little calorific value. Before and during World War II, there was a heavy demand for hydrogen and ammonia, therefore, an upsurge in the recovery of hydrogen from coke oven gas. The advantage for this route also gave a means for recovering ethylene, especially if the capacity of the coke ovens was large enough. A considerable number of plants were built in Europe, and even a few in the United States. Coke oven gas separation has been rendered almost obsolete by SR, POX, and other processes. Nevertheless, hydrogen recovery from coke oven gas has made a comeback using noncryogenic separations as explained in Chapter 9.

8.1.6 HYDROGEN GENERATION PLANTS A potential user of hydrogen is always confronted with the same age-old question. Shall the user purchase the product from an experienced and reliable supplier or shall the user invest the capital to generate his own product? The present market for hydrogen in the United States is around 177 million Nm3=d. The largest industrial users (96%) are petroleum refiners, the ammonia industry, and methanol producers. The second tier (3.75%) includes relatively small volume consumers engaged in chemical synthesis, metallurgical processing, bright annealing of tin plate, semiconductor, and electrical component manufacture. Finally, vegetable oil hydrogenation consumes around 0.25%. In any decision for hydrogen generation, or for that matter any other industrial gas, one has to calculate present daily requirements, and include the best estimate for future needs in the next 5–10 years. The project manager has to take into account the fact that the minimum economic size for hydrogen generation from hydrocarbon feedstock is around 110 Nm3=h. Other processes such as electrolysis have a very high utility cost. This has to be balanced with the following facts: Generation of hydrogen may not be compatible with the owner’s current business activity. It may require setting up a new department, a new staff, a new organization, as well introduce serious safety problems. There is also the question of capital investment and its return over a projected period of time. Furthermore, there may not be any minimum limits to purchase hydrogen from existing reliable suppliers. The various processes to generate hydrogen are described in the following sections. 8.1.6.1

Electrolysis of Water (Figure 8.1)

This option is for small quantities. Equipment is generally supplied in prepackaged modules. Each module has a maximum capacity to generate about 28.32 Nm3=h of hydrogen at a purity of 99.9%, and simultaneously about 14.6 Nm3=h of pure oxygen at a purity of 99.7%. Energy required is approximately 140 kW for this capacity of production, and cooling water


Vents O2

Electric cells

Vent H2

H2

Water seals

Water Water seal seal Gas storage tank

First stage compression

Cooler

Pulsation drum

Vents

Purge bottle

Deoxo unit

Pipe line distribution system Separator

Only if high purity hydrogen is necessary

Heater

Air blower

FIGURE 8.1 Gaseous hydrogen system. (Courtesy of F.G. Kerry, Inc., 1977. With permission.)

requirements are 1100 L=h. As can be seen, therefore, the consumption for power is high, and this option can only be cost-effective in areas where the cost of electrical energy is very low. To offset the high cost of producing hydrogen by this method it is usual for the owner of this type of hydrogen generation to sell the by-product oxygen at a nominal price to a local industrial gas company. In general, the prospective buyer of the by-product oxygen may be happy to purchase it at the right price, in order to eliminate a possible competitor. In the generation of hydrogen by electrolysis the following parameters for safety should be observed during construction and operation of the plant: 1. Adequate building ventilation is needed to prevent accumulation or trapping of hydrogen and oxygen. 2. Building and construction materials that are flammable should be avoided. 3. All rooms containing hydrogen should be properly sealed to prevent the gas entering the unprotected areas. 4. Measures should be taken within the hydrogen area to prevent a spark, flame, or other source of ignition. 5. Purity of electrolytic product gases should be checked continuously. 6. For any system containing hydrogen, devices should be employed to ensure a positive pressure at all times, and to stop operations if pressure should fall in order to exclude air from entering the system.


7. Piping should be arranged and identified so as to make impossible the introduction of hydrogen into an oxygen line, or oxygen into a hydrogen line. 8. Oxygen should never be permitted to contact flammable materials, such as petroleum products. 9. Only nonsparking tools should be used for maintenance work. 10. Hydrogen burns with a pale blue flame almost invisible, so that special precautions should be taken for recognition. 11. Hydrogen piping, valves, and accessories should never be assumed to be entirely leakproof. 12. The above rules are not exhaustive, and should be included in the general rules for safety by the owner and the operator of the plant. 8.1.6.2

Thermal Cracking of Ammonia

The selection of this process is limited, more or less for very small applications, for example, to combust residual oxygen in the recovery of pure argon from an air separation plant. In this application, however, the hydrogen product may carry over some residual nitrogen that may require an increase in liquid nitrogen in the final refining for pure argon. If the hydrogen product is used for the combustion of residual oxygen, then the product of the combustion chamber should be passed through a mechanical chiller, then treated for the adsorption of water. Figure 4.5 shows the typical system layout. If the required production is sufficiently large, the use of VPSA may be considered for a direct separation of cracked ammonia in its gaseous phase. Similar application with a VPSA unit has produced hydrogen with a purity of 99.999% (see Chapter 9 for details). 8.1.6.3

Treatment of Hydrocarbon Feedstock for Hydrogen Recovery

Processes available include catalytic SR of hydrocarbons, thermal dissociation of natural gas, POX of hydrocarbons including residuals as well as petroleum coke, recovery from refinery off-gases, and coke oven gas. Of all the previously mentioned options, however, catalytic SR has become the most frequently used process. It is more practical and more efficient. It can handle natural gas, liquefied petroleum gases, and heavier fractions, such as naphtha. Heavier residuals and petroleum coke, however, have to be treated by POX. 8.1.6.4

Small Steam Reforming Plants (150–1000 Nm3=h)

As the main process requires the feed to be in a gaseous phase, vaporization step may be necessary if it involves the use of liquid, such as propane or naphtha. The former will require the use of a steam vaporizer, the latter the use of a fired heater. Natural gas will only need to be preheated to the specified desulfurization temperature. Assuming that the process feed contains some organic compounds of sulfur, the latter must be removed as they poison the catalysts used in the process. After using the reformer effluent process stream for preheating if necessary, the process gas undergoes treatment to convert the organic sulfur to hydrogen sulfide over a cobalt–cadmium catalyst. This treatment also converts any undersaturated hydrocarbons in the feed to saturated hydrocarbons. The sulfur in the process feed is reduced to less than 0.2 vppm by nonregenerative adsorption on a zinc oxide catalyst, using two beds arranged in series. The sulfur-free process stream is mixed with steam and heated in a preheat coil in the convection section of the reformer. Then the heated mixture passes through the catalyst-bearing tubes in the main reformer furnace. The reactions are


(a) Reforming stage Cn Hm þ nH2 O ¼ nCO þ (m=2 þ n)H2

(8:1)

CO þ H2 O ¼ CO2 þ H2

(8:2)

(b) Shift stage

The reformer reaction is endothermic, the shift reaction is exothermic, but the total of both reactions is endothermic. Therefore, heat to be supplied by the combustion in the radiant zone of the furnace is controlled very carefully. Hot exhaust gases from the combustion chamber are used (a) to preheat the process feed entering the reformer; (b) to generate steam; and (c) to preheat the boiler feed water. As shown in Equation 8.2, the process gas enters a shift converter where additional hydrogen is produced. The reaction is limited by equilibrium and exothermic, favored by a low temperature. The high temperature shift converter involves an iron–chromium catalyst that converts about 60% of the incoming CO. Therefore, the feed stream is cooled and passed through a low temperature shift, which achieves a 93% plus conversion with a copper–zinc catalyst. The stream is then cooled sequentially by deaerated boiler feed water, deaerated feed water, and cooling water. Condensate is sent to the deaerator. Finally the process stream is sent to the hydrogen purifier. The purifier removes impurities such as methane, carbon monoxide, carbon dioxide, etc. It involves a specially designed PSA unit, which recovers a final pure product ranging from 99.0% to 99.9999%. The option for the application of a PSA unit for the final purification of hydrogen from SR of hydrocarbons is that it replaces the use of an MEA system, a methanator, and a dryer. Equipment typically involves four vessels containing the adsorbents, interconnecting piping, and necessary valves. It can be preassembled, and for small plants the entire unit can be prepackaged and sent to the site as a complete unit. The PSA system is reliable and has a wide turndown capability (see Chapter 9 for details). 8.1.6.5 Large Hydrogen Generation Plants (over 1000 Nm3=h) The basic principles already outlined for small plants are the same as for large plants with some variations in design. As outlined for small plants, generation of up to 1000 Nm3=h, four PSA vessels have been recommended. As the main process feed goes up in volume, so does the number of required PSA vessels. At 100,000 Nm3=h of feed, the number of vessels may go up to 12 and even higher. The effects of pressure, temperature, control, and switching system must be studied closely. For hydrogen recovery from a SR gas, the optimum temperature is between 283 and 288 K. Desorption of some components is difficult if not impossible. Purge gas pressure is very important. If the required pressure is too high, over 1.3 barA, there is a drop in hydrogen recovery and a loss to the purge gas. If vacuum is required to drive off certain impurities, a standard PSA system cannot be applied. The design of the catalytic SR must be matched with the design of the PSA system for optimum efficiency of the overall process. In a typical example of using catalytic SR for the production of 33,000 Nm3=h of hydrogen, the raw gas is stripped of sulfur compounds by normal techniques using a cobalt–nickel catalyst for hydrogenation, and zinc oxide for adsorption of hydrogen sulfide. The feed gas is then reformed with steam at 1150 K, at a pressure of 28 bar. The resulting mixture of hydrogen, carbon monoxide, and carbon dioxide is cooled in a waste heat boiler where steam is generated at 45 bar. After this, the process stream enters the shift converter phase and leaves with a hydrogen content of 70% by volume. Once the condensate is removed at 308 K, the feed enters the purification system involving PSA separation, and leaves with a purity of 99.995%.


8.1.7 8.1.7.1

SYNTHESIS GAS, PARTIAL OXIDATION History

For the past 100 years, engineers have been engaged in the production of a synthesis gas that could be used for the production of chemicals. Originally, it started off with the production of water gas from coal, but this process was very inefficient, around 56%, because its fuel value was wasted with the accumulation of inert during the blow period. In 1926, Vandaveer and Parr2 two engineers at the University of Illinois, did experiment with the use of oxygen in place of air, with the following result. They postulated that the reaction of oxygen and carbon to form carbon monoxide is the controlling reaction when carbon, oxygen, and steam are used for making water gas because it produces heat for the decomposition of steam, and at the same time yields a combustible gas. The reactions between carbon and oxygen are (a)

2C þ O2 ¼ 2CO þ 110,150 kJ

(8:3)

(b)

2CO þ O2 ¼ 2CO2 þ 258,280 kJ

(8:4)

(c)

C þ O2 ¼ CO2 þ 184,210 kJ

(8:5)

(d)

CO2 þ C ¼ 2CO þ 71,430 kJ

(8:6)

And the reactions between carbon, oxygen, and steam are (a)

H2 O þ C ¼ CO þ H2 74,800 kJ

(8:7)

(b)

2H2 O þ C ¼ CO2 þ 2H2 75,220 kJ

(8:8)

At the time it was suggested that theoretically, if reaction (a) is tripled before adding to reaction (e), enough heat could be produced to maintain the carbon, oxygen, steam reaction continuous, and the result would be 7C þ 3O2 þ H2 O ¼ 7CO þ H2 þ 255,650 kJ

(8:9)

In other words, the reaction of carbon, steam, and oxygen will result in an increase in the ratio of carbon monoxide to hydrogen, which is a distinct advantage in the production of a synthesis gas for certain chemicals. Economic calculations made at the time (1926) indicated, moreover, that if oxygen were available at a price of US$ 12.00=t, the proposed process could compete on an equal basis with air. Unfortunately, this idea had to wait for another 25 years for the industrial gas industry to catch up and to design an air separation unit large enough to produce oxygen at a cost close enough to the postulated figure. 8.1.7.2

Partial Oxidation Process

In the early 1950s, the Texaco Development Corporation introduced its POX process for the manufacture of ammonia. Basically, it followed the concept suggested by Vandeveer and Parr, but on a more industrial scale. It was a very versatile process as it could use any source of carbon such as pulverized coal, liquids such as naphtha, natural gas, heavy hydrocarbon distillate residuals, or petroleum coke. Equipment was very simple, easy to operate and what was very important, it could operate at very high pressures, 40–65 bar, even higher, which is a definite advantage for ammonia synthesis. It did though require an air separation unit to


supply the required oxygen, which represented an important part of the investment and energy. This unit, however, could also supply liquid nitrogen for the liquid nitrogen scrubbing unit necessary for the removal of residual carbon monoxide, as well as pure nitrogen for stoichiometric make-up for the ammonia product at minimal extra cost and energy. This more than compensated for the overall cost of the oxygen unit. The synthesis gas unit has to be operated in a steady-state condition with long onstream periods to minimize consumption of oxygen and fuel. Composition of the product gas does not vary with pressure. The ratio of H2 to CO is a function of the C=O ratio in the fuel, and the amount of steam added with the feed. The H2=CO mole ratio is nearly unity, and about 4.5 mol% CO2 is present in the effluent gas. Approximately 95% of the sulfur in the feedstock is converted into H2S with trace residuals of COS and CS2 that are troublesome as well as dangerous. Argon from the oxygen will also show up in the product gas, which has been a problem in the nitrogen wash column. Oxygen used in this process should have a minimum purity of 98%, as anything lower will increase the possibility of an explosion in a liquid nitrogen scrubbing unit. This has already happened in the past. Heavy molecules of argon have a tendency to lower the ignition temperature of hydrocarbons and especially hydrogen3. The use of natural gas is convenient in areas where it is available and low cost. Its use, however, requires a little extra oxygen. On the other hand soot production is almost nonexistent. In other areas, where naphtha is the cheaper fuel, its use is very convenient as it simplifies the process and requires a little less steam. Its use, however, does produce some soot, which can be recovered by its direct transfer from the quench water into the primary fuel. 8.1.7.3 Ammonia Synthesis The integrated process for ammonia production is as follows: the air separation unit supplies oxygen at a minimum purity of 98%, and a high purity stream of nitrogen at 99.999%. Oxygen is fed to the reactor for the production of synthesis gas. In the shift reactor, CO is reacted with steam in the quenched gas stream to produce additional hydrogen and carbon dioxide, which must then be removed by either an MEA, Rectisol, or other system. Finally, the resultant effluent is fed to the liquid nitrogen scrubbing unit for the removal of any final traces of CO, CH4, and argon, leaving nitrogen as its sole impurity. This stream is then joined with an additional stream of high-purity nitrogen at the correct stoichiometric ratio for ammonia synthesis, and fed into the ammonia reactor. 8.1.7.4 Hydrogen Recovery from Ammonia Synthesis Plants The increasing cost of energy has already pushed ammonia producers to recover hydrogen wasted in the purge stream of very large capacity plants. In one specific example using a cryogenic process and treating 8558 Nm3=h of purge gas, the recovery included 155 Nm3=h of ammonia that was sent directly to storage, 5917 Nm3=h (86% H2) at 70 bar sent to the inlet of the compressor wheel operating at 66 bar, and 2486 Nm3=h at 4.3 bar of tail gas to burners. This represented a 97.2% recovery of the purged hydrogen, and a payout of 1.4 years, in terms of extra ammonia production (Figure 8.2). 8.1.7.5 Other Uses for Synthesis Gas In addition to ammonia, the synthesis gas is used for the production of methanol, in petroleum refining, iron ore reduction, oxo synthesis, and production of nickel carbonyl catalysts.


67 kg

Recycle gas

66 kg 26 kg

134 kg Coolers and separators

147 kg Synthesis

+20⬚C

⫺20⬚C

Hydrogen recycle

NH3

70 kg and ⫺25⬚C

Desulfuration

1000 T/D (5917 Nm3 h)

At 67 kg

Purgal separation

Purge gas feed (8558 Nm3 h) 140 kg and ⫺23⬚C

4.3 kg

(2486 Nm3 h)

Tail gas to burners

(155 Nm3 h) Ammonia recovery—50 kg and − 50⬚C

FIGURE 8.2 Purgal system synthesis and refrigeration with purging and purgal added for hydrogen recovery. (ß Air Liquide, all rights reserved, 2006. With permission.)

8.1.7.6

Fuel Cells

The world production of hydrogen is approximately 50 million tonnes per year, but 90% of hydrogen generation goes into the synthesis of ammonia, methanol, or the use in petroleum refinery operations. The balance 10% is supplied for the propulsion of NASA’s space vehicles. Nevertheless, there will be a growing market (about 2% of the world’s production) in the very near future in the use of fuel cells for vehicular propulsion. Fuel cells are considered seriously as a power alternative that requires hydrogen as a key energy component. These devices transform the potential chemical energy of hydrogen and oxygen into direct current power, heat, and water. Obviously, it is a clean and more efficient method of producing both stationary and vehicular power. As fuel cells require hydrogen, there is ongoing study for a low-cost generation and portability of hydrogen from various processes, and storage in various forms, such as gas, liquid, or hydride. Presently, there are several demonstration automobiles operating satisfactorily in United States, Japan, and Europe. Unfortunately the cost of power per kilowatt is still too high, and has to be lowered to a more reasonable level to be acceptable for the average consumer.

8.2 CARBON MONOXIDE Carbon monoxide is a highly toxic, colorless, and flammable gas occasionally found in the atmosphere in trace amounts as a by-product of inefficient combustion engines, and the incomplete conversion of carbonaceous fuels used in industry. Its physical properties are given in Table 8.3. It also has a condensation point of 81.15 K, and a freezing point of 74.15 K. It is slightly soluble in water, and in general its physical properties are close to that of nitrogen. Its high toxicity is due its reaction with the hemoglobin of red blood cells, which prevents them from taking up oxygen. Since carbon monoxide is used in highly industrialized chemical and metallurgical complexes, project engineers should take every precaution to design the


TABLE 8.3 Physical Properties of Carbon Monoxide Molecular weight Normal boiling point Vapor pressure Latent heat of vaporization Molar specific heat capacity of gas at constant pressure in the range of 0–1 bar at 298 K Thermal conductivity of gas Gaseous density real Liquid density Critical temperature Critical pressure Flammability range

28.01 kg=kmol 81.6 K at 101.325 kPa 154 mbar at 68.1 K 6.05 MJ=kmol at boiling point and 101.321 kPa 29.07 kJ=(kmol K) 0.024 W=(m K) at 273.15 K 1.250 kg=m3 at 273.15 K and 101.325 kPa 0.79 kg=L at 81.6 K and 1.013 bar 132.3 K 34.96 bar 12.5%–75% in air, 15.5%–94% in oxygen


proposed area for the production and supply of the gas in the safest possible way. It should be kept in mind that the toxic effect of carbon monoxide is cumulative over time, even though its amount in the atmosphere at the working area may be in trace quantities. According to the National Bureau of Standards Technical Paper 212, the effects of exposure to carbon monoxide at a concentration of 400 vppm are given in Table 8.4. The effects naturally depend on the physical condition of the individual. Another property of carbon monoxide that must be kept in mind is that it is flammable in mixture with air. Its flammability limits with air are 12.5% to 75%, compared with 4% to 74% for hydrogen. In other words, in the production and handling of pure carbon monoxide, in gaseous or liquid phase, equipment must be electrically grounded, and sparking equipment within 5 m should be explosion-proof or purged continuously with gaseous nitrogen (Table 8.5).

8.2.1 SOURCES Carbon monoxide has become a vital chemical raw material for the synthesis of monomers, polymers, ethanol, acetic acid, and oxo alcohols. It can be produced in several ways: 1. Coal gasification by traditional ways using coal and steam producing a H2=CO ratio to up to 1:2, which is chemically advantageous. This process, however, also produces a high content of nitrogen, which must be removed by a subsequent cryogenic application. It also involves ash and tar removal problems. This process is limited to areas, which do not have low-cost natural gas or liquid hydrocarbons. TABLE 8.4 Effects of Exposure to Carbon Monoxide Exposure 0 to 45 min 45 min to 1 h=30 min 1 h=30 min to 2 h=15 min Over 2 h=15 min

Effect Not perceptible Perceptible Headache and nausea Dangerous to life

Source: Courtesy of Natural Bureau of Standards, USA, 2005. With permission.


TABLE 8.5 Comparison Chart of Properties Carbon Monoxide

Nitrogen

28.01 81.6 6.05 154 1.249 0.79

28.01 77.3 5.55 at boiling point and at 1.0132 bar 126 at its triple point 1.249 at NTP 0.81 at boiling point and 1.013 bar

Molecular mass (kg=kmol) Boiling point (K) Latent heat of vaporization (MJ=kmol) Vapor pressure (mbar) Density of gas (kg=m3) Density of liquid (kg=L)


2. SR of natural gas or naphtha. This will produce a hydrogen to carbon monoxide ratio of 3:1 and may require a cryogenic application for the separation and upgrading of carbon monoxide. 3. POX of heavy oil that will produce a hydrogen to carbon monoxide ratio of 1:1. With an associated CO2 recycle it can increase the CO production to a higher level.

8.2.2 8.2.2.1

CARBON MONOXIDE RECOVERY General

Carbon monoxide can also be recovered from waste gas from nitrogen scrubbing units used in the production of ammonia synthesis gas, producer gas, and water gas 3,4. From Table 8.6 it is obvious that SR of natural gas or light oil (naphtha), and POX of heavy hydrocarbons are the preferred choices. Water gas may be a third choice if the process equipment already exists near the site. Otherwise its selection should be made only if no natural gas or petroleum were available. Producer gas has the disadvantage that it uses air for the combustion of coal, therefore increasing the quantity of nitrogen in the product gas, lowering the calorific value of the gas, and making the recovery of CO more difficult.

8.2.3

GENERAL PROCESS

OF

RECOVERY

The cryogenic separation of hydrogen from carbon monoxide from either SR or POX is relatively easy. It is simply a matter of cooling the mixture sufficiently to condense the carbon TABLE 8.6 Typical Composition in % of CO Containing Gases Item

SR of Natural Gas

POX of Natural Gas

Nitrogen Scrubbing Unit

Producer Gas

Water Gas

H2 N2 CO A O2 C1 C2 CO2

70.58 0.16 23.53 — — 0.33 — 5.4 100.0

61.8 0.3 35.0 0.4 — 0.5 — 2.0 100.0

3–6 50–70 7–35 0–5 — 0.2–18 — — 100.0 ?

14.4 51.1 26.3 — 0.2 1.8 1.0 5.2 100.0

48.7 3.4 43.1 — 0.5 0.6 — 3.7 100.0

Source: From Jost, W. in Explosion and Combustion Processes in Gases, McGraw-Hill Publications in Aeronautical Science, New York, 1946. With permission.


monoxide, then separating the liquid carbon monoxide from the gaseous hydrogen. Any hydrogen dissolved in the liquid carbon monoxide can be removed by expanding the liquid to a lower pressure in a flash drum. The hydrogen will evaporate out of the liquid, leaving liquid carbon monoxide free of hydrogen. The hydrogen flashing out of the carbon monoxide will be of a relatively small stream and can be discarded, but if a high recovery of carbon monoxide is important, the stream should be recycled and brought back into the plant with the feed. With feed gases containing a high percentage of hydrogen, it is necessary to produce a relatively pure hydrogen stream to obtain a reasonable recovery of carbon monoxide. This must be done even if there is no use for the hydrogen. Producing a pure hydrogen stream, free of carbon monoxide is, however, a more difficult problem than producing carbon monoxide free of hydrogen. To obtain a good recovery of carbon monoxide it is necessary to cool the gas to about 73 K at a pressure of 22 to 28 bar. This low temperature can be produced by expanding the hydrogen product in an expansion machine, sometimes with the use of a twostage expansion. This method has the advantage that the expansion machine can produce enough refrigeration to counterbalance all the heat gain from the overall process. The normal atmospheric boiling point of nitrogen is 77.35 K, that of carbon monoxide 81.55 K, and that of argon 87.25 K. As this proximity of boiling points indicates, the relative volatility between nitrogen– carbon monoxide and carbon monoxide–argon is very small. High reflux ratios are required, therefore, to achieve any satisfactory separation by fractionation. A plant in which carbon monoxide must be separated from either nitrogen or argon requires the circulation of large recycle streams. This recycle can be kept to a minimum, however, by using a double column system as in air separation units in which a high-pressure column provides reboil for the low-pressure column, whereas the low-pressure column provides reflux for the high-pressure column. Fortunately, in most commercial applications for carbon monoxide recovery small concentrations of nitrogen or argon are not harmful. They are merely inert, which can be occasionally purged from the system. Argon is of course a valuable product, and if the feed gas contains an appreciable quantity its recovery for sale should be considered. The gases containing carbon monoxide in commercial quantities rarely contain oxygen except as a minor impurity. It can and should be removed from the feed gas by catalytic hydrogenation. Oxygen should not be allowed to concentrate in any low-temperature separation unit, which treats hydrogen, carbon monoxide, or methane. The safest procedure is to remove the oxygen before the feed enters the low temperature unit. Methane and carbon monoxide can be easily separated by distillation. Their relative volatility is of the order of 10, therefore, it is possible to produce without too much trouble both a methane-free carbon monoxide stream and a pure methane stream. In actual practice the purity of the methane stream is limited more by other considerations, (heat transfer, the maximum desirable operating pressure of the feed gas or recycle streams) than by distillation limitations. Hydrocarbons heavier than methane if present in small quantities can be removed with the methane stream. Larger concentrations should be removed from the system as they are condensed out of the feed gas. Acetylene is always a dangerous compound to introduce in any low-temperature separation unit unless the feed gas also contains enough ethylene to absorb the acetylene. If there is not enough ethylene for the purpose, the feed gas should be hydrogenated to convert the acetylene to ethylene.

8.2.4 BASIC CRYOGENIC RECOVERY PROCESSES4 A simplified schematic is shown in Figure 8.3 of a plant producing pure carbon monoxide from natural gas. The feed for the low-temperature separation unit is prepared from natural gas either by SR or by POX with high-pressure oxygen. Typical compositions of these feed


1 psig Impure H2 n Brine

Warm exchanger

LP Column 4 psig

1 psig

CO

4

Pure CO Pure H2

Separator

Expansion engine

H2

Final exchanger

Methane bleed Dehydration drums

CO H2

HP Column 400 psig H2

FIGURE 8.3 Basic cryogenic recovery of CO. (Courtesy of F.G. Kerry, Inc. With permission.)

gases are given in Table 8.6. They have a high concentration of hydrogen, a low methane concentration, and minor amounts of nitrogen and argon. In this case no attempt has been made to remove the nitrogen and argon, so these components are the main impurities in the carbon monoxide product. A fairly high concentration of carbon monoxide is left in the methane fraction. The small quantity of methane in the feed gas results in only a small loss of carbon monoxide, but at the same time allows the methane to be evaporated by the feed gas without resorting to excessive feed gas pressure. Feed gas is compressed to 22 to 28 bar and treated for the removal of carbon dioxide and water by molecular sieve adsorption. If concentration of carbon dioxide is on the high side, however, it may be more economical to use either an amine or hot carbonate solution that can be regenerated, followed by activated alumina dryers. The dryers may be regenerated in turn by heating a part of the hydrogen product in a steam heater, and passing it through the dryer bed that is not in service. After the bed has been thoroughly regenerated, the steam heater is bypassed, allowing the hydrogen to cool the bed to its operating temperature. The hydrogen used for dryer regeneration is cooled and rejoins the main hydrogen product stream. The feed gas passes in series through the primary warm and cold exchangers where it is cooled to about its dew point by countercurrent heat exchange with the product gases. It flows through the high-pressure pot into the condenser–vaporizer where most of the methane, carbon monoxide, and nitrogen mixture is condensed. The uncondensed feed gas leaves the top of the condenser–vaporizer and enters the final condenser where nearly all the remaining nitrogen, carbon monoxide, and methane condense, leaving a hydrogen product containing about 98% hydrogen. The hydrogen product leaves the top of the final condenser and enters another pass of that exchanger where it is warmed by countercurrent heat exchange with the feed gas. The hydrogen product is cooled by expanding it in the first stage of the reciprocating expansion engine, and it is again heated in another pass of the final condenser. Expansion in the second stage of the expansion engine cools the hydrogen again so that it can be heated in another pass through the final condenser. Passing the hydrogen product three times through final condenser provides enough refrigeration to condense nearly all the carbon monoxide entering that exchanger in the feed gas. The refrigeration produced by the expansion engine is


also sufficient to balance all other heat gains into the unit; therefore, no other external source of refrigeration is required. The hydrogen leaving the final condenser is warmed by countercurrent heat exchange with the feed gas in the cold exchanger and the warm exchanger. The liquid, which condenses out of the feed gas in the final condenser, flows back into the condenser–vaporizer where it mixes with the liquid condensed out of the feed gas in this exchanger. Together these liquids flow into the high-pressure pot. This liquid contains all the methane and nearly all the carbon monoxide and nitrogen in the feed gas. This liquid leaves the high-pressure pot and is expanded into the flash drum. Any hydrogen, which has remained dissolved in the liquid, evaporates, leaving a liquid completely free of hydrogen. This hydrogen flash gas is warmed by countercurrent heat exchange with the feed gas in the cold exchanger and the warm exchanger. It is sent to the suction of the feed compressor. Returning to the process thus increases the recovery of the hydrogen and carbon monoxide it contains. The liquid leaves the bottom of the flash drum and enters the recycle condenser where it is partially vaporized. From the recycle condenser it is fed into the carbon monoxide column where it is separated into a pure carbon monoxide stream and a methane stream. Reboil for the column is provided by the condenser–vaporizer. Methane in the bottom of the column is vaporized by heat exchange with the feed gas in this exchanger. The methane stream is withdrawn as a liquid from the bottom of the column. It is vaporized and heated in the cold and warm exchangers. To provide sufficient temperature difference to operate the condenser–vaporizer, the methane stream contains about 50% carbon monoxide. Because of the small quantity of methane present in these gases, this represents only a small loss of carbon monoxide. Pure carbon monoxide withdrawn from the top of the carbon monoxide column is split into two streams: the carbon monoxide product stream and the carbon monoxide recycle stream. The carbon monoxide product is warmed in the cold and warm exchangers where it helps to cool the entering feed gas. The carbon monoxide recycle stream is warmed in the recycle warm exchanger by countercurrent heat exchange with high-pressure carbon monoxide recycle. The low-pressure carbon monoxide recycle is compressed to about 4.15 bar in the recycle compressor. This high-pressure carbon monoxide is cooled in the recycle exchanger and liquefied in the recycle condenser. It is expanded into the carbon monoxide column where it provides reflux for the carbon monoxide–methane separation. The carbon monoxide produced by this process contains less than 0.05% hydrogen and less than 0.05% methane. 8.2.4.1 Methane Wash Cryogenic Recovery5,6 (Figure 8.4) A further development to the previous cryogenic process is the recovery of carbon monoxide using a methane wash by absorption followed by fractionation. This process cycle is of interest because it produces both high-purity carbon monoxide as well as hydrogen with less than 1 vppm of CO. It does require, however, a methane to hydrogen ratio of 1=99 in the process feed. Two fractionation columns are also necessary, one for the separation of hydrogen (K-1) and the other for the separation of carbon monoxide (K-2). The process feed is first cooled in heat exchangers E-1 and E-2; and after further cooling in heat exchanger E-4, is partially condensed in separator V-1 at 90 K. The bulk of the condensed liquid in V-1 is flashed through an expansion valve JT-1 to a lower pressure, and led to the CO=CH4 column K-2. The vapor from V-1 is led into the high-pressure methane wash column K-1, and used to strip CO from the high-purity hydrogen. The product hydrogen is reheated in heat exchangers E-1 and E-2. A portion of the returning product hydrogen is expanded to provide added refrigeration to the process before leaving the cycle at a lower pressure. The liquids from the sumps of K-1 and V-1 are flashed (expansion) and fed to heat exchangers E-5 and E-4. When the latter liquids reach the sump of K-2 they are combined,


LP H2 product JT-4

HP H2 product Feed gas CH4

K-1 E-1

JT-3

E-5

K-2

E-6

E-3 E-7

V-1 E-2

E-8

C-2

E-4 JT-1

JT-2

C-1

P-1 CO product

FIGURE 8.4 The methane wash process for H2=CO recovery. (From Isalski, W.H., Separation of Gases, Oxford University Press, London, 1989. With permission.)

removed, pumped out, and partly withdrawn as fuel gas; the balance is subcooled in heat exchanger E-5 and used as reflux in K-2. The vaporized CO leaving the top of K-2 as pure CO free from any CH4 is combined with the vapor leaving heat exchanger and reboiler E-3. After passing through E-6 and E-8, the total is led to a compressor, which serves both as a process recycle and product compressor. The high-pressure cycle CO is cooled in E-8, combined with E-6, and used as a refrigerant in E-3 and partly as a reflux in K-2 after being flashed in expansion valve JT-4 to complete the cycle. As noted, the process cycle is easy, but complex in operation. It has been replaced by a more simplified process cycle, incorporating a PSA and membrane technology and is described below. 8.2.4.2

Simplified Carbon Monoxide Recovery

With the rapid development of noncryogenic separation of gases by the use of permeable membranes, or adsorption PSA, or VPSA, the recovery of carbon monoxide has been greatly simplified, and made less expensive (see Chapter 9). It was mentioned previously that to increase the recovery of carbon monoxide one has to improve both the recovery and purity of the hydrogen component normally found in the same feed gas with carbon monoxide. Because of the large differential between their boiling points, namely 61.2 K at 101.325 kPa, the separation can be carried out by simple condensation, followed by either a permeable membrane or adsorption. This separation can be carried out to a very high degree, especially ˚ , and for by applying a recycle. (The diameter of a hydrogen molecule is above 2.4 A ˚ , nearly equal to that of carbon dioxide.) (see carbon monoxide molecule it is around 3.7 A Chapter 9.) As to the small content of methane in the main feed gas, the final separation of methane from the carbon monoxide can be carried out very easily by stripping, as the differential in their boiling points at 101.325 kPa is around 30.1 K. A recently proposed process cycle7 is as follows: the process and equipment involves a primary heat exchanger system consisting of brazed heat exchangers (BHXs), which can withstand a pressure of 65 bar. The main feed


leaves the cold end of the BHXs at a temperature of around 98.15 K, entering a condenser vessel (A), wherein the cold vapor, mostly hydrogen and some CO return through the BHXs, where they give up their refrigeration in countercurrent with the incoming feed. Once out of the BHXs, it passes through a permeable membrane, or adsorbent unit, letting the hydrogen pass through, and the retained carbon monoxide at a lower pressure (20 bar) is recycled to the main feed after undergoing external compression. The remaining liquid in the sump of the condensing vessel (A), which is primarily CO with some hydrogen and all of the contained methane, is expanded to 20 bar, and led into a stripping column (B), which vaporizes the remaining hydrogen, concentrating the CO with all the methane. The more volatile hydrogen, stripped of the entire CO, rises to the top and leaves the cryogenic system through the BHXs. At this point, the hydrogen can either be compressed and recycled to the main feed for maximum recovery of CO, or join the hydrogen feed at the entrance of the permeable membrane for maximum recovery and purity of hydrogen, as well of those of CO. This latter option depends on the requirements of the process cycle. The remaining liquid in the sump of stripper (B) is then passed through a second stripper (C), where the residual methane is completely stripped from the product CO. This gaseous product CO is then extracted from the cycle after giving up its refrigeration capacity at the BHXs system. The residual methane is collected as a liquid in the sump of vessel C, and after being revaporized passing through the BHXs is sent out either as a fuel, or to flare. Another variation of the basic process is the use of an extra heat exchanger system where part of the hydrogen is expanded in an expansion machine with either one or two stages. This latter variation will save on external compression energy, but entails the addition of extra machinery with its ancillary piping and controls.

8.2.5 COMPRESSION

AND

CONVERSION MACHINERY

FOR

CARBON MONOXIDE8

Because of its high toxicity, any compression machinery used for the delivery of CO in gaseous state must be carefully selected, and should be from a very reliable source. No process gas leakage into the atmosphere can be tolerated especially in areas involving workforce. A double multiple seal arrangement, or even a three-stage seal must be considered. This must also be backed up with nitrogen injection to block leakage of CO. For transportation purposes, liquid carbon monoxide may be considered, provided that extraordinary precautions are taken during its conversion from liquid to gas at the customer’s station. This demands extreme care in designing and operating the cryogenic pump. 8.2.5.1 Requirements for a Liquid CO Pump A pump to handle liquid carbon monoxide may be a standard cryogenic pump, with the following modifications: 1. Electric switches should be mounted in the pump control cabinet at the back of the pump in nitrogen-purged compartments. 2. The pump motor should be totally enclosed, fan-cooled, and provided with a nitrogen purge. A pressure switch should be interlocked with the starter switch to prevent starting until purge pressure is correct. 3. All possible leakage points and safety valve outlets should be extended to a high vent stack with a nitrogen-blanketing inlet at the bottom. 4. A circuit breaker protected main switch should be located 5 m minimum from the pump and should be connected to the pump with sealed explosion-proof conduit and fittings.


8.2.5.2

Conversion from LCO to Gas

Liquid carbon monoxide (LCO) should pass from the bottom of the liquid storage vessel through a short length of vacuum-insulated piping and a vacuum-insulated valve to the pump intake. During initial start-up, the liquid valve to the pump and a vapor return valve from the pump should be opened approximately 30 and 15 min, respectively, before the pump is started. This will allow the pump’s components to achieve thermal equilibrium with cryogenic conditions so that all seals and moving parts are at the same working tolerances. This procedure is more conservative than that generally recommended for nontoxic gases to keep seal wear and clearances to a minimum. High-pressure liquid expelled by a positive displacement piston pump should be passed through a high-pressure vaporizer selected especially for safe operation. The tubes used in this exchanger should be designed for 200 bar, and hydrostatically tested to 310 bar. Highpressure vapor should be carried through high-pressure piping to a tube storage bank. The high-pressure line to the storage bank should be from extra heavy brass in place of Schedule 80 steel. This eliminates the possible danger from low-temperature brittleness should cryogenic liquid enter the line before the pump is shutdown. 8.2.5.3

Safety of LCO Transport

Safe design and operation of the semitrailer carrier of liquid carbon monoxide should eliminate any possibility of uncontrolled leakage, spillage, or formation of flammable mixtures with air in any part of the system. The design of the trailer should include minimum heat gain, and filling should involve a top-fill spray line to keep pressure buildup to a minimum. The latter procedure condenses the vapor remaining in the trailer, thus lowering the trailer pressure. The trailer should also be provided with an excess flow valve, and failure occurs in the transfer hose. Any excess CO vapor emitted during the filling operation should be vented to a flare, assuring that low-temperature liquid is in the trailer when it leaves the filling station.

8.2.6 SAFETY EQUIPMENT

IN

GENERAL

Safety equipment such as CO indicators, face masks, gas masks, and portable breathing apparatus should be made available at the loading stations on the transport trailer and at the customer loading station. In addition, a carbon monoxide resuscitation apparatus should be readily available in areas where carbon monoxide is used.

8.3 HIGH-PURITY NITROGEN 8.3.1 GENERAL CHARACTERISTICS Nitrogen ranks sixth in cosmic abundance. Apart from making up 78% by volume of the Earth’s atmosphere, it is also found in free form in meteorites, in gases from volcanoes, in mines, mineral springs, stars, and in combined form in some minerals. The natural nitrogen on Earth consists of two stable isotopes, nitrogen-14 (about 99.63%) and nitrogen-15 (about 0.37%). There are also three other radioactive isotopes and are as follows: nitrogen-12, nitrogen-13, and nitrogen-16. Nitrogen has an atomic number of 7 and the physical properties of nitrogen are given in Table 8.7. Nitrogen also forms some very important oxides (Table 8.8). Nitrogen gas is chemically inert at ordinary atmospheric temperatures, which is a desired property for many applications in the chemical and metallurgical industries. Because of its inertness and its relative low cost


TABLE 8.7 Physical Properties of High-Purity Nitrogen Molecular weight Normal boiling point Vapor pressure Latent heat of vaporization Molar specific heat capacity of gas at constant pressure in the range 0–1 bar at 298 K Thermal conductivity of gas Gaseous density real Liquid density Critical temperature Critical pressure Flammability range

28.01 kg=kmol 77.347 K at 101.325 kPa 126 mbar at 63.2 K 5.55 MJ=kmol at boiling point 77.347 K and 101.325 kPa 29.07 kJ=(kmol K) 0.024 W=(m K) 1.251 kg=m3 at 273.15 K and 101.325 kPa 0.81 kg=L at 77.347 K and 1.013 bar 126.10 K 33.94 bar Inert


and ease of availability, it is employed as a blanket to exclude the deteriorating and dangerous effects of oxygen and moisture. In its liquid phase, it is very suitable for freeze-drying food and as a refrigerant during the transportation of food. Liquid nitrogen is also used extensively for the storage of human as well as animal semen.

8.3.2 RECOVERY About 90% of commercial nitrogen is recovered by fractionation of liquefied air. The balance is produced by adsorption or by passing air under pressure through hollow permeable membranes. Though the latter processes are less efficient than cryogenic separation, they are lower in investment and less costly in operation.

8.3.3 APPLICATIONS

FOR INERTNESS

In the early 1940s, engineers at Air Liquide Canada had just finished installing a pipeline to supply oxygen to Dominion foundries in Hamilton, Ontario. In fabricating tin plate, the steel mill used cracked natural gas as a composite-inert gas medium. This operation was not very efficient; the cracking units were in constant need of repair and the percentage of rejects was high. It was natural, therefore, that engineers thought of replacing cracked gas with waste nitrogen from the same plant, which supplied oxygen. Nitrogen containing 1.5% oxygen was TABLE 8.8 Basic Properties of Oxides of Nitrogen Dinitrogen monoxide (nitrous oxide) (Mono) nitrogen monoxide (nitric oxide) Dinitrogen trioxide Nitrogen dioxide Dinitrogen tetroxide Dinitrogen pentoxide Nitrogen trioxide

N2O, colorless gas, boiling point 184.65 K NO, paramagnetic, boiling point 121.35 K N2O3 (blue solid), melting point 172.45 K dissociates reversibly into gas phase NO and NO2 NO2 (brown gas, paramagnetic, dimerizes to N2O4) N2O4 (colorless liquid dissociates reversibly to NO2) N2O5 (colorless solid sublimes at 305.5 K to unstable molecular gas) NO3 (unstable paramagnetic radical)

Source: From Pauling, L. in General Chemistry, W.H. Freeman & Company, San Francisco, 1953. With permission.


TABLE 8.9 Removal of Hydrogen and Carbon Monoxide from Ultrapure Nitrogen Nitrogen flow Impurities Catalyst recommended Residence time Normal guarantee

5000 Nm3=h at 5 bar and 473 K Hydrogen 5 ppm; carbon monoxide 5 ppm E 221 P=D, 0.5% 2.5 s, volume 0.48 m3 5 vppb for each of hydrogen and carbon monoxide


found to be unacceptable. Nitrogen with less than 100 vppm was then used successfully but only when small additions of hydrogen and carbon monoxide were made. Finally, high-purity nitrogen (less than 10 vppm oxygen) proved successful with only small additions of hydrogen. How much H2 had to be added, however, did not obey stoichiometric ratio, but had to be determined empirically for each particular application. Depending on specific alloys and temperatures there also had to be a different ratio of H2 to produce water vapor to prevent the breakdown of water back to molecular oxygen and hydrogen. Unfortunately, because of the pressures of wartime production, the development of the process had to be abandoned. With the termination of the war and the rapid growth of petrochemicals, there was an urgent and growing need for a low-cost inert gas for blanketing toxic and dangerous chemicals. Nitrogen, therefore, could offer a wide range of services. As will be detailed in Chapter 12, there are three key elements to initiate an ignition and fire: a flammable material, the presence of oxygen, and a source of energy. The successful removal of one of the key elements deters the possibility of fire and explosion. All flammable materials have wide explosive ranges (Table 8.9). In order to maintain a mixture from becoming flammable, its vapor concentration in air must be kept outside the explosive range, which is not always easy to do. Ignition, the next key element, can be triggered by an increase in temperature, sparks, static electricity, friction, and improper grounding. It is not always possible to anticipate and prevent this possibility. The third element, namely oxygen, however, can be controlled quite safely with the reduction of its concentration to a safe level by using an inert gas, such as composite gases (which may be by-products of combustion), carbon dioxide, or nitrogen, which is generally used and most cost-effective. 8.3.3.1

Nitrogen as a Preservative

Chemicals, polymers, and food products degrade through reactions with oxygen or moisture in air. Nitrogen, therefore, can be used to protect them against degradation by purging, pressure transferring, and blanketing. In the use of blanketing, dry nitrogen replaces atmospheric air inside a storage tank or reactor, thus reducing the risk of formation of potentially hazardous products. The nitrogen can be introduced and maintained with a slight positive pressure to exclude the entrance of outside air. The use of inert blanketing also prevents the infiltration of outside moisture, which in turn prevents corrosion. 8.3.3.2

Nitrogen as an Emissions Controller

Nitrogen can be used as an inert blanket in the headspace of a vessel to minimize vapor and valuable product losses, as well as to protect plant personnel in the plant area. In a blanketed storage vessel, the vapor can only be vented when a preset relief valve blows.


8.3.3.3 Nitrogen Use in Sparging Product streams reaching a reactor vessel or storage tank contain considerable quantities of dissolved oxygen, which may be detrimental to product quality. Nitrogen sparging dissolves or infuses nitrogen into liquids to strip dissolved oxygen. 8.3.3.4 Nitrogen for Pressure Transferring Whenever pumps are not available, when their use might entail contamination or degradation, or when handling abrasive or corrosive materials, pressurized nitrogen can prove valuable as a means of transferring products from one location to another in the same facility. In such cases, consumption of nitrogen will be a function of the density of the material to be moved, the length and configuration of the transfer line, the flow rate and difference in levels between transfer points. 8.3.3.5 Liquid Nitrogen for Vapor Recovery Nitrogen in its liquid phase is very useful as a refrigerant to condense and to recover chemical vapors for environmental or economic reasons. Such systems use cryogenic nitrogen to recover these vapors as they are vented from trucks, railcars, or other vessels. A shell- and tube-heat exchanger is employed as the vent condenser. Liquid nitrogen flows through the tubes whereas the solvent-laden air from the carrier is led by a duct to the shell side of the exchanger. The solvent vapors cool and condense into a liquid holding tank. The remaining clean vapor stream contains very little residual vapor, and can be vented to a suitable location. Nitrogen exhaust may be reused elsewhere in the facility. 8.3.3.6 Liquid Nitrogen Makes Worn Rubber Tires Profitable Liquid nitrogen has been used to freeze and embrittle worn automobile tires making them easier to shred into small particles. These rubber particles are now used as a foundation in place of coarse gravel for the construction of concrete highways. So far the tests have proven satisfactory and cost-effective.

8.3.4 PROCESS

AND

EQUIPMENT OPTIONS

Nitrogen in various purities is a very commonly used commodity in the chemical, petrochemical, and food industry. Only sulfuric acid is produced in greater quantities than nitrogen in the United States. Traditionally, nitrogen has been produced and supplied in three modes: 1. For small requirements, by liquid nitrogen delivered by transport carriers, and stored on-site in vacuum-insulated vessels. Converted into gas using ambient or heated vaporizers as required by the client. 2. For larger volumes, by on-site cryogenic air separation units that are owned and operated either by the industrial gas supplier or by the client (Figure 8.5). 3. For very large volumes, by transmission pipeline from the supplier’s central production facility to the client’s site. In the case for relatively small quantities, less than 1000 Nm3=h, where ultrahigh-product purities are not involved, as in case (1), noncryogenic units, such as hollow fibers membrane permeation, PSA, and VPSA, have taken an increasing portion of the market. In fact, industrial gas suppliers actually rent such units to clients as a temporary measure to tie up the client with a contract until demand increases to the point where it merits supply from a larger cryogenic plant.


Cold

box

Pure Liquid nitrogen nitrog en

Air filter

el

l pan

o Contr

Silencer l trica Elec ch it sw gear After

coole r

Surface tank

or

ress

omp

Air c

FIGURE 8.5 High-purity nitrogen (HPN) generator series. (Courtesy of F.G. Kerry, Inc. With permission.)

8.3.4.1

Cryogenic Process Cycle

As the market for high-purity (oxygen-free) nitrogen increased geometrically after World War II, the cryogenic industry answered the call in the middle 1960s by designing low-cost, completely packaged automatically monitored units ranging in capacity from 350 to 5000 Nm3=h. Purity range was less than 10 vppm of oxygen or even lower, and with a product dew point of less than 200 K. These units were designed with a lower operating cost than composite gas generators or any other means of supply of oxygen-free gas. This process suited market requirements of the chemical, petrochemical, aviation, and the burgeoning electronic industries. These cryogenic nitrogen generators have proven outstanding for several reasons: using a low-pressure process cycle, supplying the product stream at a pressure of 3 to 4 bar continuously, controlled for product purity and automatically monitored. Only one compressor and one turboexpander are used. Originally, impurities such as water and carbon dioxide were removed by reversing heat exchangers with the usual silica gel adsorption filters to safeguard process and product. There was no corrosion of internals to cope with, and no refractories to deteriorate, when compared with composite gas generators using natural gas or other fuels. With the advent of prepurification using zeolite molecular sieves to remove impurities from air, the process was modified to include completely automated external prepurification. This modification required regeneration gas: a waste stream from the main condenser, which could include a 45% oxygen. Experience from field tests proved that this was a safe stream for use as a low-pressure regeneration cycle. The flow sheet in Figure 8.5 refers to a typical high-purity nitrogen generator. Such units could also produce up to 10% liquid nitrogen with the same product purity, but at the usual


expected loss of gaseous product at a ratio of 3:1. A small backup liquid nitrogen storage tank was always recommended as part of the overall supply, which could be used to store the liquid nitrogen product, or serve as a liquid assist to keep the plant in full operation should the air supply or the expansion machine stop unexpectedly. In this plant, only the high-pressure (lower) column was used, fitted with a high number of fractionation trays. As the number of trays increased, the purity of the nitrogen improved. In fact, field results showed that reaching an oxygen impurity of slightly less than 1 vppm, and even 0.1 vppm was realistic. But to reach less than 10 vppb, the standard that was being imposed by the electronic industry, the design had to include an additional column as well (see Section 8.3.4).

8.3.4.2 Permeable Membrane Separation Process Permeable membrane separation (PMS) systems, to be described in more detail in Chapter 9 covering noncryogenic separation, began to arrive on the gas separation scene around the middle of 1970s. Individual elements consist of hollow fibers, which separate gas components according to their individual permeabilities. These operate quite simply. They have no moving parts; therefore, require no mechanical maintenance. On the other hand, they can only separate nitrogen with a purity, which may range from 95% to 99.5% depending on the volume treated. Nevertheless, these purities may be satisfactory for many applications. This includes purging lines and vessels, blanketing storage vessels, bunkers, and cargoes to prevent degradation or explosion, inerting atmospheres for curing, heat treating, and protecting food and grain from insects, rodents, and spoilage. The PMS system is ideally suited for inert gas generation in geographically remote areas. The units are reliable, they can tolerate temperature excursions over 383 K, and pressure surges up to 50 bar. Though the individual elements are small, they can be assembled into larger systems and can be made portable. Each element is expensive to replace, however, if damaged either mechanically or by the gases being processed. (Some suppliers claim that their units can surpass the temperatures and pressures cited previously.) The system consists of a compressor, an inlet air coalescer and filter, the PMS modules assembled and packaged into a complete unit, a manual flow control valve, and an oxygen analyzer with an automatic shutoff valve. The product has the same pressure as the main compressor less than the pressure drop through the system, which may be quite high. 8.3.4.3 Pressure Swing Adsorption PSA systems have been on the commercial market for over 35 years, and recent improvements in the performance of the carbon molecular sieve (CMS) at the heart of the PSA unit have improved the technology. This application is often referred to as CMS. As described in Chapter 9, the operation and process schematic is similar to that of a prepurification unit designed for air separation. Compressed atmospheric air is treated to remove dust particulates, water, oil, and carbon particles. Treated air enters a receiver, acting as a buffer, then after passing through a control valve enters a vessel containing specially treated carbon. This carbon has been imbued with the property of selectively adsorbing oxygen faster than nitrogen. The process continues until the carbon is saturated with oxygen. The air is switched over to a standby vessel also filled with a CMS, and the process continues. The switched-off bed is depressurized to release the adsorbed oxygen, and regenerated with nitrogen. It is thus prepared for the next cycle. The process is repeated in cycles of approximately 1 min. Purities of 99.9% are generally cost-effective relative to delivered liquid nitrogen and even higher purities of 99.95% have been realized in practice.


8.3.5

ULTRAHIGH-PURITY NITROGEN

During the past 10 years, the manufacturers of electronic semiconductor wafers have been demanding, imposing very stringent nitrogen purities not only for the complete removal of oxygen but also for the elimination of other trace impurities, such as carbon monoxide, hydrogen, methane, moisture, and carbon dioxide. The industrial gas companies, in turn, have developed cost-effective process technologies to meet these demands, within the limitations of analytical instruments available on the market. The suppliers, moreover, had to be concerned with any contamination arising from the cryogenic and ancillary equipment, in terms of leaks, retrodiffusion, outgassing, and minute metallic and nonmetallic particles from fittings and welds. In 1995, at least two cryogenic plants were built by Air Liquide delivering ultrahigh-purity nitrogen with an oxygen content of 1 vppb directly from the cold box without supplementary purification units. Now, clients’ specifications for oxygen range anywhere from 1 to 10 vppb depending on the specific application. For greater purities, less than 1 ppb, it is necessary to install additional purification beyond cryogenic distillation involving catalytic oxidation and adsorption. With the latter supplements, it is possible to achieve a level of ppt for contaminants. For this purpose, one must use a high-precision analytical procedure such as an atmospheric pressure ionization mass spectrometer (APIMS), or equal. There are a number of systems supplying ultrahigh-purity nitrogen: 1. Truck delivery from a central bulk plant delivering ordinary high-purity liquid nitrogen to the site where a part of the product is vaporized and passed through a supplementary purifier, converting it to ultrahigh purity. 2. Truck delivery from a central bulk plant, which does itself produce an ultrahigh-purity nitrogen. This is vaporized at the site. This system, however, can invite contamination from unnecessary extra handling and equipment. 3. On-site standard high-purity nitrogen generator, which may also include a supplementary purifier if the client requires two nitrogen products: one with a standard high purity, and the other with an ultrahigh purity. The latter option is more cost-effective, and quality control is more reliably achieved. 8.3.5.1

Removal of Outside Impurities

Design of impurity removal system depends entirely on the impurities and level of elimination desired. For example, compounds such as hydrocarbons heavier than methane, carbon dioxide, and water can be removed by adsorption in a standard prepurification system of activated alumina and molecular sieve 13. It is possible to eliminate helium and neon by a combination of liquid flash and distillation. Oxygen, and to a certain extent, hydrogen may be removed by distillation but the latter may require supplementary catalytic oxidation. It can also be removed by adsorption as it is standard in the majority of air separation plants. Carbon monoxide and methane may also be eliminated by adsorption or catalytic oxidation. One possibility is to use a catalytic converter immediately following the main air compressor before any aftercooler, taking advantage of the heat content of the process airflow. If the air compressor is highly efficient, however, one may have to add a special heater to achieve heat of reaction, especially to remove hydrogen. In such a case, it may be necessary to add a refrigeration machine to lower the temperature downstream if cooling water temperature is not sufficient. Another possibility is to include a reacting catalyst as part of the standard prepurification system. The reacting catalyst layer should be inserted between the activated alumina and the


molecular sieve. A factor to be kept in mind is that the temperature required for the combustion reaction may be much higher than that needed to regenerate the alumina and the molecular sieve. In other words, the integral process may be costly in terms of energy. Another drawback to the use of combustion catalysts, especially for the removal of hydrogen, is their sensitivity to aging and poisoning. In a recent communication received from the Degussa Company on the removal of both hydrogen and carbon monoxide, the following example is of interest. Incidentally, hopcalite, which is another name for manganese dioxide, has been found in certain conditions to promote the air oxidation of relatively low concentrations of hydrocarbons, such as acetylene, and at relatively low temperatures (Chemical Engineering Progress, 57(4) 1961). Presently there is a patented catalyst operating at close to ambient temperature to purify nitrogen. Whatever the specific requirements, the process engineer must study the optimum combination of cryogenic and, if necessary, supplementary purification technology for the final product. 8.3.5.2 Process Cycle for Ultrahigh Purity Nitrogen7 As far as the cryogenic process is concerned, the removal of impurities is limited to those that are more volatile than nitrogen, such as helium and neon, which are easy to remove. Hydrogen, although more volatile than nitrogen, shows up with traces (vppm) in the product nitrogen, and can be removed catalytically by combustion. In general, hydrocarbons can be removed by normal adsorption using molecular sieve. Methane, and to a certain extent, ethane can only be adsorbed around 90% allowing 10% to slip through, mostly with the oxygen, which in this case goes to waste. In the case of ultrahigh-purity nitrogen, however, some of the methane, and possibly ethane may contaminate the nitrogen product and have to be removed by catalytic combustion. A variety of advanced cryogenic process cycles have been developed over the past 10 years to improve the purity of product nitrogen and yield. Yield is important to improve the energy consumption. The main objective has been to increase oxygen concentration in the waste stream, improving the reboil, and waste recirculation (mostly oxygen). The most recent developments include the use of a high-yield heavy component separation process, with an added column to remove light components. The use of an additional rectification column has achieved a product nitrogen purity of 1 ppb of oxygen. In fact, one designer has achieved 1 vppb with only a single distillation column. As for the removal of any traces of hydrogen, carbon monoxide, and hydrocarbons, the use of a final catalytic purifier is inevitable. With the combination of both features, it is possible to achieve a purity down to parts per trillion (vppt). Hydrogen has been removed completely and successfully by the addition of a distillation column, but this also complicates the operation. 8.3.5.3 Outside Factors in Contamination8–23 8.3.5.3.1 Contamination Due to Outside Factors While with the advanced process features mentioned previously, one may achieve an ultrapure nitrogen, there are certain outside factors that may contaminate the final pure product: 1. Leaks and retrodiffusion due to a partial pressure diffusion from various components such as fittings, welds, valves, and flowmeters that may allow oxygen and moisture to leak into the product. To avoid these problems, the following action should be taken: .

The use of stainless steel bellow seal valves for larger diameter piping should be avoided as they are risky. They can contaminate with moisture and particles.


.

.

.

. .

. . .

Use stainless steel safety valves (type 304L or 316L) without manual actuators to minimize leakage. Use either tongue or groove flanges with a polytetrafluoroethylene gasket, or ringjoint gasket flanges. For additional safety install the flange under a commercial nitrogen seal. In piping, use a bending radius of more than 5 times the internal diameter as a minimum for cryogenic equipment within the cold box; and more than 10 times the internal diameter for bending warm piping outside the cold box, to achieve an internal smooth surface free of irregularities and particles. All temperature probes shall be installed with thermowells. Analytical sample lines should be electropolished with connections and diaphragm valves specified for ultraclean systems. Electroplated piping for product flows. Avoid dead-ends in piping, as it takes ages to purge completely. Control valves for 2 in. (50 mm) or under should be specified as stainless steel bellow seal valves, and for larger sizes use stainless steel commercial valves with a casing to ensure a nitrogen atmosphere.

2. Dead volume diffusion because of improper start-up procedures allowing moisture, oxygen, and carbon dioxide to enter the apparatus. This situation may be due to purge lines, defrost lines, analysis lines, and complicated product distribution lines. Generally, it occurs because of a prolonged start-up, resulting in leaks and mostly in retrodiffusion showing up during termination of the dead volume. To cope with this problem, one should purge the lines before start-up with a high-velocity turbulent flow, namely of at least 2000 Reynolds number, through the lines that may hold the dead volume. 3. Outgassing that comes from gases entrapped in the basic material during manufacture. This will allow hydrogen and moisture to be transferred to the product nitrogen. To overcome the problem of hydrogen contamination, a serious work was carried out by Succi16 of L’Air Liquide’s Centre de Recherche Claude Delorme in France. In regard to moisture contamination the reader is referred to the work carried out by McAndrew et al.17 also of Air Liquide. 4 Another factor is the possibility of particle contamination from improper surface treatment of generation equipment or from the shedding of nonmetallic garnitures used in fittings. This problem cannot be overcome easily because it is very difficult to measure the number of minute particles per standard cubic liter if such a guarantee is specified. At very low flows, no particles will be detected. At very high flows and at high velocities, the sample line itself may generate particles, which have to be rejected in the count. The sample line, therefore, must be electropolished and static free. The critical orifice for the sample line should be designed to reduce sample pressure to atmospheric and maintain isokinetic flow. It is necessary to choose suitable highly efficient filter, which should be installed in a temperature-controlled area. Shedding of the filters may occur but can be controlled by installing electropolished piping and diaphragm valves downstream. The advanced cryogenic process cycles employed with present-day nitrogen generators and equipped with catalytic combustion add-ons as previously described, are more than adequate to provide ultrahigh-purity nitrogen to satisfy the high-technology electronic industry with sub-ppb purities. The challenge is for the manufacturers of analytical instruments to keep up with—or to better—the new standards set by APIMS.


8.3.6 OTHER ATMOSPHERIC NITROGEN COMPOUNDS 8.3.6.1 General Diatomic nitrogen that makes up about three-quarters of the Earth’s atmosphere is a very stable molecule. The bond between its two atoms can normally be broken only at very high altitudes when exposed to ultraviolet radiation, and at or near the Earth’s surface by lightning or other high-voltage electrical discharges. Its oxides, however, can be extremely reactive and can confront the cryogenic engineer on almost a daily basis. Nitrous oxide finds wide use as a mild anesthetic, and nitric oxide in extremely minute concentrations is involved in critical metabolic reactions. Though these compounds can be beneficial and even life supporting, however, nitric oxide and other oxides of nitrogen can also pose severe danger. Nitrous oxide, for example, forms an explosive compound with acetylene. Nitric oxide and nitrogen tetroxide are aggressive and toxic gases and should also raise a red flag to the cryogenic designer. To help process and project engineers a brief description of the several nitrogen oxides are described in the following sections. 8.3.6.2 Dinitrogen Monoxide (N2O) or Laughing Gas The English chemist, Joseph Priestley, discovered dinitrogen monoxide in 1772, and another English chemist, Humphrey Davy, worked on its physical, chemical, and physiological properties. Davy named it nitrous oxide, and indicated that when the gas was inhaled, it produced laughter, almost to hysteria, then insensibility. When mixed with either air or oxygen, it has become an important anesthetic in surgical operations of a short duration, for example, in dentistry. Another popular and extensive application of nitrous oxide is its use as a propellant in pressurized canisters that hold anything from aerosol disinfectants to whipped cream. In this application, the gas is dissolved under pressure in the substance within the canister. When the pressure is released it expels and expands the contents. Nitrous oxide is produced by the decomposition and heating of ammonium nitrate: NH4 NO3 ¼ 2H2 O þ N2 O Solid

Water Gas

(8:10)

The gas produced is compressed into high-pressure cylinders to be delivered to users. The physical properties of dinitrogen monoxide are given in Table 8.10. 8.3.6.3 Applications of Nitrous Oxide Aside from its well-known application as an anesthetic for minor operations it is also used as an aerosol propellent for pharmaceuticals, for food and beverages. It is also used as a comburant for flame in atomic adsorption spectrophotometry, in the calibration of gas mixtures for petrochemical industry, environmental emission monitors, and trace impurity analyzers. It is also used as a source of oxygen in electronic manufacturing for chemical vapor deposition of silicon oxynitride or silicon dioxide. 8.3.6.4 Dangerous Side of Nitrous Oxide In the 1950s, it was reported that an explosion took place in the main condenser of an air separation plant24. After an exhaustive investigation the final report indicated that it may have been caused by a chemical combination of traces of nitrous oxide and acetylene in the atmosphere. In 1955, moreover, it was found that nitrogen dioxide, sometimes called nitrogen


TABLE 8.10 Properties of Dinitrogen Monoxide (Nitrous Oxide) Molecular weight Normal boiling point Vapor pressure Latent heat of vaporization Thermal conductivity of gas Gaseous density Liquid density Critical temperature Critical pressure Compressibility factor, Z Specific gravity Specific volume Heat capacity, Cp Heat capacity, Cv Ratio of g Viscosity, m

44.013 kg=kmol 184.65 K at 1.013 bar 58.5 bar at 293 K 376.14 kJ=kg at boiling point and 1.013 bar 14.57 mW=(m K) at 1.013 bar and boiling point 3.16 kg=m3 at 1.013 bar and 273.15 K 1.222. kg=L at 1.013 bar and boiling point 309.55 K 72.45 bar 0.9939 at 1.013 bar and 288 K 1.53 at 1013 bar and 294 K 0.543 m3=kg at 1.013 bar and 294 K 0.038 kJ=(mol K) at1.013 bar and 288 K 0.029 kJ=(mol K) at 1.013 bar and 288 K 1.3022256 at 1.013 bar and 288 K 136 micropoise

Source: Courtesy of F.G. Kerry, Inc. With permission.

peroxide, is a potent catalyst in the reaction of carbon monoxide mixtures, even in small quantities. It produces a reaction as the partial pressure increases from 0 to 0.3 mm and upward it produces a reaction when the temperature reaches 5378C, which is well below the normal ignition point of carbon monoxide25. For this reason engineers should take good care to investigate the complete analysis of the ambient air especially if the location of the plant is near a chemical plant or refinery.

8.4 CARBON DIOXIDE 8.4.1 GENERAL CHARACTERISTICS Carbon dioxide is a colorless, odorless gas, 1.5 times as dense as air. It is not classified as toxic but a large concentration will result in suffocation because it will displace oxygen, which is necessary for human survival. Whereas carbon dioxide may be an annoyance in cryogenics for the process designer, it is a useful and profitable industrial substance both in its gaseous and its solid phase. Its chemical and physical characteristics are given in Table 8.11. In solution with water it forms carbonic acid which is slightly acidic (H2CO3), and is used extensively for carbonated drinks. In its gaseous phase it is used for fire extinguishers, and also for enhanced oil recovery (EOR). In its solid phase, called dry ice, it is used as a refrigerant for food products. Atmospheric carbon dioxide is in dynamic equilibrium with that dissolved in water, and with that bound as carbonate in the Earth’s crust. A very important property of carbon dioxide is that with sunlight and chlorophyll serving as an unconsumed catalyst, green plants convert carbon dioxide and water into sugar (starch), and more importantly, oxygen. The reaction is called photosynthesis, as it uses the energy of light. In fact, the thick forests of the Amazon jungle supply approximately 40% of the Earth’s oxygen. The simplified overall equation for this reaction is as follows: 6CO2 þ 6H2 O ¼ C6 H12 O6 (starch) þ 6O2


(8:11)

TABLE 8.11 Properties of Carbon Dioxide Molecular weight Normal boiling point Vapor pressure Latent heat of vaporization Molar specific heat capacity of gas at constant pressure in the range 0–1 bar and 298 K Thermal conductivity of gas Gaseous real density Liquid density Critical temperature Critical pressure Flammability

44.01 kg=kmol 194.7 K it sublimes 5180 mbar at 216.6 K 25.3 MJ=kmol at boiling point and 101.321 kPa (sublimes) 36.13 kJ=(kmol K) 0.015 W=(m K) at 273.15 K 1.977 kg=m3 at 273.15 K and 101.325 kPa 1.18 kg=L at 216.6 K triple point 304.2 K 73.70 bar Inert


Conversely, carbon dioxide is returned to the air by the respiration of plants and animals.

8.4.2 SOURCES

OF

CARBON DIOXIDE

There are a variety of sources of carbon dioxide, and among them the following are the important: .

. . .

Combustion of hydrocarbons with air or oxygen, or with a combination of oxygen and steam, such as SR, POX of hydrocarbons Fermentation of grain for the production of beer, or ethanol for spirits Off-gases from petroleum refineries, oxidation of ethylene, and automotive combustion Natural gas wells

8.4.3 RECOVERY PROCESSES

FOR

CARBON DIOXIDE

Though there are many sources of carbon dioxide, the choice of recovery process depends on desired product purity. If the ultimate use is for EOR, or for use in fire extinguishers, a product purity of 95% may be sufficient. On the other hand if the use imposes a food-grade level, then the recovery of carbon dioxide needs to be more thorough in order to strip all components, which may be harmful to human ingestion. 8.4.3.1 Food Grade Recovery from Petroleum Off-Gases The process cycle used to treat this feed gas is very close to the one proposed for the recovery from the fermentation of grain; but one must keep in mind that the former raw gas may have certain contaminants such as hydrogen sulfide, other sulfur compounds, or other substances, which must be removed beforehand. In one specific case where the feed gas volume required an oil lubricated compressor, it was necessary to install a system of coalescers to remove the oil contamination completely. In this case, moreover, it was also necessary to install a packaged refrigeration machine using ammonia as the refrigerant for the reflux in the distillation column. The feed gas and process cycle for this application is described in Figure 8.6.


Tail gas Recycle gas Carbon bed for H2S removal

Main compressor Precooler

Predryer

Distillation column

Mechanical refrigeration

Separator Condenser

Feed Aftercooler Separator

Storage Recycle cooler Drain

FIGURE 8.6 CO2 recovery from refinery off-gas. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

8.4.3.1.1 Example for CO2 Recovery Referring to Figure 8.6, the feed gas is passed through a water precooler, then a separator, and enters the main compressor where it is boosted to 19 barA. After compression the feed passes through the aftercooler and separator, and then enters the adsorption bed filled with activated carbon for the removal of any sulfides, which may have entered with the feed. The gas then enters the twin molecular sieve beds charged with molecular sieve Grade 4A for the complete removal of water down to less than 1 vppm. Once purified in terms of water and trace sulfide contaminants, the feed enters the cryogenic system, which is supplied with an external single-cycle refrigeration with ammonia operating at 236 K (378C). The pure product is then withdrawn from the bottom of the distillation column and transferred to the storage tank. A bypass is also supplied to take care of any off-specification product, which may have been produced by operational error. Off-specification product is transferred to a separate tank. Residual gas is then sent to a common flare where combustion is maintained by the various hydrocarbons also rejected by the refinery. The molecular sieve dryer is regenerated by a portion of the product gas, which is heated and sent through the depressurized vessel for regeneration. The effluent of the gas used for regeneration is then sent to the common flare of the refinery. Meanwhile, a portion of the dry effluent is sent to a cooler and recycled to the main feed compressor to lower the service capacity of the dryer. Liquid product carbon dioxide is stored in a vessel at a pressure of approximately 15 barA and at a temperature of 248 K. 8.4.3.2

Food Grade Recovery from a Fermentation Source

The process cycle for recovering food grade carbon dioxide from a fermentation source is identical to the previously described process design, with the exception that zinc oxide and activated charcoal beds are employed to purify the gases generated by fermentation. 8.4.3.3

Nonfood Grade Carbon Dioxide

For industrial applications other than food, such as EOR, a product purity of 98% to 99% is generally sufficient. Because of its inertness and molar mass (44.01 kg=m3) carbon dioxide has


found wide use in stimulating petroleum production. For small applications, liquid CO2 is trucked in, then the liquid is vaporized to the required pressure. For very large volumes, carbon dioxide is generated at the site with the combustion of carbonaceous material, purified to the correct purity, and compressed to the desired pressure.

8.4.4 DRY ICE: FOOD GRADE Carbon dioxide, below its triple point at 4.165 bar, sublimes to a vapor instead of melting. This property makes the use of carbon dioxide very advantageous in refrigeration because it avoids wet surfaces and the necessity of drainage. Moreover, dry ice at a pressure of 101.3 kPa has an equilibrium temperature of 194.26 K, which makes its use very favorable for cooling down perishable food. Its use should be avoided, however, if any of the products are damaged by freezing. Its application is in general use for cooling, freezing, and shipping, whenever mechanical refrigeration is unavailable or too expensive.

8.4.4.1 PRODUCTION

OF

DRY ICE

The production of dry ice involves a three-stage compression to a level of 62–63 barG. During compression, the CO2 is purified after the first stage. Compression is followed by a three-stage isenthalpic expansion. The expansion produces a liquid in the final liquid receiver just above the triple point. This liquid is finally expanded into a snow press; expansion begins below the triple point so carbon dioxide turns into snow. Toward the end of the expansion cycle, pressure in the press is allowed to rise above the triple point. At this pressure the liquid carbon dioxide turns into snow. As soon as the right quantity of snow forms, the expansion valve is closed and the hydraulic press is activated. The wet snow solidifies into a solid block, which is stored, then cut into the desired commercial sizes. The actual process of turning carbon dioxide into dry ice has approximately a 50% efficiency, but gaseous product is recycled into the compression system.

8.4.5 APPLICATIONS

OF

CARBON DIOXIDE

One of the largest and oldest applications of gaseous carbon dioxide is for the carbonation of beverages. A food grade purity is required for this purpose, ranging from 99.995% to 99.9995%. This practice is also used for fruit juices to enhance customer appeal. In the chemical industry, the off-gas of both SR and POX is mostly carbon dioxide, which is often wasted. It is more profitable, however, to react it with ammonia to form urea, an important fertilizer. Urea is helpful both at a small scale for the normal home gardener, as well as at a much bigger scale for huge farms. The reaction is as follows: 2NH3 þ CO2 ¼ NH2 CONH2 (Urea) þ H2 O

(8:12)

Carbon dioxide is also used in gas-cooled nuclear reactors. Several hundred tonnes of CO2 are needed per reactor. It is also used in portable fire extinguishers, and for pressurizing aerosols, replacing the now nonacceptable Freons. It is also well known in the welding industry for the inert-arc welding of normal carbon steels, but is not recommended for the welding of special alloys or aluminum where the inert gas used is argon or helium. CO2 is also used as a carrier for the spraying of paints and vegetable oils in cooking. In the latter case it has to be of food grade. Used as an industrial carrier and dispersant, CO2 has almost taken over the field from halogen aerosols, which have been banned as a cause of breakdown of the Earth’s invaluable ozone layer. Dry ice is also used as a mechanical cleanser. For this application the snow produced by Joule–Thompson expansion is extruded through a plate to form hard round pellets at 194 K.


When the pellets are blasted against a contaminated surface by compressed air, they achieve a velocity of around 300 m=s, shock-freezing the surface in a fraction of a second. Their kinetic energy punctures the contamination on the surface and removes the dirt. Yet, because the Mohs hardness of the pellets is only 2, such cleaning is nonabrasive and surface quality is maintained. The pellets themselves sublime, leaving no residue. This novel process finds applications as diverse as cleaning forms and molds without disassembly, conveyor belts while still in operation, cleaning motor parts, gears, transport containers and house facades, and removing paint.

8.5 8.5.1

OZONE GENERAL

Human beings have always assumed that the air which composes the Earth’s atmosphere is an inexhaustible supply in terms of quantity and quality. Modern science has taught us, however, that over the centuries, its quality has been subject to human and industrial intrusiveness. If animal and vegetable life is to exist on Earth, the atmosphere must maintain a proper balance among its components of oxygen, nitrogen, carbon dioxide, water vapor, and its other constituents, especially ozone. A layer of ozone must be maintained in the upper stratosphere in the areas around the north and south poles, to screen out the harmful effects of the ultraviolet light of the sun’s rays. In addition, the accumulation of toxic materials in the air from industrial off-gases, which may destroy the ozone, must be kept to a practical minimum.

8.5.2

PROPERTIES

OF

OZONE

The characteristic odor of ozone associated with electrical machinery had been reported as early as the latter part of the eighteenth century, but the gas was studied more intensively in 1840 by C.F. Schoenbein who recognized it as a new gas, and named it ozone. This same gas was also produced at the anode during the electrolysis of water. Ozone is a triatomic allotrope of oxygen. In other words, it has three atoms instead of the normal two of the molecular oxygen. Allotropy is the existence of an element in two forms. Nevertheless, ozone is considered as elementary oxygen and not a compound. Ozone (its name is from the Greek ozein to smell) is an irritating, pale blue gas, and explosive at high concentrations and highly toxic even at low concentrations. It occurs naturally in small amounts in the Earth’s stratosphere (10–50 km above sea level) where, as noted, it absorbs solar ultraviolet radiation, which otherwise could cause severe damage to living organisms on Earth’s surface. In certain conditions, photochemical reactions between nitrogen oxides and hydrocarbons in the lower atmosphere can produce ozone in concentrations high enough to cause irritation of the eyes and mucous membranes. The result of such reactions is commonly known as smog. Since ozone has three oxygen atoms, its molecular mass is 1.5 times that of oxygen. It condenses at 161.15 K as a dark blue liquid, which freezes at 21.75 K. Ozone gas decomposes rapidly at temperatures above 373 K, and in the presence of certain catalysts at ambient temperatures. Ozone contains more energy than oxygen, which evolved when 48 kg of ozone decomposes to oxygen is 375 kWh. This is the equivalent energy that must be supplied by an electric discharge when ozone is formed. Because of its greater potential energy, ozone is more reactive than oxygen. It converts mercury and silver into oxides, and readily allows iodine from potassium iodide, whereas oxygen does not cause these reactions at room temperature. Ozone also converts olefins into aldehydes, ketones, or carboxylic acids.


8.5.3 ATMOSPHERIC OZONE LAYER There is a region in the Earth’s atmosphere about 15–48 km in altitude, which contains very small quantities of ozone. This is called ozonosphere. Nevertheless, even this small quantity is sufficient to absorb most of the sun’s ultraviolet light, which would cause skin cancer to humans and damage other organisms. In 1970, Crutzen of Holland warned that the decomposition of the ozone layer by nitrogen oxides such as NO and NO2 acting as catalysts would prove harmful to human beings in the long run. Then in 1974, Rowland and Molina of United States published their report that the popular use of chlorofluorocarbon (CFC) gases were transported to the ozone layer, and with the influence of ultraviolet light destroyed ozone molecules. The work carried out by Crutzen, Rowland, Molina, and other concerned scientists led to the Montreal Protocol in 1987 in which industrialized countries agreed to phase out the production of CFCs. Halons (chlorofluorobromine-compounds) can also destroy ozone if diffused into the ozonosphere.

8.5.4 GENERATION

OF

OZONE

Ozone is generally produced by passing oxygen or dry air through an electric discharge. The resulting product is a mixture of ozone and original gases. Recovery increases if oxygen instead of air is used. This procedure is suitable for most industrial applications, although purer ozone can be produced by a variety of methods. For example, ozone can be produced by using a cryogenic cycle resulting in a liquid oxygen–ozone mixture, which will separate into two layers. The denser one will contain around 75% ozone. But due to the extreme instability and high reactivity of concentrated ozone, this procedure is very hazardous and is not recommended.

8.5.5 APPLICATIONS Being a strong oxidizing agent, ozone has the power of converting many colored organic substances to colorless products. It finds use as a bleaching agent for oils, waxes, starch, and flour. It is a strong germicide and has also been used instead of chlorine for sterilizing drinking water. The largest water purification project in the world using ozone is now being built in Dallas, Texas. It will treat approximately 1,703,250 m3=d of potable water. The ozone generation system will include nine individual ozone generators, each with a production capacity of 76 kg=h. The total system capacity of 684 kg=h will be produced at high concentrations of more than 10% by weight using oxygen in place of air to maximize efficiency. Because ozone can selectively oxidize NOx to higher oxides of nitrogen, which are readily water-soluble and can be removed in a wet scrubber, it also finds application in pickling stainless steels.

8.6 METHANE The gaseous production of methane in global terms has been going on for millions of years due to decaying plants, and more recently in the Earth’s history from sewage, silage on farms, and landfills. Recent studies have shown that a significant portion of methane present in the atmosphere is generated in the rumen of cattle and similar beasts and expelled by belching. Coal, petroleum, and associated gases have been produced in rich sediments over millions of years. These natural gas products include methane, ethane, propane, and heavier hydrocarbons as well as carbon dioxide, some quantities of nitrogen, and even helium. As methane is the most abundant product worldwide its study merits priority. Table 8.12 lists the physical properties of methane.


TABLE 8.12 Physical Properties of Methane Feed gas mol % Hydrogen Nitrogen Carbon monoxide Argon Oxygen Methane Carbon dioxide Water Pressure Temperature

3.354 0.191 0.191 0.191 0.191 0.192 91.489 4.201 19 barA 313 K

Product gas Carbon dioxide N2=O2=Ar CH4 H2O

Mol % 99.9995% Less than 1 vppm Less than 3 vppm Less than 1 vppm

Source: Courtesy of F.G. Kerry, Inc., Report, 1980. With permission.

8.6.1

PROPERTIES

OF

METHANE

Methane, or as it is more commonly known as natural gas, is a fascinating substance that is ideal for cooking, space heating, and industrial applications. It can be thermally dissociated by SR or POX into synthesis gas, hydrogen, and carbon monoxide, which in turn is converted into useful products such as ammonia, urea, and oxo chemicals. The main problem with the natural gas, however, is that the large known reserves in the world are often faraway from industrial consumers eager to use the gas either for energy generation or for the production of profitable chemicals. Whenever natural gas reservoirs were found in the same landmass they were quickly connected to consumer centers by large diameter pipelines to supply ever-increasing demands. This practice is still taking place in North America, South America, and Europe, both east and west. In other cases where industrial development cannot be connected directly to a gas field by pipeline, enterprising consortiums have been formed to bridge the maritime gap between supplier and ultimate consumer. As it is explained in more detail later, the latter procedure include the liquefaction of the natural gas, the extraction of heavy components such as propane and butane, and the maritime shipment of the supercooled gas to distant ports. The tankers used for this purpose have been designed not unlike giant thermos bottles and use the in transit vaporized gas as fuel.

8.6.2

HIGH-PURITY METHANE FOR CHEMICALS

The application of cryogenic systems includes the extraction and purification of methane from natural gas pipelines. In this application methane is completely stripped of nitrogen, ethane, and propane. Higher hydrocarbon impurities can be held to less than 25 vppm. This very high-purity methane is used for the manufacture of chlorinated hydrocarbon products. Low-temperature separation is integrated perfectly with methane production equipment to produce exacting purity specifications. For example, the direct chlorination of methane has replaced older methods for the production of many chlorinated methane products, such as


Warm exchangers

Methane warm exchangers

Cold box limit Oil adsorber Cooler Separator Typical high purity methane units Methane recycle compressor

FIGURE 8.7 High purity methane. (Courtesy of F.G. Kerry, Inc. With permission.)


Subcooler

Pure methane column

carbon tetrachloride, chloroform, methyl chloride, methylene chloride, and perchloroethylene. Although the use of some of these products may be limited by environmental laws, their chemistry and manufacture are of much interest. Each of these products acts as a starting point for the production of other products, which give rise to a variety of small chemical industries. In chlorination equipment methane is combined with chlorine under controlled conditions in a reactor to simultaneously produce chlorinated products together with by-product hydrogen chloride. Direct ratios of the specific products may be adjusted by changing reactor conditions. Condensables are then separated from hydrogen chloride with subsequent separation of chlorinated products by distillation. By-product hydrogen chloride is absorbed in water for use or sale as hydrochloric acid. Trace impurities present in methane before the chlorination stage are themselves chlorinated, however, and pose a problem. The difficulty lies in the fact that the boiling points of the chlorinated impurities lie so close to the boiling points of the desired products that purification by distillation is not practical, and other means such as prepurification by molecular sieve adsorption have to be employed. Before the feed gas reaches the low-temperature separation unit impurities such as oil, carbon dioxide, hydrogen sulfide, and water are completely removed (see Figure 8.7). In the low-temperature section, the final high-purity methane product is obtained by cooling the natural gas in a heat exchanger system, and by rectifying it in a distillation

column. The high-purity methane reflux essential for this column is obtained by recompressing some of the product methane and sending it, after oil removal if necessary, to the cold box where it is cooled and condensed in methane exchangers. It is then expanded by a valve into the distillation column, serving as a reflux liquid. This same stream acts as a closed refrigeration cycle for the entire process in the cold box, overcoming the heat gain into the box, and the energy loss due to the small positive temperature approach at the warm exchangers. An independent cycle such as this provides a large degree of freedom for operating at different capacities and with different feed gases. The tower bottoms are expanded to approximately 3.45 barG, warmed to ambient temperature in the exchanger system, and removed from the cold box as a source of fuel gas.

8.6.3 NATURAL GAS PEAK LOAD SHAVING In local pipeline distribution of natural gas, the problem of high peak demands daily or seasonally poses serious problem to distribution companies. They have, therefore, employed liquefaction and storage of a precalculated portion of the liquefied gas to take care of this problem. As each case is different in terms of storage capacity, there is no standard cryogenic solution to be used. For example, if the storage capacity is approximately 16 MM Nm3 to take care of a liquefaction rate of around 67,000 Nm3=d, and the line pressure is about 50 bar, a modified cascade cycle using a single multicomponent refrigerant may be a good choice. An expander system may not prove very efficient, and the standard cascade cycle is too complex for such a relatively small plant, particularly if the feedstock is lean in terms of ethane and propane (Figure 8.8 and Figure 8.9). If the natural gas line includes sufficient ethane for recycling, and if propane can be purchased locally at a low price, then the choice of a classical cascade process cycle may become economically viable (see Figure 8.9). The main consideration of this latter process is to make sure that the enthalpy–temperature curve of the mixture evaporating at the low-pressure level of the cycle over the entire temperature range (311–105 K) will approach as closely as possible, but not touch the enthalpy–temperature curve for the natural gas and recycle gas mixture, which is condensing and subcooled at the design liquefaction pressure. A good design may include a temperature differential of 1–5 K. In order to hold the expected entropy increase to a minimum, the produced liquids should be subcooled before expansion in a throttle valve. As to impurities to be found in the gas stream, the use of a molecular sieve absorber is strongly ˚. recommended, preferably one with a pore size of about 4 A

8.6.4 BASE LOAD NATURAL GAS PLANTS (LNG) This application merits special treatment by itself because it involves liquefaction and transportation of natural gas by maritime tanker, barge, trailer truck, and ultimately local distribution by pipeline. Millions of cubic meters are liquefied each day by using a variety of process cycles, but in general the choice includes a modified cascade system with mixed refrigerants involving propane, ethane, methane, and sometimes nitrogen. A variety of cascade process cycles exist each with its own distinct character depending on the engineering company’s ingenuity in design. A short description of a few designs may be of interest. Whatever the case, they all use similar basic principles of cascade refrigeration, condensation, and vaporization. 8.6.4.1

Cascade Cycle with Mixed Refrigerants in General

The cascade cycle, which has been explained in Chapter 2, sounds simple enough at first glance, but to obtain optimum process efficiency one must use a minimum of two, three, and even four


High pressure pipeline

Low pressure pipeline

Precooling Vaporizer fuel

Peak shaving gas

Purification Vaporizer

Turbo compressor

Heat exchangers

Joule–Thomson valve

Turbine expander

Liquid separator

Joule–Thomson valve

Flash compressor

LNG storage

LNG pump

FIGURE 8.8 Simple expander cycle. (Courtesy of F.G. Kerry, Inc. With permission.)

stages for each refrigerant to achieve thermodynamic reversibility for minimum power consumption (Figure 8.10). In actual practice, the classical cascade cycle employing propane, ethylene, and methane in closed independent loops can be ruled out because of several factors: external and costly storage facilities; a variety of external compressors, expensive in investment, maintenance and control; costly piping systems; and finally operational problems of control with changes in process gas composition and cooling water temperature. The first modification examined by process engineers was to use ethane instead of ethylene, as there was not much to lose in terms of refrigeration potential, and was readily available as a component in natural gas. A dual process cycle was then studied in detail. The proposed design would consist of a standard propane cycle for precooling purposes; to be followed by a separate refrigeration cycle consisting of a well-balanced binary ethane– methane mixture. Though this concept was an improvement over the standard cascade cycle, a problem was encountered because the natural gas product would leave the lowest refrigeration level with inadequate subcooling, resulting in a substantial flash loss. To overcome this problem, moreover, it would be necessary to add another compressor to reliquefy the unused portion with fresh feed. The next concept was to design an efficient mixed refrigerant process, which would employ a single process compressor to handle all the refrigerants in the same frame and driver, with the possibility of employing a single as well as more efficient axial compressor.


Pipeline Peak shaving gas

Precooling Vaporizer fuel

Purification Heat exchangers

Cooler

Vaporizer

Feed compressor

Recycle compressor Recycle gas

Expander and generator

Flash compressor

Liquid separator Joule– Thomson valve Joule–Thomson valve

LNG storage

LNG pump

FIGURE 8.9 Recycle expander cycle. (Courtesy of F.G. Kerry, Inc. With permission.)

Pipeline

C3

Vaporizer fuel

C2 C1

Vaporizer Refrigerant

Coolers

Compressors

Compressor fuel

Flash gas

Purification

Natural gas booster compressor

C3 vaporizer

C2 vaporizer

C1 vaporizer

Flash drum

LNG storage

FIGURE 8.10 Classical cascade cycle. (Courtesy of F.G. Kerry, Inc. With permission.)


LNG pump

Peak shaving gas

8.6.4.2 ARC Process Cycle26 The first liquefied natural gas (LNG) base load plant to use a modified cascade cycle began operation as a pilot plant in 1964 in Nantes, France designed by L’Air Liquide in cooperation with Gas de France. It treated a small quantity of natural gas from a field located nearby (Figure 8.11). The liquefaction unit was designed to treat approximately 29,000 m3=d (1106 cfd). The cycle involved a continuous refrigeration transfer from cooling water to decreasing temperature levels using a single mixture refrigerant with a composition such that partial condensation at any one stage corresponded exactly to the refrigeration requirement of the next stage. The partial condensation of the refrigerant took place at one pressure level, and corresponded to the high pressure of the process cycle, whereas the vaporization occurring at a single low pressure corresponded to the cycle compressor suction pressure. This cycle was named auto-refrigeration cycle (ARC). Figure 8.12 and Figure 8.13 show the difference in the work required between a standard cascade cycle and the mixed refrigerated cycle. This was the first industrially cost-effective process cycle, which opened the door to the maritime transportation of natural gas from one continent to another. The actual application of the process showed that the natural gas at a pressure of 41 bar was first treated for the removal of H2S, CO2, and water, then passed through a first exchanger section at a temperature of 311 K. In a second exchanger section, the process stream was further cooled to 214 K where the liquid phase containing heavy hydrocarbons, which could result in plugging at a lower temperature, was removed in a specially designed separator. The gaseous phase from the separator entered another section of the combined group of heat exchangers, that is, cooled by countercurrent heat exchange with the cycle refrigeration gas. The process stream, now at 110 K, condensed and sufficiently subcooled after expansion, was sent to storage with almost no flash occurring. The cooling cycle stream containing a mixture of nitrogen and hydrocarbons supplied the refrigeration at the necessary temperature levels required in the modified cascade. The cycle gas left the discharge of the compressor at 308 K, was taken down to a second level at 205 K, to a third level of 169 K, and finally to the final level of 107 K. The optimum value of temperature difference is obtained by the economic balance between the heat exchange surface and the operating energy. Pipeline C1 + C2 + C3 Refrigerant compressor

Vaporizer fuel

Cooler

Vaporizer

C3 Rich separator

Purification

Natural gas booster compressor

C3 vaporizer

C2 separator

Compressor Flash gas fuel C2 Vaporizer

C1 vaporizer

Flash LNG drum storage

FIGURE 8.11 Autorefrigerated cycle. (Courtesy of F.G. Kerry, Inc. With permission.)


LNG pump

Peak shaving gas

Ambient temperature variation

Enthalpy (W/h)

B

A

105

Temperature (K)

311

FIGURE 8.12 Work of liquefaction: ARC cycle. (Courtesy of F.G. Kerry, Inc. With permission.)

If one adds an appropriate quantity of nitrogen to a mixed refrigerant cycle it is possible to subcool the LNG product, eliminating any handling problem associated with its use as an automotive fuel. If LNG is to be used for distribution to factories and homes; however, where local regulations impose a minimum calorific content per unit volume, any excess nitrogen in the gas may have to be ejected. The principle advantage of a modified cascade cycle is the possible use of a single axial radial compressor withimportant savingsin investment, operation, and maintenance. The advantage of the ARC cycle is that the total quantity of refrigeration is always available in different sections of the plant. Any variation in composition or pressure of the natural gas may be compensated by modifying the composition of the refrigerant cycle. This procedure results in a redistribution of available refrigeration to the various sections of the plant. A major industrial application of this cycle was engineered and constructed at Arzew, Algeria by Technip and is still in operation. (At a later date a further modification was made to the process to recover helium.) 8.6.4.3

Further Development of Mixed Refrigerant Cycles27–30

The various modified cascade cycles proposed and applied later on included a precooling propane cycle for desuperheating and partial condensation of the mixed refrigeration cycle. This was followed by the main mixed refrigerant cycle for natural gas precooling and


Ethylene Methane

Propane cycle

cycle

cycle

W/h

A

105

Temperature (K)

311

FIGURE 8.13 Work of liquefaction: standard cascade cycle. (Courtesy of F.G. Kerry, Inc. With permission.)

condensation. The precooling cycle consists mainly of propane and some ethane is compressed, sent to the first section of the heat exchanger, and expanded in a throttle valve where it precools the compressed liquefaction cycle consisting of ethane–methane–nitrogen. The precooling refrigerant is collected at the bottom of the section and recycled into the first stage of compression. The precooled liquefaction cycle leaving the first heat exchanger enters a flash separator. The bottom liquid enters the main heat exchanger and is used to cool the incoming natural gas. The gaseous phase leaving the flash separator also enters the main heat exchanger, but is throttled at the top of the main heat exchanger where it condenses the main process stream. The latter stream leaves the main heat exchanger and is pumped into storage after it passes through a holding drum allowing any flash gas to leave. After the first precooling of the main process stream, heavies such as propane, butane, other heavies and makeup are collected, fractionated, and sent to the refinery for reuse or sale. Large spiral-wound heat exchangers are used for the liquefaction section. Methane, ethane, and propane make-up requirements are supplied from the natural gas by the light ends fractionation system. 8.6.4.4 Heat Exchangers The main point of any LNG liquefaction process is the design and fabrication of the large heat exchangers capable of transferring heat in the most efficient manner with small temperature


differentials. First the classical system was used to fabricate spiral-wound exchangers horizontally using copper coils. It soon became obvious that the weight of the final product caused large bending moments, resulting in high deflections and difficulties in handling. To alleviate the situation the use of aluminum coils replaced copper, and the coiling was executed along a vertical axis. With the recent development of large brazed aluminum plate fin heat exchangers that can withstand pressures up to 70 bar, one cannot arbitrarily state that spiral-wound heat exchangers are advantageous for all LNG projects. An intensive economic and technical review has to be made for each project before making any definitive decision. 8.6.4.5

Propane Refrigeration System

The propane refrigeration system precools the main process stream upstream of the prepurification section, and is sometimes used to precool the mixed refrigerant stream. It is essentially a standard refrigerator unit consisting of one multistage compressor case and a driver. This system includes a normal desuperheater, propane condenser, accumulator, and suction drums that supply propane to the high-, medium-, and low-level propane evaporators. The propane system supplies the precooling stage as in the Linde and Claude process cycles outlined for air separation in Chapter 2.

8.6.5

PRITKO PROCESS CYCLE

The Pritko process cycle was designed and realized by Pritchard Engineering and Kobe Steel. The process is very similar to that described for the mixed refrigeration cycle but it has a fixed composition stream, which in turn depends on the analysis of the natural gas. Another difference is that the mechanical design makes extensive use of brazed aluminum heat exchanger (BAHX) and fin heat exchanger in place of spiral-wound heat exchangers. 8.6.5.1

General Process Cycle

Feed gas first passes through a separator and then enters a dryer for dehydration. The dry feed gas then enters directly into the primary heat exchanger system where it is cooled down to liquefaction and passed into storage. The refrigeration stream uses the same principle as the mixed refrigerant cycle in that the high-pressure separator at the discharge of the compressor supplies a vapor- and liquid-phase refrigeration streams to the main exchanger systems. Then these streams are combined and returned to the exchanger system for high-level refrigeration. The composition of the refrigeration stream is adjusted externally by the addition of butane, propane, ethane, and nitrogen. This flexibility results in an easier control of plant operation. From a process point of view, the Pritko process shows a general simplicity in operability, but one has to remember that the design of BAHX is very critical for a two-phase flow.

8.6.6

FINAL PRODUCT PURIFICATION

When natural gas product is delivered to consumers, homes, or factories, the calorific content of the gas must conform to a minimum calorific standard set by federal or state law. If diluents such as nitrogen and helium are present they may have to be removed to raise the calorific level to the minimum standard. 8.6.6.1

Nitrogen Rejection

As a rule, no more than 10%–20% of the nitrogen in the main feed can be removed by any classical or modified cascade process cycle. To increase the quantity of nitrogen to be rejected,


it is necessary to use a fractionation column at the final phase of the process. This application will increase the investment, but the extra cost is minimal compared with the overall cost of the plant and product profitability. With this method, it is possible to remove over 80% of the nitrogen in the feed gas. Nitrogen rejection, moreover, will improve the calorific value of the extracted liquids such as butane and propane, also that of the residual gas, which can then be used for firing the plant boilers without start-up problems. 8.6.6.2 Helium Recovery Another important application of cryogenic liquefaction on natural gas is the possible recovery of helium. This application is economically viable even if the gas contains as little as 0.4%–2.0% by volume. The feed gas is compressed, dried, and cooled in a series of cryogenic stages until all the components, except helium and nitrogen are condensed and removed. The final overhead product is further condensed and separated into a crude helium containing about 80% helium and a nitrogen stream containing 0.5% t or less of methane. A recovery of 99% of helium in the feed is possible economically by this process. Helium recovery is accomplished while maintaining the process stream at high pressure, therefore minimizing the cost of recompressing the tail gas.

8.6.7 NATURAL GAS PREPURIFICATION Before any natural gas is liquefied using a cryogenic process, one must analyze it for impurities and remove them to avoid operational problems. 8.6.7.1 Acid Gases (CO2, H2S) Carbon dioxide and hydrogen sulfide can be removed by amine treating, hot carbonate scrubbing, sulfanol treating, water scrubbing, etc. If the carbon dioxide content is less than 1.5% consideration of adsorption with a molecular sieve may be preferable as well as more reliable. 8.6.7.2 Water Removal (2H2O) The process selected for water removal is generally based on economics, but since plant downtime for defrosting is costly, operating reliability should override investment consideration. Glycol systems have proven troublesome if operating temperatures are below 240 K, and methanol systems for low temperatures will result in expensive losses. They are also inadequate for nitrogen rejection plants. A fixed bed desiccant (alumina) dryer system will provide a very dry gas (down to 1 vppm) eliminating operating problems due to hydrate formation. Drying temperature should be slightly above 330 K to prevent formation of hydrates. 8.6.7.3 Mercury Contamination Many natural gas streams have been found to contain metallic mercury up to 180 mg per normal cubic meter. Analysis should be made by the use of a dual-beam ultraviolet photometer. Below 1 mg, however, there is very little effect from corrosion on aluminum heat exchangers. The removal of mercury can be effected by a molecular sieve adsorbent, but a study should be made for any possible long-range degradation of the aluminum exchangers.


8.6.7.4

Mercaptans

Mercaptans can be removed quite easily with the use of UOP molecular sieve type A. 8.6.7.5

Butane

A gas containing 1.7% C4þ will form solids at 144 K and at 8 barA. But a gas containing C4þ and heavier constituents will form solids at a much lower temperature. If solids do form they may cause plugging in pumps operating at 111 K for LNG storage. 8.6.7.6

Propane and Ethane

Propane and ethane are both soluble in all proportions in liquid methane at 90 K. Only the heavier hydrocarbons and carbon dioxide are of primary concern.

8.6.8

ECONOMICS31

The various energy users may be classified as follows: 1. The natural gas system including sweetening, dehydration, heavy product extraction, shipment to storage, and boil-off recovery. 2. Liquefaction process train including heat exchangers, refrigeration compressors, refrigerant coolers, and condensers. 3. Mechanical energy and utility production including drivers, cooling system, and generation of utilities to be used. 4. Since the liquefaction plant per se consumes slightly more than 50% of the total energy it merits a complete energy analysis including the thermal efficiency of the heat exchangers; the pressure drops due to piping and throttle valves; the choice of either centrifugal or axial compressors (the latter are more efficient, 85% compared with 78%); and the selection of steam, electricity, or natural gas for drivers. Finally, there is selection of cooling systems either once through seawater, closed loop circulation of fresh water operating with a cooling tower, or closed loop air coolers. Every system in the LNG project should be studied in terms of exergy, that is, to extract the maximum potential of energy from each unit in every system, and to balance each gain in energy with any corresponding increase in investment. As with all projects, the final study will result in an economic optimization between energy consumption and investment. Maritime export terminals at the delivery end including large cryogenic liquid storage tanks, and pumping systems to transfer the product to the tankers. Maritime import terminals at the receiving end including large liquid storage tanks, pumping systems to transfer the liquefied product to storage tanks and regasification units if necessary. Apart from the actual liquefaction unit the overall worldwide LNG distribution concept includes export terminal storage to supply ships with supercooled natural gas; a fleet of specially designed tankers to carry and distribute the natural gas product with the minimum evaporation loss; an import terminal to receive the liquefied gas and store it in cryogenic vessels for final distribution either by liquid trucks or by pipeline in gaseous form. In France, the imported LNG is evaporated and pumped into a pipeline, the refrigeration recovered from the evaporated LNG used to operate a nearby air separation plant.


8.6.8.1 LNG Economics LNG plays an important role in the world’s ever-growing demand for energy, and to maintain this economic growth industry must depend on reliable sources of energy. As Lee Raymond, Chairman and CEO of ExxonMobil, stated in his address to the LNG Ministerial Summit Washington, D.C. in 2004, the next 20 years natural gas will capture 30%–35% of the increase in energy to be supplied to consumers (see Table 8.13). Moreover, according to the US Geological Survey, approximately 14 quadrillion cubic feet (14 1015) of natural gas are available worldwide. This supply, if properly taken care of, could last the world for the next 175–200 years. As already noted some of the sources are too far away to interconnect by pipeline. Therefore LNG is playing and will continue to play a very important role in its distribution to world industry. Currently the Asian countries riding an industrial boom and with no national sources of fuel for energy are looking at the importation of LNG to maintain their industrial expansion. China is planning to build as many as 10 LNG import terminals soon. This will also force China to compete with Japan, South Korea, and Taiwan to conclude attractive deals with Russia and the Middle East for LNG. According to Mr. Raymond, in an effort to reduce production costs, the liquefaction unit capacity of his company’s project has grown from slightly over 2 to 8 MMTA (million metric tons per annum), and the reduction of costs per metric ton has dropped from $350 to slightly over $100. In 2004, the consortium, which includes ExxonMobil, was building two large liquefaction trains at Qatar with a total liquefaction capacity of 15.6 MMTA. As noted, such a project requires a very large investment, which demands the need for a strong contractual understanding between suppliers and end-users.

8.6.9 SAFETY32 (FIGURE 8.14 AND FIGURE 8.15) The ever-increasing worldwide consumption of energy has impelled a greater use of natural gas resulting in new pipelines, and a greater demand for LNG from outside sources. The latter subject, however, has also created a vocal anxiety from some groups who fear that the construction of import terminals, and the use of large tankers hauling LNG in local docks, may lead to unforeseen accidents or even attract acts of terrorism. To relieve any anxieties about safety from all sources, one has only to review the past history of the liquefaction and transportation of liquefied methane (i.e., LNG). TABLE 8.13 Physical Properties of Methane Molecular weight Normal boiling point Vapor pressure Latent heat of vaporization Molar specific heat capacity of gas at constant pressure in the range of 0–1 bar and 298 K Thermal conductivity of gas Gaseous real density Liquid density Critical temperature Critical pressure Flammability range Source: Courtesy of Lotepro Data Book, 1975. With permission.


16.04 kg=kmol 111.7 K at 101.325 kPa 117 mbar at 90.7 K 8.18 MJ=kmol at boiling point and 101.325 kPa 35.72 kJ=(kmol K) 0.030 W=(m K) at low pressure and 273 K 0.717 kg=m3 at 273.15 K and 101.325 kPa 0.43 kg=L at 111.7 K and 101.325 kPa 190.6 K 46.04 bar 5.3%–14% in air, 5%–59% in oxygen

Design of maritime LNG transport

A cross-section of the LNG ship’s hull and containment system— in total of more than six feet in width Primary insulation Primary membrane Ship’s hull

Water ballast Secondary membrane Ship’s inner hull

Secondary insulation

FIGURE 8.14 A cross-section of the LNG ship’s hull and containment system—in total of more than 6 ft in width. (Courtesy of BP Corporation, 2006. With permission.) Maritime terminals of LNG A cross-section of the storage tank wallsin total about five and one-half feet thick Reinforced concrete

Typical liquefied natural gas storage tank with double walls

Perlite (balls) insulation Inner tank (walls and base) 9% nickel steel alloy Blocks

Stainless steel secondary base Base insulation foam glass Heating ducts to prevent ground freezing

FIGURE 8.15 A cross-section of the storage tank walls—in total of about 5.5 ft thick. (Courtesy of BP Corporation, 2006. With permission.)


TABLE 8.14 US Energy Consumption and Electricity Generation in 2003 3847.91 (Billions of kWh) Approximate distribution All sources Fossil fuels Coal=coke Natural gas Petroleum Nuclear Hydro pump storage Renewable Hydro Biomass Geothermal Solar Wind Other

100%

51.25 16.64 3.07 19.85 0.23 7.15 1.55 0.34 0.01 0.28 0.13

Fossil fuels separately (%)

Renewable separately (%)

72.21 23.46 4.33

76.56 16.64 3.66 0.15 2.99

Source: Courtesy of Department of Energy, National Energy Information Center (NEIC), and Energy of Information Administration (EIA). With permission.

In 2004, there were four import terminals in operation on the eastern American coast, and five more were approved for construction by the Federal Energy Regulatory Commission (FERC). An up-to-date report published by the Center for Liquefied Natural Gas indicates that from the time the first plant went into successful commercial operation in 1964, over 30,000 t of LNG were liquefied and transported throughout the world without any incident. Opponents have stated that a terrorist attack on an incoming tanker may trigger a spill into a harbor, and the spilled product upon contact with the warmer water will vaporize. Then in some circumstances a spark or outside energy could initiate ignition, leading into an uncontrollable and catastrophic conflagration. No critic has yet defined some circumstances in wellfounded scientific terms. LNG when vaporized into gaseous form (methane) and when mixed with ambient air at atmospheric pressure (1.013 bar) is inflammable when ignited, but does not explode. The opponents, moreover, ignore the fact that LNG tankers have been designed with utmost care and safety. The liquefied product is contained and separated in completely isolated compartments within the tanker so that in case of a pierced compartment, a total tanker spill is impossible. Furthermore the tanker’s hull involves two separate and parallel shells separated by approximately 2.5 m of protective material. The safety codes followed by the designers of the LNG tankers are those prescribed by National Fire Protection Agency (NFPA), and those issued by the US Department of Transportation (DOT). The Center for Liquefied Natural Gas states, moreover, that the Maritime Transportation Security Act of 2002 required all LNG ships and import terminals to submit security plans to the US Federal Government before the end of 2003. In addition, the US Coast Guard carefully screens LNG tankers that enter US waters, and if necessary, boards ships before the entry into US waters. Finally, US Federal Regulations authorize security zones to safeguard vessels, harbors, ports, and waterfront facilities against unauthorized entry and terrorism attacks. Apart from what government agencies have prescribed for safety in regard


to export and import storage tank terminals, the LNG industry is constantly improving the design to avoid any possible accidents and harmful acts due to terrorists.

REFERENCES 1. Hugill, T., and F.G. Kerry. 1955. Hydrogen recovery and purification of refining off-gases, API, Division of Refining, May 11, 1955. 2. Vandaveer, F.E., and S.W. Parr. 1934. Synthesis gas (partial oxidation). In Gas engineers handbook, 291–293. New York: McGraw-Hill. 3. Jost, W. 1946. Explosion and combustion processes in gases, 295–296. New York: McGraw-Hill. 4. Nelson, W.L., and F.G. Kerry. High purity carbon monoxide production. AIChE paper. 5. Foo, K.W., and I. Shortland. 1976. Compare CO production methods. Hydrocarb Process, May 1976, pp. 149–152. 6. Isalski, W.H. 1989. Separation of gases (Monogram on Cryogenics, No. 5), 171–175. Oxford University Press, pp. 171–175. 7. L’Air Liquide. US patent 6,173, 585 B1. 8. Dr. W. Ing Maxmilian of BASF, A.G. Ludwigshafen, and R. Schmidt, Schroeder of Messer Griesheim GmbH, Production of Carbon Monoxide. 9. Venet, F.C., E.M. Dickson, and T. Nagamura. 1993. Understand the key issues for high purity nitrogen production. Chem Eng Prog, January 1993, pp. 80–85. 10. Cheung, H. Hybrid nitrogen generator with auxiliary reboiler drive, Union Carbide and Chemicals. US patent 4,594,085. 11. Patel, S., T. Cormier, and K. Wilson. Nitrogen generator cycle. US patent 4,400,188. 12. Parker, C., and R. Flostello. Process and apparatus for producing N2 from air. US patent 4,966,002. 13. Okada, H., and S. Urata. N2 Production method, Nippon Sanso KK. US patent 4,617,037. 14. Ha, B., et al. Procede de Production d’Azote, E.P. Patent EP 0 413 631 AL. 15. Agrawal, R., et al. 1991. Efficient processes to produce ultra high purity nitrogen. In XV111 International Congress of Refrigeration, Paper No. 32, August 10–17, 1991. 16. Succi, M., SAES, 9th International Symposium on Contamination Control, September 26–30, 1988. 17. McAndrew, J., et al. Establishing moisture test and methods for process gas distribution systems, Micro Contamination, CRC Liquid Air. 18. Palier, V., et al. 1991. New simulation tools for designing and producing ultra pure gas distribution systems, Semi-Con., EUROPA, Conference Proceedings 1991. 19. Wang, H.C. 1989. Evaluating elements of a particle test sequence for point of use filters, MicroContamination, December 1989, CRC Liquid Air. 20. Wang, H.C. 1991. Evaluating elements of a particle test sequence for point of use filters, MicroContamination, January 1991, CRC Liquid Air. 21. Hinds, W.C. 1980. Aerosol technology, New York: John Wiley & Sons. 22. Jain, R. Purification of gases, The BOC Group, European Patent 0 438 282 A1. 23. Faith, W.L., et al. 1956. Nitrogen oxides: A challenge to chemical engineers, CEP, August 1956, pp. 342–344. 24. Rotzler, R.W., et al. 1960. Oxygen plant reboiler explosion, vol. 56 (6), CEP, June 1960, pp. 66–73. 25. Lewis, B., and G. Von Elbe. 1951. Sensitization and inhibition by additives, combustion, flames and explosions of gases, 93. New York: Academic Press. 26. Salama, C., and D.V. Eyre. 1967. Advantages of multiple refrigerants in the liquefaction of natural gas. New Developments in Natural Gas Session 28, 61st National Meeting, February 19–23, 1967, Houston TX, paper 28 C. 27. Bourguet, J.M., R. Garnaud, and M. Grenier. 1970. Large capacity LNG installations, LNG-2, October 1970. 28. Chen-Hwa, C. 1978. Evaluate separations for LNG plants. Hydrocarb Process, September 1978, pp. 266–272. 29. Inman, H., and R.I.J. Soetopo. 1979. A run LNG plant on stream in Indonesia. Pipeline Gas J, June 1979.


30. Dufrene, J.P., H. Jaradowski, and E. Cretani. 1980. Energy Savings in Natural Gas, LNGConference, Kyoto, Japan, April 7–11, 1980. 31. Bourgeut, J.M. 1981. LNG cold still a problem. Hydrocarb Process, January 1981, pp. 167–172. 32. University of Houston Law Center’s Institute for Energy, Law and Enterprise. Introduction to LNG: an overview on liquefied natural gas (LNG), its properties, the LNG industry, safety considerations, and LNG Safety and Security. 33. Kerry, F.G., Chemical engineering progress, Vol. 57, No. 4, April 1961.

FURTHER READING ON LNG Crawford, D.B., J.D. Cronk, and J.C. Norenberg. 1969. Economic factors of cryogenic gas processing. Natural Gas Processors Association, Dallas, Texas, March 13, 1969. Fundamentals of the Global LNG Industry, Published by Petroleum Economist, March 2001. Johnson, W.D. 1983. On entropy, efficiency and process design. Hydrocarb Process, February 1983: 61–64. Kenney, W.F. 1989. Current practical applications of the Second Law of Thermodynamics. Chem Eng Prog, February 1989: 57–63. LNG Today: The promise and the pitfalls, published by the Energy Publishing Network in Cooperation with Gas Strategies, 2002 edition. LNG Today: The promise and the pitfalls, published by the Energy Publishing Network in Cooperation with Gas Strategies, 2004 edition.



9

Noncryogenic Separations

Just after the termination of World War II, the CEO of one of the largest industrial gas companies advised his engineering department not to be too complacent in regard to the expanding sales of industrial gases produced by cryogenic systems. Some day an enterprising engineer would develop a machine in which air will be fed through one end, and its components will come out separately from the other. Little did he realize at the time that his offthe-cuff statement was prophetic. About five years later, a young engineer delivered a technical paper at an AIChE meeting in New York City, demonstrating that it was possible to separate air into oxygen and nitrogen by passing compressed air through a series of porous membranes. Admittedly, the purity of the oxygen was not up to the desired standards, but it could be improved by adding more membranes in series. The audience at the meeting, though mildly interested, opined that the proposed process could not compete with the cryogenic distillation processes already in use. They were wrong. Although the new idea has not displaced the cryogenic processes completely, noncryogenic separation and purification of gases has become a serious player in the industrial gas industry over the last several decades.

9.1 PERMEABLE MEMBRANE SEPARATION Permeation was the first process considered for noncryogenic separation of industrial gases.

9.1.1 GENERAL PRINCIPLES Permeation is the diffusion of a substance in solution through a barrier. Permeability, on the other hand, is the capacity of a porous material for transmitting a fluid. The standard unit of permeability is the darcy, equivalent to the passage of 1 cm3 of fluid (having a viscosity of 1 cP) per second through a sample of 1 cm2 cross-sectional area, under a pressure of 1.013 barA=cm of thickness. In the gas industry, the feed gas is labeled the permeator, and the selected component to be separated is known as the permeate. The membranes used are thin, dense, and continuous films formed from cellulose acetate or polymers. The separation of a component in a gas mixture is carried out in three steps: the component must dissolve in the membrane wall, diffuse through the membrane material, and be desorbed on the opposite side of the membrane wall. This procedure may be defined by Henry’s law of solubility—The solubility of a gas in a liquid is proportional to the partial pressure of the gas—and by Fick’s law of diffusivity as expressed with the following equations. The feed gas solubility in the polymer is ziF ¼ Si PF xi

(9:1)

Fick’s Law inside the membrane material: Qi =A ¼ Di (xi yi )=Z


(9:2)

Permeate solubility in the polymer: ziP ¼ Si PP yi

(9:3)

where ziF ¼ concentration of component i in the membrane on the feed side ziP ¼ concentration of component i in the membrane on the permeate side xi ¼ partial pressure of component i in the feed gas yi ¼ partial pressure of component i in the permeate gas PF ¼ feed gas pressure PP ¼ permeate gas pressure Qi ¼ flow of component i across the membrane Si ¼ solubility coefficient of component i Di ¼ diffusivity coefficient of component i A ¼ membrane area Z ¼ membrane thickness The operating performance of any polymeric membrane is characterized by two factors: (a) permeability, which defines productivity, the transport rate for a given species in the feed stream, and therefore, the cost of the system; and (b) selectivity, which in turn defines recovery or purity of the selected stream. For example, according to Fleming and Dupuis,1 expressions for permeability and selectivity can be derived from the basic Equation 9.1 through Equation 9.3 where hydrogen is the more permeable component and methane is the less permeable component: H2 ¼ ½(DH2 SH2 )=Z ¼ (QP yH2 )=A(PF xH2 PP yH2 )

(9:4)

PCH4 ¼ (DCH4 SCH4 )=Z ¼ QP (1 yH2 )=A½PF (1 xH2 ) PP (1 yH2 )

(9:5)

Permeability ratio a ¼ PH2=PCH4 (also known as separation factor) ¼

yH 2 (1 yH2 )f½PF (1 xH2 ) PP (1 yH2 )=½PF xH2 PP yH2 g

(9:6)

where PH2 is the hydrogen permeability, PCH4 the methane permeability, and a the ratio of hydrogen permeability to methane permeability. Furthermore, Equation 9.2 may be used to calculate the hydrogen concentration in the product (permeate) stream: 1=2 yH2 ¼ K K 2 PF =½(a 1)PP xH2

(9:7)

where K ¼ 1=2 þ ½PF =2(a 1)PP þ ½PF =2PP xH2 The operating principle of selective permeation is based on the difference of partial pressure of say hydrogen, a fast gas, and methane, a slow gas. Gases with a high permeation rate (fast gases) diffuse rapidly through the membrane, and are directed with the permeate stream. Gases with a slow permeation rate (slow gases) flow around the walls of the membrane and


TABLE 9.1 Comparison of Permeation Rate of Industrial Gases with Helium the Fastest and Propane the Slowest He, H2, H2O, NH3, CO2, H2S, CH3OH, CO, O2, Ar, N2, CH4, C2H6, C3H8 with helium having the fastest permeation rate, and propane the slowest. Source: From Fleming, G.K. and Dupuis, G.E., Hydrocarb. Process., April 1993. With permission.

remain in the residue stream. One may state that ‘‘the driving force for separation is the difference in partial pressures of each individual molecular species on opposite sides of the membrane wall.’’ The various industrial gases may be classified, in terms of permeation rate or separation factor as given in Table 9.1. As shown, the difference in permeability enables the membrane to separate gas species, but it must be kept in mind that any membrane will not produce high-purity products or selectively remove only one of the species. Each gas has a finite permeability in the membrane. Enrichment is achieved because of relative permeabilities, and not because of any zero permeability for one of the species. For example, it is fairly easy to attain a high recovery of hydrogen from propane and heavy hydrocarbons. To recover hydrogen from ammonia purge gas, however, ammonia content must be removed by a thorough water wash and followed by a drying phase, as both NH3 and H2O have permeation rates too close to H2. Purity and recovery are inversely related depending on how close one species is to the other. In some cases where both maximum purity and recovery are desired, membrane separation must be followed either by adsorption or even cryogenic separation. Weller and Steiner2 studied the influence of total pressure ratio as well as separation factor on permeate purity by modifying Equation 9.3 and converting it into Qi =A ¼ Pi=l(pi2 pi1 )

(9:8)

This equation shows the ratio of the permeation flux for component i to that of a component j. Applying the definition for partial pressure for the separation factor as indicated in Equation 9.6, and that for a binary mixture yj ¼ (1yi), the same equation can be written as Qi =Qj ¼ yi =(1 yi ) ¼ aij ½(P2 =P1 )yi2 yi1 =½(P2 =P1 )(1 yi2 ) (1 yi1 )

(9:9)

where P2=P1 is the ratio of total pressures on the respective sides of the membrane. This equation is solved for the permeate composition y1 using a quadratic formula. Figure 9.1 shows the influence of total pressure ratio and separation factor on permeate purity and the influence of feed composition and separation factor on permeate purity, respectively. Obviously, permeate composition is not changed significantly above a pressure ratio of 6, and permeate purity is not influenced significantly by a separation factor greater than 20. If the feed gas is a multicomponent mixture, it must be understood that the membrane treats it as an equivalent binary mixture, separating it into a group of fast gases and another stream of slow gases. Then an equivalent selectivity between the averages of the two streams is obtained. Another problem arises from the fact that during operation the concentration of the very fast gas, say hydrogen, decreases as it diffuses into the bore of the collection tube, making it necessary to repeat the calculations by dividing the permeator into radial increments3. The concentrations xi and the partial pressure pi will change along the stream’s flow path. Therefore, since the flux equation is a point function, it must be integrated along the entire flow path to determine the amount of species that permeates. Moreover, a flux


100 Pressure ratio = 10 95 6 90

Product purity (mole %)

85 Pressure ratio = 4 80

75

70

65 30% Feed concentration 60

55

50 10

20

30

40

50

60

70

80

90

100

Separation factor

FIGURE 9.1 Effect of total pressure ratio and separation factor on product purity. (From Weller, S. and Steiner, W.A., J. Appl. Phys., 21, 279, 1950. With permission.)

equation can be written for each permeating species such that the integrating process must be carried out for each, while ensuring that the sums of partial pressures equal the total pressure at each point. Fortunately, membrane suppliers have databases and simulation procedures that are available gratis to serious clients. With the possible exception of vendor selection, the process designer has no control over membrane selectivity, because this parameter depends primarily on the interaction of the gas categories being separated with the polymer supplied by its vendor. Typical separation factors for commercial membrane separators are shown in Table 9.2. Finally, as in all fluid flow, the pressure drop at the bore of the membrane must be calculated, and is proportional to the permeate flow rate per fiber, which in turn depends on the cross-section area of the membrane. The pressure drop is calculated using the following formula4. Dp ¼ Cf lw2=rd 5 ‘

(9:10)

where Dp is the calculated bore pressure drop, C the constant function of the units selected, f the friction factor, l the length of fiber, w the mass flow, r the density of permeate, and d the inner radius of fiber.


TABLE 9.2 Typical Separation Factors for Membrane Separation Separation

Factor Range

H2–CO H2–CH4 H2–N2 CO2–CH4 O2–N2

35 %–80 % 50 %–200 % 50 %–200 % 10 %–50 % 3 %–12 %

Source: From Stookey, D.J. et al., Chem. Eng. Prog., November, 1986. With permission.

As indicated in Figure 9.1, permeate purity is not greatly influenced by separation factors greater than 30. Hence, the separation factor alone will not be the determining factor in the selection of the membrane apparatus. Figure 9.1 confirms that permeable membranes will not produce high-purity products, or selectively remove only one of the gases. To upgrade the purity, one may employ multiple membranes in series or combine their use with other separation technologies, such as adsorption, absorption, or even cryogenic separation. It has also been established that many industrial applications do not necessarily require very high purities. The main target of the equipment designer is to achieve maximum permeability (productivity) with optimum selectivity (purity). The selective permeation characteristic of a desired membrane is to permit fast gases such as hydrogen, helium, and carbon dioxide to be separated from such slow gases as nitrogen, methane, and other hydrocarbons. As noted, the separation efficiency for a given membrane with a specific binary gas mixture will depend on gas composition, pressure ratio between the feed and the permeate gas, and the separation factor. The higher the separation factor, the more selective the membrane will be, and a greater separation efficiency will be obtained. From relationships shown between pressure ratio, gas composition, and separation difference of the more permeable gas, one may deduce the following: 1. Increasing the gas permeability at the same separation factor decreases the membrane cross-sectional area requirement proportionally. 2. Increasing the feed-gas pressure at constant permeate pressure decreases the membrane cross-sectional area, increasing permeate selectivity (recovery). 3. Increasing the separation factor may decrease the number of membrane stages required to achieve a given enrichment, or allow for a higher permeate pressure. In terms of economics, the two most important factors are feed or permeate compression that translates into energy, and membrane cross-sectional area that means investment. Because they are inversely proportional to each other, there is an optimal operating design for a given set of conditions for each application that must be determined beforehand.

9.1.2 MECHANICAL DESIGN

OF

MEMBRANES

Membranes originally made from cellulose acetate can be manufactured in flat sheets with a plate and frame configuration or incorporated into a spiral wound design. The latter shape may be preferable for industrial applications, as the membranes can be inserted into pressure cylinders in series. These tubes can be skid mounted in either a series or a parallel arrangement depending on the specific system dynamics. On the other hand, the alternative choice is to employ a hollow-fiber configuration made from polymer material that has a greater


packing density (membrane area per packaging volume) than flat sheets. It should be kept in mind, however, that the mechanism for gas separation is independent of any membrane configuration, and depends only on the simple principle of physics that certain gases permeate more rapidly than others, as previously explained. Spiral wound sheets and hollow-fiber processes differ only in membrane material type and system performance (Table 9.1). It may be noted, moreover, that the use of hollow-fiber polymer membranes outnumbers the spiral type made from cellulose acetate by a wide margin. In regard to the use of hollow-fiber membranes, the driving force for separation is the difference between the stream components’ partial pressure on the outside of the fibers (shell side), and that of the inside (bore side). These units may be configured either vertically or horizontally depending on space requirements. The membranes in question are hollow-fiber polymers, each permeator containing thousands of these fibers, and each approximately twice the diameter of a human hair. Composition and dimensions are specifically designed for each application, and the devices can operate at temperatures up to 300 K. They can also withstand pressure differentials of 110 bar. They provide a convenient and cost-effective way to recover and recycle hydrogen from hydrogen-rich synthesis gas or purge streams. They are widely used in oil refineries for the recovery of catalytic reformer off-gas, hydrocracker and hydrotreater purge streams, hydrodealkylation recycle gas, isobutane dehydrogenation gas, and hydrogen plant feed. On the other end of the scale, membranes are used to recover valuable helium from the helium– argon–nitrogen mixture used in underwater diver systems because of the permeation rate. The use of separately layered zeolites can be selective; and operating at a higher level of product purity, their use may be very competitive with cryogenic systems. For example, present day cryogenic systems can produce ultra pure nitrogen at a purity level of nine nines (an impurity of 1vppb), at a flow of 24,000 Nm3=h (720 t=d). In comparison, an up-to-date membrane system is limited to a purity level of 99.9% at a flow rate of 200 Nm3=h, and even up to flows of 4000 Nm3=h for custom cases. Membrane separation used jointly with a catalytic system-scan, nitrogen purities as high as 99.9995% can be obtained. Another newcomer that has entered the membrane separation field is the employment of ceramic (inorganic) membranes. Moreover, ceramic membranes are being developed presently such that they can produce ultra-high purity of up to six nines, 99.9999%. Unlike polymeric membranes ceramic membranes are stable at high temperatures and resistant to solvents. Because of these advantages ceramic membranes are now being tested in a variety of gas separations at temperatures up to 1273 K. If successful, such applications would be of tremendous benefit to coal gasification processes.

9.1.3 GENERAL APPLICATIONS The major industrial gas companies have not ignored the competition from membrane separation. On the contrary and without exception, they have made this noncryogenic system a serious part of their process technology for gas separation and recovery. In fact, this upstart process has become an increasing part of their development and marketing agenda. In 2004, the noncryogenic systems and equipment sales have surpassed a figure of 20% of their total sales. Because of their portability and simplicity, membrane separation systems are the obvious choice for remote locations such as ships and offshore oil platforms. In the latter use, they can operate at high pressures up to 23–24 bar, and with a nitrogen purity of 95% to 99.5%, which is more than adequate for that purpose. One must keep in mind, however, that inert nitrogen generated from permeate membranes may not be completely compatible in terms of flammability for certain materials that it is intended to protect. The oxygen content of the nitrogen product must be lower than the acceptable lower flammability limit specified by code (Table 9.3).


TABLE 9.3 List of Typical Critical Oxygen Concentrations Acetaldehyde Acetone Allyl chloride Ammonia Benzene 1,3 Butadiene Butane 1-Butene Ethane Ethanol Ethylene Gasoline (octane 100) Heptane Hexane Isobutane Methane Methanol Pentane Propane Vinyl chloride

12 11.6 12.6 15 11.2 10.4 12.1 11.4 11 10.6 10 11.6 11.6 11.9 12 12.1 9.7 12.1 11.4 9

Source: From Lewis, B. and von Elbe, G. in Combustion, Flame, and Explosion of Gases, Academic Press, New York, 1961. With permission.

9.1.3.1 Nitrogen Separation As shown, the use of polymer membranes has almost taken over nitrogen separation. In this application, the permeate from the air feed includes a mixture of oxygen, carbon dioxide, and water vapor, leaving behind as a by-product the slow gas, namely a dry nitrogen (dew point 224–200 K), with traces of oxygen and argon, within the specified limits of noncombustibility, and with particulates less than 1 mm. This quality of nitrogen is generally useful to enhance safety from combustible products, prevent toxic emissions, improve product quality, and eliminate odors from chemical processes. Its high dryness has also helped prevent problems associated with moisture contamination of processed materials. Standard systems provide flows from 0.5 to 200 Nm3=h and custom systems have been designed for flows up to 4000 Nm3=h. Figure 9.2 shows diagrammatically a possible flow schematic for nitrogen separation. In the figure, a surge tank is indicated, if required by the application, also a back-up liquid nitrogen storage tank. The latter tank is needed in the event to take care of incidents that may stop the flow of nitrogen. It is also possible, and highly recommended, to incorporate a computerized control system as most of these units are generally unattended in operation. The entire system can be easily arranged to handle variations in feed-gas flow and composition. A typical membrane system consists of a pretreatment skid and a group of membrane modules. The overall system is self-contained. Site preparation is minimal, generally requiring a level concrete pad, and the ancillary piping and wiring connections can be prepackaged on the same skid as a complete unit. The entire system involves static components and needs little operator attention or maintenance. The life expectancy of the membrane system is 5–10 years.


1.0

O2 in product gas (%)

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0

10

20

50 30 40 Product gas flow (m3/h)

60

70

80

FIGURE 9.2 Effect of product gas flow on oxygen content in nitrogen. (Courtesy of Carbo Tech Engineering GmbH, 2006. With permission.)

9.1.3.2

Disadvantages of Membrane Separation

The main disadvantage of membrane separation of nitrogen is that it may require a catalytic deoxygenation unit to reduce the oxygen content to the same level as other separation systems. For example, a membrane separation unit is competitive at nitrogen purities of 95%–99.5%; but beyond that purity, one may have to consider the use of a pressure swing adsorber (PSA) system. Moreover, its applications are cost effective only for relatively small capacity units. It has been recommended, however, for inerting commercial aircraft fuel tanks, gas liquid contactors (switches), and for drying of air. 9.1.3.3

Hydrogen Recovery

Apart from nitrogen, recoveries of hydrogen may be obtained up to 80% –90% in most applications (the recovery of hydrogen in oil refineries for use in hydrotreating crudes that become lower in quality over time). As it has been classified as a ‘‘fast gas,’’ hydrogen can be recovered quickly and economically from various chemical purge-gases such as ammonia synthesis purge gas, ethylene off-gas, methanol synthesis purge gas, etc. As previously noted, membranes have been accepted in oil refineries for the recovery of catalytic reformer off-gas, hydrocracker and hydrotreater purge streams, hydrodealkylation recycle gas, and isobutane dehydrogenation gas. At one refinery, a membrane system has been used to upgrade 14,000 Nm3=h of hydrogen at 85% content up to 99.5% with a recovery of 85%. One application of polyamide membranes involves natural gas sweetening. It provides safe and efficient removal of CO2, H2S, SO2, and water vapor from natural gas wells, especially in geographically remote areas. Furthermore, it has been shown that polyamide membranes deliver a higher productivity and separation efficiency than cellulose acetate membranes, resulting in higher gas purities and lower production costs.


9.2 GAS SEPARATION BY ADSORPTION 9.2.1 GENERAL OVERVIEW Although the technical principles of adsorption have been described in Chapter 5, its original use in gas separation was for the prepurification of process air prior to separation. Once the use of activated alumina and artificially prepared zeolites for gas adsorption was established and universally adopted, the next and obvious step was to use the same principle for the recovery of pure, or nearly pure, products from mixtures of gases as well as from air. 9.2.1.1 Adsorption Processes Studies With the ever-increasing demand for hydrogen, especially for the petrochemical industry, the industrial gas companies studied and promoted the advantages of PSA and vacuum pressure swing adsorption (VPSA). These studies fulfilled the demand for recovering hydrogen from a wide variety of mixtures including coke oven gas. Depending on gases to be adsorbed, there are various materials that have been used as adsorbents (see Table 9.4). As noted in Table 9.5, there is qualitative variation in the qualitative order of adsorption (adhesion). For example, helium and hydrogen have weaker forces of adsorption than either methane or nitrogen. Moreover, nitrogen, oxygen, and argon have almost the same order of adsorption making them difficult to separate. Because of this difference in adsorption it is possible to separate a wide variety of industrial gas mixtures into their various components. 9.2.1.2 Regeneration of Adsorbent Once saturation of the adsorbent is reached, regeneration is carried out by either applying heat or by lowering pressure. The former procedure is generally named Temperature Swing Adsorption (TSA), and if the pressure is lowered before heat is applied it may be called Pressure Temperature Swing Adsorption (PTSA) (see Chapter 5). Regeneration may also be carried out by only lowering the operating pressure. This is called PSA. If vacuum is applied, followed by elution, a higher product purity can be achieved. In the latter case, the process is called VPSA. PSA and VPSA are advantageous for the recovery and upgradation of hydrogen, especially if the hydrogen content is at least 50%. It has been successfully applied in treating such gas mixtures as (a) catalytic reformer off-gas, (b) coke oven gas, (c) electrolysis gas, (d) ethylene off-gas, (e) hydrotreatment purge gases, (f) hydrogen from H2=CO cold box=(g) NH3 TABLE 9.4 Variety of Materials Used for Adsorption Material Configuration Main components Particulate size (mm) Bulk density (kg=m3) Surface (m2=g) ˚) Pore diameter (A

Silica

Alumina

Activated Carbon

Molecular Sieves

Beads Granules SiO2 1–5 700–800 600–850 20

Beads Al2O3 2–10 700–850 250–350 30–35

Pellets Granules C 1–5 300–600 700–1200 10–10000

Beads Pellets SiO2Al2O3 1–5 600–900 500–1000 3–13



TABLE 9.5 Qualitative Variation in Order of Adherence Strength Adsorption Strength Very Low

Medium

High

Very High

Ar O2 N2

CO CH4 C2H6 CO2 C3H8 C2H4

C3H6 C4H8 C5þ H2S NH3 BTX H2O

He H2


plants off-gas, (h) refinery purge streams, (i) syngas from partial oxidation, (j) syngas from steam reformers (natural gas, naphtha, methyl alcohol), and (k) styrene off-gas. Generally the PSA and VPSA processes can be applied without any preconditioning of feed stocks, but entrained liquids such as water or hydrocarbon condensates may damage the adsorbents, therefore, a separator, mist eliminator, and steam tracer may be desirable. In hydrogen recovery, one must bear in mind that helium cannot be removed. 9.2.1.3

Hydrogen Recovery from Coke Oven Gas

In this category two plants have been built and in operation since 1985, one in Belgium producing 4000 Nm3=h of hydrogen, and one in Canada with a capacity of 2700 Nm3=h of hydrogen. The feed-gas composition affects the design of the adsorber vessels, as well as selection of adsorbents (Table 9.6). The optimization leads to multiplying the number of the superimposed beds, each designed for a specific impurity.

9.3

NITROGEN RECOVERY

The pressure swing adsorption (PSA) system has been readily adopted as simpler as well as more cost effective for the application of inert gas blanketing, and nitrogen-based controlled atmospheres. The same process can also be employed to recover oxygen. In the latter case and with the addition of vacuum desorption, the use of VPSA avoids the need for cryogenics with all the latter’s costly ancillary equipment.

TABLE 9.6 Material Used in Separate Layers to Remove Specific Gases Alumina H2O NH3

Carbon Prefilter

Activated Carbon

Molecular Sieve

CxHy H2S BTX

Ar O2 CH4 CO2

N2 CO



9.3.1 CARBON ADSORBENT (CARBON MOLECULAR SIEVE) (CMS) The use of specially treated permeable carbon, sometimes called carbon molecular sieve (CMS), has an advantage over the use of various synthetic zeolites, especially for inert atmospheres5. Even though this material adsorbs oxygen and nitrogen in comparable quantities in equilibrium conditions, the rate of adsorption of oxygen is considerably faster than that of nitrogen. In fact, within a few minutes the oxygen equilibrium load reaches a value of 80% whereas that of nitrogen is less than 5%. The diffusion determined volumetrically between the two shows a significant difference.5 During the adsorption of oxygen, moreover, the water vapor and carbon dioxide contained in the feed air are also adsorbed without affecting the adsorption capacity of the carbon (Figure 9.3). The process is similar to any adsorption cycle using two adsorption vessels, but includes a vacuum pump, two buffer pumps, one downstream of the feed compressor, and the second after the adsorption system in order to minimize any fluctuations in the flows (Figure 9.4). Loading and desorption are generally adjusted to have identical periods, around 60 s each, and the desorption pressure is usually 93 mbar. The oxygen waste product contains about 35% oxygen, the balance nitrogen as well as carbon dioxide and water vapor. Adjusting air flow rate, however, the oxygen concentration of the nitrogen product may be lowered to 0.1% by volume, though this also lowers nitrogen product recovery. Product recovery may be increased by raising the feed pressure or flow, but this also has a limit, because it increases the oxygen impurity in the product nitrogen (Figure 9.3 and Figure 9.4). Systems with production of up to 2500 Nm3=h have been designed and operated, and the oxygen impurity can be lowered by means of a catalytic purifier unit. Caution: if one uses a lubricated feed air compressor, it will be necessary to include one or even two coalescers in series downstream of the compressor in order to remove all traces of entrained oil. Using permeable carbon for nitrogen generation is economically advantageous in small capacities (1000–2500 Nm3=h), and in geographical areas remote from large industrial suppliers. The same systems have also been designed for the production of oxygen, but the purity of

60 0.5 vol % O2

Product-gas flow (m3/h)

50

40

0.2 vol % O2

30

20

10

0 1

2

3 Operating pressure (bar)

4

5

FIGURE 9.3 Influence on operating pressure oxygen content in nitrogen product. (Courtesy of Carbo Tech Engineering GmbH, 2006. With permission.)


1.0

Loading factor, L/Le

O2

0.5

0

N2

30

60

90

Time (min)

FIGURE 9.4 Oxygen and nitrogen adsorptivity of molecular sieve coke. (Courtesy of Carbo Tech Engineering GmbH, 2006. With permission.)

the product is limited to 80%. Using a CMS for oxygen production has not made much of a dent in the industrial market even for small units.

9.3.2

HIGH-PURITY HYDROGEN RECOVERY6,7

With the increasing demand for pure hydrogen, industry has been looking for more costeffective processes to obtain the product. Adsorption at approximately ambient temperatures has the answer. PSA, especially when combined with VPSA, has lowered the cost of recovering hydrogen by 35% and simplified the process operation. PSA and VPSA have been extremely helpful to the food industry for the hydrogenation of oils, the glass industry for the manufacture of float glass, the electronic industry for transistors, and the metallurgical industry for bright annealing tin-plate. In this application, the molecular sieve used contains a sodium ion that can easily be substituted by any of the species K, Li, Ba, Ca, or Mg depending on the specific impurities to be removed. One or more of the various elements can be used. Desorption of the adsorbent may be accomplished either by vacuum or elution or a combination of both. The additional use of vacuum improves the yield of the hydrogen product. The operating range of the switch valves is very short, anywhere from 2 to 20 min. This short cycle serves to limit heat oscillations due to adsorption and desorption to a few degrees. PSA and VPSA processes have been tested and studied for a wide variety of applications ranging from the production of high-purity hydrogen (99.999%) to the recovery of helium–neon. In the latter case, pure helium (99.95%) was recovered from a mixture of He and 10% air saturated with moisture. By using the most suitable adsorbent such as activated alumina, silica gel, activated carbon, or a variety of synthetic molecular sieves in any combination, it is possible to remove impurities simultaneously, for example, H2O, CO2, N2, CO, CH4, C2H6, NH3, and Ar, thereby eliminating the usual cumbersome series of individual purifying units involving conversion, drying, decarbonation, etc. Furthermore, it enables the process engineer to treat


purge gases from the production of ammonia and methanol recovering valuable fuels such as CO and H2 that can be used to assist firing furnaces or boilers.

9.3.3 OXYGEN SEPARATION

AND

VACUUM PRESSURE SWING ADSORPTION

9.3.3.1 Process Description For noncryogenic generation of oxygen, the use of synthetic zeolite adsorbents in conjunction with VPSA is the preferred process. To increase the adsorption capacity for nitrogen and improve N2–O2 selectivity, the use of CaA zeolites, and the new LiX series has been found more efficient than the older NaX series8. These improvements have led to a significant reduction in adsorption volume as well as increased efficiency. Furthermore, vacuum desorption for larger capacity units has led to a saving of energy. For industrial plants from 50 to 185 t=d the energy consumed is 350–120 kW=t of contained oxygen, but at a purity of 93%, which is the current possible maximum. The balance is argon and some nitrogen.6 With the use of activated alumina upstream of the molecular sieve adsorbent, compressed air is fed into the vessels, where the corrected selected adsorbents retain nitrogen to a greater degree than oxygen, and also remove water, carbon dioxide, and dangerous hydrocarbons entrained by the air feed. The usual sequence for a normal VPSA is as follows: a. During adsorption, the pressurized air flows into the vessels. b. During desorption, the pressure in the vessels is reduced below atmospheric pressure, and the residual gas is exhausted by a vacuum pump while the adsorbent is being regenerated. c. After desorption, the required adsorption pressure is built up again in the vessels. As a rule, two vessels plus a buffer vessel are provided to enable the plant to produce an optimum continuous product oxygen flow. Some designers use three vessels operating in series because it increases the number of equalizations, thereby maximizing product recovery by reducing purge gas losses8. This design also lowers equipment costs because the vessels have thinner walls cutting fabrication costs and the switching valves, though higher in number, are smaller, and therefore, much lower in total cost. A few designers even offer a single vessel followed by a very large buffer vessel in order to keep investment to a bare minimum. This practice is not recommended except in extreme circumstances, and always with the prior permission of the ultimate owner or user. Another development that has given a marked impetus to the noncryogenic separation of air is the increased use of lithium-based zeolites and lowering the use of sodium type. A greater use of the former has been found to offer a more efficient separation.

9.3.4 ENGINEERING DESIGN 9.3.4.1 Basic Principles The basic principle of the design is to select one adsorbent or a combination, offering a high net loading of the unwanted components (productivity), as well as a high selectivity of the desired component (purity) and pressure (economy). Because of the different variables involved in adsorption, it is very important to examine closely the properties of a preferred adsorbent, and how it behaves during the process conditions of the proposed cycle. Another point to keep in mind is that the adiabatic nature of the adsorption process may create undesirable temperature gradients that have to be avoided or at least reduced to a minimum.


This can be effected with the use of multilayered adsorbents. With such predetermined multilayer configurations, it is possible to reduce recycle requirements, overall operating pressure, and produce other economies. In regard to the actual design of the vessels, the basic requirement is high throughput, which has been addressed largely by increasing the horizontal diameter of the adsorbent, reducing the bed height, and by increasing the number of cycles per unit of time. Although this concept may avoid a costly pressure drop across the adsorbent bed, it also increases the risk of fluidization, which should be avoided as much as possible. Moreover, the design of very large diameter vessels also hampers transportation from the fabrication shop to the site. Some of the problems involving air flow distribution have been minimized, but not completely eliminated, by the use of inert ceramic spheres ranging in diameter from 6 to 50 mm spaced at precalculated horizontal layers in the adsorbent bed. These problems have been almost completely overcome by the development of the radial flow adsorbers, which have been described in Chapter 5, and these adsorbers are widely used for prepurification of air in cryogenic plants since their development in the 1970s. The basic principle involved has been the horizontal circulation of the feed air through the vertically layered adsorbents, thus avoiding problems of attrition or lifting the bed. The latter problem is even more critical because the reduced cycle time associated with VPSA units is combined with high gas velocity, as well as adsorbent granules of small diameter. During the production phase, air enters the lower section by radially crossing the activated alumina bed, adsorbing all moisture and carbon dioxide. The purified process air stream continues upward and flows horizontally through the adsorbent beds. During the regeneration phase, the nitrogen eluted from the zeolite by vacuum pumping sweeps at low pressure across the alumina bed, desorbing the previously retained water and carbon dioxide. Special studies have been made to ensure a uniform circulation of these flows to minimize pressure drop and dead air volume, thereby saving energy. Vacuum-assisted regeneration of the adsorbent beds is normally used by most if not all process designers to reduce nitrogen contamination of the oxygen product. 9.3.4.2

Disadvantages of Adsorption

The application of either PSA or VPSA processes requires the use of switching valves that may present inherent problems. Basically, each vessel used in either process operates as a batch process, and the switching valves operate at close cycles, anywhere from 2 to 20 min in conventional applications. An additional vessel such as a surge tank may also be used in the cycle to reduce variations in stream flow. To optimize the overall operating efficiency of the process, therefore, a fairly complex system of valves is involved to switch gas flows between the adsorbent beds. These valves are switched by programmed logic controller (PLC) solenoids. To be cost effective, the switching valves must be reliable in performance and have a long operating life between maintenance. To avoid particle abrasion, gas velocities through the adsorbent beds are also limited by fluidization problems. On the other hand, slow operation of switching cycles to avoid particle erosion leads to large pressure vessels and ancillary equipment, resulting in increased material and erection costs. Moreover, because of fluidization constraints adsorbent particle size is limited. In normal industrial applications, particle size used is approximately 2 mm, resulting in high adsorbent inventories and slower switching cycles. A high number of switching valves with their ancillary piping, instrumentation, and control systems add to the dead volume of the process, thus reducing product yield and efficiency. Furthermore, unless very reliable in quality and operation, solenoid-actuated switch valves may prove to be very expensive in the long run. Field reports have indicated a high rate of failure and replacement of internals, not to mention costly down-time.


9.3.4.3 Economics Generation of oxygen from a VPSA plant will yield a purity of anywhere from 90% to 93% maximum. It should be kept in mind, however, that the upper purity of 93% proposed by sales representatives may be slightly optimistic; because during contractual discussions of warranties, suppliers have been more conservative and prefer to talk of purities closer to 91%. This alternative to the cryogenic process has found an economically viable niche both in the medical as well as the industrial field, and in generation capacities between 13 and 150 t=d of contained oxygen. Well over 100 noncryogenic units have been constructed worldwide during the past 15 years. The commercial advantages are quite clear to industries that do not demand a minimum purity of at least 98%. This includes the medical field, production and smelting of primary metals, pulp and paper, and glass products. It should be kept in mind that certain states in the United States of America, such as California, have strict laws regarding the amount of NOx produced by the industry. To avoid this dangerous by-product from any oxy-combustion process, the oxygen product must have a minimum purity of 98%, keeping nitrogen impurity in the oxygen product to less than 1000 vppm, and the balance of the impurity being argon. In hydrogen recovery, the economics depend on the value of hydrogen to the user. If recovery is as low as 50% the enriched hydrogen is usually recycled using a booster-pressure compressor that adds to power consumption. Furthermore, the calorific value of the hydrogen should also be added to its recovery cost to reflect its total loss to the plant’s fuel bill.

REFERENCES 1. Fleming, G.K., and G.E. Dupuis. 1993. Hydrogen membrane recovery estimates. Hydrocarbon Processing, April. Gulf Publishing Company: 83–86. 2. Weller, S., and W.A. Steiner. 1950. J Appl Phys 21:279. 3. Weller, S., and W.A. Steiner. 1950. Chem Eng Prog 46:585. 4. Stookey, D.J., et al. 1986. Membranes separate gases selectively. Chem Eng Prog (November): 36–40. 5. Knoblaugh, J. 1972. Pressure swing adsorption: Gases for small volume users. Chem Eng (November): 87–89. 6. Watson, A.M. 1983. Use pressures wing adsorption for lowest cost hydrogen. Hydrocarbon Processing (March):92. Gulf Publishing Company. 7. Eluard, R., and G. Simonet. (Sept.) 1970. La purificacion de hydrogene par adsorption a temperature ambiante. Chimie et Industrie 15:3–8. 8. Monreau, C. 2000. The air liquide compact VSAy, 147–157.

ADDITIONAL READING ON NONCRYOGENIC SEPARATIONS 1. 2. 3. 4.

Hairton, D. 2000. Membranes put the squeeze on cryogenics. Chem Eng (March):33–39. Monsanto. 2005. Prism separators. Bulletin EPD-5-073. DuPont=Air Liquide. 2004. Medal membrane separation system. Bulletin ABCISS=XANADU France. Schell, W.J., and C.D. Houston. 1982. Process gas with selective membranes. Hydrocarbon Process (September):249–252. 5. Shaver, K.G., G.L. Poffenbarger, and D.R. Grotewold. 1991. Membranes recover hydrogen. Hydrocarbon Process, (June):77–80. 6. Webster, J., and A.E. Hodel. 1986. PSA technology cost effective in meeting plant nitrogen needs for critical process inert gas blanketing. Chem Process (July):97–98. 7. Smolarek, J., et al. 1997. Single bed pressure swing adsorption for recovery of oxygen from air. US Patent 5,658,371, August 19, 1997.


8. Knaebel, K.S. 1995. For your next separation consider adsorption. Chem Eng (November):92–102. 9. Ackley, M.W., F. Notaro, and J. Smolarek. (April) 1999. Recover industrial gases via adsorption. Chem Eng (April):70–76. 10. Linde, A.G. 2005. VPSA for adsorption plants for production of oxygen. Process & Engineering Division Bulletin. 11. Linde, A.G. 2003. PSA plants for production of nitrogen. Process & Engineering Division, Bulletin. 12. Air Products & Chemicals. 2005. The next generation of PSA technology, Publication No. 522–8914.


10.1

Cryogenic Equipment, Materials, and Machinery

The separation of air involves more than thermodynamic calculations. It also involves the use of equipment such as heat exchangers, valves, compressors, expanders, pumps, etc. The individual operability of these items in terms of process efficiency and operating reliability is of fundamental importance in the choice of any cryogenic process cycle. Moreover, much of this equipment must operate in a reliable and efficient manner at temperatures well below 173 K (1008C).

10.1.1 HEAT EXCHANGERS 10.1.1.1 GENERAL In a general overview of cryogenic gas separation processes, one must give top priority to the importance of the heat exchanger system. Many process engineers spend a great deal of time and effort on the design of the distillation columns, but they should also remember that the external energy provided to reach very low temperatures required for the distillation process must be recovered through an efficient and a thoughtfully calculated heat exchange system to arrive at an economical process. Credit must be given to William Siemens,1 who as far back as 1857 conceived the idea of heat recovery and its application in metallurgical operations. He developed the use of heat recuperators in the smelting of steel, and this concept was later applied in the form of heat exchangers for other industries. Credit must also be given to Hampson who conceived the idea of a spiral-wound design for an efficient heat exchanger in his laboratory. This was further developed by Heyland and was finally put into industrial practice by Linde. As the early pioneers of air separation found out, an efficient heat exchanger system was mandatory for a cost-effective cryogenic process.

10.1.2 PARAMETERS OF DESIGN Innumerable textbooks and articles have been published on the subject of heat transfer and its equipment, so there is no need to dwell too long on this subject. Nevertheless, it may be wise to present some basic facts as applied to air separation. Heat exchangers operate on the basic principle of heat transfer from one fluid to another across a common wall. The transfer of heat across this common wall is analogous to an electric circuit. Ohm’s law states that I ¼ E=R, or I the current (heat transfer) is directly proportional to E, the voltage drop (temperature differential), and inversely proportional to the electric resistance, or the thermal resistance of the material forming the common wall. It should be noted, however, that the heat-transfer literature uses the reciprocal of thermal resistance, and this thermal conductance is then


reported as the product of a coefficient of heat transfer multiplied by the surface area between the streams exchanging heat. I ¼ E ¼ R ¼ Or I

heat transfer in kcal=unit of time T1T2 in Kelvin thermal resistance in kcal=unit of time ¼ (T1 T2)=R, or heat transfer ¼ DT thermal conductivity

In the design of air separation, and other cryogenic separation processes, engineers strive for a maximum heat transfer between incoming and outgoing fluids across a common wall separating them. Therefore, they calculate for a minimum temperature differential (T1 T4) at both ends of the exchanger, and within economic limits select an exchanger material with a maximum coefficient of thermal conductivity. In such calculations, the coefficient of heat transfer of the subject fluids has to be taken into consideration. In summary, the following parameters must be studied to reach an optimum combination of process efficiency and capital cost: . . . .

Small temperature differences between inlet and outlet streams to increase efficiency Large surface area to material volume ratio to minimize heat gain High heat transfer to reduce surface area Low pressure drop to reduce compressor energy to a minimum

Consider Figure 10.1.1 operating at constant pressure and NB receiving heat from NA; one shall accept the fact that according to the first law of thermodynamics NA (hA2 hA1 ) þ NB (hB2 hB1 ) þ Q ¼ 0

(10:1:1a)

The second law also demands, however, that if B stream receives heat then the following inequality must be satisfied at all cross sections of the heat exchanger, or tA > tB

(10:1:1b)

The above ratio is self evident, yet in some calculations it is easy to overlook a violation of the second law. This violation can happen whenever there is a change of phase in one of the streams, or even when there is no change in phase in either of the two streams. More details will be found in Section 10.1.7 on operability.

NA Cold end NB

T A2

Hot end

T B1

NB

T B2

N A T A1

FIGURE 10.1.1 Typical classical exchanger. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


10.1.3 BASIC PRINCIPLES Consider a simple countercurrent heat exchanger with m0 moles of warm air (A) traversing the inner wall of the shell of the exchanger, and m00 moles of cold nitrogen (N) passing through the tubes countercurrent to the air stream (Figure 10.1.2). The air stream being warmer at T1 decreases in temperature to T2. The nitrogen gas on the other hand increases in temperature from T3 to T4. In these conditions, therefore, the quantity of heat absorbed by the nitrogen is equal to the quantity of heat extracted from the air stream (neglecting all losses). The air feed is cooled from T1 to T2, therefore Q (heat extracted) ¼ m0 cp DT ¼ (h1 h2 )m0

(10:1:2)

where h1 is the total heat or enthalpy of m0 moles of air at T1 and h2 is that at T2, and cp the specific heat does not vary with temperature. Moreover, the nitrogen of m00 moles is heated from T3 to T4 and has absorbed Q00 ¼ (h10 h200 ) m00 (kJ=kg) of energy. And neglecting all losses (h01 h02 )m0 ¼ (h004 h003 )m00

(10:1:3)

One may conclude, therefore, that the total quantity of heat received by one fluid is equal to that extracted from the fluid in counterflow. The above equation is not only valid for the exchanger as a whole, but also for any point or cross section along the heat exchanger. And since for a specific pressure the enthalpy (h) is a sole function of temperature (T), the calculations can be carried out by using temperatures only, or (T1 T2 )m0 ¼ (T4 T3 )m00

(10:1:4)

when the two streams have the same specific heats. It follows, therefore, that if three out of four temperature points are known, either the fourth point, or m0=m00 can be determined. The following factors and correlations are used in the process design of heat exchangers in the cryogenic industry:2

Th1

dO

Tc2 Temperature

dT h Th2 dT C Tc1 dA

Area

FIGURE 10.1.2 Basic principle of classical exchanger. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)


1. Prandtl number, Npr ¼ mcp=kt where m is the fluid viscosity; cp is the fluid specific heat at constant pressure; kt is the fluid thermal conductivity. 2=3 2. Colburn J factor, jH ¼ (hc=Gcp) NPr where hc is the film coefficient of heat transfer; G ¼ m=Aff is mass flow rate per cross sectional (free-flow) area. 3. Nusselt number NNu ¼ hcDe=kt where De is equivalent diameter of the flow passage through which the fluid is flowing. In a circular tube, it is the internal diameter. 4. Reynolds number NRe ¼ DeG=(m defined in 1) 5. Friction factor f ¼ (Dp=L) (G22gcrDe) where Dp=L is the frictional pressure drop per unit length; gc ¼ 1 kg m=N s2; r is the fluid density. The Prandtl number will appear in many equations, but as it involves viscosity (m), specific heat at constant pressure (cp), and fluid thermal conductivity (k), it can be readily and easily calculated from data retrieved from available thermodynamic tables. Most tables include the Prandtl number for a given temperature. G is defined in factor 2 above. For heat exchangers with flow inside circular tubes the following correlations can apply: For laminar flow (NRe is less than 2300) with De as inside diameter of tube, and L as the tube length (Hausen, 1944),2 and where fluid properties are evaluated at the mixed–mean temperature of the fluid, i.e., 1=2(T1 þ T2), according to Barron:2 NNu ¼ 3:658 þ [0:0668 (De =L) NRe NPr ]={1 þ 0:04 [(De =L)RRe NPr ]2=3 }

(10:1:5)

for turbulent flows, where NRe is greater than 3000, De is the inside diameter and Dh is the diameter of the helix of a wound-coil exchanger (Colburn, 1933),3 and where the fluid properties are evaluated at the mean film temperature Tm ¼ 1=2(Tb þ Tw), which is the average between the bulk fluid temperature Tb, and the mean tube-wall temperature Tw: 0:2 [1 þ 3:5(De =Dh )] jh ¼ 0:023NRe

(10:1:6)

For straight and long circular tubes, Dh is infinite, therefore De=Dh ¼ 0. Since calculations for heat exchangers involve varied equipment, material, and configurations, theoretical formulas are modified to fit the fabricated element. Therefore, some empirical formulas are given below for calculating heat transfer and corresponding pressure drop for concentric tubes. Straight tubular pipe Flow inside, no phase change, NRe < 2300 h ¼ {3:658 þ [0:0668(De =L)NRe NPr ]=(1 þ 0:04)[(De =L)NRe NPr ]2=3 }k=De Dp=DL ¼ 32G2 =NRe gc De r Straight tubular pipe Flow inside, no phase change 2,100 < NRe < 10,000


(10:1:7) (10:1:8)

2=3

h ¼ 2:439 106 (NRe 125)[1 þ (De =L)2=3 ](cp =De )(mf =mb )0:14 0:2 Dp=DL ¼ 0:158G2 =(NRe gc De r)

(10:1:9) (10:1:10)

Straight tubular pipe Flow inside, no phase change NRe > 10,000 2=3

0:2 NPr h ¼ 0:023cp GNRe

0:2 Dp=DL ¼ (0:092G2 )=(NRe gc De r)

(10:1:11) (10:1:12)

Spiral wound tubular pipe Flow inside, no phase change NRe > 10,000 2=3

0:2 NPr h ¼ 0:023cp GNRe

[1 þ 3:5(De =Dh )]

0:2 Dp=DL ¼ 0:092G2max [1 þ 3:5(De =Dh )]=(NRe gc De r)

(10:1:13) (10:1:14)

In all the above equations, De is the inside pipe diameter; for annulus De ¼ D2 D1 where D2 is the outside diameter, and D1 is the inside diameter; Dh is the diameter of helix (wound coil); all fluid properties are evaluated at mean film temperature Tm ¼ 1=2 (Tw þ Tb), where Tw is the wall temperature, and Tb is the bulk temperature of fluid. G is the mass flow per unit of area; h is the enthalpy. For noncircular piping De ¼ 4(AcL=Aw) where Ac is the inside free-flow cross-sectional area, L the length of pipe, and Aw the heat transfer or wetted tube surface area; mf the viscosity at mean fluid temperature, and mb the viscosity at bulk fluid temperature. Constants for equations are with SI units. For flow inside smooth tubes, the following correlations apply for friction factor, f. For laminar flow NRe less than 2300 f ¼ 64=NRe

(10:1:15)

For turbulent flow, NRe is greater than 3,000 but less than 50,000 0:25 f ¼ 0:316NRe

(10:1:16)

0:20 f ¼ 0:184NRe

(10:1:17)

For a flow, NRe of more than 50,000

In an exchanger with noncircular tubes, the above mentioned correlations for turbulent flow apply if the following equivalent diameters are used instead of the circular-tube diameter. For heat-transfer correlations De ¼ 4Aff L=A


(10:1:18)

where Aff is the tube free-flow (cross-sectional)area, L is the tube length, and A is the area wherein heat is transferred. For example, for a square tube heated on all four sides De ¼ 4Aff L=Aw

(10:1:19)

where Aw is the wall surface area that is wetted by the fluid. For a flow normal to banks of tubes, the following correlations apply: For banks of staggered tubes (Colburn, 1933)3 0:4 jH ¼ 0:33NRe

(10:1:20)

The Reynolds number is defined as NRe ¼ Do Gmax=m, where Do is the outside diameter of the tubes; Gmax ¼ mAff; Aff is the minimum free-flow area between the tubes, the latter being the minimum free-flow area between the tubes. The fluid properties are evaluated at the mean film temperature Tm ¼ 1=2(Tb þ Tw) where Tb is the average fluid bulk temperature, and Tw is the wall temperature. For banks of tubes in line, as in spherical-wound exchangers (Colburn, 1933),3 the Reynolds number is defined as in Equation 10.1.20, and the fluid properties are evaluated at Tm 0:4 jH ¼ 0:26NRe

(10:1:21)

For flow across the tubes, the friction factor is different than for flow inside tubes. For flow outside the tubes, where N is the total number of tubes in line across which the fluid flows, f1 ¼ DpN=G2max 2gc r

(10:1:22)

0:16 f 0 ¼ [1 þ 0:470(XT 1)1:08 ]NRe

(10:1:23)

For staggered tubes,

where XT is the transverse pitch–tube outside diameter (Jacob 1938).4 For in-line tubes (Jacob, 1938), 0:15 f ¼ [0:176 þ 0:32XL (XT 1)n ]NRe

(10:1:24)

where XL is the longitudinal pitch–tube outside diameter, n ¼ 0:43 þ (1:13=XL ) See Figure 10.1.3 for the definition of the longitudinal and transverse pitch for (a) tubes in line and (b) staggered-tube arrangement.

10.1.4 TYPICAL EXAMPLE FOR DESIGNING TUBULAR HEAT EXCHANGERS As an example in the use of the aforementioned parameters for the design of tubular heat exchangers, assume that 60 g=s of cold nitrogen at a temperature of 94 K, is flowing through a straight tube with an inside diameter of 25 mm and with a wall at a temperature of 106 K. The


Longitudinal pitch Flow direction

Transverse pitch

(a) Longitudinal pitch Minimum flow area

Transverse pitch Flow direction

(b)

FIGURE 10.1.3 Definition of longitudinal and transverse pitch for (a) tubes in line and (b) for staggeredtube arrangement. (From Barron, R.F., Cryogenic Systems, 2nd ed., Oxford University Press, New York, 1985. With permission.)

internal fluid pressure is at 150 kPa. What are the heat-transfer coefficient, the heat-transfer rate, and the pressure drop? First of all, one has to look over the properties of the fluid at its mean temperature, which is Tm ¼ 1=2(94 þ 106), or 100 K from Table D-4 in Appendix D, which are: . . . . .

cp ¼ 1.067 kJ=kg K kt ¼ 9.33 mW=m K NPr ¼ 0.797 r ¼ 3.484 kg=m3 m¼ 6.98 m Pa s _ _

Specific heat Thermal conductivity Prandtl number Density Viscosity

Then follow this procedure: 1. Determine mass flow=unit of free-flow area, or G ¼ m=Aff G ¼ 0.060=(p=4)(0.025)2 122.231 kg=m2 2. Then determine Reynolds number from NRe ¼ DeG=m (0.025)(122.231)=6.98 mPa ¼ 437,790 (turbulent) 0.2 3. Calculate Colburn factor from jH ¼ 0.023NRe [1 þ 3.5(De=Dh)] from Equation 10.1.5, but since Dh is infinity (straight tube) jH ¼ 0.023(437,790)0.20 [1 þ 0] ¼ 0.0017 2=3 4. jH ¼ hcNPr =Gcp, therefore, hc ¼ (0.0017)(122.231)(1067)=(0.797)2=3 ¼ 257.92 W=m2 K


5. Heat-transfer rate to the nitrogen per unit area is Q ¼ hc(Tw Tb) ¼ (257.92)(106 – 94) ¼ 3095 W=m2 6. The friction factor for a turbulent flow from Equation 10.1.17 0.20 f ¼ 0.184NRe , or (0.184)(437,790)0.20 ¼ 0.0137 7. r of the gas at 150 kPa and 100 K ¼ (3.484)(150=101.325) ¼ 5.1577 kg=m3 8. Therefore the pressure drop=unit length is Dp=L ¼ fG2=2gcrDe, from Equation 10.1.12 ¼ (0.0137)(122.231)2=(2)(1) (5.1577)(0.025) ¼ 793.7 Pa=m, or 0.7937 kPa=m

10.1.5 BRAZED ALUMINUM HEAT EXCHANGERS As noted in the design of tubular heat exchangers, the process engineer has to determine the heat-transfer coefficient, the heat-transfer rate, and the friction factor, so as to specify a suitable heat-transfer surface and configuration. The same parameters are also involved for the design of brazed aluminum heat exchangers (BAHX). Additional problems present themselves, however, because the fabrication of these exchangers is restricted by their cubiform geometry, and their highly specialized manufacture. In this regard, the process engineer has to work in concert with the supplier, selecting a suitable design that will offer the optimum Colburn factor jH, and friction factor f in correlation with the Reynolds number. Depending on the operating pressure and temperatures involved, the selection of a suitable material is also crucial, but limited to either an aluminum alloy, or a stainless steel. Once these parameters are decided, the designer can proceed with the calculation of the specific heat, the rate of heat transfer, the pressure drops, and the final total heat-transfer surface. Although the process design of the BAHX follows, in principle, the same process criteria as those used for tubular exchangers, it also imposes a few other important factors, such as . . . .

Limitation of a maximum usable pressure More accurate calculation of heat transfer Precise arrangement of the metallic layers for optimum heat transfer Correct use of distribution of fluids for single and two-phase streams

In general, the process involves the heat-transfer coefficient and pressure drop based on the experimentally correlated Colburn coefficient jH, and friction factor F , which are then correlated against the Reynolds number, as has been indicated previously for the tubular exchangers. Understandably, moreover, the manufacturers of these exchangers employ both past experience as well as a highly advanced computer program to achieve their results. The same procedure is carried out for brazed aluminum exchangers as for the tubular type. Once the process designer agrees with the supplier of the exchangers on the specific Colburn factor jH, the coefficient of heat transfer, the friction factor, and the correlation with the Reynolds number that will apply, then the specific geometry of the plate fin can be selected. Moreover, the cross-sectional area of the fin is also to be used as the (p=4De2) previously used for tubular exchangers. Then, the cryogenic fluid properties at the given temperatures from standard tables are applied as before.

10.1.6 EFFECTIVENESS («) The effectiveness of a heat exchanger may be defined as « ¼ actual energy transferred=maximum possible energy transfer The maximum possible temperature change between a cold and a hot stream passing through a common heat exchanger is governed by the second law of thermodynamics, Th1 Tc1,


which states that the cold stream cannot achieve a higher temperature than that of the warm stream at the inlet; and that the warm stream cannot cool down to a temperature lower than that of the cold stream at the outlet. The effectiveness «, therefore, cannot be greater than unity. The influence of heat exchange effectiveness cannot be over emphasized in the performance of any cryogenic process. It can be shown with the use of a temperature–entropy diagram that if the effectiveness of a heat exchanger is below design, the temperature of the outgoing fluid will be lower. This condition will result in a substantial decrease in refrigeration recovery, and in a higher temperature differential at the warm end of the exchanger. The latter condition will lead to a loss of refrigeration, and therefore, to a higher use of energy. It may be stated that the maximum temperature change occurs in the fluid with the minimum capacity rate, or Cmin ¼ (mcp). In other words, the heat lost by the hot fluid equals the heat gained by the cold fluid, or (mcp )c (Tc2 Tc1 ) ¼ (mcp )h (Th2 Th1 )

(10:1:25)

where Tc refers to the cold fluid and Th refers to the hot fluid. Assuming an ideal theoretical situation, the temperature of the cold fluid leaving the exchanger could increase to that of the hot fluid entering the exchanger, or the temperature of the hot fluid leaving the exchanger at the other end could decrease to that of the cold fluid entering the unit at the same end. In other words the maximum temperature change is equal to Th1 Tc1, and the maximum possible energy transfer rate is Cmin (Th1 Tc1), where C ¼ mcp or capacity rate. It is possible, therefore, to express heat exchanger effectiveness in reference to either the hot or cold fluids, respectively, as indicated by Barron2: For hot liquids « ¼ Ch (Th1 Th2 )=Cmin (Th1 Tc1 )

(10:1:26a)

« ¼ Cc (Tc2 Tc1 )=Cmin (Th1 Tc1 )

(10:1:26b)

and for cold liquids

and if the hot fluid has been assumed to have the minimum transfer capacity, the effectiveness can be reduced to « ¼ (Th1 Th2 )=(Th1 Tc1 )

(10:1:27)

In a normal practical application, the validity of any cryogenic process cycle is directly influenced by the effectiveness of the heat exchangers employed in the system. Consider a simple high-pressure Linde liquefier cycle using nitrogen as the liquefaction medium (Figure 10.1.4), and with a heat exchanger system that is not ideal. In this example, points 1, 2, 3, and 4 are the state points for the 100% effective system; and points 10 , 20 , 30 , and 40 are the state points for the same system, but with the nonideal heat exchanger system. The recycled cold nitrogen will leave the heat exchanger at a condition represented by state 10 , instead of the warmer condition represented by state 1. If one defines Cmin ¼ Cc, and if one sets T1 ¼ T10 , the heat exchanger effectiveness becomes « ¼ Cc (T10 Tg )=Cmin (T1 Tg ) ¼ (h1 hg )=(h01 hg )


(10:1:28)

2750 2500

Enthalpy change (chu per lb. mole of air)

2250

Minimum ratio of N2 to air

2000 1750

Nitrogen heating curve

1500 1250 1000 750 500

Air cooling curve

250 0

50

75

100

125

150 175 200 Temperature (K)

225

250

275

300

FIGURE 10.1.4 Violation of the second law of thermodynamics in a heat exchanger in which no phase change occurs in either fluid. (From Dodge, B.F., Chemical Engineering Series, McGraw-Hill, New York, 1944. With permission.)

where the h is the enthalpy at the given state denoted by the associated subscript, and the f and g represent the saturated liquid and saturated vapor states, respectively. The resulting liquid yield y, therefore, now is y ¼ (h01 h2 )=(h01 hf ):

(10:1:29)

By solving for h1 in Equation 10.28, and substituting the resulting relation into Equation 10.29, the liquid yield may be expressed in terms of the heat exchanger effectiveness as y ¼ [(h1 h2 ) (1 «)(h1 hg )]=[(h1 hf ) (1 e)(h1 hg )]

(10:1:30)

where hf and hg are the saturated liquid and vapors, respectively. The ineffectiveness of the heat exchanger system will increase the energy required for the system by the amount of DW ¼ m(h1 h01 )

(10:1:31)

Solving for h1 in Equation 10.30, and substituting into Equation 10.31, the additional work becomes DW ¼ m(h1 hg )(1 «)


(10:1:32)

Example: Refer to Figure 10.1.4 describing a high-pressure plant producing liquid using nitrogen as the working medium for the production of refrigeration. Find the additional work required in nonideal conditions. Inlet pressure into the compressor is 1.013 bar; outlet compressor pressure is 202.7 bar; inlet compressor temperature is 300 K; effectiveness, « is 96.5%. From table of properties, enthalpy for nitrogen is h1 h2 hg hf

¼ ¼ ¼ ¼

462 J=g at 1.013 bar and 300 K 432 J=g at 203 bar and 300 K 229 J=g at 1.013 bar at saturated vapor 29 J=g at saturated liquid

Since « ¼ (h10 hg)=(h1 hg) h01 ¼ hg þ (h1 hg ) or 229 þ 0:965(462 229) ¼ 453:84 kJ=kg and the yield y from Equation 10.1.29 is (453.84 432)(453.84 29) ¼ 0.0514. In other words, the additional energy required is directly influenced by the heat exchanger effectiveness.

10.1.7 OPERABILITY As noted, the driving force in a heat exchanger depends on the temperature differential between the incoming and the outgoing fluids at the same point, or cross section of the exchanger. Cryogenic heat exchangers are generally designed to operate with relatively low temperature differentials for reasons of maximum heat (energy) recovery. It is important, therefore, to review the overall parameters of ‘‘operability’’ of the heat exchanger before any decision is made in its selection.5 In the design of a simple countercurrent heat exchanger (refer to Figure 10.1.1) operating at a constant pressure and with only two fluids, two cardinal laws of thermodynamics must be obeyed. The first law is the standard rule of heat exchange and can be stated as m0A (hA2 hA1 ) þ m00B (hB1 hB2 ) þ Q ¼ 0

(10:1:1)

The second law demands, however, that if A is the fluid receiving the heat, then the following inequality be satisfied at all cross sections of the exchanger TB > TA

(10:1:2)

that is TB TA must always be positive. It is not enough, therefore, for a designer to calculate the temperature differentials at the warm and cold ends of the heat exchanger. The differentials at the ends may be positive, but the exchanger may be inoperable. An enthalpy–temperature curve when plotted may indicate a negative temperature difference, called a crossover, at some point or cross section within the exchanger. Such a crossover may occur, moreover, whether or not there is a change of phase in one of the fluids. In the design of critical heat exchangers, therefore, the plotting of an enthalpy–temperature curve is mandatory. This is especially true in the cryogenic industry where process designers strive for very close temperature approaches. As a general rule, one may state that


Enthalpy change (chu per lb. mole of gas at 10 atm)

1600 1400 Warm end Δt

1200 1000 .

m

0 t1

800 O2

600

at

.

tm

a t1

a

O2

a

Cold end Δt

400 200 0 100

120

140

160

180

200

220

240

260

280

300

Temperature (K)

FIGURE 10.1.5 Violation of the second law of thermodynamics in a heat exchanger (case of a phase change in one of the fluids). (From Dodge, B.F., Chemical Engineering Series, McGraw-Hill, New York, 1944. With permission.) When the fluid receiving heat, say product oxygen, in an air separation system, has the greater heat capacity, only the cold end DT can be made to approach zero. When it has the smaller heat capacity, only the warm end DT can be made to approach zero.

Refer to Figure 10.1.4. Dodge has also stated5 that the same possibility of a temperature crossover may apply due to a violation of the second law of thermodynamics (Figure 10.1.5).

10.1.8 EFFICIENCY (h)5 The efficiency of a heat exchanger may be defined on the basis of the temperature differential, or DT, at one end of the exchanger. The closer the approach, the higher the efficiency. When one succeeds in bringing both temperatures to DT apart, then the efficiency will be 100%; but this is impossible given the limitations of the second law. The efficiency of a heat exchanger may be defined by the following expression: h ¼ the actual temperature rise of the outgoing fluid=the maximum possible rise of temperature, or from Figure 10.1.5. h ¼ (TA2 TA1 )=(TB2 TA1 )

(10:1:3)

One may state, therefore, that if the enthalpy lines of the two fluids are essentially the same, then: a. If the heat capacity of the fluid receiving heat is less than that delivering heat (fluid A), the warm end can be made to approach zero, and h ¼ (TA2 TA1 )=(TB2 TA1 )


(10:1:33)

b. If the heat capacity of the fluid receiving heat is greater than that delivering heat (fluid A) the cold end can be made to approach zero, and

h ¼ (TB2 TB1 )=(TB2 TA1 )

(10:1:34)

10.1.9 INDUSTRIAL APPLICATIONS The original heat exchanger design as applied to cryogenics was a straight shell and tube design with intermediate baffles to improve heat transfer. By directing the warmer fluid (air) at right angles to the tubes carrying the colder product fluid (nitrogen or oxygen), there was a gain of heat transfer. These heat exchangers were made of copper and lead soldered together. Thanks to Hampson, such exchangers soon gave way to the spiral-wound coil design (Figure 10.1.6). With the latter design, which is still in use today for certain applications, it is possible to create an exchanger of very long tubes within a relatively small outside volume. This design also permits the warmer fluid to flow always at right angles to the tubes without using special baffles. It also conforms to the principle that heat transfer improves when the tubes are longer and their diameter is smaller. This concept, however, also increases investment. The spiral-wound coil exchanger also involved copper and lead soldering. As the art of lead soldering declined, especially after World War II, the number of leaks increased, especially when the components became larger. In the early 1950s, however, this mode of assembly was quickly replaced by the use of silver soldering, which proved more reliable against leaks.

FIGURE 10.1.6 Hand coiled heat exchanger for LNG project. (ß Air Liquide, all rights reserved, 2006. With permission.)


However, as the plants grew larger and larger, the weight of the tube bundle when in a horizontal position during fabrication became a problem, resulting in unacceptable deflections. For this reason, and for very large base load liquefied natural gas plants, wound coil exchangers were and are still being built, but using aluminum tubing in place of copper. For very large exchangers, the coiling is carried out around a vertical axis.

10.1.10 DEVELOPMENT OF BRAZED ALUMINUM HEAT EXCHANGERS During the 1950s the use of copper exchangers became less prevalent, and these were replaced to a large degree with extended surface plate and fin exchangers made of brazed aluminum alloys. The latter, however, also have the following limitations as those prescribed by the existing codes (Table 10.1.1).

10.1.10.1 PRESSURE LIMITATIONS Maximum working pressure limitations vary in accordance to the metal alloy used for the fabrication of brazed heat exchangers, whether it is an aluminum alloy or stainless steel, and for the application of the exchangers. For example, for industrial gas production, the pressure may range from 1 to 60 barA. For petrochemical production and natural gas processing including nitrogen rejection and helium recovery, the pressures may range from 1 to 75 barA, and even up to 100 barA. This design consists of thin aluminum sheets or plates separated by a series of corrugated aluminum sheets forming the passages. These passages may be separated and manifold so that any number of fluids can be handled in one complete assembly or core as it is generally called. For example, in an air separation plant the air feed passes countercurrent to pure oxygen, pure nitrogen, argon, and waste nitrogen. Each product has its own manifold at each end of the exchanger (Figure 10.1.7a and Figure 10.1.7b). According to Chart Industries,6 these exchangers are manufactured by the use of aluminum alloy sheets having a width of up to 1300 mm, and a length up to 6.1 m. They are separated by specially designed fins that have a thickness ranging from 0.15 to 0.8 mm. The separations between the parting sheets vary between 3.8 and 15.7 mm. The fins can be plain, perforated, serrated or lanced, or herringbone. Prior to the assembly, all material is checked for quality and is cleaned with the utmost care. The core is then clamped together and placed in a high-vacuum furnace where it is brazed with an aluminum brazing compound at approximately 873 K. After the brazing operation, the core is removed and the various attachments, such as nozzles, headers, etc., are welded together by BTAW or GMAW techniques. Posttesting procedures on the finished core include x-raying, pressure and helium tests, and, if requested, flow testing. Depending on

TABLE 10.1.1 Aluminum Alloys Used for Brazed Heat Exchangers Alloy

Temperature Limitations ASME (K)

AD-MERKBLATT=VdTuv (K)

3003 5083

4=477 4=338

4=338 4=353

For more details on metallurgical analysis and specific application, the reader is referred to ‘‘Courtesy of The Standards of the Brazed Aluminum Plate-Find Exchanger Manufacturer’s Association’’, 1994 Edition. With permission.


Plain fins

Plain-perforated fins

Serrated fins

Herringbone or wavy fins

(a)

Cross-counterflow with internal turnaround (Serpentine)

(b)

Cross-flow

FIGURE 10.1.7 (a) Variety of horizontal fin pattern types. (b) Cross-section of two brazed aluminum heat exchangers: Cross-counterflow and Cross-flow. (Courtesy of Chart Energy & Chemicals, Inc., 2006. With permission.) ß 2006 by Taylor & Francis Group, LLC.

the specifications sent by the client, a multiple of these cores can be welded together as a complete prepackaged assembly for cost-effective shipping and erection purposes. The fact that these exchangers can be designed and built in a wide variety of configurations and applications has made them a first choice for low-temperature processes. They are used not only for countercurrent and crossflow heat exchange, but also as main condensers, auxiliary vaporizers, and liquid subcoolers. They are still being used as reversing exchangers for cryogenic pre-purification, but to a rapidly diminishing degree. This situation has been explained in more detail in Chapter 5 covering pre-purification and adsorption. Originally, the brazing was carried out in a salt bath. After brazing, the core was removed and cleaned out with clear water. This procedure, however, was not found to be completely foolproof, and after 5–10 years of operation, complaints of leaks due to corrosion arose from the field. This situation was also aggravated by the fact that the principal application of these cores was for freezing out impurities in air separation plants by the use of reversing valves. The reversal of pressure took place every 15 min, which meant approximately 35,000 reversals per year. This mode of operation imposed a serious strain on the thin passages of the core, which already may have been weakened by the effects of corrosion. In these conditions, a reversing exchanger core had an average operational life of approximately 7 years, which was definitely not acceptable by users. An intensive study on this problem was carried out by the various companies engaged in the cryogenic industry, and the reader is referred to a comprehensive review made by both the International Coal Refining Company, and Air Products and Chemicals, Inc. The paper was presented at the International Corrosion Forum sponsored by the National Association of Corrosion Engineers during April 6–10, 1981, at Toronto, Canada. One of the hypotheses suggested by this study was that ‘‘hypochlorous acid is the most probable species responsible for the initiation and propagation of corrosion on aluminum in air separation plant reversing heat exchangers.’’ By changing over to high-vacuum brazing, however, manufacturers have overcome the problem, and the cores manufactured by the new process have had so far an excellent record in the field, even for pre-purification with reversing exchangers. Nevertheless, the majority of the air separation plants built over the last 25 years have employed and are opting for adsorption for pre-purification, the process which has been found to be more reliable.

10.1.11 VACUUM BRAZED HEAT EXCHANGERS As noted, the free assembly of the various aluminum sheets and fins is heated to brazing temperature in a vacuum furnace by means of radiant heat applied to the surface, and by heat conduction in the interior of the assembly. The aluminum solder required for the brazing operation is roller-coated on both sides of the parting sheets. It melts at approximately 873 K. All the assembly elements are wetted by the fused solder. The entire operation is carried out in a vacuum to make sure that no aluminum oxides are formed during brazing at the high temperature. During the cool-down period, the aluminum solder that has penetrated into the main material solidifies and become unidentifiable with the various component parts. The resultant completed assembly is then removed and is welded to the various collectors and ancillary piping manifolds. The extended surface plate and fin exchanger cores, being light in weight, compact, highly efficient in transferring heat, cost effective, and which offer the possibility of vertical installation, provide many advantages in the air separation, chemical, and petrochemical industries.


10.1.12 MECHANICAL CONSTRUCTION The cores have a high heating surface density because of the compactness of the plates and fins. A unit core can pack up to 1000 m2 of heat exchange surface per cubic meter of volume, and up to a maximum of 7000 m2 of heat exchange surface, compared with a shell and tube heat exchanger that has a heat exchange surface capacity of 100 m2=m3, and a maximum surface of 2000 m2 per unit. Moreover, with the use of stainless steel in place of aluminum, operating pressures up to 102 bar have been achieved. Admittedly, the selection of BAHX involves the use of complex design calculations, a rigid configuration, and the use of relatively clean process streams. On the other hand, they offer the advantages of close temperature approaches of less than 58C; true countercurrent flow (even crossflow when needed); and the capability of exchanging heat with multiple streams. The problem of the demand for higher mathematics in their design has been overcome to a certain extent with the use of computerized programs, the use of simulation techniques, plus guidance from the availability of the publication of ‘‘The Standards of the Brazed Aluminium Plate-Fin Heat Exchanger Manufacturers’ Association’’ (1994).7 The corrugations or fins as they are called serve both as heat-transfer surface and structural support. As shown in Figure 10.1.7a, the corrugations are made with a variety of configurations depending on the requirements of the process to be used, as well as on the specific application of the element. The height and density (fins per decimeter, or per meter), will be a function of the process fluid characteristics and the operating pressure. The composite heat exchanger itself is divided between the distributor and the heat-transfer area, as indicated in Figure 10.1.8. The distributor areas are plate fins that direct the fluid from the nozzles to the heat-transfer areas. They are designed for a very low pressure drop, less than 25% of the total allowed for the entire element, to avoid maldistribution that may affect the overall performance. Therefore, as stated by Isalski,8 the design of the nozzles and headers is very crucial. As noted, fins that compose the heat-transfer area are plain, plain-perforated, serrated, herringbone, or wavy. Their individual height may vary from 3.8 to 15.7 mm, the fin spacing from 24 to 100 per decimeter (240 per meter). With the use of a suitable stacking arrangement of the layers, one can diminish the pressure drop to a prespecified minimum. The stacking arrangement of the layers may become quite complicated if multiple fluids are used for the same layer. Nevertheless, even with this complicated arrangement, the unit still retains the same pressure drop division between the distributor and the heat-transfer area. In specifying the BAHX one must be careful to avoid a ‘‘secondary heat transfer,’’ especially when handling three streams, such as in air separation. To avoid this possible problem it is necessary to carry out the layering of the aluminum sheets with the utmost care. It should be kept in mind that waste nitrogen passing through the primary (front end) heat exchanger system has a higher capacity to transfer cold than product oxygen and any other smaller pure product, i.e. argon, according to Isalski.8 Another point to consider is that for handling single phase products, the distribution of the fluids in each layer must be done efficiently and with care in order to avoid dead zones. This can be done by designing the central core to operate with a slightly higher pressure drop than that at the end distributor. This will have a flushing effect, thus avoiding any dead zones. Careful attention also has to be paid to the distribution of fluids across the width of each layer, especially when two-phase flow is involved. For a two-phase system it is necessary to distribute both phases evenly into each layer as well as across each layer. This may be carried


Nozzle Header

Wear plate (rubbing plate)

Distributor fin Support angle

Side bar Cap sheet Parting sheet

FIGURE 10.1.8 Composite unit of brazed aluminum heat exchanger. (Courtesy of Chart Heat Exchangers. With permission.)

out by the use of impingement plates in the inlet header, which can add an extra pressure drop and promote a better distribution into the various layers, or alternatively the same can be done by using injection tubes (Figure 10.1.9 through Figure 10.1.11).

10.1.13 LIMITATIONS There are, however, some distinct disadvantages in the use of BAHX. For example, they should not be used for fluids that may contain an alkali with a pH higher than 8, or acids with a pH lesser than 5; copper containing media; mercury containing media; and of course, sea water. The BAHX, as it is generally referred to, should not be used for fluids that have particulate impurities, such as fine dust. These impurities, once they enter, are almost impossible to expel. Even if the exchangers are employed in a nonreversing mode, downstream of pre-purification adsorbers, the latter should have very efficient post filters, and the necessary valving to avoid a sudden high-speed flow of process-air pass through the adsorbent, which may pick up dust, and entrain it to the exchangers. About 10 years ago, this did happen, and although the nonreversing exchangers are still operating, they do so with a high pressure drop and little possibility of improvement. Another limitation is the use of sea water in any upstream piece of equipment that may find its way into the exchangers.


FIGURE 10.1.9 Large size brazed aluminum heat exchanger. (Courtesy of Chart Heat Exchangers. With permission.)

10.1.14 OPERATION AND MAINTENANCE6 As with any piece of thermal equipment, and especially in cryogenic operations, the heat exchangers must be brought to or from operating temperatures and pressures gradually. A minimum of 30 min must be allowed for lowering the heat exchanger temperature from ambient to cryogenic operating temperatures, or for warming up the exchangers from cryogenic to ambient temperatures. Valves must be opened and closed gradually to avoid any sudden change in temperature or pressure (unless the unit is specifically designed for this type of service).

FIGURE 10.1.10 Large multi component module BAHX hauled by truck. (Courtesy of Chart Heat Exchangers. With permission.)


FIGURE 10.1.11 Large component of BAHX raised during construction. (Courtesy of Chart Heat Exchangers. With permission.)

FIGURE 10.1.12 An endoscope reveals a clear exchanger. (Courtesy of Chart Heat Exchangers. With permission.)


FIGURE 10.1.13 An endoscope reveals a blocked exchanger. (Courtesy of Chart Heat Exchangers. With permission.)

Precautions must also be taken to prevent the transmission of operating fluid pulsations or vibrations (emanating from pumps, compressors, etc.) to the heat exchanger. If a prolonged or indefinite shutdown is scheduled, the unit should be mothballed. A low dew point should be maintained in the exchanger system to prevent any galvanic corrosion, especially if any traces of rust or other foreign material have gained access during operations. A small positive pressure should be maintained internally and continuously with the use of dry nitrogen. If any leak or fouling does develop, no attempt should be made to repair it without the assistance of a service team from the manufacturer of the exchanger, or by specialized service organizations. Fouling can be avoided by the use of a filter with a 177 mm (80 Mesh Tyler Standard) screen upstream of the exchanger. The use of a double filter system allowing each filter to be cleaned without leading to a plant shutdown will be the cost-effective choice. Retro-blowing (puffing) has also been used to clear up any fouling occurrence, but this has to be carried out with care to avoid exceeding the design pressure of the exchanger. Incidentally, the use of an endoscope is very helpful to determine if any passages are blocked or not. Figure 10.1.12 shows a clear passage and Figure 10.1.13 shows a blocked passage.

REFERENCES 1. Gomonet, E. 1952. Les basses temperatures, production et emplois, 54–60. J.B. Bailliere et Fils. Paris, France. 2. Barron, R.F. 1988. Cryogenic systems (Monogram on Cryogenics), 113–130. Oxford University Press, New York. 3. Colburn, A.P. 1933. Trans. Am. Inst. Chem. Eng. 33, pp. 174–210. 4. Jacob, M. 1938. Trans. ASME. Vol. 60, pp. 384–386. 5. Dodge, B.F. 1944. Chemical engineering thermodynamics, 365–369. McGraw Hill, New York. 6. Chart Industries ‘‘Users Operating Manual.’’ 7. The standards of the Brazed Aluminum Plant–Fin Heat Exchanger Manufacturers’ Association. 1944. 8. Isalski, W.H. 1989. Separation of gases (Monograph on Cryogenics), 261–262. Oxford: Clarendon Press.


FURTHER READING Diery, W.S. 1988. Potential use of vacuum-brazed exchangers, Linde Reports on Science and Technology, No. 44. Lunsford, K.M. 1996. Advantages of brazed aluminum heat exchangers. Hydrocarbon Processing, (July):55–63.


10.2

Expansion Machines

10.2.1 EXPANSION MACHINES 10.2.1.1 GENERAL Georges Claude successfully employed an expansion machine in air separation for the first time in 19021. The principle employed was well known at the time and was based on performing work during the expansion of a gas. With the extraction of energy from the gas, there is a drop in enthalpy of the expanded gas with little or no increase in entropy. Refer to Figure 2.6. The actual drop in temperature of the gas is a function of the gas component or mixture being expanded, the pressure ratio of expansion, the temperature and pressure at the inlet of the expander, and the mechanical efficiency of the expansion of the system. The efficiency of the expansion machine is a function of the rate of expansion, gas leakage, heat gain, and the pressure drop ratio through the machine. With small expansion ratios, the losses represent an appreciable percentage of the available enthalpy drop. In these conditions, expansion turbines are more efficient than reciprocating expanders because their heat gains are lower and valve losses are nonexistent. Normal adiabatic efficiency for an expansion turbine for a pressure ratio of 5:1 will be around 80%, whereas in this region, the efficiency of a reciprocating expander will be around 70%. If the pressure ratio is high, say 7:1, an expansion turbine becomes inefficient, and a two-stage expansion machine may have to be considered. On the other hand, a reciprocating expander has no limit, and the higher the enthalpy drop, the lower the effect of losses and the higher the efficiency of the machine. For example, the reciprocating machine has an advantage over turbines in the expansion of hydrogen for cryogenic systems. When expanding a gas from a very high pressure to an intermediate pressure of say 8 bar, and at a relatively warmer temperature, an expansion engine will attain efficiencies of 70% to 80%; but between an intermediate to a very low pressure (atmospheric), to arrive at the lowest possible temperature, the efficiency drops considerably. For example, in expanding hydrogen to atmospheric pressure to attain a temperature below 73 K the efficiency drops to 50%. When operating an expansion machine, one should always keep in mind that the power developed by an expansion machine is not an extra bonus. The power generated, whether it be electricity or compression, should be equal to heat-gain absorbed by the equipment in the cold box or the quantity of liquid products or both, if any are removed from the plant. A well-designed cryogenic plant should require a minimum power generation by the expansion machine. For example, a 350 t=d oxygen plant designed in 1949 developed 180 kW at the generator of the radial expander. Yet, in 1954, designed only four years later, an oxygen plant of a similar capacity generated 75 kW, indicating the advance in the design of cryogenic units.


10.2.2 RECIPROCATING EXPANSION ENGINE A reciprocating expansion engine may be used either as a single or a double acting unit depending on the process cycle. The design and operation is similar to a steam engine. The cylinder assembly is completely surrounded by insulation (mineral wool) and covered by an outer frame of cast iron. The valves are actuated by rocker arms operated by a system of cams mounted on external cam shafts directly geared to the crankshaft. High-pressure gas is admitted into the cylinder and allowed to expand freely forcing the piston back through the stroke. The pressure of the gas in the cylinder at the end of the stroke depends on the volume (mass) of the high-pressure gas introduced into the cylinder. This volume (mass) is determined by the percentage of the stroke through which the inlet valve remains open. An expansion engine is normally equipped with four cams on each of the two inlet valves corresponding to 20%, 25%, 30%, and 35% of the stroke. The power produced by the expansion is connected to either a mechanical brake or an electrical generation unit, and is usually dissipated as it generates only a small percentage of the total operating power. By changing the amount of braking, the speed of the machine can be changed; consequently, the flow of gas through the engine is adjusted accordingly. The brake normally used is an asynchronous (induction) motor, whose speed is adjusted by a rheostat. When the pressure of the feed air reaches 5 bar (in an air separation plant), the generator is started. Always make sure that the ammeter indicates that power is being generated and returned to the electrical circuit. During a plant start-up, and in order to produce the maximum refrigeration in the shortest possible time, the inlet valves are set on the smallest cam (20%); and the engine, if it has a mechanical brake, is operated at a fairly high speed (around 180–185 rpm), with the inlet gas pressure set as high as possible. As a rule, the machine is protected by an overspeed device. The latter item is especially important for the electric generator in case of a power failure. When the inlet and outlet temperatures have reached the desired values, the amount of refrigeration generated may be decreased by lowering the operating pressure of the process, or by decreasing the efficiency of the expander (operating on a larger cam, i.e., at a lower speed, say between 140 and 150 rpm). More recently manufactured reciprocating machines used Teflon rings that were impregnated with bronze filings for added resistance to wear. When used for fairly large liquid producing plants, in the range of 50–100 t=d, the reciprocating expansion engine has proven very efficient (above 70%). In these conditions, air feed is compressed to 175 bar, with about 65% of the process feed passing through the machine and expanding down to 6 or 7 bar and at a temperature of 107 K.

10.2.3 RADIAL EXPANSION MACHINES The application of an expansion turbine for the cryogenic separation of gases is almost synonymous with the name of Kapitza,2 a Russian scientist who conceived the idea in the 1930s. In fact, Kapitza selected a radial inflow machine because of its simplicity of operation and suitability of capacity control by inlet guide vanes.2 His first practical attempt included a small reaction rotor, 8 cm in diameter made out of Monel metal, rotating at 40,000 rpm and with a clearance of 0.3 mm. It used a water pump as a brake, and had an efficiency of 79%–83%. Needless to say, the industry was impressed. The first industrial application occurred in the late 1930s with the design of a very large air separation plant, equivalent to an oxygen plant with a capacity of 300 t=d. This plant was built by Linde AG for the sole purpose of recovering krypton from air and supplying it to the lamp bulb manufacturers of Europe.


Although the application of an expansion radial turbine in an air separation cycle was new, the concept of the machine was not, especially in the field of hydraulics. Moreover, the engineer should not confuse an ‘‘impact’’ wheel with either an ‘‘impulse’’ or a ‘‘reaction’’ wheel. The first implies a fast-moving water stream striking a slow-moving surface creating a shock, whereas in the other two, the velocities, angles, and speeds of the surfaces contacted are adjusted to produce a gradual retardation. To divert the water (fluid), a pressure is required from the vane on the water (fluid). The equal and opposite reaction pressure of the water (fluid) drives the rotor. In this sense, all modern expansion rotors including ‘‘impulse’’ rotors3 are reaction rotors. The original water radial turbines were designed in the early part of the nineteenth century and received water around the entire inner circumference from a whole circle of adjacent nozzles resulting in an outward flow, similar to a modern water sprinkler. In 1826, Fourneron, a French engineer, improved it with the addition of guide vanes inside the rotor. Howe patented an inward-flow turbine in 1838, which was further improved by Francis in 1849 that set the standard in its class for years to come. The Pelton wheel was developed and patented in 1889. In view of the configuration of its buckets, it too may be classified as a reaction wheel, or more correctly, an impulse–reaction wheel. Figure 10.2.1, Figure 10.2.2, and Figure 10.2.3 show vividly the configuration of each rotor. Figure 10.2.4 also indicates graphically, the relationship in efficiency between the various hydraulic turbines compared with that of the radial inflow turbine normally used in air separation plants. The development of the radial expander, especially after World War II, has progressed to the point that the unit has replaced the reciprocating expansion engine almost completely, at least for air separation, and even for liquid producing plants. Radial expanders are being coupled directly to radial compressors (boosters), which are being used for very high-pressure ratios, even up to 55 bar, and a speed of 45,000 rpm. In the latter case, these units are used primarily for liquefiers serving air separation plants having a liquefaction capacity of well over 200 t=d. The energy of a compressed gas can be converted to that of a gas flow in several ways: (a) it can be fully expanded in a stationary nozzle, and the high velocity of the stream can be directed to strike a wheel with impellers or blades as with an ‘‘impulse–reaction expansion turbine’’; (b) the feed gas stream may be completely expanded directly between the impellers without any preexpansion as with a ‘‘reaction turbine’’; it may be more advantageous, however, to combine

Inlet

Impulse turbine

Discharge

FIGURE 10.2.1 Impulse reaction wheel. (From Swearingen, J.S., Engineer’s Guide to Turbo Expanders, Hydrocarbon Processing, April, 1970. With permission.)


Inlet Reaction turbine

Rotating blades

Discharge

FIGURE 10.2.2 Axial reaction. (From Swearingen, J.S., Engineer’s Guide to Turbo Expanders, Hydrocarbon Processing. With permission.)

options (a) and (b) by pre-expanding the gas stream at the nozzle to an intermediate pressure, then expanding the stream to its final pressure through the impellers of the turbine. This latter option is more common in practice.

10.2.4 PROCESS APPLICATIONS Over the years, process engineers have been designing air separation plants more and more efficiently, with the result that the quantity of gas to be expanded for make up of refrigeration lost by either heat gain or liquid removal or both has been getting smaller and smaller.

Inlet

Discharge

FIGURE 10.2.3 Radial inflow turbine. (From Swearingen, J.S., Engineer’s Guide to Turbo Expanders, Hydrocarbon Processing, April, 1970. With permission.)


Range of cryogenic expanders 1

Francis

Pelton

Kaplan

Efficiency

0.8

Radial turbines

0.6

(908)

Axial turbines 0.4 o ns = n y Δ His3/4 0.2

0

1

100

10

1000

Specific speed, Ns

FIGURE 10.2.4 Comparison of efficiencies of a variety of cryogenic expansion machines. (From Swearingen, J.S., Engineer’s Guide to Turbo Expanders, Hydrocarbon Processing, April, 1970. With permission.)

Presently, the amount of gas to be expanded for a standard oxygen plant with little or no liquid production may be slightly less than 10% of the total process air feed. In fact, the majority of plant designers are moving away from the use of motor–generator sets for expander loading because of their higher maintenance, and of possible local electrical problems, especially those of power factor. Preference is given to the use of a radial booster compressor–blower coupled directly to the expansion turbine. The latter system is less expensive, avoids maintenance problems, and gives the designer a variety of options in the use of this extra compression to improve the efficiency of the process cycle. The use of motor– generators cannot be completely ignored, however. In situations where the cost of power is near or above $40=MkWh, the use of a motor–generator may very well prove economically viable, depending on the production capacity of the plant.

10.2.5 OPERATIONAL FACTOR (AIR SEPARATION PLANTS) Normally, about 10% –20% of the incoming process air traversing the primary heat exchangers is withdrawn around the midpoint of the exchangers at a predetermined temperature level, but at the same pressure as the process air, and is then sent to the machine for expansion. This expanded air, at almost atmospheric, is fed into the upper (low pressure) column for added recovery of product oxygen. This side air stream reserved for the expander is generally known as the Lachman air stream named after the inventor. Refer to Figure 3.18 as it diagrammatically shows the use of the Lachman stream. Another option, used more and more frequently, is to employ a radial compressor–blower directly mounted on the same shaft as the expander. The Lachman air stream will then be diverted from the main air-feed upstream of the primary heat exchangers, and sent to the radial compressor–blower where it will be boosted to a higher pressure. This higher pressure stream, after precooling, will then be sent to the inlet of the primary heat exchangers. After being cooled to the right temperature, the side stream is withdrawn and sent to the expansion


side of the turboexpander. After expansion, the stream is sent to the upper column as previously mentioned. The extra compression ratio of the stream to be expanded adds to the efficiency of the process cycle. In the thermodynamic design of a turboexpander, the important parameters are the isentropic enthalpy drop across the expander, and the volume flow-rate at the expander discharge. The first determines the rotor speed, and the latter determines the expander wheel flow area. These two, within stress and mechanical limits, determine the configuration of the hydraulic channel, and therefore, the expander efficiency. If high-purity nitrogen is not a major priority, the designer has the option of expanding nitrogen or a predetermined mixture of oxygen and nitrogen from the lower (high pressure) column and not simply air, because the former products can be used more efficiently in terms of a lower entropy increase. In the first option, a nitrogen stream can be withdrawn from the lower (high-pressure) column and at the same pressure as the air feed, warmed to a predetermined higher temperature by sending it countercurrent to the incoming air stream in intermediate heat exchangers, then sent to the expansion machine for final expansion. After expansion, this cold stream of nitrogen will be totally removed from the cryogenic process after it first passes through the primary exchangers where its high refrigeration capacity is used to cool down the incoming process air. With this option, by warming the nitrogen to a higher temperature, it is possible to obtain a higher temperature drop across the expander, which means a higher enthalpy drop, and consequently a higher process efficiency. This procedure, however, is limited by the quantity of high-purity nitrogen desired as a final product. As the requirements for pure nitrogen product increase, the availability of nitrogen for expansion decreases.

10.2.6 REFRIGERATION AVAILABILITY There seems to be a general misconception about the use of an expansion machine. Many potential operators of air separation plants, who have little or no experience with the unit’s application, often believe that any quantity of liquid products can be produced from the plant by simply increasing the capacity of the expansion machine. In a standard air separation plant, with no additional external compression cycles, about 10% of the total products can be extracted in liquid form, but the latter quantity will include any combination of oxygen, nitrogen, and argon. In this situation there will be a simultaneous loss of gaseous oxygen product in the approximate ratio of 3:1, or even lower. In other words, for every tonne of liquid products produced and removed, in any combination, there will be an approximate loss of 3 t of gaseous oxygen if the latter is the main product. Furthermore, the plant will have to be supplied with two 100% capacity turboexpanders operating simultaneously. For the production and removal of larger quantities of liquid products, a special liquefier will be required, added to or integrated with the air separation plant (see Chapter 6). It is also prudent to remember that the expansion machine has a very important function, namely, to fine tune the refrigeration balance of the cryogenic process cycle; in other words supplying just the right refrigeration to make up for the losses due to heat gain, a warm end temperature differential, and a liquid product removal. If the refrigeration generated by the expander is below the calculated quantity and quality (in terms of temperature), the main condenser will warm up, the oxygen in the main condenser will begin to boil off beyond the prescribed quantity, and the liquid level will begin to get lower. This condition means that an excess of the gaseous oxygen product will boil off, be entrained by the nitrogen, and carried off as waste, unless the withdrawal of gaseous oxygen is increased, thus decreasing the efficiency of the recovery required. It also means that the hydrocarbon content, if any exists, will remain in the main condenser, and will increase in terms of percentage, which is a very dangerous


Waste

GOX product

Insufficient refrigeration High reboiling Low LOX level High GOX content in waste Very high GOX product purity High CxHy contaminants in main condenser Corrective action: Initially increase GO2 output Reduce LOX withdrawl Later increase slowly the loading of the turbines

FIGURE 10.2.5 Effects of insufficient refrigeration. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

condition. This situation should be monitored continuously by discerning any increase in the quantity of oxygen in the waste nitrogen or any increase in the purity of the oxygen product or both beyond the designed amount, and observing the liquid level in the column sump (see Figure 10.2.5). If, on the other hand, the expansion machine is overloaded, especially at emergency startups, the extra cold will prevent the nitrogen, and especially the argon, from boiling off. This condition will result in a drop in purity of the oxygen product as well as a loss of valuable argon, which is undesirable. Unfortunately, this has happened more than a few times. If this condition arises, the operator or supervisor should not panic. All one has to do is unload the expansion turbine at predetermined increments, allowing at least 2 h intervals to permit the process cycle to reach thermodynamic equilibrium at each unloading. During this procedure (which requires patience), the oxygen purity and the argon recovery will begin to increase (Figure 10.2.6). An important principle for the control of the aforementioned conditions (excess, or shortage of refrigeration) is the control of gaseous oxygen flow and withdrawal of liquid from the column sump. That is, one must keep a steady liquid level in the main condenser.

10.2.7 PROCESS TECHNOLOGY4,5 Although the radial expansion turbine is simplicity itself in operation and control, its design is neither simple nor easy. It demands a fundamental understanding of the Euler equation in terms of change of momentum, and the first two laws of thermodynamics, as well as their application in single- and multicomponent fluid processes. The principal parameters that govern the design in terms of geometry and efficiency are given in Table 10.2.1:


Waste

Excess refrigeration Placid reboiling High LOX level Low GOX product purity Low CxHy contaminants in main condenser

GOX product

Corrective action: Initially reduce GO2 output Increase LOX withdrawl Later increase slowly the loading of the turbines

FIGURE 10.2.6 Effects of excess refrigeration. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

From a review of these equations, it is obvious that both the process designer of the purchaser and the turbine specialist of the supplier must work in close concert to arrive at an optimum design in terms of efficiency and reliability. The process designer must specify the factors of Dh, Q3, Pr, and n2 at the very least, and the turbine specialist will select the appropriate Ns, D2, and w3 as well as the ancillary equipment. Incidentally, in the design of this type of turbine, because of the pressures and the high tip speeds involved, the Reynolds number NRe may be considered almost negligible. Normal tip speeds employed in the design of the impellers in radial expansion machines used in air separation plants are in the range of 275–335 m=s. The maximum design of tip speeds are given in Table 10.2.2. On air compressors with stainless steel wheels, the tip speeds may be in the range of 430–488 m=s. TABLE 10.2.1 Principal Parameters Used in Design of Radial Expansion Turbines Specific speed, Ns

Ns ¼ N(Q3 )1=2 =(Dho )3=4 (10:2:1)

Specific diameter, Ds

Ds ¼ D2 (Dh)1=4 =(Q3 )1=2 (10:2:2)

Turbine exhaust or relative mach number, Mw3

Mw3 ¼ w3 =a3

(10:2:3)

Total pressure ratio, Pr

Pr ¼ P01 =P03

(10:2:4)

NRe ¼ (U2 D2 )=v2

(10:2:5)

Turbine Reynolds number, NRe 3

Ns ¼ shaft speed (rpm), Q3 ¼ turbine exhaust volume flow (m =s), Dh ¼ isentropic process enthalpy differential (kJ=kg), D2 ¼ turbine tip diameter (mm), P01 ¼ total inlet pressure (barA), P03 ¼ total exit pressure (barA), w3 ¼ relative velocity at turbine discharge (m=s), a3 ¼ speed of sound at turbine discharge (m=s), U2 ¼ tip speed (m=s), v2 ¼ kinematic viscosity of process gas at turbine wheel inlet (m2=s). Source: From Swearingen, J.S., Engineer’s Guide to Turbo Expanders, Hydrocarbon Processing, April, 1970. With permission.


TABLE 10.2.2 Maximum Design of Tip Speeds for Turbine Wheels Type of Wheel Open wheel Closed wheel

Material Aluminum 6061 (m/s)

Material Aluminum 7075 (m/s)

365 350

430 395

Source: From Swearingen, J.S., Engineer’s Guide to Turbo Expanders, Hydrocarbon Processing, April, 1970. With permission.

10.2.8 EXPANSION TURBINE EFFICIENCY Including the behavior of the mechanical components of the machine, the efficiency of an expansion turbine can be measured by the actual enthalpy drop (kJ=kg, or kcal=m3, at the specified conditions) realized by the expanded gas, divided by the isentropic (adiabatic) enthalpy drop that is theoretically possible between the intake pressure and temperature points, and the final discharge pressure point preset by the process design: Efficiency h ¼ (h1 ha ) actual=(h1 h2 ) isentropic

(10:2:6)

As it is impossible to achieve a 100% isentropic enthalpy drop, the process designer normally sets the intake conditions for flow, pressure and temperature, and the final discharge conditions for temperature. He also assumes a machine efficiency for the required enthalpy drop based on past experience, say 80%. The supplier of the machine, on the other hand, will run these figures through his computer and offer an expansion turbine that is close or superior to the designer’s estimated efficiency for enthalpy drop. As a matter of principle, however, the process designer should use an efficiency slightly lower than that offered by the supplier in order to avoid any unpleasant surprises during operation in the field. It should be remembered, however, that the expander supplier will present the efficiency of the rotor of the machine, not necessarily the finally installed unit, complete with all its ancillary piping and valves. Isentropic efficiency also depends on the process cycle in question, and the gas to be treated. Some process designers prefer to use an enthalpy point just above the dew point, and others prefer to dip a little below the dew point, regardless of the occurrence of some liquefaction. In fact, according to tests made by Swearingen4, operating the turbine at a few degrees below the dew point results in a slightly higher efficiency than when operating at a few degrees above the dew point. Normally, radial expanders have had problems with condensed liquid because the liquid tends to be centrifuged back toward the rotor, resulting in cavitation and possibly erosion. Swearingen has stated that the problem can be solved by shaping the rotor blades in the expander rotor so that their walls are at every point parallel to the vector resultant of the forces acting upon suspended fog droplets. In this design, the suspended fog particles do not deflect toward the rotor walls to collect and interfere with performance, nor do they create any erosion or vibrations of the blades. This design, however, does result in a slightly lower efficiency. Example (a) for the calculation of turboexpander efficiency using air as the expansion medium. Inlet pressure of 494.37 kPa, inlet temperature of 125.35 K, h1 of 70.7 kcal= std m3, outlet pressure of 134.45 kPa, outgoing temperature of 94.75 K, ha of 62.6 kcal=std m3, adiabatic enthalpy at 134.45 kPa is 60.0 kcal. Efficiency h ¼ [(70:7 62:6)=(70:7 60:0)] 100 ¼ 75:7% from Equation 10:2:6 ß 2006 by Taylor & Francis Group, LLC.

Example (b) for the calculation of efficiency using an oxygen (33%) and nitrogen (67%) mixture as the expansion medium. Inlet pressure 619.17 kPa, inlet temperature 133.65 K, h1o (oxygen) ¼ 80.7 0.33 ¼ 26.63 kcal=m3, h1n (nitrogen) ¼ 71.6 0.67 ¼ 47.97, total h1oþ1n ¼ 74.6, outlet pressure 138.2 kPa, outlet temperature 99.25 K, h2o (oxygen) ¼ 71.7 0.33 ¼ 23.66 kcal=m3, h2n (nitrogen) ¼ 62.7 0.67 ¼ 42.01, total h2oþ2n ¼ 65.67, Adiabatic enthalpy at 138.2 kPa, h3o (oxygen) ¼ 66.6 0.33 ¼ 21.98, h3n (nitrogen) ¼ 58.7 0.67 ¼ 39.33, and total h3oþ3n ¼ 61.31: Efficiency h ¼ [(74:60 65:67)=(74:60 61:31)] 100 ¼ 67:2% from Equation 10:2:6 It is obvious from these examples that the optimum design of a turboexpander demands close study to combine the correct inlet and outlet pressures, and the right temperatures, as well as to select the right expanding medium.

10.2.9 EXPANSION TURBINE LOSSES One has to bear in mind that all irreversible losses during the expansion process will result in an increase in entropy, and will result in lower overall efficiency. In discussing efficiencies with the manufacturer of turbomachinery, it is necessary to bring up and review the following possible losses: 1. Inlet guide vanes creating friction and turbulence. 2. The design of the inlet portal of the impellers. 3. The design of the flow pattern of the impellers to avoid friction losses. There is a difference of about 3% in efficiency between a closed wheel and an open wheel. 4. Wheel disc friction. 5. Discharge losses due to the formation of a vortex at the outlet portal of the machine. The outlet velocity of the gas stream should be very low. 6. Volumetric losses due to labyrinth leakage. Always check the seal clearances. 7. Misalignment in the expander–blower assembly. A few millimeters off may reduce efficiency seriously. 8. Ancillary piping and valves interconnecting the machine to the distillation equipment may introduce additional pressure losses. The piping configuration may also cause a strain on the rotating shaft and impellers.

10.2.10 MEASURING EFFICIENCY (FIGURE 10.2.7) The efficiency of a turboexpander depends on the isentropic enthalpy drop in kJ=kg or in kcal=m3 (at 158C) across the expander impeller, and the volumetric flow rate in actual m3=s at the discharge. The former determines the shaft speed in rpm, and the latter determines the flow area of the impeller. These two factors, within the constraints of the tip speed, determine the final overall efficiency of the machine. To measure the efficiency of a turboexpander in the field is not an easy matter. Recording the readings of the inlet temperature and inlet pressure, as well as of the discharge temperature and pressure of the gas being expanded is easy enough, but to interpret the readings in terms of enthalpy is another matter. One must first find an enlarged version of a Mollier (enthalpy–entropy) diagram for air or nitrogen. The diagram, so chosen, should be as accurate as possible, especially around the area between the


Inlet pressure, kPa Inlet temperature, K

Actual process

Enthalpy H

Δ H1s

Δ Ha

Discharge pressure, kPa Discharge temperature, K

Ideal isentropic process

Entropy increase Note:

Efficiency

H1s = Isentropic enthalpy drop H a = Actual enthalpy drop = Δ Ha Δ H1s ⫻ 100

FIGURE 10.2.7 Measure of efficiency. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

saturated vapor and the dual phase zone below the dome. One is not sure, moreover, of the exact phase of the expanded gas at the discharge of the expander. Most operators prefer to expand the gas stream from the upper pressure to the discharge of the expansion machine while still in its gaseous phase. This procedure avoids condensation within the impellers, which may give rise to cavitation. Two different operators may arrive at two different efficiencies. A small difference in the reading of enthalpy may alter the efficiency considerably. Admittedly, the basic work required to calculate the enthalpy at various pressures and temperatures, and to enter them in the computer entails a great deal of labor, but once the work is completed and placed in the computer, the efficiency of any turboexpander can be calculated very quickly and accurately. At the same time, the choice of a correct equation of state is most important. There are several equations available, also some accurate gas tables on the thermodynamic properties of air, such as those developed by Keenan, Chao, and Kaye, published by John Wiley and Sons (1940 or latest revision). Using similar tables, computer models can be prepared in advance.

10.2.11 VARIOUS EXPANSION TURBINE SYSTEMS From the first attempt to use an expansion turbine, the cryogenic engineer has progressed from the use of the basic impulse type to the axial radial unit, and finally to the radial inflow design that has become the general standard for the industry. It was quickly realized that a high-speed turbine can offer a high mechanical and thermodynamic efficiency with a low


input-pressure (5–7 bar). Basically, there are five configurations that accommodate specific applications: 1. For small expansion units that may generate less than 75 kW the loading device generally used is an oil brake or its equivalent. The system dissipates expandershaft energy extracted from the process in a controlled manner for operating stability. It is used where energy recovery is neither feasible nor economically viable (Figure 10.2.8a). 2. For larger values and wherever energy is at a premium, an electric generator coupled to the turbine shaft by means of a speed reducing gear is a cost-effective alternative. In this case, however, the project engineer should first investigate the electrical system of the plant in order to make sure that the power generated by the expansion turbine can be supplied to the main plant without any disruption of the power factor (Figure 10.2.8b). 3. In the past few years, popular application has been the combination of a compressor and an expansion turbine on a common shaft, especially in applications of air separation units. This not only simplifies the mechanical configuration of the machine, but also increases the enthalpy drop with little increase in entropy (Figure 10.2.8c). 4. Finally, the integration of an expansion turbine as an extra stage coupled with an integral gear compressor is a preferable combination. Although this choice is highly desirable, it also requires an optimum matching of the various gears in order to arrive at a high efficiency (Figure 10.2.8d). 5. Two expansion turbines are used in series for expanding high-pressure gaseous feed stocks, such as natural gas. The enthalpy drop of a turbine is higher when the machine is operated at higher temperatures, thus producing more refrigeration. On the other hand it is required to produce a certain amount of outlet gas at as low a temperature as practicable. There is a contradiction between the two requirements. This is why using two expansion turbines at two different temperature levels is such a great improvement in the overall efficiency of the process (Figure 10.2.8e, also refer to Figure 6.5). Basically, integral gearing also permits efficient multistaging with up to four stages (two stages per pinion) on one gear box. In this type of application, however, the project engineer should seriously review the compressor manufacturer’s proven experience, as well as his shop facilities. The project engineer should be wary of the usual paper guarantee issued by the manufacturer that may prove inadequate to compensate any financial disaster, resulting from a failed guarantee in the field.

10.2.12 MECHANICAL DESIGN PARAMETERS 10.2.12.1 GENERAL As noted in the case for liquefiers, in Section 6.2.4.3, two turboexpanders directly coupled to compressor–boosters are operated in series. One is placed in the warm end of compression and the second unit follows in the second stage boosting the refrigerant medium, generally nitrogen, to a higher pressure. It should be kept in mind that there are two ways of increasing the liquid production capacity of a cryogenic plant: either by increasing the flow rate and keeping the pressure constant, or by increasing the pressure and maintaining a constant flow. In the former case, the energy required increases arithmetically with the increase in flow. In the latter case, the energy required increases with the logarithm of the compression ratio, which is less expensive in terms of energy consumption.


(a)

(b)

FIGURE 10.2.8 (a) Expander with fluid brake. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.) (b) Direct-drive generator-loaded expander. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.) continued

Process designers should keep in mind, if they wish to upgrade the recycle compressor– booster pressure of say 60 bar and higher, that there is a limitation in terms of the discharge capacity of any radial compressor. The minimum discharge flow for any radial compressor is around 5.36 actual m3 (200 acfm), and even at this flow-rate compressor efficiency falls very


(c)

(d)

FIGURE 10.2.8 (continued) (c) Compressor-loaded expander for air separation applications. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.) (d) The expander on the upper-left driving the third stage booster on the upper-right. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.) continued


(e)

FIGURE 10.2.8 (continued) (e) Integrally geared expander. Two expansion turbines geared in series. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.)

low due to high seal losses. Process and project engineers must accept a compromise between what they wish to design, and what modern technology can deliver with an iron clad warranty. If a liquefier is designed with two turboexpanders with boosters in series, the efficiencies as listed in Table 10.2.3 should be expected. In general, turboexpanders should be specified mechanically within the following parameters: . . .

To operate with a variable flow while maintaining a high efficiency To tolerate a minimum of dust and some condensation High bearing strength to avoid any damage from rotor unbalance


TABLE 10.2.3 Typical Efficiencies Expected for Expanders Coupled with Boosters Warm End

Expected efficiency at operating rpm (%)

Cold End

Expander

Booster

Expander

Booster

85

76

82

75

Note: For overall process efficiency the booster efficiency is much less significant than that of the turbine. Source: Courtesy of F.G. Kerry, Inc., Report, 1980. With permission.

. . .

High efficiency with an acceptable high speed (rpm) Very high mechanical reliability Positive shaft seals

10.2.12.2 OPERATIONAL CONTROL During the operation of an air separation plant, it may be necessary to adjust the load on the expansion turbine to vary the enthalpy drop that in turn regulates the quantity of refrigeration available to the process. This is accomplished by the use of a system of variable inlet area nozzles. Manufacturers of expansion machines generally offer adjustable area nozzles to provide continuous control between 20% and 150% flow to achieve optimum flow patterns and entry angles. This system of control can be computerized locally or linked to an existing computer network (Figure 10.2.9).

FIGURE 10.2.9 Variable nozzles to control flow. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.)


(b)

(a)

FIGURE 10.2.10 (a) Typical open-wheel design. (b) Typical closed-wheel design. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.)

10.2.12.3 SHAFT SPEED (RPM) Depending on the nature of the gas to be expanded, the shaft speed may vary for the same enthalpy drop. For nitrogen, an enthalpy drop of 100 kJ=kg is reached at an expansion ratio of 8–10 to 1. With natural gas, the same enthalpy drop can be reached with an expansion ratio of 3–4 to 1. The higher the speed the better the chance one has to achieve an ideal isentropic expansion5.

10.2.12.4 IMPELLER DESIGN (FIGURE 10.2.10A AND FIGURE 10.2.10B)6,7 The design and the fabrication of the turboexpander wheel are critical to a successful operation. In an expansion machine, expansion of the gas occurs partly in the inlet nozzle and partly in the impeller. In the impeller, the gas will have its direction changed while expanding and with this change of direction its thrust will also be changed. If the impeller is properly designed, the resultant vector of velocity will be oriented at right angles to the rotor blades, which is an ideal situation. One has to select a suitable blade design that will minimize both static as well as dynamic stresses, permitting a high-energy extraction in the expander. A higher shaft speed means a higher enthalpy drop, but the latter is limited by the tip speed of the impeller blades. Too high a speed may result in turbulence, with a resulting resonance at the impeller tip with disastrous results in terms of inefficiency. It will also create fissures at the edge of the blades. Tip speeds in excess of 100 m=s may result in resonance frequencies that may prove harmful to the impeller. The radial inward flow design, however, has indicated lower stresses at a given tip speed, resulting in a higher efficiency than other designs. Although radial inward flow expanders are preferable for most process cycles, the axial reaction turbine may be found more desirable for a multistage expansion because it allows for a much easier flow path from one stage to the next.

10.2.12.5 MATERIALS

OF

CONSTRUCTION

Machining wheels from a solid forging will provide the maximum mechanical integrity by the elimination of any welded joints, as well as provide the uniform high strength properties of a forging.5


10.2.12.6 BEARINGS Until recently, all turboexpanders have been equipped with forced feed oil lubricated bearings similar to those used for turbocompressors. The oil flows through a labyrinth seal, but is impeded from entering the cryogenic section by an air–nitrogen buffer stream flowing in the opposite direction (Figure 10.2.11a and Figure 10.2.11b). During the past five years or so, the

Buffer gas injection

Vent, inlet, etc.

Vent to atmosphere Bearing

Floating oil seal Oil drain (a)

Vent or buffered port Vent, flare, etc.,

Carbon ring seal cartidge

Floating oil seal Beal

(b)

FIGURE 10.2.11 (a) Buffered labyrinth seal. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.) (b) Buffered carbon ring seal.


use of magnetic bearings has been gaining ground and is steadily replacing oil-lubricated bearings for air separation plants. Field reports for this design have been very encouraging. The use of air bearings has also come to the fore, especially for machines of a relatively small size, but their use for air separation plants of a serious size is relatively new. In this design, the bearing is formed by a thin film of air between the moving parts. Shaft speed of an air bearing expander is quite high. In fact, a shaft speed of 100,000 rpm has already been tested. The one serious disadvantage to the air or gas bearing design is the possibility of an inadvertent intimate contact between the rotor and the other surfaces during an interrupted gas flow even for a small fraction of a second. This situation can cause serious damage to the entire turboexpander system. Yet, in one application involving a 300 t=d oxygen plant, a helium gas bearing has been in use for the past 10 years. According to Swearingen,8 where high-speed rotating machinery is involved, the thrust bearing is very often the point at which trouble begins. The thrust bearing must carry the net thrust load, but unfortunately its calculation is not carried out with any high degree of accuracy. Its balance is subject to change due to seal wear or pressure variation. 10.2.12.6.1

Lubrication System

In the case where an oil lubrication system has been chosen, the design should include an efficient seal gas system to keep the lubricating oil or fluid from entering the section where the gas is expanded or compressed by the booster–blower if the latter is mounted on the same shaft. In the latter case, the lubrication unit must be monitored on the following: . . . .

Seal gas flow Seal gas supply pressure Seal gas manifold pressure Expander impeller back pressure

The lube oil system should be monitored for the following: . . . . . .

Oil level Temperature upstream of the oil cooler Temperature downstream of the oil cooler Bearing temperature Differential pressure across the oil filter Bearing oil supply pressure

10.2.13 INSTRUMENTATION AND CONTROL 10.2.13.1 PROCESS CONTROL Because of the high speed of rotation encountered with turboexpanders, it is mandatory to monitor the critical points. The machine should be equipped with the following items: . . . . . .

Overspeed protection High-speed emergency shut-off valves (at high speeds) Pressure control valve Temperature control valve Temperature indicator on bearings and alarms Continuous flow switch


10.2.13.2 INSTRUMENTS REQUIRED In regard to process control, the following items should be supplied: . . . . . . . . . . .

Suction filter with differential pressure indicator Inlet and outlet pressure indicator Outlet and inlet temperature indicator Guide vane actuator Rotational speed indicator Overspeed protection Vibration probes and monitors Emergency shut off valves Pressure control valve Temperature control valve Temperature indicator on bearings

10.2.13.3 COMPUTER CONTROL (DCS) Presently it has been possible to link the turboexpander controls directly to the distributed control system (DCS) so that all parameters including the inlet guide vanes can be monitored and controlled from the central operating desk.

10.2.14 SPARES In the past, the standard rule was to supply two turboexpanders, one in operation and the other as a complete standby spare. The reason generally given was that the use of two machines operating simultaneously could accelerate the start-up of a very large plant that usually has an enormous heat gain, or of a plant that is located in a region far from any viable maintenance services. This reasoning may not be valid anymore because turboexpanders have been proven very reliable, especially with the introduction of spare plug-in rotor assemblies, which are now readily available. A damaged rotor can be replaced very quickly with a new rotor assembly, if readily available from the spare-parts department of the plant. The changeover can take place during a shift, or say, in less than 8 h. A plug-in rotor assembly is much less expensive than a completely installed spare turboexpander. In fact, for liquefiers, especially of the integrated type, it is not uncommon to use a common system of turboexpanders to supply the total refrigeration for both the liquefier and the air separation plant with no spares. In this situation, the common expander system liquefies products for storage purposes, and also supplies liquid nitrogen reflux for the upper column of the air separation plant. A straight gaseous product air separation plant supplying product gas to a customer via a pipe line on a guaranteed continuous basis must, of course, have a spare installed expansion turbine already piped and hot-wired, unless there is a liquid storage tank backup.

10.2.15 GENERAL APPLICATIONS FOR EXPANSION MACHINES9,10 Aside from their major use in cryogenic gas separation of gases, turboexpanders are also employed in the following industries: .

Gas processing for the recovery of natural gas (LNG), as well as for the recovery of ethane from natural gas


. . .

Cogeneration for energy conservation Ethylene recovery Gas pressure let down

REFERENCES 1. Gomonet, E. Les tres basses temperatures production et emplois. Librairie J.B, Bailliere et Fils, 19 Rue Hautefeuille, Paris. p. 89. 2. Gomonet, E. Les tres basses temperatures production et emplois. Librairie J.B, Bailliere et Fils, 19 Rue Hautefeuille, Paris. p. 123. 3. Schoder, E.W., and F.M. Dawson. 1927. Hydraulics, 303–307. New York: McGraw Hill. 4. Swearingen, J.S. 1970. Engineer’s guide to turbo expanders. Hydrocarbon Processing, April. 5. Lindhart, H.D. 1971. New developments in hot gas expander compression systems, Turbo machinery handbook. Hydrocarbon Processing, pp. 62–65. 6. Bergmann, D. and S. Mafi. 1979. Selection guide for expansion turbines, 83–86. Hydrocarbon Processing, August. 7. Holm, J., and J.S. Swearingen. 1977. The application of turboexpanders for energy conservation, 1–8. ASME publication, June 8. 8. Swearingen, J.S., and S. Mafi. 1969. Experimental investigation of vibrations in high-speed rotating machinery, ASME publication, presented at Vibrations Conference, Philadelphia PA, March 30–April 2. 9. Ershaghi, B. Power recovery in process plants using radial inflow expansion turbines. Adapted from a presentation at 1986 Annual Meeting Gas Processors Ass., San Antonio Texas, Atlas Copco Co. 10. Swearingen, J.S. 1969. Design consideration in LNG expansion liquefaction cycles. Oil & Gas Journal, October 6.

SUPPLEMENTARY READING Paturiaux, J.C. Turboexpanders in air separation. Atlas Copco Energas, Gmbh Timmerhaus, K.D., and T.M. Flynn. 1989. Cryogenic process engineering, 261–263. Science Business and Media, Dordrecht, The Netherlands.



10.3

Compressors

10.3.1 COMPRESSORS 10.3.1.1 GENERAL When the industrial application of gas separation began, especially with air separation by Linde, it initiated an interest in high-pressure reciprocating air compressors with a target of 200 bar. This target was lowered to 40 bar, and even to 25 bar when the expander was introduced by Claude in 1902. In the 1930s, when air separation plants began to get larger in capacity, centrifugal machinery began to be used, which did away with the main problem of oil entrainment by reciprocating units used previously. Since the 1950s, even for small plants, centrifugal compressors have replaced reciprocating machines almost completely for basic air separation. Reciprocating machines of the labyrinth, oil-free type, are still being used, but only for small, high-pressure systems for oxygen compression, or for gases that cannot tolerate contamination. Subsequently, there has been a perennial tug-of-war between compressor manufacturers and gas separation designers in trying to fit existing compressors to preset designs and the demands of new process cycles that change ever so quickly.

10.3.2 DEFINITIONS*1 In order to set an acceptable standard for terms normally used by the compressor industry, the following terms are defined: Mass is slightly different from weight. Mass is a quantity of matter. Weight, the force exerted on a given mass by the attraction of gravity, will vary with its distance from the center of the Earth. The difference may be small, but may be important in some cases. Absolute temperature is related to Celsius by the equation K ¼ 8C þ 273.15 K (Kelvin). Barometric pressure is the absolute atmospheric pressure existing at the surface of the Earth. It varies with altitude and weather (see Appendix). Capacity is the quantity of gas actually delivered when operating between specified inlet and discharge pressures. For all compressors (excluding ejectors) it is the volume measured at the conditions of pressure, temperature, gas composition, and moisture content existing at the compressor inlet flange. Compressibility factor (Z) is defined as the ratio of the actual volume of the gas to the volume determined according to the ideal gas law. Because many gases do not obey this law precisely, a compressibility factor Z is applied as a multiplying factor: v ¼ ZRT =p *With courtesy of Ingersoll-Rand Company Limited.


(10:3:1)

Compression efficiency is the ratio of the theoretical work requirement for a stated process to the actual work required to be done on the gas for compression and delivery. It includes leakage and fluid friction losses, and thermodynamic variations from the theoretical process. Compression ratio is the ratio of the absolute discharge pressure to the absolute intake pressure. It may apply to either a single stage compressor or to a complete multistage unit. Discharge pressure is the total gas pressure (static plus velocity) at the discharge flange of the compressor. This only applies to dynamic compressors. The discharge pressure should be expressed only in absolute terms unless the local barometric pressure is also stated. Dry bulb temperature is the local ambient gas temperature. Fixed compression ratio is the design (built-in) compression ratio for a rotary unit. Theoretical horsepower is the work theoretically required to compress and deliver a given gas quantity in accordance with a specified process. Indicated horsepower is that obtained by indicator card analysis of compression or expansion in a cylinder of a reciprocating unit. Brake horsepower is the total power input required including all friction losses. Relative humidity is the ratio of the actual partial vapor pressure in an air–vapor mixture to the saturated vapor pressure at the existing dry-bulb mixture temperature, expressed in percent. Specific humidity is the ratio of the weight of water vapor in an air–vapor mixture to the weight of dry air, expressed in kilograms per kilogram of dry air. Mach number is the ratio of the actual gas velocity at a given point to the velocity of sound in the same gas at the conditions existing at this point. These are also known as local conditions. Mechanical efficiency is the ratio, in percent, of the indicated horsepower to the actual shaft horsepower. Kilomole is the weight of gas in kilograms numerically equal to the molecular weight of the gas, or to the pseudo–molecular weight of a gas mixture. Molar heat capacity, or molar specific heat, is the heat in calories required to raise the temperature of 1 mole of gas by 18C or 1 K. Isothermal process is one where heat is constantly removed to minimize the temperature rise and maintain it at the initial level p1 V1 ¼ p2 V2 ¼ constant

(10:3:2)

Isentropic process is one wherein no heat is removed. It is a reversible adiabatic process. Adiabatic process is one in which there is no change in temperature. No heat is transferred to or from the working substance: p1 V1k ¼ p2 V2k Polytropic process is one in which changes in gas characteristics during compression are considered (Figure 10.3.1) p1 V1n ¼ p2 V2n

(10:3:3)

Since perfectly isothermal or adiabatic processes are unrealistic, dynamic units generally use the polytropic process. The exponent n is experimentally determined for a given type of machine, and may be below or higher than the adiabatic exponent k. In positive displacement and internally cooled dynamic units, n is generally less than k. In uncooled dynamic units it is


E

D

C B

C⬘ AC⬘—polytropic uncooled dynamic AB—adiabatic AC—polytropic cooled reciprocating

Pressure

AD—isothermal

Theoretical no clearance

F

A

3-A G

H

Volume

I

FIGURE 10.3.1 Theoretical indicator card showing various gas compression processes. (Courtesy of Ingersol Rand Company Limited, 2006. With permission.)

usually higher than k due to internal gas friction. The value of n changes during compression, but the effective (average) is calculated from experimental information. The exponent n is seldom required but the quantity (n1)=n is frequently needed. This can be obtained from the following equation, although it is necessary that the polytropic efficiency hp be known or approximated from a prior test. Assuming that the k value of any gas or gas mixture is known, then (n 1)=n ¼ [(k 1)=k] 1=hp

(10:3:4)

See Figure 10.3.2. Saturation occurs when water vapor is at the dew point, or the saturation temperature corresponding to its partial pressure. The degree of saturation is the ratio of weight of vapor existing in a given space to the weight that would be present if the space were saturated at the space temperature. Saturated air–vapor mixture is one in which the space occupied by the mixture is saturated with water vapor at the temperature of the mixture. In stating the pressure of a gas, because the measurement is generally taken by a manometer or gauge, it is necessary to add to it the local barometric pressure in order to arrive at the true pressure of the gas, namely in absolute pressure. The universal unit is kPa A, or kiloPascals Absolute. At sea level, 1 atm is equivalent to 101.325 kPa, where 100 kPa ¼ 1 bar, which is generally used by most engineers if the pressure is over 1 atm. Saturated vapor pressure is the pressure existing at a given temperature in a closed vessel containing a liquid and the vapor from that liquid, after equilibrium conditions have been reached. It is dependent only on temperature and must be determined experimentally. Saturation pressure is another term for saturated water-vapor pressure. Saturation temperature is the temperature corresponding to a given saturated vapor pressure for a given vapor. Slip is internal leakage within a rotary compressor.


0.8

t

en

rc

hp

0.7

50

55 60

0.6

65 70

0.5

75 80 85

0.4

pi ntro Ise 0 10 hp =

n−1 n Value of

=

Pe

c

0.3

0.2

0.1

0 1.0

1.1

1.2

1.3 Value of k

1.4

1.5

1.6

3-B 1.7

Cp Cv

FIGURE 10.3.2 Ratio (n 1)=n versus adiabatic exponent k. (Courtesy of Ingersol Rand Company Limited, 2006. With permission.)

Superheated air–vapor mixture is the one in which the space occupied by the mixture is at a temperature above the saturation temperature. Volumetric efficiency is the ratio in percent of the actual delivered capacity (measured at inlet temperature, pressure, and gas composition) to the piston displacement. Wet bulb temperature is the temperature recorded by a thermometer whose bulb has been covered with a wetted wick, and whirled on a sling psychrometer. Taken with the dry bulb temperature, it permits determination of the relative humidity of the atmosphere. Wet gas is any gas or gas mixture in which one or more of the constituents are at their saturated vapor pressure. The constituent at saturation pressure may or may not be water vapor. Compressor Staging: All compressor elements have limiting operating conditions, regardless of type. The limitations involve the parameters listed in Table 10.3.1. Reciprocating units require a separate cylinder for each stage, with required intercooling between stages. Centrifugal units may have several staging elements in the same casing, although the diffusers and diaphragms may sometimes be water-cooled. Then the gas may be sent to an external intercooler before being returned to another casing for further compression. This may occur several times before the final compression stage, which is then followed by a final after-cooling system.


TABLE 10.3.1 Basic Operating Limitations of Compressors Discharge temperature Pressure rise Compression ratio Effect of clearance (reciprocating type) Power effectiveness Source: Courtesy of Ingersol Rand Company Limited, 1980. With permission.

It should be kept in mind that for positive displacement units the final temperature is very close to or more than that of an adiabatic process. When using labyrinth-type oil-free units for compression of oxygen, high temperatures that have often distressed operators have been observed. This feature, however, is quite commonplace, and should be accepted without fear. As for power saving, assuming ideal intercooling with no pressure loss between stages, theoretically it is wise to use the same pressure ratio in all stages. If rs is the ideal compression ratio per stage, rf the overall compression ratio including all stages, and s the number of stages, then p p

rs ¼

p ffiffiffiffi s r f

For example, for three stages, rs ¼

p ffiffiffiffi 3 rf

(10:3:5)

10.3.3 CENTRIFUGAL COMPRESSORS Centrifugal machines are sometimes referred to as radial units. Regardless of type or nomenclature, this classification is different from positive displacement machines. First of all, the rotating blade impeller is designed to produce two components of flow—a circular stream around the axis, and an inline stream through the axis. The circular flow is the result of a centrifugal force, hence the name centrifugal.

10.3.3.1 GENERAL PARAMETERS

OF

DESIGN (PER STAGE)2

Applying the first law of thermodynamics, the sum of the work delivered by the rotor and the heat generated is equal to the total enthalpy increase between the inlet and the outlet of each stage, or w þ q ¼ Dh. In addition to efficiency, the characteristics of each impeller involve the following dimensionless coefficients. Enthalpy coefficient (Yh) or the ratio of total enthalpy increase to the impeller’s specific kinetic tip energy: Yh ¼ Dht =[u22 =2]

(10:3:6)

where u2 is the impeller tip speed. Isentropic pressure coefficient (Cy) or the ratio of isentropic (or adiabatic) head to impeller’s kinetic tip energy: Cy ¼ Ys =[u22 =2]


(10:3:7)

Efficiency (hs) or mass flow isentropic head, divided by mass flow total enthalpy increase: hs ¼ Cg=Yh

(10:3:8)

Flow coefficient (f), or the ratio of the effective inlet volume flow compared to a theoretical flow product p1med by the impeller’s outer diameter (OD) area its tip speed: f ¼ Vo ={[p=4]=D22 } U2

(10:3:9)

where Vo is the effective volumetric inlet flow (m3=s), and D2 is the OD (m). Specific speed, is another stage parameter. It is formed from f and Yy: Ns ¼ f 0:5=Yy0:75

(10:3:10)

Mach number of the impeller tip, or the ratio of the impeller’s tip speed to the inlet’s sonic velocity: p

Mau2 ¼ u2 =[kpo =ro ]

(10:3:11)

where k is the isentropic exponent of the gas, and po=ro is the pressure divided by density of the inlet flow. The Mach number does not impact on the design values, also on the shape of characteristic. The Reynolds number NRe also has an influence on the stage characteristics, shaped by the impeller outlet width and the impeller tip velocity. Its impact is of special importance when Reynolds numbers are higher than 106. For normal turbocompressor applications, it suffices to take into account only the OD of the impeller. When a centrifugal machine is operating at a given speed, a single compressor wheel will generate a polytropic head that varies with the inlet capacity with a curve typical of one design (Figure 10.3.3). The same impeller blade will produce almost the same curve when handling different gases, and the same degree of variation depending upon its relative Mach number and its volume reduction through the impeller. A variation in impeller geometry will affect the stable operating range, the head rise from design to surge point, and the polytropic head at design capacity.

10.3.3.2 SPEED VARIATIONS Speed changes are normally used for capacity control of centrifugal machines. The change in efficiency is very little with small variations in speed; and in such conditions, the energy (BHP) will change as the cube of the speed for theoretical power (Figure 10.3.4).

10.3.3.3 SURGE LIMITATIONS

AND

PUMPING

As can be seen from Figure 10.3.5, the dotted line is the limit of steady flow. At this speed, and for any capacity, compressor operation becomes unstable. Depending on the impeller design, the number of stages, the shape of the head-capacity curve, and the gas being compressed, the surge curve will vary anywhere from 50% to 90%. In Figure 10.3.6a, one can observe the variation between the stability limits of multistage centrifugal units with lowpressure ratios, and those of single stage units with high-pressure ratios. The former have a longer operating range before reaching instability (Figure 10.3.6b). When the capacity of the machine is reduced below its peak design, the pressure at the discharge piping exceeds that produced by the machine, and the flow reverses for a moment.


50

100

Inlet volume (m3/min) 200 400 600

1000

2000

76

Approx. polytropic efficiency

75

74 73

72

71 70 3-E 1

2

3

4

5 7 10 15 20 30 Inlet volume—thousands of CFM

40 50

70

100

FIGURE 10.3.3 Approximate polytropic compression efficiency of a dynamic compressor versus inlet capacity. (Courtesy of Ingersol Rand Company Limited, 2006. With permission.)

Application of theory Axial flow

Efficiency

Centrifugal

Specific speed

FIGURE 10.3.4 Typical relation of specific speed and efficiency for dynamic compressors. (Courtesy of Ingersol Rand Company Limited, 2006. With permission.)


Hp-Polytropic head

Constant speed

Design point

4-Q q—inlet CFM

FIGURE 10.3.5 Typical head capacity curve of a dynamic compressor. (Courtesy of Ingersol Rand Company Limited, 2006. With permission.)

If these pulsations continue it is called pumping or surging, indicating that the machine is in distress. No centrifugal machine should be operated near its surge line. Surge points of the compressor should be determined for each proposed capacity after installation, and prevention controls should be installed.

High-pressure ratio

80

ility li Stab

Stab

Head-percent of rating

ility li

mit

100

mit

120

Low-pressure ratio

60

40

20

0 0

20

40 60 80 Inlet volume—precent of rating

100

9-AC 120

(a)

FIGURE 10.3.6 (a) Comparative characteristic curves of single- and multistage centrifugal compressors. (Courtesy of Ingersol Rand Company Limited, 2006. With permission.) continued


120

100 BHP—percent of rating

Centrifugal

Axial

80

60

40

20

0

9-AH 0

20

(b)

40

60

80

100

120

Inlet volume—percent of rating

FIGURE 10.3.6 (continued) (b) Comparison of horsepower characteristics of centrifugal and axial with variable inlet volume. (Courtesy of Ingersol Rand Company Limited, 2006. With permission.)

10.3.3.4 EFFECT

OF

MOISTURE3

Since gas mixtures, and especially air, contain moisture, it must be included in the calculation of the gas property. The moisture content in a gas affects the molecular weight, specific gravity, density, R, k, and n. The polytropic head is altered for a given pressure rise; therefore, the inlet volume must be corrected. Special charts are available for this purpose, converting relative humidity into specific humidity, at a stated ambient temperature (see Figure 10.3.7 and Table 10.3.2).

10.3.3.5 EFFECT

OF

ALTITUDE

Because a centrifugal compressor performance is dependent on inlet gas density, and density will vary with the altitude of the installation, the basic design and selection of the compressor will depend on a precise knowledge of the elevation of the installation. See Table 10.3.I in Appendix-5.

10.3.3.6 COMPRESSOR RATIO CHANGES Centrifugal compressors are as a rule ‘‘constant pressure, variable volume’’ machines. Therefore, any anticipated requirements for changes in compressor ratio must be discussed with the supplier beforehand. If any variations in capacity or speed or both are imposed on the proposed unit, it may involve special controls.

10.3.3.7 MULTISTAGING Centrifugal machines seldom involve the same reasons for multistaging as reciprocating units. The polytropic head per stage is limited, however, in order to restrict the stress on the highspeed impellers and for other mechanical reasons. The number of stages for a centrifugal


10

20

30

40

50

60

70

80

90

0.6 0.7 0.8

0

100

1.0


Inlet pressure p1 (bar A)

Relative humidity j1 (%)

27 2

Wet molecular mass Mf (kg/kmol)

1.

4

1.

6

1.

0

2.

28

5 2. 0 3. 0 4. 0 5. 0 . 7 .0 10.0 20.0 30 330 (K)

29 0

0.05

0.10

0.15 260

280

270

290

300

310

320

Absolute humidity x (−) ⫺10

0

10

20

30

Inlet temperature T1 (K), t1 (⬚C)

40

50

(⬚C) 0584 002

FIGURE 10.3.7 Determination of the absolute humidity x and the molecular mass Mf of the wet air. (Courtesy of MAN Turbo AG, 2006. With permission.)

TABLE 10.3.2 Performance Data on Radial Isotherm and Axial–Radial Isotherm Compressors Item Mass flow mf (dry) Suction pressure (p1) Suction temperature (T) Relative humidity (f1) Discharge pressure (p2) Cooling water temperature (tw) Dry molecular mass (Mf) Calculation from Figure 10.3.7 Absolute humidity (x) Wet molecular mass Mf using wet gas constant Rf ¼ 8315=Mf Calculation of wet mass flow mf ¼ mf (1 þ x) Actual suction volume V1 ¼ [mf Rf T13600]=P3600 ¼ actual discharge temperature from Figure 10.3.8

Figure 10.3.7 Type RI (3R)

Figure 10.3.7 Type ARI (6A 1 3R)

17.29 kg=s 1 bar 308 K (358C) 60% 9.8 bar 293 K (208C) 28.96 kg=kmol

136.48 kg=s 1 bar 308 K (358C) 60% 7.6 bar 293 K (208C) 28.96 kg=kmol

0.021 28.58 kg=kmol 290.94 J=kg K

0.021 28.58 kg=kmol 290.94 J=kg K

17.29 1.021 ¼ 17.65 kg=s

136.48 1.021 ¼ 139.35 kg=s

56,950 m3=h 371 K (988C)

449,520 m3=h 336 K (638C) Figure 10.3.9

Note: Industrial charts are available for obtaining the molecular weight, specific gravity, k, and n, Rf (wet) may be obtained by dividing 8315 (molar gas constant) by the wet molecular weight. Also see Figure 10.3.9 for more information. Source: Courtesy of MAN Turbo AG Publication, 2005. With permission.

machine is generally kept to a minimum, consistent with good design practice. Generally, the first stage will be subject to a somewhat higher load. The design of a multistage radial compressor is carried out on a per-stage basis. At a known mass flow m the effective volume and gas state at the inlet of each stage are calculated from the outlet state of the previous stage. Therefore, by compression and interstage cooling, the stage flow coefficients become smaller with each stage.

10.3.3.8 COOLING EFFECT Cooling of the gas during compression reduces energy consumption by lowering the average gas temperature within the unit by increasing the average density, and by decreasing the polytropic head required for a given discharge pressure. Cooling allows a reduction in speed or the number of stages or both. The reasons may involve process limitations on the discharge temperature, or economics of the machine itself. The gas may be cooled within the casing or in external exchangers (external intercoolers). Cooling within the casing may involve water-cooled diaphragms between successive stages, or direct liquid injection into the gas (see Figure 10.3.8 and Figure 10.3.9).

10.3.3.9 SPECIFIC SPEED4 The design and performance of centrifugal, mixed- and axial-flow compressors are fundamentally affected by the variables of rotative speed, capacity, and work input (head). One method of correlating these three variables is the use of ‘‘specific speed.’’ It is a function of the impeller proportions, and is a constant for any series of impellers having the same


17 16 15 14

Pressure ratio

13 12 11 10 9 8 7 6 0

5

10

20

30

40

Cooling water temperature C

4 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 180 190 Discharge temperature Determination of the discharge temperature.

FIGURE 10.3.8 Discharge temperature of RI(3R) machine. (Courtesy of MAN Turbo AG, 2006. With permission.)

11

10

Pressure ratio

9

8

7

6 0

5 30

10

40

20

30

40 Cooling water temperature °C

50 60 70 80 Discharge temperature °C

90

100

FIGURE 10.3.9 Discharge temperature of type ARI(6Aþ3R) machine. (Courtesy of MAN Turbo AG, 2006. With permission.)


proportions and angles, or for any one specific impeller operating at any speed. Specific speed or Ns is not related to rpm but is a dimensionless number. It is normally expressed mathematically as noted in Section 10.3.3.1. Note: Since variations of Ns are used, one must clarify the formula to be used before any discussions with the proposed compressor supplier take place. The design of a multistage radial compressor is carried out on a per-stage basis. At a known mass flow m the effective volume and gas state at the inlet of each stage are calculated from the outlet state of the previous stage. Therefore, by compression and interstage cooling, the stage flow coefficients become smaller with each stage.

10.3.3.10 STONEWALLING This phenomenon may occur when the velocity of the gas at some point within the machine approaches that of sound (sonic velocity, or Mach I) in the gas at the specific point in consideration within the machine. Normally, the velocities of gases in turbocompressors are well below the sonic speed, but with heavy molecular gases the designer must consider this possibility. To overcome stonewalling one may use either impellers with a different geometry or more stages.

10.3.3.11 BEARINGS Radial bearings for in-line centrifugal machines are generally of the multisegment tilting type with four tilting pads with oil inlets on both sides (Figure 10.3.10a). They should be designed for forced feed lubrication and for reverse rotation. Axial thrust bearings are usually of the Kingsbury type with self-equalized pads with directed lubrication as shown in Figure 10.3.10b. They too should be designed for forced feed lubrication and for externally induced loads.

10.3.3.12 SEALS Seals are of the commonly used labyrinth type, using dry inert nitrogen as the sealing gas and venting to the atmosphere.

10.3.3.13 LUBRICATION SYSTEM The lubrication system should follow API-614 or its equivalent. It should involve dual pumps, dual coolers, and a continuous-flow switch valve. The oil line to the compressor should be of stainless steel. Only one cooler may be used if it has a large enough capacity.

10.3.3.14 INLET GUIDE VANES Inlet volume control for any centrifugal compressor is accomplished with inlet guide vanes as indicated in Figure 10.3.11. The inlet fans can be adjusted to reduce the flow from 100% to 70% or 60%. It should be kept in mind, however, that any saving in energy is reduced proportionally only down to 70%. If the flow is below 70% of design capacity, the power saving does not follow a straight line, and is not an economic procedure.

10.3.3.15 DIFFUSERS A diffuser is a stationary passageway in which the gas velocity energy imparted by the impeller is converted into static pressure. The diffuser may be vaned or vaneless. The use of vaned diffusers is not always recommenced for oxygen compressors.


Adjusting plate

0582 0008

Oil inlets (both sides)

0582 0006

(a)

Oil injection

(b)

Adjusting plate

0582 0010

FIGURE 10.3.10 (a) Split journal bearings. (Courtesy of MAN Turbo AG, 2006. With permission.) (b) Actual thrust bearings. (Courtesy of MAN Turbo AG, 2006. With permission.)

10.3.4 AXIAL–CENTRIFUGAL COMPRESSORS5 (FIGURE 10.3.12A) A standard radial isotherm compressor referred to as RI or RIK, and shown in Figure 10.3.12a, may be combined with axial stages to offer a viable solution for very large volume flows. For equal aerodynamic loading, and for the same tip diameter of the rotor blade, an axial stage will handle a volume flow about twice that for a wide centrifugal impeller. According to MAN Turbo AG, if a centrifugal unit is preceded by an axial booster, the latter having a pressure ratio of about 2, the centrifugal section is correspondingly reduced in size and its optimum speed will coincide with that of the axial part. The two sections, therefore, can be combined in one machine, with one single rotor running on two bearings only, a design principle well known from the industrial single-shaft gas turbine. Stage and cooler optimization for the predominant pressure ratios, for air between 6 and 8, leads to a compact axial–centrifugal unit with six axial and three centrifugal stages and three pairs of intercoolers. This configuration results in optimum overall isothermal efficiency because of the higher efficiency of the axial part, and the high stage efficiency of the subsequent three wide impellers of the centrifugal section. The axial–centrifugal machine (ARI) as shown in


120 y ilit e ab id St h gu t wi

Head—percent of rating

100

t limines va

Normal stability limit Guide van es

ra

dia

l

80

g

sin

60 s

ne

de

clo

va

ui

40

G

20

0 0

20

40 60 80 Inlet volume—percent of rating

100

9-AE 120

FIGURE 10.3.11 Effect of inlet guide vane rotation on the characteristic of a constant speed compressor. (Courtesy of MAN Turbo AG, 2006. With permission.)

Figure 10.3.12b has a definite cost-effective place in air separation plants, requiring an inlet air flow over 180,000 N m3=h (equivalent to 1200 t=d of oxygen).

10.3.4.1 BEARINGS

AND

SEALS

Bearings and seals are almost identical to those used for the straight centrifugal machines with the exception of the journal bearings on the larger frame sizes, which operate at a moderate speed. In the latter case, the journal bearings are of the two-lobe type.

10.3.4.2 LUBRICATION SYSTEM This follows the same parameters as indicated for standard centrifugal machines.

10.3.4.3 INLET GUIDE VANES The system for volume control also follows the same principle as for the standard centrifugal machine (Figure 10.3.11).

10.3.5 AXIAL COMPRESSORS6 Axial compressors fit into the market for a very large capacity, and low to moderate discharge pressures. Their capacity ranges from a possible low of 50,000 m3=h to a high of 1,700,000 m3=h (12–467 m3=s). Pressure ratios involved may be as low as 2–5, and with a maximum of 7 in 1 casing. The head capacity curve is much steeper than that of a standard centrifugal machine. Therefore, the operating range from normal to surge is somewhat restricted. It may be extended, however, with the use of adjustable stator blades (Figure 10.3.12c).


FIGURE 10.3.12 (a) Standard RI. (Courtesy of MAN Turbo AG, 2006. With permission.) (b) (5Aþ3R) Axial radial machine. (Courtesy of MAN Turbo AG, 2006. With permission.) continued


FIGURE 10.3.12 (continued) (c) Axial machine. (Courtesy of MAN Turbo AG, 2006. With permission.)

In general, these machines have higher efficiencies than the centrifugal units but they require more stages for a given pressure rise. They are also more sensitive to erosion and corrosion. In view of the specific, rather restrictive operating characteristics, this machine has seen very little use in gas separation so far, and will continue to be left on the side-lines, at least in the foreseeable future. Therefore, it will not be reviewed in detail in this chapter. (Axial machines have been discussed in the chapter on liquefied natural gas.)

10.3.6 INTEGRALLY GEARED CENTRIFUGAL COMPRESSORS (API-STANDARD-672) 10.3.6.1 GENERAL OVERVIEW Since the 1950s or so, integrally geared centrifugal compressors have been rapid growing devices. As such, they deserve a detailed study. In integrally geared centrifugal compressors, as indicated in Figure 10.3.13, it is possible to configure either one or two stages on the same shaft (pinion). At present it is feasible to arrange a maximum of four pinion shafts. Speeds and impeller diameters may be selected at will. It is possible, therefore, to project optimum flow coefficients for all stages. According to the designers and suppliers of these compressors, the advantages may be stated as follows: (a) stages can be designed for optimum flow coefficient ranges due to optimal speeds and impeller diameters; (b) by cooling at each stage, an optimum isothermal efficiency is possible; (c) with the use of inlet and variable diffuser guide-vanes, a variety of geometries can be used at each stage; and (d) maximum stage efficiencies may be obtained with the use of an axial impeller at the inlet.

10.3.6.2 FUNCTIONAL COMPONENTS AND 10.3.6.2.1

THEIR

DESIGN7

Gas Side

On the gas side, operating components are the same as those an in-inline centrifugal, namely, the impeller blades; the shroud (blade channel contour); the diffuser, which may be open or vaned; the volute (collector); and the stage seals, which may be of the standard labyrinth type or the floating carbon ring seal or the dry face seal type. The dry face seal has now become the state of the art for speeds up to 45,000 rpm. It consists of a combination of stator–rotor rings arranged axially and penetrated by the leakage gas in a centripetal direction. The two inner


Fixed diffusers Cross section shows the configuration of impellers, seals, gearing and bearings in a typical four-stage H-series compressor

Radial tilting-pad, tapered-land thrust bearings

Optional adjustable inlet guide vanes allow up to 9% energy savings

Dual low-speed thrust absorption AGMA Q-13 gearing

Labyrinth seals

Groove cut into gearbox accommodates O-ring to eliminate oil leaks Horizontally split gearbox

FIGURE 10.3.13 Shows design of components on gas side of integrally gear compressor. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.)

sealing rings are separated by about 5 mm. The leakage gas loss is less than 1 kg=h. However, a sealing gas filter is mandatory. 10.3.6.2.2

Mechanical Power Side

On the mechanical power distribution side, the concept of the integrally geared centrifugal compressor departs completely from the in-line centrifugal design. It is on the gear side that


the power and speed of the driver are passed on to the single compressor stages. The bull gear is the driving gear. All gearing has to meet AGMA and DIN standards. 10.3.6.2.3

Bearings

A characteristic of integrally geared compressors is the overhung arrangement of the impellers on the pinion shaft ends. With this arrangement, separate compressor bearings are not necessary. The bearings of the pinion shafts are identical to those of the compressor rotor, which results in a substantial reduction in mechanical power losses. On the other hand, the typical integral-gear rotors with one or two overhung masses, e.g., impellers, relatively small bearings with optimized power loss, little space for required installation of the shaft seals between the bearing and the impeller, and a small number of pinion teeth, may lead to a supercritical operation of the compressor during start-up and coast-down. The most practical way to minimize this problem is to use journal bearings of the multiple tilting-pad type. In regard to axial bearings, Werner Bosen et al. of Atlas Copco Energas GmbH have assumed that the sum of the resulting axial gas forces, and possible axial forces from the single helical teeth, are either absorbed by one thrust-bearing per pinion shaft, or by specially designed thrust collars located close to the tooth mesh. The bull gear shaft is supported by multilobe sleeve bearings. One of the bearings is a combination radial and thrust bearing featuring tapered land thrust faces that allow rotation in either direction. 10.3.6.2.4

Seals

Seals employed for integrally geared compressors to avoid any contamination from lubrication oil follow the same labyrinth design as those for in-line centrifugal machines. If one uses an integrally geared machine for the compression of oxygen, it is recommended that the seal indicated in Figure 10.3.14 be employed.

10.3.6.3 ECONOMICS In regard to costs, this design of the integrally geared compressor enables specific speed requirements to be adopted more easily. Its design geometry lends itself to more frequent use of high-speed impellers, rather than having the speed limited for mechanical reasons. Impeller diameters can be kept as small as possible, as this design keeps the casing diameters small, thereby lowering the overall costs for casings and the connecting integral gearbox. What is more important, moreover, is the capability to combine two or even three separate compressor services in one gearbox. Each service can receive its own dedicated and independent process and surge controller loop. In air separation, this grouping may combine the main air compression, the high-pressure recycle nitrogen stream for liquid production, and even the expansion turbine, so long as the total number of pinions does not exceed 4. Furthermore, the overall design lends itself quite easily to prepackaging, important for the reduction of field costs (see Figure 10.2.8d). In summary, the main advantages of the integrally geared compressor are (a) the availability of more than one speed, (b) an uninterrupted axial inlet into each impeller, and (c) a compact design made possible by the integration of the compressor and the gear. One word of caution, however. The attractive efficiency of the integrally geared compressor depends entirely on how well the individual gears are fabricated, and how accurately they mesh with each other. The purchaser must keep in mind, therefore, that design, fabrication, and factory testing of the individual components must be carried out with a high degree of accuracy, and intensive supervision and inspection. Otherwise, the promised field performance may not be


Vent to compressor inlet Regulated buffer gas Atmospheric opening Vent Vent

Oil drain to reservoir

Labyrinth seal

FIGURE 10.3.14 Seal used for oxygen compressor. (Courtesy of Atlas Copco Energas GmbH, 2005. With permission.)

achieved. (This observation is the result of the review of a number of projects involving intense competition.)

10.3.7 PRODUCT OXYGEN COMPRESSORS (RIO) FIGURE 10.3.15 Although the basic design of a product oxygen compressor is similar to that of a standard centrifugal air compressor, there are several important differences necessary to achieve high safety. More compression stages and more intercooling are needed to maintain low compression temperatures, and materials of construction must be chosen with care to provide high heat conductance to avoid overheating and prevent ignition.

10.3.7.1 GENERAL OVERVIEW Because product oxygen was originally delivered in high-pressure cylinders (over 200 bar), the original machines used were water-lubricated units that had to undergo maintenance almost every week. During World War II, when the gas industry had to consider pipeline supply to large users such as steel mills, it began to use a three-stage reciprocating compressor fitted with carbon rings made of graphite. Even so, discharge pressure was limited to 30 bar. Reciprocating units were quickly replaced after the war, either by liquid oxygen pumps filling high-pressure cylinders, up to 200 bar, or with oil-free labyrinth-type reciprocating compressors supplying oxygen at a pressure of 40 bar for partial oxidation of natural gas in the production of ammonia. Finally, when the supply of oxygen reached larger and larger volumes, the gas industry began to employ oil-free centrifugal machines.


FIGURE 10.3.15 Radial machine (RIO) for oxygen. (Courtesy of MAN Turbo AG, 2006. With permission.)

After some years, however, the application of centrifugal compressors for oxygen gave rise to a series of fires that were impossible to contain, resulting in a disastrous loss of machinery, fatal injuries to operators, as well as financial loss of production. To avoid similar catastrophes, several organizations began a series of ongoing studies to set up standards for materials and operating procedures. These organizations are Compressed Gas Association Inc. (USA); American Society for Testing and Materials, International Committee GO4 (USA); Industrial Gases Committee (Europe); and the Federal Institute for Material Testing (BAM), Berlin, Germany. Since 1965, all safety minded organizations, machine manufacturers, as well as users have worked hard in unison to raise safety to a high standard, and to reduce accidents to a minimum.

10.3.7.2 GENERAL SAFETY PARAMETERS Improper cleaning at the factory: Proven by subsequent field inspection at the site and before start-up. Unsuitable inlet filter: The filtering material should be of Monel with a mesh size of 100–150 mm, and the casing should be of stainless steel. The process piping should be of stainless steel, but if the pressure exceeds 50 bar only Monel should be used. The process valves should be of Monel or stainless steel with a Monel trim. The recycle oxygen valve should be of Monel only. A study has shown that past fires have started with unstable operation of the recycle valve. All intercooler bonnets should be of stainless steel, and a recycle cooler should be supplied for the compressor with the same specifications as for the normal coolers. As for vibration, the design should follow API-670 as well as the IGC (Europe) latest edition. It is recommended that two proximity probes (Bently Nevada) be mounted at 908 to each other, located at each radial bearing, and axial position detectors should be provided at each thrust bearing. Phase angle probes should be located to measure each speed of the


compressor–gear train. It is also recommended that the DCS system include a jury system of three independent probes monitoring vibration. If two out of three sensors indicate stoppage, the machine stops operation, but if only one sensor indicates stoppage the unit does not stop. Diffusers should preferably be of the vaneless type unless the selected supplier can prove that the proposed vaned diffusers have achieved long-term safe operation in similar type and size machines. The piping and valves to supply the sealing gas to the rotor should be of stainless steel. The impellers and rotor assembly should be of stainless steel, and the shaft should preferably be a single forged piece. If the compressor includes a high-pressure stage that operates at higher than 50 bar, the assembly for that stage should be made of a bronze alloy with a wall thickness sufficient to provide adequate strength. If the oxygen compressor operates above 50 bar, the section handling the higher pressure should include all piping and valving made from Monel. The entire oxygen compressor should be enclosed within a concrete wall barrier at least 2 m high, capable of withstanding a pressure of 5 bar, and open at the top. It should also include a fire-proof door that is kept locked, requiring the permission of the supervisor or plant manager to have it unlocked. Obviously, operating controls have to be located outside the safety wall. If the compressor or any part of its ancillary equipment such as valves, piping, etc. remains inoperative for any prolonged length of time after its arrival and installation at the site, then 1 week prior to initial operation, it should be completely disassembled and inspected for absolute cleanliness. Several unfortunate experiences in the field have proven that this inspection has been necessary. Even reliable suppliers of top class compressors have had unfortunate mishaps at the factory, which were corrected only in the field by such last minute inspections.

10.3.7.3 SAFE OPERATION

OF

CENTRIFUGAL OXYGEN COMPRESSORS

A source of danger is the possible deflection of the rotor. A shaft deflection can create damage to seals, labyrinths, and bearings, and in pure oxygen cause a fire. To avoid such incidents the compressor should undergo two serious tests before being shipped from the factory: 1. Aerodynamic tests to verify design calculations to determine performance at design point, at 120% capacity, at surge, and at one point between design and surge points. 2. Mechanical tests on the machine and coupled-to-gearing shaft vibration measurements at bearings should be mandatory to determine shaft vibration from 0%–110% of rated speed. When completed, the casing should be opened, and the machine inspected for labyrinth rubs and clearance. Prior to installation, the surrounding area should be cleaned thoroughly from oil and debris accumulated during construction. Control valves should be fully stroked assuring their functioning. All alarm, shut-down switches, and ancillary devices should be checked by simulating operating conditions. Prior to start-up the oil system should be flushed, and before the auxiliary oil pump is started, seal gas pressure should be established to compressor shaft seals. The vent ejector on the oil reservoir drain section should be adjusted to maintain a slight vacuum in compressor bearing housings. Once preparations for start-up are completed the machine should be operated with oil-free nitrogen to ensure mechanical integrity of the unit. To simulate the gas density of oxygen, an additional inert gas such as CO2 may be added to reach the molar mass of oxygen. These conditions will permit a check for rotor instability and adjustment of antisurge control system


prior to introducing oxygen in the machine. After this test the bearings and suction screen should be disassembled and inspected. Once preparations and inspections are completed and found satisfactory, the unit may be restarted with nitrogen as before, and operated until it reaches normal operating temperature. If everything performs in a satisfactory manner, the throttle valve may be opened slowly, allowing oxygen product to enter the machine while simultaneously cutting back on the nitrogen flow. Oxygen flow should be gradual over a period of several hours to prevent any sudden oxidation and ignition of small particles. During this period, vibration levels, bearing temperatures, and oil and gas temperatures should be closely monitored. Rotor instability should be closely watched.

10.3.7.4 ULTRAHIGH-PRESSURE OXYGEN COMPRESSORS8 In the early 1976, because of the promising market for synthetic fuels, a consortium decided a few years ago to develop an industrial-scale high-pressure oxygen machine, without going through laboratory development. The test pressures selected were 62 barG at the inlet, and 124 barG at the outlet. The oxygen volume was 1675 N m3=h. The machine developed was the RB28-4, driven by a GE-2500 HP motor. Polytropic efficiency was given as 66% at peak efficiency, but the average was 60%. The machine had oval bearings, and the speed increasing gear was a BHS gear. The coupling was a solid type, and the aftercooler was a shell and tube type with the oxygen in admiralty metal tubes. Tube sheets were of Monel equipped with four separately located bursting disks set at 19 barG. All ancillary piping was stainless steel, but all elbows and valves were of Monel. For all tees the impinged wall had a Monel shield. The intake filter was of (100 mesh) sintered bronze. The process design employed is given in Table 10.3.3. There was no intercooler. Seal losses at 6.9 and 13.8 barG were recovered and sent to existing pipelines at equivalent pressures. The labyrinths were heavily lined with silver and had Monel strips. In the first test run of 300 h, the following procedures were carried out: . . .

Surge at 138 barG Trip at 138 barG Steady load at 124 barG

After the first test run, the machine was opened and a few material deformations were observed in the inner casing, and part of the sealing section of the balance piston. It was decided, therefore, to thicken the cross section of the casing, and some changes were made to the assembly of the strips. After the final run of 300 h, the casing was opened up for public inspection. There were no visible defects on any of the parts exposed. The final test was considered completely successful. The vibration limits (peak to peak) were very small even TABLE 10.3.3 Design of High-Pressure Stage of Test Compressor Design flow Design pressure Hydraulic test pressure Discharge pressure Isentropic efficiency

17 kg=s 165 barG 248 barG 138 barG 59.4%

Source: Courtesy of MAN Turbo AG Publication, 2005. With permission.


TABLE 10.3.4 Comparison of Efficiencies of Various High-Pressure Compression Machines Centrifugal air compressor Reciprocating on nitrogen Centrifugal oxygen compressor Experimental machine

76% 73% 70% 66%


during trip, or passing through the critical. At the most it was 1 mil. During the normal run it was around 0.25 mil. The observations made in comparison of compression efficiencies are given in Table 10.3.4. In an approximate analysis comparing costs and power for an oxygen plant with a capacity of 1500 t=d and at a pressure of 124 barG, the following may be noted:

Gas plant with oxygen compressors LOX plant with LOX pumps and nitrogen recycle

10.3.7.4.1

Relative Power

Relative Costs

100 110

100 102

Summary

Although the market for oxygen compression in the range of 124 barG has not yet developed because of international economic conditions, the technology of achieving ultrahigh oxygen pressures over 100 bar has already been exhibited as a strong possibility.

10.3.8 HIGH-PRESSURE LABYRINTH PISTON COMPRESSOR9 10.3.8.1 LABYRINTH PISTON COMPRESSOR 10.3.8.1.1

General

This compressor was conceived and built in 1935, and its design was based on the principle to seal the piston and the piston rod gland, so as to keep the gas being compressed as pure as it was when it entered the compressor. In 1949, this design was used for the first time for compressing oxygen from atmospheric to 150 bar. Since then, many similar units have been used not only for product oxygen and nitrogen compression, but for such gases such as helium, carbon monoxide, and ethylene, wherein high purities must be maintained without losses through leakage.

10.3.8.2 BASIC DESIGN

FOR

ACHIEVING OIL-FREE OPERATION

As indicated in Figure 10.3.16 and Figure 10.3.17, the machine is upright, which eliminates any possible misalignment; and the upper guiding element of the unit is not placed in the cylinder, but in the oil-lubricated crankcase. The design achieves a separation between the oillubricated crankcase and the oil-free cylinder. This is accomplished by using a distance piece that is longer than the piston stroke, and an oil scraping system for the piston rod. During the up-stroke, the oil scraper rings efficiently remove the oil from the piston rod before its entry into the distance piece.


Labyrinth piston

Piston rod gland

Piston rod gland

Oil scraper rings

Oil scraper rings Upper guiding element

Non-lubricated

Piston with upper guiding element

Lower guiding element

Oil-lubricated

Lower guiding element

FIGURE 10.3.16 Principle of labyrinth piston oil-free compressor. (Courtesy of Burckhardt Compression, 2006. With permission.)

10.3.8.3 SEALING SYSTEMS The necessary throttling effect is achieved by having rings arranged in series, and are considered a small number of throttling points with a very narrow gap between the rings and the counterpart. Each ring has two leak paths requiring static and dynamic sealing capabilities. If the gap has to be zero for a long period of time, a certain amount of wear ensues, which means that the rings may require regular maintenance.

10.3.8.4 LABYRINTHS Normally this involves an extensive series of throttling points. In the case of this design, it has a larger number and a slightly wider gap between the labyrinth tips and the counter-surface. For the machine in question, there is no permanent mechanical friction between the sealing labyrinth and the counterpart.

10.3.8.5 INTERNAL OPERATING ELEMENTS Figure 10.3.18 shows the internals of the machine. The casing and cylinder block are made of cast iron for standard machines; and the crankshaft and connecting rod are, apart from a few exceptions, made of forged steel. Depending on the specific application, the piston rod is made of either stainless steel or nitrided steel. The piston rod–guide bearing is combined with the oil scraper rings at the top end of the bearing, and serves as the upper guiding element for the piston rod. The bearing casing is


Design features of the D-and E-type compressors with open distance piece

1 2 4

5

6

8

7

Gas delivered Cooling water Oil 1 Cylinder 2 Labyrinth-piston 3 Compressor valve 4 Labyrinth-piston rod gland 5 Extra long distance piece

6 Oil-lubricated guide bearing with oil scrapers 7 Lubricating-oil pump 8 Crankshaft seal

FIGURE 10.3.17 Internal details of the labyrinth compressor. Design features of the D- and E-type compressors with open distance piece. (Courtesy of Burckhardt Compression, 2006. With permission.)

made of cast iron, and the splash-lubricated replaceable bushing consists of bronze with a thin babbitt layer. The piston rod–gland seals between considerably different pressure levels by means of floating labyrinth rings. The leaking gas is collected in a ring chamber at the lower end and, in a standard compressor, is piped back to the suction side first stage, hence, no leakage to the outside. The double acting labyrinth piston involves three pieces only, the two piston covers, and the piston skirt with the labyrinth grooves.



FIGURE 10.3.18 Internal operating elements. (Courtesy of Burckhardt Compression, 2006. With permission.)

10.3.9 COMPRESSOR DRIVERS10 Not only must the driver operate the machine at its design rating, but it must also start the unit from rest, accelerate it to full speed, and then maintain operation under any condition of capacity and power called for by the process design. At present, there are a variety of drivers available for this work: electric motors, high-pressure steam turbines where steam is available as a by-product, and gas turbines in cogeneration plants. In regard to electric motors, which are the most widely selected, they should not be chosen for any other than full-load continuous operation. Frequent interruptions in operation will lead to overheating and deterioration of the insulation. For very large motors, say over 1000 hp, start-ups should be restricted to no more than two per day. As for insulation and ratings see Table 10.3.5.

10.3.9.1 MOTOR TORQUE An important characteristic of the electric motor to be selected is the motor torque, which is the ability to start the compressor from complete rest and accelerate it to full operating speed. This must be considered apart from the 100% full load operating torque. The formula for its calculation, in Newton meters is Torque (N m) ¼ [9543 kW]=rpm

(10:3:12)

The starting torque is the critical value because it occurs at zero speed when the normal voltage and frequency are applied. The breakdown torque occurs when speed drops in an induction motor, and the same torque is called a pull-out torque for a synchronous motor. In the case of a synchronous motor, which starts and accelerates to around 97% speed as an induction motor, there is also a pull-in torque. This is the torque necessary to accelerate the TABLE 10.3.5 Drivers for Stationary Compressors

Type Enclosure Open, including drip proof, splash proof, pipe ventilated, and weather-ventilated Totally enclosed, fan-cooled, and water–air cooled Motors with encapsulated windings Totally enclosed, nonventilated

Temperature Rise of Insulation Windings (8C)

Temperature Determ.

Class A

Class B

Class F

Class M

RTD

60

80

105

125

RTD

60

80

105

125

RTD RTD

65 65

85 85

110 110

135

Typical ratings follow: Class B insulation 1. 908C with 1.15 service factor at service factor load. 2. 808C with 1.0 service factor at rated load. Class F insulation 1. 1158C with 1.15 service factor at service factor load. 2. 1058C with 1.0 service factor at rated load. Atmospheric conditions materially influence insulation life. The motor manufacturer should be advised of all local conditions, such as: higher ambient temperatures; high humidity; high torque requirements; and the presence of hazardous atmosphere, water, sand, and other corrosive elements. Source: Courtesy of Ingersol Rand Company Limited, 1980. With permission.


unit to synchronous speed against the existing load. The pull-in torque must be considered in all compressor applications.

10.3.9.2 ENCLOSURES A proper enclosure is a matter to be discussed among the purchaser, insurance company, and the local safety-enforcing authority. There are a variety of enclosures available for induction motors, such as open, drip proof, weather-protected, weather-protected-fan-cooled, and totally enclosed explosion-proof and certified. For synchronous motors, the variety of enclosures also include open, drip proof, splash proof, outdoor, but weather protected, self-ventilated (pipe-type), force ventilated, totally enclosed, water–air cooled, and totally enclosed and inertgas filled. For synchronous motors with collector rings there are enclosures available for the rings.

10.3.9.3 POWER FACTOR The choice of the driving motor will be influenced by its effect on the overall plant power factor. The power factor is the cosine of u (phase angle), and may be either negative (lagging) or positive (leading), depending on whether the current (I) curve lags behind or leads the voltage (E ) curve. A lagging power factor may indicate a reduced ability of the plant to carry extra capacity; more current per kilowatt or higher line losses; or a reduction in voltage, hence a reduction in plant efficiency. The overall effect on the power bill may cause the power supplier to impose a penalty. On the other hand, a leading power factor may indicate the possibility of adding a new compressor without any major changes. Some power suppliers may even offer a bonus if the resultant power factor is above the minimum specified.

10.3.10 OPERATING RELIABILITY VERSUS CAPITAL COSTS OF COMPRESSORS The question of using a single air-compressor or two half-size machines for air separation plants is as old as the gas industry itself. Economy dictates the use of a single unit, whereas production reliability points to the use of two machines. To resolve the question, a visit was made to a large oxygen plant complex near Houston, Texas, which operated a 1500 t=d oxygen plant producing argon and other rare gases, as well as pure nitrogen. The unit was driven by a single integrally geared-type compressor. This advantage, however, had to be counterbalanced by the fact that the large oxygen plant was only one out of a complex that included other air separation units as well as large cryogenic LOX storage tanks. To arrive at a rational conclusion, large international suppliers were interviewed to discuss the perennial problem of compressor purchasers. Needless to say, both parties very strongly defended the reliability of their machines. At the same time, both made the same recommendations for following certain operating procedures regardless of capacity. In the first place, the machine should be factory tested at full speed using its own bearings, as well as fully assembled with its own housing, bolts, and all ancillary equipment. Then, after it is delivered, starts-up for operation, and has run for a period of one year, the entire machine has to be dismantled completely and inspected thoroughly. The impellers have to be cleaned and then checked for hairline cracks; the complete rotor assembly has to be checked for proper balance; the oil system, filters, and pumps have to be examined. The overall instrumentation and vibration probes have to be checked, and the electrical system and field control panel have to be maintained. Once the overall machine has been given a clean bill of health, the unit can be restarted.


After the machine has operated for two additional years, it has to be stopped again and the entire procedure repeated without exception. Once this is carried out, the operating department of the owner can decide if the operating time interval between inspections can be extended beyond the normal two-year period recommended by the suppliers and prescribed in their operating manuals. Nevertheless, onstream operation of the compressor should never exceed four years, and this only in exceptional cases where ambient air is clean. This is never the case in metallurgical areas, and hardly ever in petrochemical areas. It may be of interest to note that when the author in a consulting capacity was present at a steel mill in Brazil during a stoppage for planned maintenance on a single compressor, he learned that the machine had been in continuous operation for nine years because production could not be interrupted. During this long interval, the upstream bed of activated alumina in the molecular sieve pre-purification system had been changed twice without stopping the system. Remarkably, the molecular sieve downstream of the activated alumina was still in operable condition. The bull gear was also in good shape. The radial bearings were slightly worn but were earmarked to be sent to the mechanical shop for repair, and replaced in the spares bin. The thrust bearings, however, were completely beyond any possible repair. This unusual case should not tempt anyone, however, to allow a compressor to go this long. The result could be a disaster. Major compressor suppliers have complained that many customers spend a great deal of money to purchase a high-quality machine, and then try to save pennies on the purchase of a cheap ineffective air filter. The last stage of a suitable air filter for an air separation plant should remove at least 99% of all particles larger than 1 mm. Returning to the principal question on how long it takes to accomplish a complete planned maintenance program as recommended and printed in the operating manuals of compressor suppliers, according to both Demag and Sulzer, it takes about two to three weeks, depending on whether one uses union or nonunion labor. This has been confirmed from the field experience of major industrial users of air separation plants. Merchant plants owned and operated by the gas industry and generally concentrated in highly industrial areas are adequately supported with large liquid storage back-up systems. If an unforeseen breakdown occurs, the plant owner can easily obtain an LOX supply from a competitor in the same area due to an unwritten agreement within the gas industry. Furthermore, with a multiple plant system, there is very little need for more than one air compressor. If, on the other hand, a merchant plant supplies customers by pipeline, then conditions change. To service customers reliably, the supplier immediately considers multiple compressors, multiple air separation units, a pipeline network, a large LOX-product storage for back up, or any combination of these possibilities. For metallurgical plants, especially in the steel industry, there is no such thing as down time. Blast furnaces keep going almost forever. Molten hot metal from the blast furnace cannot be stored economically. For this reason, steel mills operating their own oxygen plants have either multiple separation plants or multiple compressors. In fact, several steel mills in Brazil have spare 100% air compressors piped and hot-wired in line, so that one machine can stop for programed preventive maintenance while the spare starts up. In North America and in Europe, where the majority of steel mills purchase oxygen via pipeline, the supplier makes sure that at no time the supply of oxygen will stop. This forces the use of multiple plants connected by a network of pipelines or the use of multiple compressors if only one plant is in operation. Petrochemical and chemical plants, including those for syngas production, generally operate continuously for one to two years. Down time is not tolerated. On the other hand, they do have a turnaround after two years for a general preventive maintenance program that can last from two weeks to one month depending on the complexity of the main process. This is


always true if partial oxidation is used. In such circumstances the use of a single air compressor is also a viable choice, given the present reliability of centrifugal machines.

10.3.10.1 RECOMMENDATIONS There was one common thread in the review of all cases. Although small emergency repairs could be executed in a few days, the program for a planned and complete overhaul every two, three, or four years as recommended by air compressor suppliers is sacrosanct. The reliability of the radial compressor in both the in-line centrifugal and the integral geared machine has improved over the years. So has the rectification process of a well-designed air separation plant. Thus, it has been possible to operate a 1200 t=d oxygen plant in an isolated metallurgical area, at 50% capacity with only one half-size air compressor, and only one halfsize product oxygen compressor, and with a high efficiency in energy consumption. One sole compressor can service a single air separation plant operating at 50% – 45% capacity, but at a higher power cost. On the other hand, in an industrial area that includes a mechanical service outlet for the compressor within a radius of 150 km, it very much possible to operate with a single radial machine with complete reliability. It is assumed, however, that the quality of the machine, its testing during manufacture, and its installation, are given top priority.

10.3.11 APPLICABLE COMPRESSOR CORRECTION FACTORS Because compressors are major consumers of energy, they form an integral part of performance warranties and liquidated damages if a process does not perform according to contractual guarantees. Generally, after a performance test, actual instrument readings of energy consumption are compared with the design figures as promised in the contract. Actual readings taken in the field are then corrected in accordance with variations in ambient temperature, barometric pressure, humidity, and cooling water temperature. To simplify the exercise, an example is set forth involving an air separation plant with gaseous and liquid productions including a separate booster compressor, and an adsorption system for process air pre-purification upstream of the cryogenic unit. Outlining the compressor system the following design and actual readings are used to determine the correction factors (K). Note that cryogenic and other pumps do not undergo any correction. If they are included in the overall as-read power consumption they should be removed during the application of correction factors, and then reinstated afterward for a final total energy consumption.

VARIABLES Wcc Wcb KWcc KWcb Pod Podb Pid Pia xd xa Td Ta Twd

compressor power of main air compressor (MAC) compressor power of booster air compressor (BAC) as-read power of MAC as-read power of BAC design discharge pressure of MAC ¼ 579 kPa A design discharge pressure of BAC ¼ 979 kPa A design barometric pressure ¼ 98.36 kPa G actual barometric pressure ¼ 95.833 kPa design humidity ¼ 0.012 kg=kg of dry air actual average humidity from local airport (60.02%) design ambient temperature ¼ 293 K average temperature during test ¼ 284.81 K design cooling water temperature ¼ 294 K


actual cooling water temperature ¼ 284.81 K approach on first intercooler of MAC ¼ 7.28C approach on second intercooler of MAC ¼ 2.58C (higher efficiency because of plate exchanger) approach for BAC intercooler ¼ 3.38C (calculated using graph of cooling water tower with a factor of refrigeration of 104 K) motor efficiency taken as 99.5% or factor ¼ 0.995 same as previous but actual reading during tests was ¼ 5.68C instead of 3.38C approach after passing through direct contact water system ¼ 2.58C, which only affects the booster compressor

Twa DTw DTw DTw Mh DT DTw

Example of applying electrical correction factors. Let us consider a variable load air separation plant with an oxygen production capacity of 308 t=d for a metallurgical operation producing 98% oxygen. This includes a small quantity of high purity oxygen and nitrogen including LOX or LIN (20t=d) to be stored for sale in the area. The plant will include : 1. 2. 3. 4. 5. 6. 7.

A variable load cold box. A standard water cooling tower. A three stage air compressor. An expansion turbine directly coupled to a booster compressor. A direct contact water cooler. An evaporative water chiller to drop the process air temperature to approximately 108C. A molecular sieve unit for elimination of dangerous hydrocarbons and drying purposes.

In these conditions what is the actual power consumption of the main air compressor and the booster compressor individually? The overall correction equation is as follows: ½ðK2 þ K3 þ K4 þ K5 Þ=3 K1 0:995 for the main air compressor: K6 K7 0:995 for the booster compressor where K1 ¼ the log of the design pressure ratio divided by the log of the actual ratio pressure of main compressor K2 ¼ the weight of water as a fraction of dry air (from design) K3 ¼ the correction due to the difference of ambient temperatures. K4 and K5 ¼ 7.58C is the approach temperature of the intercoolers of the main air compres sor; 2.58C is the approach temperature of the highly efficient intercooler of the booster compressor; 3.38C is the approach temperature of the water cooling tower and 5.68C is the approach temperature of the water cooling tower during the actual operating tests. K6 ¼ 2.58C indicates the influence of the water cooling tower on the booster compressor. K7 ¼ the correction factor due to the difference in the design and actual pressure of the booster compressor. 0.995 ¼ there is a loss of 0.005% in power between the compressor motor and the actual meter.


SUMMARY OF FACTORS K1 ¼ log (579=98:36)= log (579:95=95:833) ¼ 0:9846 K2 ¼ (1 þ xd )=(1 ¼ xa ) ¼ (1:012)=1:0097) ¼ 1:0023 K3 ¼ Td =Ta ¼ (293)=(284:81) ¼ 1:0288 K4 ¼ (Twd þ DT2 )(Twa þ DT2 ) ¼ (294:1 þ 7:2 þ 2:5 þ 3:3)=(284:1 þ 7:2 þ 2:5 þ 5:6) ¼ 1:0257 K5 ¼ (Twd þ DT3 )=(Twa þ DT2 ) ¼ (294:1 þ 8:3 þ 2:5 þ 3:3)=(284:1 þ 7:2 þ 2:5 þ 5:6) ¼ 1:0270 K6 ¼ ðTwd þ DT4 )=(Twa þ DT4 ) ¼ (294:1 2:1 þ 3:3)=(284:81 2:1 þ 5:6) ¼ 1:0242 K7 ¼ log (979:0=98=:36)= log (979:0=95:833) ¼ 0:9888

APPLICATION OF FACTORS KWMAC ¼ [(K2 K3 þ K4 þ K5 )=3] K1 0:995 3838=4156 (ratio MAC= total KW) ¼ [(1:0023 1:0288 þ 1:0233 þ 1:0270)=3] (0:9846 0:995 3838=4156) ¼ 0:9293 K6 K7 KWBAC ¼ [(1:0242 0:9888 0:995 318Þ=4156] (ratio BAC=total KW) ¼ 0:0771

REFERENCES 1. Loomis, A.W. 1980. Compressed Air and Gas Data, 3rd ed, 2-1–3-23. Ingersol-Rand. 2. Bosen,W., Wagner, B., Ispas, I. Integrally Geared Technology Lowers Capital and Operating Costs. Atlas Copco Energas Gmbh, 2nd ed, 3–4. October 2000. 3. Sulzer Escher Wyss. Sulzer Turbocompresssors, 11–12. Bulletin 28.14.10.40 RIK, ARI. 4. Hans Gartmann. De Laval Engineering Handbook, 3rd ed, 6–76. 5. Sulzer Escher Wyss. Sulzer Turbocompressors. 11, 12, 22, 23. Bulletin 28.14.10.40, ARI. 6. Sulzer Escher Wyss. Sulzer Turbocompressors, 4–5. Bulletin 25.01.10.40. 7. Bosen, W., Wagner, B., Ispas, I. Integrally Geared Technology Lowers Capital and Operating Costs. Atlas Copco Energas Gmbh, 2nd ed, 7–10. October 2000. 8. For further information contact Sulzer Turbocompressors International, 2901 Wilcrest Drive, Suite 450, Houston TX 77042 USA. 9. Labyrinth-Piston Compressor for Oxygen Service, Sulzer Burckhardt, Report No. 0911W=Ve=ja, 10.03.88. 10. Loomis, A.W. 1980. Compressed Air and Gas Data, 3rd ed, 15.1–15.15. Ingersoll-Rand.



10.4

Valves and Valve Stations for the Cryogenic Industry

10.4.1 GENERAL DESIGN AND MATERIALS The proper selection of valves is a major consideration in the design and operation of a cryogenic process system. Basic valve-designs may appear no different than those available for other chemical processes, but the very fact that cryogenic valves have to operate efficiently, smoothly, and with no leakages at temperatures between 100 K and down to 4 K imposes exacting specifications in the design and materials of fabrication. Obviously, during start-up of any project, ambient temperatures will be close to 300 K. Furthermore, capital costs cannot be ignored. The overall cost of valves runs between 20% and 30% of the total piping costs. But undersizing magnifies pressure losses across the valve. The latter is unacceptable in the expansion (throttling) of cryogens where precision is necessary to avoid vaporization, and achieve a specific precalculated temperature drop downstream of the valve. In the cryogenic industry, valve sizes have increased to 48 in. (DN1200) and cycle speeds have increased to fractions of a second. Valve materials must maintain their ductility at temperatures of 77 K (liquid nitrogen) and well below to 4 K (liquid helium). Selection of shafts, seats, and seals should also meet the compatibility of individual cryogens, especially in high-pressure oxygen service. Valves should also be required to operate over a wide range of excursions, and go from ambient temperatures of 300–77 K in a matter of a few minutes. Seals should be liquid-tight in shutoff over the complete operating temperature range. As cryogen production involves high-energy consumption, valves should be designed with a low thermal mass. Finally, valves should be easy to disassemble for cleaning and maintenance. In general, valves operating at 77 K and below use a vacuum-jacketed wall to minimize the quantity of heat transfer and boil-off of cryogens. Certain control valves operating within the cold box allow the complete removal of the seat and trim outside the casing for easy maintenance. Bonnet extensions should be rugged to absorb stresses associated with the torque of the valve and the weight of a fail-safe or spring actuator. Finally, the valve seat should be of Kel-F (PCTFE) as the preferred material because of its minimal thermal contraction and indicated good mechanical properties during operation at low temperatures. It is, however, very expensive compared with conventional Teflon.

10.4.2 GENERAL DESIGN IN SIZING It is useful to express the capacity and flow characteristics of a valve in terms of a flow coefficient known as Cv or Kv (the latter is used in Europe):


Cv ¼ Rate of flow of water in US gallons per minute at 608F at a pressure drop of 1 lb=in.2 across the valve. Kv ¼ Rate of flow of water in cubic meters per hour (m3=h) at a pressure drop of one kilogram force per square centimeter (kgf=cm2) across the valve. 1 kg=cm2 ¼ 0.98065 bar (exactly); and if a kilopound is used: 1 klb=cm2 ¼ 1 kg=cm2. Cv ¼ 0:0694Q

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi [r=Dp(999)] (in US gallons)*

(10:4:1)

Q ¼ Rate of flow in L=min r ¼ Density of flow in kg=m3 Dp ¼ bar Converting the above equation in terms of pressure drop: Dp ¼ (Q=Cv )2 r 4:7952 US gal=min

(10:4:2)

Generally, instead of calculating pressure drops it is easier to use the resistance coefficient factor K, which indicates the static head loss through a valve. K varies with each type of valve and its value is given in an equivalent length of pipe diameters L=D. The values of K may be found in any valve handbook or supplier’s catalog such as Crane’s Technical Paper on ‘‘Flow of fluids through valves, fittings and pipe.’’ The latter also has a copy using the metric system.

10.4.3 SIZING PARAMETERS The above equations show that flow, density, and the desired pressure drop must be known in order to calculate the coefficient Cv. The final valve selection is important. If the valve selected is oversized it will be higher in cost, and will operate too close to the seat for good control. On the other hand, if the valve is undersized in order to save money, the increase in pressure drop will prove costly in energy, and it will have to operate almost fully open with very little extra capacity to control impacts of surge. It will be prudent to keep in mind that at very low temperatures gases may not behave as ideal gases. Also, one has to take into account the g factor (ratio of specific heats). Furthermore, cryogenic liquids may be only partially subcooled and flashing may occur within the valve. In the latter case, trying to control the process will be very difficult, especially if the two-phase fluid in the valve varies from minute to minute. The process valve will be handling a two-phase fluid, although designed and selected for a single-phase fluid. This condition will upset the steady-state operation of the process. Very often plant designers supply the valve manufacturer only with a range of flow rates and pressure drops, leaving it up to the supplier’s application engineer to come up with the right answer. This is a common source of trouble. In the final selection, the following factors must be observed: 1. Process efficiency in terms of product recovery, purity, and steady-state operability. 2. Mechanical integrity against valve failure and back-up requirements. 3. Minimum leakage in terms of lost product and safety. A leakage rate of 1 g=min is the equivalent of 500 kg=y (1=2 t=y). *As yet there is no international agreement for the SI equivalent.


10.4.4 VALVE CATEGORIES Because the majority of the valves used in cryogenic projects handle gases or fluids at comparatively low pressures (less than 35 barA), selection with regard to pressure is either Class 150 or Class 300. Liquefiers are the exception, however, because they may operate up to a pressure of around 55 barA. The selection with regard to materials, therefore, is more exacting. Valves may be divided roughly into three categories: 1. Warm valves that are part of the operating process, but are located upstream of the cold box. In this category, the valves may be either plug or butterfly valves, and may be of carbon steel. In this service, however, butterfly valves are subjected to a high torque so that the disk and stem–shaft may require the use of stainless steel for large diameters (12 in. and over). If pure oxygen is involved or if the plant is near a sea coast with a saltladen atmosphere, the use of Monel or other alloys may have to be considered. 2. Cryogenic isolation (block) valves, which are fully open or fully closed. In this category both ball and gate valves are used. The ball valve requires stainless steel for body and ball; but for gaseous oxygen conditions, irrespective of temperature, the ball should be made of Monel. The gate valve should have a stainless steel body but the disk–wedge should be of bronze. 3. Cryogenic process control valves that may be either pneumatically or electronically operated. In this category, the valves used may be globe, butterfly, or angle valves. They may be of stainless steel (preferably), aluminum, or bronze if the latter is applicable. Gate valves should not be used for throttling services as erosion may set in or chattering may be induced by a high-velocity fluid flow, both resulting in leakage. Chattering may also result in severe damage.

10.4.5 NONMETALLIC MATERIAL With regard to nonmetallic material for valves, the engineer has a wide variety of choices. In the field of cryogenics, however, the choice becomes a little more restrictive especially if one of the products is oxygen. Some of the principal nonmetallic materials are fluoropolymers, such as PTFE, PCTFE, and FEP. These materials have outstanding physical and chemical properties. They have high-temperature stability, low-temperature flexibility, low coefficient of friction, abrasion resistance, nonflammability, good weatherability, and a wide temperature service range. As noted previously, Kel-F or its equivalent has certain advantages in spite of the higher cost. It should be kept in mind that PTFE, however, if used for valve seats or seals may result in leakage over a period of time because it has a tendency to creep. That is, it lacks elastic memory. PCTFE has given better results for valve seats and seals. The materials may include fillers such as powdered glass, ceramic, metals, and metal alloys impregnated in the basic material. It is always prudent to discuss with the valve manufacturer the specific conditions within which the valve must operate before arriving at a final choice. If flanged valves are used, a good quality of compressed asbestos will be found more useful than most of the fluoropolymers available for gaskets.


10.4.6 MANUFACTURERS OF FLUORINATED POLYMERS A list of fluorinated polymers and manufacturers is given in Table 10.4.1. TABLE 10.4.1 Fluorinated Polymers and Manufacturers PTFE Teflon Fluon Halon TFE Polyflon Hostaflon Teflon PCTFE Kel—F 81 Aclar, Aclon Daiflon Voltalef FEP Teflon FEP Teflon PFA

Polytetrafluoroethylene DuPont ICI Ltd Allied Daikon Hoechst Mitsui Polychlorotrifluoroethylene 3 M Company Allied Co Daikin-Kogyo Ugine-Kuhlman Fluorinated ethylene–propylene DuPont DuPont

Source: Courtesy of Manufacturers’ Literature. With permission.

10.4.7 WARM END SWITCHING VALVES Before any gas can be liquefied, it must be purified for the removal of water, carbon dioxide, and other hydrocarbon contaminants. As noted in previous chapters, the case of atmospheric air separation, the process of pre-purification generally used is the temperature swing adsorption (TSA). Actually it should be called pressure temperature swing adsorption (PTSA), because the pressure is lowered to almost atmospheric before temperature is applied. This system involves two vertical and parallel vessels, each holding a combination of activated alumina (upstream of the main process flow), and a layer of molecular sieve adsorbers downstream to take care of carbon dioxide and hydrocarbon removal. For radially designed vessels the alumina is installed around the perimeter of the molecular sieve adsorbent (see Chapter 5). In the PTSA system, there is a flow of process air through one adsorber unit until it becomes saturated, and then the full flow of air is switched through the second adsorber unit while the first unit is being regenerated. This means that the full flow of air at about 5–8 barA has to be switched every 2–8 h depending on the size of the plant and the procedure recommended by the process engineer. The process pressure in the saturated vessel is turned off to near atmospheric, and hot dry nitrogen is sent as a scavenger gas. At the very worst case, say for a plant capacity of 2000 t=d of oxygen, the switching valves go through about 5000 cycles per year. For this operation, butterfly valves are most commonly used. They are specified for mechanical strength at the shaft because the torque is very high. They are generally made of carbon steel, and should be of the high-performance tricentric type, so that the disk is parallel to the seat before closing. The packing material and lubrication material are also important because of dusting from adsorber and alumina fines. Leakage is unacceptable; therefore metal-to-metal seats cannot be used. Moreover, valve material has to withstand a temperature range from 278 K (temperature of the process air) to 410–475 K


(temperature of the waste gas regenerating the adsorbate), at every cycle of say, 2–4 h or more for small plants.

10.4.7.1 WARM END REVERSING VALVES If the plant is designed to pre-purify the air stream using reversing exchangers, the system is more commonly known as ‘‘Revex.’’ In this system, the reversing exchangers operate in temperature ranges from 313 to 100 K, and the problem of reversing valve selection becomes more acute. Normal and dangerous impurities are frozen out by the temperature changes of the reversing exchangers from 310 K at the warm end, to100 K at the cold end. For about 10–15 min, the air at a pressure of 5–8 barG enters one series of passages in the same exchanger core. The process air is switched through another series of passages in the same core while the original passages are cleaned up by a counterflow of waste nitrogen at a pressure of about 1 barG. This means that the process-air flow and the countercurrent waste nitrogen flow are switched every 10–15 min, or 4 reversing cycles per hour, or 50,000 cycles per year, regardless of plant size or capacity. Normally, reversing valves for a Revex system should be specified to withstand a minimum of 150,000 cycles without failure. In the past, only the eccentric-type plug valve has met and even surpassed this minimum requirement. This type of valve, however, has been found heavy, cumbersome, and costly, especially in very large plant capacities (1000 t=d and over of product oxygen). It should also be noted that during the start-up of a Revex plant, though the front-end system is operating for the removal of water and carbon dioxide, the cycle is very short, somewhere between 1 and 1.5 min so that a greater number of cycles are necessary. With the steady increase in plant capacity, the use of the simpler and less costly butterfly valve has become more widespread for the Revex system. In their original application, even in small sizes (12 in. diameter), there were some reports of shaft failure and leakages. Much to their credit, valve manufacturers have responded to the challenge, and more recent field reports have shown an improvement mainly due to the elimination of a metal to metal contact and with the use of more suitable PCTFE (Kel-F) in the seat of the seal ring. In some cases, the PCTFE seal ring has had an aluminum back-up ring for better results against leakage. Butterfly valves with a diameter of 24 in. are already in use for many air separation plants. In fact, valves with a diameter of 32 in. and even of 36 in. have been in operation for some years with no apparent problems.

10.4.7.2 WARM END SWITCHING VALVES FOR PSA PRE-PURIFICATION SYSTEMS Since 1990, there has been a new development in the introduction of pre-purification systems using PSA. This system, as explained in Chapter 5, eliminates the use of heat in regeneration, and relies only on the difference in pressure, using artificially modified alumina resembling a low-cost zeolite. Although this has merit as an alternative system to the PTSA system, it also has certain drawbacks. It uses more adsorbent, though lower in price. It also requires switching valves every 10–15 min, bringing the valve problem into the same category as the Revex system. In many cases, moreover, three vertical vessels are used in place of two, resulting in an extra number of valves and consequently additional maintenance problems and additional capital investment up front.

10.4.8 FLOW CONTROL CHECK VALVES The application of these valves is restricted to air separation plants using the Revex system for pre-purification of air plants. The check valves are located immediately downstream of the


Revex system and operate in sequence with the switching valves located immediately upstream of the Revex unit. The mechanical requirements for these valves are similar to those of the main switching valves, but added to them are the very low temperatures involved, namely, 100–102 K. An inexpensive choice for this valve is a bronze spring-loaded check valve. Field reports, however, have indicated consistent failures in the spring mechanism, resulting in unplanned plant stoppages and loss of production. In this operation a better choice is the use of a vertical free travel poppet valve made of stainless steel, which has been a reliable standard since 1950.

10.4.9 CRYOGENIC PROCESS VALVES (GENERAL) The selection of process control valves depends on 1. the type of control required such as flow control, pressure reduction or throttling, isolation, etc.; 2. the type and degree of automation in order to obtain the best combination of valve and actuator; 3. the operating conditions, such as liquid phase, gaseous phase, dual phase, and temperature; and 4. individual preference for facility of inspection and maintenance versus initial valve costs. All cryogenic process valves operating within the insulated casing (cold box) should have an extended bonnet and stem protruding from the casing, so that the hand wheel and the actuator are located outside, and are under ambient conditions of operation. It is also recommended that even automatic valves have a manual over-ride for emergency purposes. In general, the materials of construction for body, bonnet, stem, and plug–disk should be either 316 stainless steel, K-Monel, manganese bronze, or even of aluminum (excluding system and plug). The packing should preferably be of PCTFE (V-ring type), and the gasket of PTFE.

10.4.10 HAND-OPERATED CRYOGENIC VALVES The number of hand-operated valves used for process control is getting lower and lower. They are still useful, however, for specific applications. For example, in changing over to a new set of operating conditions, for isolating individual components or circuits requiring defrosting such as adsorbers, auxiliary vaporizers, etc., and for the complete defrosting of the entire plant. Valves used for this purpose are generally gate valves for liquid phase in large diameters, but ball valves have also been used for liquid phase in somewhat smaller diameters (6 in. or less), or for gaseous phase in larger diameters. Both types of valves have the advantage of giving an unrestricted flow (zero pressure drop) in the fully open position, and complete isolation (zero leakage) when fully closed. For gate valves, the design calls for a 316 stainless steel body and a manganese bronze wedge. The wedge should have a pressure relief vent hole. The packing should be preferably PCTFE (V-ring), and the gasket PTFE (ring type) (see Figure 10.4.1). Gate valves may also be of the dual disk type, with a pressure relief hole drilled in the disk. An advantage to the gate valve is that it opens and closes slowly, preventing effects of fluid hammer and damage to the piping system. These valves, however, cannot be used for throttling applications in process control. Such a use will result in high-velocity fluid flow at the gate seat, and may lead to erosion. Moreover, a partially open gate may lead to vibration, cause damage, and require costly maintenance (Figure 10.4.2).


FIGURE 10.4.1 A single gate valve. (From F.G. Kerry, Inc., 2005. With permission.)

With regard to ball valves, the project engineer should be careful in its selection. Sometimes, in their design the weld collar is too short, resulting in excessive heat conduction in the valve body and a temperature rise in the seating surface of the gasket. The end result is leakage at either the flange assembly or the ball. This has been a recurring problem over the years leading to costly valve removals and replacements. Ball valves are available in reduced, full port, and venturi patterns. These valves are easy to operate, have a high flow-capacity, and a high-pressure tolerance. If globe valves are used for the same service (liquid phase isolation at cryogenic temperatures), the same material specifications apply. Plug valves may also be used as block valves. These valves have a cylindrical or tapered plug with a hole bored through. As with ball valves, a quarter turn will stop the flow. In addition they can have a seating material of fluorocarbon or be completely lined with it. For gaseous phase isolation at cryogenic temperatures, the use of a butterfly valve may be considered, but the practice imposes strict specifications to avoid any leakage (Figure 10.4.3).

10.4.11 PROCESS CONTROL VALVES With regard to valves that actually control the process within preset operating conditions, the selection includes butterfly valves for gaseous phase, and either globe or angle valves for liquid phase. The principle of the globe valve is the perpendicular movement of the disk toward or away from the seat. This causes the annular space between the disk and seat ring to


FIGURE 10.4.2 Gate valve with double gate. (Courtesy of Crane Company, 2006. With permission.)

FIGURE 10.4.3 Cut-out of a ball valve. (Courtesy of Metso Automation, 2006. With permission.) ß 2006 by Taylor & Francis Group, LLC.

Large reservoir volume above seat

Stem guide

Plug

Curving exit

FIGURE 10.4.4 A standard globe valve. (From Kerry, F.G., Oil Gas J., December, 1985. With permission.)

close gradually as the valve is closed, resulting in good throttling capability. This characteristic also has the advantage of diminishing leakage around the seat, and can be adapted more easily for automation in throttling service. The globe valve, however, has a low flow coefficient Cv (higher pressure drop) because fluid flow changes direction at right-angles twice in its passage through the valve body. Angle valves are essentially globe valves having inlet and outlet connections at right-angles. Their design offers a slightly lower resistance to flow than globe valves, especially for throttling operations, because the fluid flow through the valve body makes only one right-angle turn. In general, these angle valves are made of aluminum, which is perfectly acceptable so long as the fabrication and testing are of a very high standard (Figure 10.4.4 and Figure 10.4.5).

Stem guide

Plug

Liner

FIGURE 10.4.5 A standard angle valve. (From Kerry, F.G., Oil Gas J., December, 1985. With permission.)


Globe valves used at cryogenic temperatures should have bodies, bonnets, stems, plugs, and seats made of 316 stainless steel. The disk–plug should preferably have a dual seal design. The packing and gasket should have a PCTFE (Del-F) insert. A more recent design of globe valve eliminates the need to enter the insulation when undergoing maintenance. The body seat that remains in the cold box is of integral stellite, threaded stainless, or threaded stainless with stellite. The plug is stainless steel and the trim has hardened or soft seat options. The entire assembly of stem and trim can be removed for servicing from the outside without disturbing the insulated casing (Figure 10.4.6). The body, bonnet, and disk of butterfly valves operating at cryogenic temperatures should be of 316 stainless steel. The seal ring should have a PCTFE seal, and the packing should be PTFE (V-ring). There should not be any metal–metal contact. If the gaseous phase is pure oxygen, it is highly recommended to use K-Monel or bronze for the body and disk. (Indeed, the latter statement applies for all valves regardless of design, in pure gaseous oxygen service.) They can be used as block valves and for throttling applications, have a high Cv (low-pressure drop), and present a minimum-wear surface, which reduces the operating torque requirement. Unfortunately, their drawback is that the disk and shaft are located in the flow path of the fluid. Some high-performance valves have an eccentric stem that is offset and causes the disk to be parallel to the seat before closing, thus eliminating leakage (Figure 10.4.7).

FIGURE 10.4.6 Control valve that can be removed and maintained externally to the cold box. (Courtesy of KOSO America, Inc. (Hammel-Dahl) 2006. With permission.)


FIGURE 10.4.7 Butterfly valve. (Courtesy of Metso Automation, 2006. With permission.)

10.4.12 PRODUCT FLOW CONTROL VALVES 10.4.12.1 GASEOUS PRODUCTS Generally, the selection involves butterfly valves activated either pneumatically or electronically from the main control panel or computer center. Materials of construction are mainly carbon steel, with the exception of oxygen lines, which require the use of a Monel disk. A PCTFE packing is recommended. As previously mentioned, no metal–metal contact is permitted.

10.4.12.2 LIQUID PRODUCTS These valves are of a comparatively small diameter (1–6 in.) and are either gate or globe, as they involve both flow control and tight shutoff. The material used for the body, bonnet, stem, disk–wedge, plug, seat etc. is 316 stainless steel. The packing is generally V-ring PTFE and the gasket should be glass filled PTFE. An integral seat is recommended.

10.4.13 VALVE CONNECTIONS Within the insulated casing enclosing the cryogenic process equipment, it is imperative to avoid leaks. For this reason, butt-welded valve connections are strongly recommended. Transition couplings are readily available on the market to permit the welding of stainless steel valves to, say, aluminum piping. Stainless steel valves are also available with split bodies so that the body itself does not have to be removed for servicing. The bonnet can be unbolted allowing the stem, disk–plug, and trim to be removed for inspection and maintenance.


FIGURE 10.4.8 Cryogenic wafer sphere butterfly valve. (Courtesy of Metso Automation, 2006. With permission.)

With regard to butterfly valves, if a single flange valve is to be used, and if of a large diameter, a wafer design is recommended. This type of valve is inserted between two pipeline flanges. The bolts pulling the mating flanges carry all the tensile stress, leaving the valve section in compression. The alternative multilug type is heavier as well as cumbersome during construction and for dismantling in maintenance (Figure 10.4.8). Aluminum valves, if offered by suppliers, have the disadvantage of requiring flanged connections. Therefore, aside from the possibility of leaks, maintenance can become a problem as the entire body has to be removed for inspection and service.

10.4.14 INSULATION AND CASING DESIGNS FOR CRYOGENIC VALVES Because all process and isolation valves operate within the confines of the insulated casing, and because they require periodic inspection and maintenance, a special design has to be developed for the casing and the insulated area around the valve body. If the casing is insulated with mineral rockwool the problem is minimal, because it is easy to remove a bolted casing panel, a tunnel through the rockwool, and work around the valve body without any fear of having the insulation collapse or run out. But the use of mineral rockwool as an insulating material has been supplanted almost entirely by the use of expanded perlite, which has better insulating properties, is less costly, and is much easier to handle. Expanded perlite, however, flows very readily and quickly (almost like water), so that the casing must be designed air-tight. Valve maintenance or inspection is unthinkable with the use of expanded perlite unless a special design is developed and applied to the casing around the valve area.


First of all, cryogenic valves are available with long bonnets and stems that extend through a porthole cut out from the casing. Therefore, the bonnet, stem, handwheel, and automatic actuator are external to the casing, and operated under ambient atmospheric conditions. The porthole itself is sealed by means of a flexible neoprene diaphragm, which has a small hole cut out, permitting only the valve bonnet–stem to protrude through. Within the casing, and around the valve body, a welded carbon-steel housing (valve box) is attached by welds or bolts to the inner wall of the main casing. The housing has special cut-outs to allow the ancillary process piping to enter the housing and to be connected to the valve body. Thus the housing protects both the valve body and its immediate connecting process lines from any intrusion by the expanded perlite (refer to Figure 7.6). During periods of inspection and maintenance the housing itself is then filled with mineral rockwool, which can easily be removed and replaced from the outside without disturbing the expanded perlite insulation. The valve boxes are usually made of expanded stainless steel metal with solid plate steel (ss) roof. An entry manway through the cold box casing is provided into the valve box. Valve boxes can house one or several valves. Other valve box designs exist, including a flexible fine wire mesh that prevents the flow of expanded perlite, and is easier to adapt around the valve body. As noted previously in Section 10.4.11, the new design of globe valve that permits its complete removal from the cold box for maintenance is the preferred choice.

10.4.15 LIQUID PURGE VALVES The valves for this service are relatively small in size (up to 2 in.). As a rule, ball valves are used. Unless properly specified and examined, however, they can be a constant source of trouble for maintenance. Depending on the operating temperatures involved in the process, the selection should include a 416 stainless steel body with a 416 stainless steel ball and stem, and a modified PTFE seat and seal. These valves have screwed end-connections and a handle for operation. They are generally hand operated. Unless purchased from a reliable supplier, these valves may prove to be a constant source of trouble from leakage and require constant servicing.

10.4.16 AUTOMATIC CONTROL OF CRYOGENIC VALVES The old hand wheel control valve where the valve body was connected by a steel rod from the outlet of the cold box to a so-called control panel 2 m away was quickly replaced by a pneumatically controlled system in the 1950s. The pneumatic actuators were then in turn replaced by the electronic system that removed the limits of distance between the cold box and the central control panel, and lowered the overall investment cost of the instrumentation. The 4 –20 mA control signal in turn is being replaced by the use of microprocessors embedded in the valve’s control unit. With regard to actuators, the state-of-the-art is changing so rapidly that anything written on the subject may be obsolete within a month. Suffice to state that the selection of an actuator for a cryogenic valve will be no different than that for any other chemical process. In fact, after 1990, merchant plants with a capacity of 2000 t=d and over are being operated automatically by means of a completely computerized program. In 1995, a 1200 t=d oxygen plant had a cryogenic valve with a programed microchip embedded in the valve’s control unit, and which controlled the flow of the liquid nitrogen reflux to the top of the low-pressure (upper) column.


10.4.17 CRYOGENIC LIQUID STORAGE VALVES Valves used in this service are primarily for flow control and isolation. They follow the same material specifications, therefore, as for those used in the main cold box. Automatic plug valves are generally used for flow control in liquid product lines carrying the product to liquid storage tanks. For certain storage vessels with a capacity of 500 t or less, however, where a low heat gain is very important, a double-wall storage tank may be used with a vacuum insulation between the walls. The insulation may also include expanded perlite or superinsulation as well as vacuum. The valves selected to maintain the vacuum within the double wall of the storage tank must have high-quality control for materials and workmanship in order to be absolutely in-leak proof. The degree of vacuum employed for industrial cryogenic storage may be 0.025 mbar or lower. These valves should have a high unrestricted flow to reduce the pumping speed of the vacuum system, and be made of high-quality materials to reduce outgassing to an industrially acceptable minimum. Various valve designs are offered for vacuum service. These include angle valves, ball valves, slide valves, and diaphragm valves. Both the angle valve and the ball valve have a high unrestricted flow, and the use of a suitable O ring can act as a good seal for the valve stems for both the angle and slide valve. The most commonly used valve in vacuum service is the weir diaphragm valve, which is simple in design and less expensive. This valve, however, has a lower unrestricted flow, and the commonly used elastomers for the diaphragm may be more permeable to gases than may be required. The diaphragm may be PTFE, specially reinforced to handle vacuum services. Stainless steel diaphragms are also widely used in these applications. Three factors affect in-leak rate: 1. Permeability of the diaphragm 2. Outgassing, or the removal of water vapor, etc., absorbed on the body and diaphragm walls during their fabrication 3. Leakage across the weir and weir flange face Specifications for a diaphragm valve for vacuum service must be very exacting. The materials used for the diaphragm and the valve body as well as the overall assembly must undergo rigid supervision for quality control and inspection. In one instance involving a liquid nitrogen storage tank with a capacity of 249,375 L at 95% full, the specification called for a vacuum pressure of 0.025 mbar to be maintained. Unfortunately, the diaphragm valve supplied had a problem, and the vacuum pressure increased to 0.27 mbar in 7 days. The vacuum pump was then restarted.

10.4.18 PRESSURE SAFETY RELIEF VALVES: OVERVIEW Safety relief valves are recommended to avoid damage to equipment and piping as a result of an unexpected and dangerous overpressure in the system. Depending on the system’s operating pressure and temperature, the manufactured valve should meet the applicable ASME, API, and ANSI Codes and Standards for material and safety. In cryogenics there are two overall classifications in common use: a spring-operated safety relief valve and a pilot-operated safety relief valve. The former can be designed to operate as a normal relief valve, and then as it reaches a preset point, it allows the full flow to pass and remains open until the process malfunction is corrected. In the pilot-operated relief valve, the fluid applies pressure at the bottom of a free-floating area piston. The pilot line also applies pressure to the top of the piston. When the preset pressure in the pilot is reached, it opens, becomes partially evacuated allowing the free piston to move upward, thus relieving the system pressure.


The usual requirements for any safety relief valve are reliability in operation, to open automatically and quickly when the pressure rises, and to close when the system returns to normal with minimum leakage. Unfortunately, safety relief valves are not without problems. After they function they may not reseat properly resulting in leaks. They may not open at the preset pressure. They may open prematurely and chatter during operation. These problems can be overcome if a predetermined time interval is established between field testing and inspection for each safety relief valve. This practice should include the use of isolation valves to allow the removal of the safety relief valves during maintenance and a better application of valve maintenance. Care should be taken where spring-loaded safety valves are located in a cryogenic system. If the location is too close to a possible two-phase system, the valve may freeze. This has happened in the past. One should be careful in the application of either pressure or safety relief valves on cryogenic distillation columns. The use of bursting disks has given problems because they have blown out at inappropriate times, and are very troublesome when undergoing maintenance in a packed column. Spring-loaded as well as pilot-operated pressure relief valves have also proven quite troublesome over the years, because even a small leak freezes up the valve making it unreliable in operation. A plant shutdown is required to correct the problem. Somehow these valves never quite properly reseat after the first or second lift. A pressure vessel containing a cryogen should include at the minimum a pressure safety valve, a bursting disk, a vent valve, and a sample valve. If the vessel is vacuum insulated, the enclosure should also include a vacuum gauge attached to a valve, a rupture disc, and an evacuation valve (Figure 10.4.9). The sizing of the orifice of the safety valve is very important, and the following equations are generally used for the calculations.

FIGURE 10.4.9 Pressure safety relief valves. (Courtesy of Tyco Valves & Controls, 2006. With permission.)


10.4.18.1 SIZING

FOR

PRESSURE SAFETY RELIEF VALVES (INTERNATIONAL UNITS) SIy

1. Spring-loaded sonic flow A¼

h ipffiffiffiffiffi pffiffiffiffiffiffiffi 1:316W TZ =CKP1 M

(10:4:3)

or pffiffiffiffiffiffiffiffiffiffiffiffi A ¼ {V MTZ }=17:02CKP1

(10:4:4)

pffiffiffiffiffi pffiffiffiffiffiffiffi A ¼ {1:316W TZ }={CKd P1 Kb Kc M }

(10:4:5)

pffiffiffiffiffiffiffiffiffiffiffiffi A ¼ {V MTZ }={17:02CKd P1 Kb Kc }

(10:4:6)

2. Pilot-operated sonic flow

or

3. Pilot-operated subsonic flow (low pressure) pffiffiffiffiffiffiffi pffiffiffiffiffi A ¼ {W TZ }={560Kd P1 F M } pffiffiffiffiffiffiffiffiffiffiffiffi or A ¼ {V MTZ }={12,510Kd P1 F }

(10:4:7) (10:4:8)

where

A ¼ Calculated orifice area V ¼ Required capacity (gas) W ¼ Required capacity, gas or steam G ¼ Specific gravity M ¼ Molecular weight (M ¼ 29 specific gravity) T ¼ Relief temperature (K ¼ C þ 2 73) Z ¼ Compressibility factor (if unknown assume Z ¼ 1) k ¼ Ratio of specific heats (k ¼ Cp=Cv) C ¼ Gas constant (if unknown assume C ¼ 315) Kd ¼ Effective nozzle coefficient for 90% of actual capacity (From National Board of Certified Testing)z K ¼ Kd P ¼ Set pressure P1 ¼ Valve inlet flowing pressure (P1 ¼ P þ allowable overpressure inlet pressure loss þ atmospheric pressure) P2 ¼ Valve outlet flowing pressure Kb ¼ Back pressure factor PA ¼ Valve inlet flowing pressure (PA ¼ P þ allowable overpressure – inlet pressure loss) PB ¼ Valve outlet flowing pressure

y

cm3 Nm3=h kg=h — — — — — — —

barG barA barA — barG barG

Above equations courtesy of Tyco Valves and Controls. Anderson and Greenwoo, Tyco International. May be given by supplier of valve. Normally it is slightly less than 1.0.

z


KC ¼ Combination correction factor for installations with rupture-disk upstream of pilot operated relief valves. (KC ¼ 1.0 with no rupture disk; KC ¼ 0.9 when no combination factor is known.) F ¼ {k=k 1[(P2 =P1 )2=k (P2 =P1 )(kþ1)=k ]}1=2 This valve is normally used for high-pressure systems (2–250 bar). In oxygen service, the inner body should consist of brass or bronze alloy with Monel seat retainer and a stainless steel spring. Teflon or Kel-F should be used to provide good tightness in seat and seals. Example: Calculate the capacity of a spring-loaded safety valve to be used on a liquid nitrogen storage vessel and valve having a diameter of 7 cm (area ¼ 38.49 cm2), the relief temperature ¼ 288 K, molar mass of nitrogen ¼ 28.01 kg=mol, compressibility factor Z ¼ 1, gas constant C ¼ 356, Kd ¼ 0.816 from NBCT for that valve, and inlet flowing pressure ¼ 7.1 bar A. Using Equation 10.4.3 38:49 ¼ {1:316 W (288 1)1=2 }=356 0:816 (28:01)1=2 7:1 38:49 ¼ W 22:34=10,915:8 W ¼ 18,807 kg=h or 5:22 kg=s

10.4.18.2 PILOT-OPERATED SAFETY VALVES (FIGURE 10.4.10) In normal operation, the system applies pressure at the bottom of the free-floating piston and the pilot supply line applies pressure at the top of the piston. When a preset higher pressure is attained, the pilot line opens and depressurizes the dome, therefore permitting the piston to be raised and allows the flow to pass through the valve. As soon as the predetermined blowdown pressure is reached, the piston moves downward, closing the valve. These valves are available for high flows (sonic); in fact, they can operate at higher set pressures than is possible with spring-loaded safety valves. They are also available for low-pressure (subsonic) flows. Pilot-operated safety valves may be preferable to spring-loaded type, especially for lowpressure service, because they can eliminate such deficiencies as frictional resistance and leakage at low operating ranges. In oxygen service, all parts in contact with oxygen should consist of brass or bronze alloy and the carbon-steel body should be cadmium plated to prevent rusting after cleaning. The seats and seals should conform to cryogenic temperatures.

10.4.18.3 PRESSURE

AND

VACUUM RELIEF VALVES

These are designed to vent large low-pressure vessels such as flat-bottom cryogenic tanks, protecting them from physical damage or permanent deformation caused by an increase of internal pressure, or vacuum break encountered during operations (Figure 10.4.12). These valves should follow API Standards 2000, or other equivalent standards. They are spring-loaded on the pressure side and dead-weight loaded on the vacuum side. The seating design keeps the valve tightly seated until the pressure or vacuum inside the vessel approaches the related valve setting. They are also designed to be vented to existing piping, thus saving valuable products. As a rule large flat-bottom cryogenic tanks are equipped with two identical pressure and vacuum relief valves as insurance if one fails to function (Figure 10.4.11).


FIGURE 10.4.10 Pilot-operated pressure relief valves. (Courtesy of Tyco Valves & Controls, 2006. With permission.)


FIGURE 10.4.11 Pressure and vacuum relief valves. (Courtesy of Tyco Valves & Controls, 2006. With permission.)

10.4.18.4 BURSTING DISKS This fixture serves as a support to the pressure relief valve. The bursting disk is usually set at 1.2 times the operating pressure. The safety valve must be reset, and the burst disk should be replaced once the system returns to normal operation. All safety devices should be installed upstream of the vent valve and all devices including the vent valve should be oriented away from any field working personnel and nearby occupied buildings. Similar disks, called rupture disks, are also used to protect double-wall vacuum-insulated vessels.

10.4.18.5 CHECK VALVES Check valves are also used in cryogenic systems to permit fluid flow in one specified direction, and to prevent back flow in the opposite direction. These valves may be of the swing, lift, or tilting type. The choice will depend on the process design, and of the materials used for the valves to be compatible with the temperatures involved in the process. One word of caution, however, on check valves. Somehow, they never receive the serious attention they deserve during routine maintenance, and apt to become less efficient and unreliable in operation due to erosion and normal wear and tear.

10.4.19 MAINTENANCE OF CRYOGENIC VALVES A valve for cryogenic service is a precision-made piece of machinery with close fitting parts, fine tolerances, and with specially selected metallic and nonmetallic materials. It is very important, therefore, that the people servicing cryogenic valves be fully aware of the operating conditions, design, tolerances, and materials of each valve. Leakage through the


seat–disk–wedge interface can be said to be the major cause of valve problems regardless of type of service. Other problems such as failure in body, stem, packing, and bonnet or gasket are minimal, albeit more vivid. Before any replacement of parts or materials is undertaken, it is important to discuss the situation with the valve manufacturer. A poor selection of materials for a valve disk and trim may induce rubbing with worse results than before. With regard to nonmetallic materials, the use of PTFE (Teflon) has been a popular choice for cryogenic valves. Although adequate for service at low temperatures, PTFE may be a poor choice for process valves operating at ambient temperatures above 273 K. Over a period of time, PTFE has a tendency to creep, and leakage may occur during shut off conditions. A better choice may be PCTFE (Kel-F) or a modified PTFE. Reassembling the valve after servicing also requires precision. This is especially true in the case of eccentric butterfly valves where a misalignment of one-tenth of a millimeter or more can be serious. Gate valves may also be a source of trouble if excessive force is used in closing the wedge. Very often the resulting leakage cannot be overcome by a simple repair, and the machining of the injured component may have to be undertaken. Pressure relief valves, so important in cryogenic processes, are often ignored or overlooked by maintenance crews, leading to disastrous results. These valves have to be serviced in strict accordance with the manufacturer’s recommendations for frequency and procedures. Finally, during assembly it is important to keep all components, however small, scrupulously free of any dust and oil. This is especially true in servicing valves for oxygen service. The use of specially recommended solvents is advisable. A simple statement in the technical specifications that a valve, or parts, shall be cleaned for oxygen service may not be enough. The presence of a company supervisor during the cleaning and packing procedures will be more prudent.

10.4.20 VALVE STATIONS: GENERAL These units include isolation valves, and may include control valves, bypass valves, and metering stations; and their purpose is to isolate a line or a section of the system for maintenance, direct manual control, measuring of flows, etc. These stations are generally above ground and always present a potential hazard because most accidents occur at these units. Furthermore, when combustion takes place the effects are dramatic as well as dangerous to personnel and property. Oxygen under pressure when released escapes into the atmosphere with a loud sound, and molten iron oxide projections fly with high velocities in all directions as far away as 30 m (100 ft). To avoid such accidents the following recommendations are in order: 1. Every valve station should therefore be located in either an enclosure or cabin and admittance to enter should be prohibited except with permission of a supervisor. Instead of a closed cabin, an open air enclosure is recommended with easy escape exits for the protection of operators. 2. Unprotected valve stations should be at least 30 m (100 ft) from roads, highways, railway lines, and public areas. 3. Protecting walls should be constructed of standard fire-proof materials (solid concrete, concrete blocks, solid brick, etc.) with a minimum height of 2.5 m (8 ft) and should project 1.7 m (5.5 ft) above the theoretical axis of the line involved. 4. Manual control of isolation valves should be performed behind a protecting wall at least 1.2 m (4 ft) wide and the hand wheel should be no more than 1.5–2.0 m (5–6.5 ft) from the valve axis. Automatic control valves may not need safety barriers since operators are not normally near them. It should be remembered, however, to protect the electric or


pneumatic circuits operating the valves because if damaged, they may prove costly to repair. Control of compensating or purge valves should also be performed behind a protecting wall. 5. Consideration should be given to the installation of safety showers for the use of any person caught by spontaneous combustion with clothing on fire. These safety showers should be approximately 25 m (82 ft) away from the valve station. This is a safe distance from the station but also close enough to reach quickly.

10.4.21 VALVE STATION DESIGN Valve stations are an important part of oxygen pipelines. They also impose an exacting study for their location, function, and design. There should be an isolation valve with a bypass at the head of every transmission line or distribution branch line. These valves should also be tied-in with a pressure indicator controller with a signal to shut off this valve should the downstream pressure of the line fall too far below the normal operating pressure, due to a fire or rupture in the line. If the pressure downstream has to be regulated or reduced, then a pressure control valve should be located at the line. A filter should be placed upstream from the pressure-reducing valve and include isolation valves both upstream and downstream to allow for future maintenance work. If flow meters are used pressure control valve should be located upstream of the meter to maintain a constant inlet pressure at the meter orifice and with a filter ahead of the control valve. Again isolation valves should be located at both upstream and downstream to permit future maintenance. Should it be necessary that oxygen flow be continuous even during maintenance, then a double-valve station should be considered. A complete mirror image station should be incorporated in parallel so that one side may operate while the other side is undergoing maintenance work. Needless to say a safety barrier should separate the two.

10.4.22 DESTRUCTION OF A PRESSURE REDUCTION STATION The following conditions existed just prior to a fire and explosion that took place at a pressure-reducing station in 1984 at a South American steel mill. The isolation valve at the point of use in the melt shop was shut off. This in turn signaled the closure of the pressure control valve at the pressure-reducing station. Isolation valves at the pressure-reducing station both upstream and downstream of the pressure control valve remained open as is normal. As the isolation valve at the oxygen generating site remained open, which is also normal, oxygen pressure started to build up to 30 bar (which is the line pressure), first in the gaseous oxygen storage tanks, then in the distributing line, and finally in the pressurereducing station upstream of the pressure control valve. Downstream the pressure remained at 15 bar. When the upstream pressure reached 30 bar, the dynamic conditions of the flow stabilized, but only for a few minutes. Shortly thereafter the following events happened. The safety valve of the pressure-reducing station immediately downstream of the pressure regulating control valve opened up completely and started to vent the line with a great deal of noise. A huge flame shot up from the concrete enclosed station spewing heat and molten steel fragments at the pipe racks nearby. Two minutes later the main isolation valve at the oxygen plant was closed manually. An examination of the damaged station revealed the following facts: . .

.

No personnel were injured. The bottom half of the off-line filter, both casing and piping, had completely disintegrated and was nonrecognizable. The upper casing of the filter had exploded and ruptured but remained in discernable form.


.

.

.

.

.

.

.

The sintered bronze filter tubes inside the upper casing either fused, melted, or vaporized and were nonrecognizable. The pressure-regulating valve was still in place but its upstream half was destroyed by external mechanical forces. No internal fire or explosion had apparently taken place. Both upstream and downstream isolation valves at the pressure-reducing station were undamaged and still in operating condition. The piping connecting the oxygen filter to the pressure-regulating control valve was partially damaged, but only near the filter side. The piping section between the oxygen filter and the upstream isolation valve of the station was not affected to any serious degree. The piping section between the control valve and the downstream isolation valve showed no visible damage. The piping on the pipe racks near the station showed considerable damage.

10.4.22.1 HYPOTHETICAL CONCLUSIONS When the isolation valves at the melt shop were closed, the pressure-reducing control valve closed, and the dynamic flow conditions should have stabilized at 30 bar upstream and 15 bar downstream of the closed control valve. However, there was always a small leakage through the seat of the control valve, but this slow pressure buildup should have been controlled and contained by a properly set relief valve—if the latter were in good operating condition. The safety valve that is normally designed to handle the full flow of the pipeline should not have opened in normal and properly adjusted conditions: .

.

.

.

.

.

The pressure-reducing control valve did not fully close and there occurred a serious leakage problem. The relief valve did not function well, either from poor maintenance or improper adjustment or from both. The off-line filter was fabricated from nonacceptable materials, mainly carbon steel in both shell and internal piping. When the oxygen pressure started to build up downstream, the relief valve did not or could not contain it, and the safety valve opened up. When the safety valve opened up, the downstream pressure reached close to atmospheric. Therefore, the pressure-reducing control valve opened up in full, and the oxygen velocity in the piping system became close to sonic. In this situation, any particulates (scale, oxides, etc.) may have carried sufficient energy to have caused ignition when striking the carbon steel parts of the oxygen filter unit at a pressure of 30 bar. The fact that the fire or explosion occurred in the filter and not in the pressure-reducing valve resulted in an explosion of a lower force than what may be expected in similar circumstances. The larger metallic mass of the filter casing may have absorbed sufficient heat and energy to lower the effects of a more serious disaster than if the energy release was concentrated in the control valve. Furthermore, the bronze filter tubes may have also absorbed sufficient heat to prevent the total disintegration of the upper shell casing of the filter as well (Figure 10.4.12).

10.4.23 RECOMMENDATIONS APPLICABLE TO PRESSURE-REDUCING STATIONS .

All isolations valves, flow control valves, and relief and safety valves should be properly maintained and inspected at frequent intervals.


Relief and safety valves

Oxygen filter Isolation valve

B

Pressureregulating valve

Flow

D

A B C D

A

C

Complete destruction, fusion, and nonvisibility Badly ruptured, but still in recognizable form Partial destruction from external forces No damage at all

FIGURE 10.4.12 Observations made after fire=explosion. (Courtesy of F.G. Kerry, Inc., 2006. With permission.) .

.

All off-line oxygen filters should have a casing of either bronze or Monel or stainless steel (Type 304) at the very least. The inner discharge piping of the filter should also be of bronze or Monel. If Monel piping is not available, the piping may be of stainless steel, but it should be protected by a bronze or Monel shield located between the stainless steel piping and the inlet flange of the oxygen filter. Relief and safety valves especially should be properly set at the required values and their operation checked at well-defined intervals.

FURTHER READING Bishop, T., et al. 2002. Ease control valve selection. Chem Eng Prog (November): 52–56. Frenck, J.P. 2001. Making the most of valves. Chem Eng (May): 66–73. Kerry, F.G. 1985. Cryogenic control valve selection is exacting. OGJ REPORT, Oil Gas J (December): 96–101. Noel, J.W., and W. Lyons. 2001. Who is in control? Chem Eng Prog (September): 38–41. Sahoo, T. 2004. Pick the right valve. Chem Eng (August): 34–39.



11

Instrumentation and Controls

11.1 OVERVIEW The ever-increasing requirement for higher purities and yields in gas separation systems has made it imperative to optimize processes and to adhere precisely to the calculated operating conditions. This objective has been made more and more achievable with rapid development in the field of microprocessors and its marriage of classical instrumentation with computer technology (see Figure 11.1). Present systems, known as distributed control systems (DCS) facilitate such tasks as process monitoring and operation, process safety in terms of alarms and shutdowns, process stability, and finally, optimum energy utilization. Process information processing is carried out in decentralized self-sufficient function units (process stations) for measuring, sequencing, and control, which are allocated to individual process stages of the gas separation system. Furthermore, communication between process stations and the other self-sufficient operator and monitoring units is effected by means of a data highway (or bus). The decentralization of the functions as defined characterizes the process control technology with process control systems located between the individual equipment and the central process computer. Although microprocessor-based single-loop controllers can be used to obtain maximum distribution of functions, they unfortunately result in higher loadings on the data highway, which may cause a reduction in the processing frequency, and often necessitate additional wiring. Therefore, it is highly recommended that a plant, regardless of capacity, should have at least two process access facilities as well as one process engineer’s unit, all connected by the data highway. The data highway itself should also be duplicated (redundant) to prevent any failure in the control system. A color monitor is recommended of at least a 21 in. size. It should also be made to display an overview of the process, the group, details, and alarms as a minimum basis. A trend curve presentation is also a very useful tool for the operating staff, permitting the study and analysis of current and previous occurrences. In planning a DCS system, always plan an ‘‘installed reserve capacity,’’ and a ‘‘reserve space,’’ for future additions. In regard to printers, they should be of the laser type. Finally, the software or configuration prepared by the process department of the supplier should be designed as an ‘‘open system,’’ so that the end user can see the interactions of control and adjustment functions at any time, and understand them clearly.

11.2 GENERAL REQUIREMENTS The supply should include a DCS for the air separation plant. The supplier should be responsible for specifying the DCS equipment configuration and engineering operator work stations (OWS), screen displays, control strategy configurations, process calculations and optimization, and data acquisition required for the operation of the oxygen plant.


FIGURE 11.1 Typical controls of an air separation plant of the 1930s. (ß Air Liquide, all rights reserved, 2005. With permission.)

The supplier should include all required field instruments, control panels, process analyzers, local control systems outside of the DCS, and safety devices for the safe and efficient operation of the plant. Local controls should be provided for all major equipment and systems throughout the plant, completely wired, piped, and installed to the maximum degree possible at the owner’s fabrication site. Field-mounted local control enclosures should be NEMA 4. A marshaling cabinet should be provided in the control room as part of the DCS scope, and should serve as the termination interface between the DCS and field devices. An uninterruptible power supply (UPS) dedicated to the control system requirements of the oxygen plant should also be provided with the DCS. The air separation unit control system design should be coordinated and compatible with the principles of control system design used project wide. Design standardization requirements should include: 1. 2. 3. 4.

Instrumentation and controls numbering and tagging Numbering of wires, cables, and terminals Power supplies, signals, and units of measurements Types and formats of drawings and other documentation


11.3 CONTROLS AND CONTROL SYSTEM PHILOSOPHY (FIGURE 11.2 THROUGH FIGURE 11.4) All prime movers should be controlled via the DCS including start–stop, interlock, surge control, and all-permissive logic. Local gauge boards should be used for local indications only. No local PLCs should be allowed. The DCS should be located in a climate-controlled building, which may also include the control office as well as human facilities. As a minimum, the oxygen plant DCS should consist of two OWS, with dual monitors; one engineer work station (EWS), with single monitor; one historian–logger station; two event printers, black and white; one historian–log printer, color laser jet; one digital CD storage; three distributive processing units (DPU) for (a) warm end system including main air compressor, (b) cold box section, and (c) product compression; one serial interface; one sequence of event card; one universal control console (UCC) for emergency stop of main air compressor and emergency stop of oxygen compressor, if required; a UPS for a 30 min back up; and a marshaling cabinet. The ability to change production modes and rates of liquid and gaseous products, as well as choose between fixed oxygen production and demand following (ramping), should be provided. The oxygen plant DCS should be provided with programming that is designed to operate the oxygen plant at minimum power consumption for current ambient conditions, gaseous oxygen rate, and liquid product rate. When in the demand-following mode, the DCS should continuously and automatically adjust the gaseous oxygen production rate in response to demand. The DCS should, in addition, control the oxygen plant to produce liquid oxygen under the following three modes of operation: 1. Minimum liquid oxygen rate 2. A preset liquid oxygen rate 3. The maximum liquid oxygen rate available for current ambient and operating conditions

FIGURE 11.2 Contemporary DCS system: dual monitors (operator=process engineer). In background, process control panel of the 1960s, 10 m in length. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)



1 4

3

Distributed control system proposal layout Not to be considered for final system requirement, illustrates system general layout.

1 2

5 2

13

14

11

7 10 8

12

4

Printer historical logs and reports

5

Work desk and printer stand or curved desk console per optional building

6

Redundant plant loop

7

System UPS UPS power distribution

8

Air compressor interface cabinet

9

Booster compressor interface cabinet

9

16

10

Rotating equipment emergency stop interface

11

INFI-90 PCU cabinet

12

Redundant power supply redundant INFI-90 MFC

13

INFI-90 termination cabinet, typical of 2–3

14

Notes: See drawing EC-5729-1002 for optional control building layout

17

Air compressor instrumentation

LAN-90 interface

Printer process alarm and log

7A

15

2 Bailey operator interface station/LAN-90

3

6 7A

Description

Item

Booster compressor instrumentation

18

TE Air purification instrumentation

ASU instrumentation FT

Cooling water instrumentation

PT

ASU instrumentation

Field instrumentation

FIGURE 11.3 Typical DCS system. (Courtesy of Descon Engineering Co. Inc., 2006. With permission.)

Gas chromatograph

15

ASU analyzer panel

16

Analyzer panel analog and digital signal/range interface

17

Nema 4 field instrument junction boxes

18

Field instrumentation

FIGURE 11.4 Typical panel for motor control center. (Courtesy of F.G. Kerry, Inc. With permission.)

In the event of insufficient capacity, gaseous oxygen shall have priority. Process parameters such as reboiler level and column pressure, and alarms such as malfunctions of the main air compressor and the oxygen compressor (if the latter is required) should be monitored with alarms from the secondary remote process area. In addition to the serial interface, the supply should include at least 20 analog points or a minimum of 50 points originating from the oxygen plant DCS. The supplier should be responsible for programming. The interface language should be ‘‘Modbus’’ or an equivalent, to be discussed with the ultimate owner or operator. All machinery and motors should be started with an operator in attendance. The main air compressor should be started from the main control station. Operating rates of the said equipment should be adjustable from the DCS as needed to vary oxygen plant production rate and mode of operation. Individual calibrated flowmeters for oxygen flows from the oxygen plant should be provided by the supplier. All operator control PID devices (i.e., flow control valves) should be controlled using the same type of control window to change the setpoint by placing it in manual, changing the output where applicable.

11.4 MINIMUM INSTRUMENTATION 11.4.1 GENERAL The following minimum requirements for instrumentation, and process and control equipment should apply: 1. All instruments and accessories in contact with oxygen, including pressure transmitters, flow transmitters, thermowells, pressure gauges, sampling devices and accessories for process analyzers, block valves, tubing and fittings, manifold valves, orifice plates, and control valves should be cleaned and inspected prior to use for oxygen service. 2. Safety and relief valves should be ASME-approved type. Materials of construction should be compatible with temperatures and pressures of the process fluids under flowing conditions.


3. Control valves and safety relief valves in the oxygen main stream should be specified with Monel bodies and Monel-wetted materials, or stainless steel bodies with Monel trim. 4. Recycle valves for the oxygen compressor should be all Monel. All gas flow measurements should be compensated for pressure and temperature. 5. Primary flow elements such as orifice plates, flow tubes, and venturies, when used on the main oxygen stream, should be specified with Monel material and certified cleaned for oxygen service. 6. All temperature sensors should be provided with thermowells. Minimum instrumentation and process control required for each major equipment or process should apply as outlined below.

11.4.2 AIR FILTER . .

Differential pressure indication or alarm Automatic filter element cleaning

11.4.3 AIR COMPRESSOR . . . . . . . . . .

Inlet guide vanes flow control Anti-surge control and instrumentation Discharge flow measurement and control Discharge pressure measurement and control Discharge pressure safety relief valve Air compressor interstage temperature and pressure indicators Vibration sensors, bearing temperature sensors, and motor winding temperature sensors DCS control All necessary pre-alarms and shutdowns including interlock logic All necessary flow gauges and temperature indicators for intercoolers

11.4.4 DIRECT CONTACT AFTERCOOLER (IF APPLICABLE) . . . . . . . .

Cooling water flow controls Cooling water level controls Chilled water flow controls Water chiller blowdown level controls Make-up water level controls Chilled water temperature controls Injection water head flow indicator controllers (FIC) Inlet and outlet air pressure and temperature indicators

11.4.5 FRONT-END PURIFICATION . . . . . . .

. .

Inlet and outlet air pressure and temperature indicators Outlet air CO2 analyzer Inlet chilled air low pressure alarm (PALL) Reactivation heater controls Recorded temperature profile of adsorbent bed Inlet air hydrocarbon analyzer (in-line gas chromatograph) Differential pressure indicator and alarm for pre-filter (if used), adsorbent beds, afterfilter (if used), and total unit Waste nitrogen reactivation heat exchanger flow, flow pressure, and temperature controls Molsieve automatic sequencing and regeneration controls, logic, and interlocks


11.4.6 AIR SEPARATION UNIT .

. . . . . . .

All instrumentation and controls required for the air separation equipment and cold box purging Cold box purge gas pressure indicator and control Required instrumentation and controls for gas expansion turbines Flow control for all products Analyzers on all products and waste nitrogen Flow, pressure, and temperature indicators on all products and waste nitrogen Steam or other heat deriming controls Liquid drain vaporizer controls

11.4.7 OXYGEN PRODUCT COMPRESSOR . . . . . . . . . . . . .

Oxygen excess flow blow-off controls Anti-surge control and instrumentation Discharge flow measurement and control Discharge pressure measurement and control Discharge pressure relief valve Compressor interstage temperature and pressure indicators Vibration sensors, bearing temperature sensors, and motor winding temperature sensors All necessary pre-alarms and shutdowns including interlock logic DCS control All necessary flow and temperature controls for intercoolers Start-up and shutdown nitrogen control system Fire detection with lube oil shutdown and fire suppression interlocks Inlet butterfly valve for flow control

11.4.8 NITROGEN PRODUCT COMPRESSOR (Same instrumentation as in Section 11.4.3 for the air compressor.)

11.4.9 LIQUID OXYGEN STORAGE TANK . . . . . . .

Level indicator Pressure indicator Temperature indicator Heat vaporizer controls and instrumentation Double-wall vacuum indication (if needed) Truck fill pressure gauge and remote level indicator Pressure or vacuum relief valve in duplicate

11.4.10 LIQUID NITROGEN

OR

LIQUID ARGON STORAGE TANK

(Same instrumentation as in Section 11.3.9.)

11.4.11 COOLING WATER SYSTEM . . . .

Pressure indicator Temperature indicator Flow controls as required Temperature controls as required


11.4.12 LUBE OIL SYSTEM . . . .

Oil supply pressure indicator and control Oil supply temperature indicator and control Filter differential pressure indicator and alarm All necessary pre-alarms and shutdowns, including interlock logic

11.4.13 ALARMS, SHUTDOWNS, AND INTERLOCKS . . .

.

Complete system to ensure safe operation Sequential start of spare machinery items excluding oxygen compression Limit switches for all valves deemed important to operation including manual valves with positions indicated in control system, and utilized in sequence circuits Separate annunciator for conditions out of range and item emergency stop

All permissive and redundant controls deemed necessary for safe operation

11.4.14 ANALYZERS The following analyzers shall be supplied after the owner’s or operator’s approval for type and manufacture:

Analyzed Gas Trace CO2 in purified air Trace O2 in nitrogen product Percent of O2 in waste nitrogen Percent of O2 in oxygen product Percent of O2 in liquid N2 reflux Hydrocarbons analysis in LOX

Preference and Range Infrared (0–10; 10–100; 100–1000 vppm). Electrochemical (0–10; 10–100 vppm) Electrochemical micro fuel cell (0–10%) Paramagnetic (90–100%) Electrochemical (90–100%) In line gas chromatograph with the following points to be analyzed at least every 8 h: rich liquid at lower column sump, main condenser–boiler, upstream of LOX guard adsorber, downstream of LOX guard adsorber

Source: From Kerry, F.G., Inc.

The chromatograph should include an alarm for any abnormal high content of the following hydrocarbons: methane, ethane, ethylene, acetylene, propane, propylene, butane, butadiene plus acetone and carbon dioxide. Trace moisture in O2 product

Portable type (if required)

All necessary field-mounted instrumentation, i.e., orifice flanges, orifice plates, meter runs, pressure gauges, temperature indicators, process switches, temperature, flow, level, and pressure transmitters should be supplied (see Table 11.1 through Table 11.3).

11.5 POSSIBLE SPECIFIC REQUIREMENTS OF OWNER OR OPERATOR 11.5.1 SCOPE All instrumentation should meet the specific requirements of this DCS, the standard instrument specification for each type of instrument, and individual instrument data sheets that should be issued as part of the project.


TABLE 11.1 Main Air Compressor—Minimum Instrumentation Function Suction pressure First casing (after filter) Final discharge pressure=temperature Suction pressure–first casing Temp at each casing outlet Temp after each cooler Comp discharge flow Anti–surge control signal to open vent valve Vent valve open Press diff inlet filter each stage Inlet guide vane closed Bearings and lube oil system Filter differential pressure Press after filter and cooler Temp at cooler outlet Temp of each journal bearing Temp of each thrust bearing Main tank level Oil flow Pump discharge press. Main tank temp control Shaft position and vibration Axial position each shaft Radial vibration of shaft (X=Y) at each bearing, including motor but not required at bull gear Miscellaneous First out annunciation system Motor stator temp (6 Pts) Motor bearing Speed (variable speed machines) Compressor fault Water level traps Emergency stops Compressor on Motor moisture detectora Motor air I=Oa Key phasor, each pinion Water discharge each cooler Water manifold flow

DCS Indicator

DCS Alarm

DCS Trip

Local

Type Instruments

— x x x x x x

— — — Hi — — x

— — — Hi — — —

x — — — — — —

PI PT=RTD PI RTD RTD PT Controller

x x x

x x x (perm)

— — —

— — —

Switch Switches Switch (closed perm)

— x x x x — — — x

Hi Lo Hi Hi Hi Lo — — —

— Lo Hi Hi Hi — — — —

x x x — — x x x —

Switch PT=PI RTD and Lo permissive RTD (2 single or 1 dual) RTD (2 single or 1 dual) Switch Sight glass on drain PI RTD control in DCS

x x

Hi Hi

Hi Hi

— —

4–20 Thrust bearing 4–20

— x x x x — x — x x — — —

x Hi Hi — x Hi — — Hi Hi — — Lo

x — — Hi — — — — — — — — —

— — — — — x x x — — — x —

DCS RTD RTD — — Switches Switch Light Switch RTD Probe only TI Switch

Note: Assure control in DCS for large compressors (above 5000 HP). Stage taps with valve shall be provided in and out of each stage for potential pressure measurement (use a PT). a Refers to TEWAC motors only. Source: Courtesy of F.G. Kerry, Inc., 2006. With permission.


TABLE 11.2 Air Compressor Tests Shop Inspection and Tests Shop inspection Hydrostatic Helium leak test Mechanical run Performance test (air) Compressor with drivera Compressor less driver Use shop lube and seal system Use job lube and seal systema Use job vibration probes, etc. Use job vib and axial displ. probes, oscillators—detectors bearing RTDsa Pressure compressor to full Operating pressurea Impeller overspeed Ringing of milled impellerb

Read

Witness

Observed

Acknowledge by Vendor (Y=N)

x x — x x x — — x — — x

— — — x x x — — x — — x

x x — — — — — — — — — —

— — — — — — — — — — — —

x x x

x x x

— — —

— — —

Note: An operational test of the lube oil system including the lube oil flushing per API 614 is mandatory, and must be witnessed. If a steel base is included for the motor, the inspector must witness the rough alignment. a This is not always possible. If so, this matter must be discussed, prior to placing an order. b Use a hammer or other method to excite the blades=hub, then record the resonance frequencies. Source: Courtesy of F.G. Kerry, Inc., 2006. With permission.

TABLE 11.3 Product Oxygen Compressor—Minimum Instrumentation Function Suction pressure-first casing (after filter) Final discharge pressure Suction filter differential pressure Suction temp-first casing Temp. of main gas stream at each casing outlet Shaft position and vibration Axial position each shaft Radial vibration of shaft at each bearing Miscellaneous Motor stator temp. (6 Pts) Motor bearing Motor cooling temp.


Indicator

Recorder

Acknowledge (Y=N)

— — — — Hi

— — — — —

— — — — —

Hi Hi

Hi Hi

— X

— —

Hi Hi Hi

— — —

— — —

— — —

Alarm

Trip

X X X X X

— — Hi Lo Hi

X X X — X

TABLE 11.3 (continued) Product Oxygen Compressor—Minimum Instrumentation Shop Inspection and Tests Shop inspection Hydrostatic Mechanical run Performance test (air) Compressor with driver Compressor less driver Use shop lube and seal system Use job lube and seal system Use shop vibration probes, etc. Use job, vib. and axial disp. probes, oscillato–detectors and monitors Press comp. to full operating pressure Porosity test Impeller overspeed

Read

Witness

Observed

Acknowledge (Y=N)

X X X X X — — X — X

— — X X X — — X — X

X X — — — — — — — —

— — — — — — — — —

X X X

X X X

— — —

— — —


11.5.2 CODES AND STANDARDS The required codes and standards to be followed should be established and agreed upon between the supplier and owner or operator prior to the signature of any contractual agreement.

11.5.3 OPERATIONAL PHILOSOPHY The instruments and control systems should be designed to achieve safe and reliable air separation operation and minimize the operator’s manual regulatory duties. To accomplish this, monitoring and operational capability of all prime movers and support equipment should be centralized at the main control room using a DCS. All pumps should have a hand-off-autolocally mounted switche to allow local control, even though normal mode of operation should be from the main control room via DCS with automatic start or stop. Auto-status from the field should be hardwired to the DCS. Each tank should have a differential-pressure-type level transmitter for level control or indication and a level switch hardwired to the pump controls for pump shutdown.

11.5.4 DISTRIBUTED CONTROL SYSTEM The process control system architecture and features should include discrete and analog process signals to be wired to the I=O termination unit to be mounted on the supplied cabinets, which should be located in owner or operator’s planned or existing central control room. All process and equipment alarms, shutdowns, and permissive logic should be scanned in a sequential order. Discrete (on–off) and analog control actions (valve adjustments) should then be transmitted through slave output modules to the appropriate equipment or control devices. Two operator interface units (OIU) should be provided consisting of SUN-based stations with dual color CRT’s, keyboard, and trackball. From the OIU, the plant operator should be able to start and stop the equipment, change set points of PID controllers, and monitor all system monitors. The DCS system should include overview displays, control group displays, individual point displays, alarm management, and fully configured operational interactive graphics. Digital


status should be controlled with these functions. PID controllers should be adjusted and manipulated and all set points, values, and field status indications should be monitored. The interactive graphics display features should permit the operator to control the process and to acknowledge alarms from a screen panel that should visually depict a portion of the process plant. All operator functions such as start–stop, change setpoint, put controller on automatic, etc. should be accessible via trackball. Each OIU should be equipped with two color monitors, each monitor equipped with its own software so that the operator may scan four process-areas per monitor simultaneously. The historian station should be equipped with ‘‘on demand’’ trending capabilities that should include averaging, standard deviation, and totalizing of variables over operator entered time periods. One of the OIUs may be used as an engineering work-station to have complete access to the program for the purpose of viewing, editing, and generating printouts of the programs. Each of two OIUs to be provided should include their own keyboard and printer for alarm logging and report generation. A color laser jet printer should be dedicated to print process trends and history grams. To ensure reliability, a second, redundant data highway should be separately routed to the devices on the data highway. Each OIU should also have a redundant OIU maintained in a hot-standby condition for critical plant control. An analyzer panel with provisions for zero and span gas calibration should be provided. The various required analyses are listed in Section 11.4.14. A panel for the liquid oxygen storage and a backup panel that includes hard-wired relay logic for the liquid oxygen backup system should be provided.

11.5.5 FIELD INSTRUMENTS Field instruments should be suitable for outdoor installation under the climatic and seismic conditions and area classification described in the site information section of the specification. General design guidelines for selection and installation are as follows: 1. All local indication devices should be clearly visible from grade or platforms. 2. All instrumentation should be capable of being isolated from process, and pressure relieved and safely removed for repair or maintenance while under operating conditions. 3. All instrumentation should be easily removable without interference from piping, conduit, and other equipment. 4. All process tubing and fittings should be manufactured of 316 s=s. Pneumatic tubing should be PVC-coated copper. Acceptable manufacturers are Swagelock and C.P. Parker. Pneumatic fittings should be made of brass. All fittings (process and pneumatic) should be compression type. 5. All instruments should have analog transmitters with local indicators in engineering units. Use of switches other than limit switches may not be acceptable. This also includes skid-mounted equipment. This may involve a prior discussion between the contractor or supplier and the owner or operator’s engineering department. 11.5.5.1 1.

Level Instruments

Level gauges (general service) Level gauges should be of flat glass type, armored completely with gauge cock valves unless approved by the instrument lead engineer. Reflex type should be used unless process conditions dictate otherwise. Maximum center-to-center distance between vessel nozzles where a single level glass can be used is 1.83 m (72 in.)


2. Level switches (general service) Level switches should be displacer, flow, or ultrasonic type. Switches should be selected based on maximum pressure and temperature rating of the vessel or tank on which they will be installed. Capacitance level switches may be used where application dictates subject to owner or operator’s approval. Any exposed steel pipe and flanges should be given a protective paint covering as per the manufacturer’s standard. 11.5.5.2

Temperature Instruments

1. Thermowells Material should be 316 s=s unless process conditions dictate otherwise. 2. Temperature switches All temperature switches should be filled systems type, and automatically reset when the process temperature falls below the set point in the case of high temperature, or rises above in the case of low temperature, unless otherwise specified. 3. Bimetal thermometers Thermometers should be selected to read approximately 50% of scale at normal temperature. 4. Resistance temperature detectors Resistance temperature detectors (RTD) should be provided as a complete assembly with heat elements and thermowells. Where thermowells cannot be used, RTDs for bearing service should be constructed so that the assembly holds positive pressure. RTDs should be 100 V platinum, three wired, and be provided on all temperature monitoring and control applications. 11.5.5.3

Flow Instruments

All orifice plates, venturis, and meter runs should be constructed and installed in accordance with the latest AGA Measurement Committee recommendation. All primary calculations should be in accordance with AGA Report No. 3 or ‘‘Flow Measurement Engineering Handbook’’ (Miller). 1. Orifice plates and flanges Meter runs are to be used where application demands above-average calibration or precise determination of inaccuracy of meter. 2. Ventury tubes, low pressure loss tubes, and flow nozzles These devices should be used in services where pumping-power energy-loss justifies their cost, or required accuracy dictates their use. This equipment, as identified by owner or operator’s engineering, should be flow calibrated to determine the discharge coefficient. 3. Annubars Should be insertion–removable type, mounted on a fullport ball valve specified to design conditions and complete with packing gland to allow removal under pressure. Wetted parts should be 316 s=s. 11.5.5.4

Valves

1. Control valves and operators Pneumatic operators should be constructed so that a positive means is provided for moving the valve to, and holding it at, a safe position in the event of air or electrical failure.


Design: The type of trim, body style, materials of construction, etc., should be determined by the application and should be suitable for the design pressure and temperature conditions. End connection: All valves 2 in. and smaller should be of screwed, flanged, or socket-weld design and all valves 3 in. and larger should be of flanged design as defined by the applicable piping specification. Sizing criteria: Control valves should be sized to pass 120% of the maximum flow or normal operating flow at the full-open position at the calculated pressure drop. Mounting: All valves should be suitable for mounting in any position. Positioners: Pneumatic positioners should be provided on all control valves. They should be equipped with bypass switches and three pressure gauges to indicate supply pressure, input signal, and output signal and should be so labeled. Noise: Control valves and accessories should meet OSHA requirements for sound levels and attenuation to a level of 85 dBa at three feet (one meter). Piping arrangements: Control valves should be furnished with block and bypass valves unless specified otherwise. The former does not apply for control valves within the cold box, and in certain warm valves. Air supply: All valves to be furnished with filter regulators with air supply pressure gauges on inlet and outlet. Air tubing should be PVC-coated copper. Trim: Where severe operating conditions exist, chrome-moly or stainless steel bodies with Stellited steel or hardened trim should be used. For high-pressure oxygen service, however, a Monel trim is mandatory. Electropneumatic transducers (I=P) should be used where applicable, be mounted on the valve, and meet the area classification. Solenoid valves should use 24 VDC power. Operator type: Pneumatic, spring, and diaphragm-piston operators may be used where pressure drop or maximum shut-off pressure exceeds the maximum diaphragm actuator size for the control valve selected. Minimum body size: 1 in. (25.4 mm) for steam and boiler feedwater services, 3=4 in. (19 mm) for water and air services. All automatic valves used in ‘‘on–off’’ service should be furnished with limit switches for both the fully closed and open positions. Limit switches should be proximity type. 2. Electric motor operated valves Refer to owner or operator’s, if any, or standard specification. 3. Relief valves (see Chapter 8 on Section 10.4.18). General sufficient pressure or vacuum relieving devices should be provided to meet the applicable API, ASME, ANSI, or other codes. Metal-to-metal seated spring-operated relief valves may be preferred, but pilot-operated and soft-seated units may be used advantageously for low pressure and other special applications. Design: As a minimum, all safety valves should be designed and manufactured in accordance with the applicable requirements of either Section I or Section VIII of the ASME Boiler and Pressure Vessel code, and to bear the ASME code stamp. ANSI codes should also be followed for certain applications. Sizing: The required orifice area and valve relieving capacity should be calculated in accordance with the applicable ASME and ANSI codes. 11.5.5.5

Transmitters

1. Pressure and differential pressure transmitters Type: Electronic unless the application makes this type unacceptable. Pneumatic transmitters should be equipped with air input and output gauges and filter. For pressure transmitters ‘‘transducers’’ are not acceptable. ß 2006 by Taylor & Francis Group, LLC.

2. Level: Differential pressure transmitters should be used in level applications. 3. Temperature transmitters 11.5.5.6

Vibration Instruments

Manufacturer should be Bently Nevada or equal. Note: As an alternate, a vendor may consider the use of Bently Nevada 3300 vibration probes and Bently Nevada 1800 dynamic transmitters, with Bentlly Nevada coaxial. The latter combination may prove more reliable as well as lower in cost. Bently Nevada 990 transmitters are not acceptable. 11.5.5.7

Local Controllers

If any are required, they should be selected with the owner or operator’s prior approval. 11.5.5.8

Pressure Instruments

1. Pressure gauges Gauges should be selected to read approximately 50% of scale at normal operating pressure. 2. Pressure switches All pressure switches should reset automatically when the process pressure falls below (in the case of high pressure), or rises above (in the case of low pressure) the set point unless otherwise specified. 11.5.5.9

General

1. DCS processors should be sized not to exceed 70% of the memory utilization during normal operating conditions. 2. Ten percent spare I=O, or one card for each I=O type used, whichever is greater should be supplied with the system. 3. The DCS system should be furnished fully configured and shop tested to provide plant monitoring, including discrete and sequential operating control. 4. Equipment and safety interlock functions should be incorporated using either hardwired or dedicated digital controls independent of the DCS. Status of the equipment and safety interlock systems should be monitored by the DCS. 5. Upon automatic switch over, the UPS should alarm in the DCS, and an automatic countdown timer should appear on the screen. 6. All electrical power consumed by the plant should be monitored and totalized in the DCS. Also, individual power and power factor meters for prime movers should be totalized in the DCS as well. 7. The serial interface should be defined as a minimum of 20 analog points or at least 50 digital points originated from the project plant’s DCS. The contractor or supplier should be responsible for programming the points ensuring the proper function of the interface. The interface language to be used should be discussed with the owner or operator. 8. Instruments should be tagged with stainless steel tags and wires. All wiring should be properly tagged at both ends using ‘‘slip-on’’ type markers. 9. The owner or operator should participate in the tagging philosophy discussions pertaining to the DCS. The supplier should submit a proposed procedure before implementing the program.



Advantages and Design Criteria

Analyzer tubing rack

Fully integrated field I/O building Single node on DCS communication highway Low installation costs Reduced check-out and start-up Reduced instrument cable tray and instrumentation cable costs Reduced analyzer tubing and tary Faster analyzer response time closer to process Modular unit

Integrated equipment

Marshalling junction box field I/O terminations

Event and historical–log printer Westation historian/logs Westation engineering/ operator Remote UPS power bypass

Design standards Nema 3R building and 4x interface junction box 250 PSF floor loads and 125 MPH wall/wind loads HVAC and fluorescent lighting UPS and power distribution panels Fire detection, protection, and emergency lighting Computer floor

DCS I/O building

Process analyzer panel and gas chromatograph

Termination area Motor control center (MCC)

Distributed control system Vibration instrumentation MCC interface Process analyzers Process gas chromatograph Hardwired safety control interface UPS

Westnet data highway under floor

Air conditioning and heating unit

Ethernet information highway under floor Westinghouse WDPF II distributed processing unit (DPU) Instrument cabling under floor Compressor vibration monitors and safety shutdown interface cabinets Computer floor

Uninterruptable power supply (UPS)

DCS control room area Battery rack Steel base

FIGURE 11.5 Proposed package control assembly.

UPS and MCC area

45⬚

12⬚

120 V AC utility power distribution LP-a

UPS 480 volt 3 phase

MCC and UPS room

(Optional)

Front

120 v ac UPS power distribution LP-D Lighting transformer

IFC 1

5˚−0˚

Control room area

Battery UPS/batt. disconect

6⬚

Optional building package including: DCS Operator consoles Printers Analyzer and analyzer panel Compressor interface cabinets MCCs UPS Batteries Utility power distribution UPS power distribution All required interconnecting cables Fully wired and tested

FIGURE 11.6 Proposed pre-packaged control room.

Gas chronato

DCS cabinet 2

IFC 2

DCS cabinet 3 DCS cabinet 4 DCS cabinet 5

Marshaling junction box

DCS cabinet 1 3 bay analyzer fanel

Field I/O termination area

6⬚

30⬚

42⬚

Remote bypass switch

Bypass trans. AC service disconnect Transformer

42⬚

30⬚

10⬚−0⬚

Front

480 V AC panel board PPC

5⬚−0⬚

Storage

42⬚

HVAC unit § 1 5 ton unit 480 volt 3 phase

5⬚

HVAC unit § 2 5 ton unit 480 volt 3 phase


40⬚−5⬚ 18⬚−0⬚

UPS 120 V AC inverter output service disc.

14"`

10. The owner or operator should review and decide on the DCS graphics proposed by the supplier. All graphics should be easy to use and free of clutter. An overall plant graphic should be provided to access all main process area windows of the plant. All graphs should be submitted for owner or operator’s prior approval during their development or on a monthly basis. 11. All operator-controlled on–off devices, i.e., pumps, etc., should be controlled using the same type of control window to start, stop, or bypass and reset where applicable. 12. The DCS to be so programmed for three levels of access: (a) operator level, (b) supervisor level, and (c) engineer level. 13. No shutdown alarm setpoint should be accessible at the operator level. 14. All DCS stations should be packaged in control room quality stations for ergonomic fit to the operator’s needs. 15. All pneumatic devices within the plant should run on instrument air at 60 psig (482.6 kPaG). As a start-up, source discharge air from the existing instrument air may be used. 16. All local instruments should be wired to a local junction box. Each type of instrument should be wired to separate terminal blocks. 17. All wiring should be in a rigid conduit to withstand a maximum liquid pressure of 2 ft (60 cm) of water in depth. 18. All tubing should be seamless stainless module; fittings should be preferably Swagelock. 19. All switches should be UE 100.

11.5.6 INTERCONNECTIONS To simplify field connections at the site, the following procedures are suggested. 11.5.6.1

Pre-Packaging

All the necessary instrumentation and controls can be connected and pre-packaged as a single transportable module as shown in Figure 11.5 and Figure 11.6. 11.5.6.2

Large Transformers

These transformers may be connected in a separate and transportable cabin and mounted at the site on concrete pilings.


12

Safety

To appreciate the importance of safety one must fully understand the consequences of mixing fuels with either air or pure oxygen. Explosive gases are mixtures of two reactants—oxygen and a fuel gas—within the limits of inflammability; but some gases such as ozone, acetylene, and hydrazine are by themselves explosive. The propagation of a combustion wave through an explosive gas occurs when initiated by a source of ignition. When a flow of oxygen and hydrogen or acetylene is properly adjusted, the flow of explosive gas mixture is maintained forming a steady, brilliant luminous cone. When the flow of gas mixture is throttled to the flash back point, however, the combustion wave enters the gas mixing chamber and develops into a detonation wave evidenced by an explosive sound. This event does not occur when the mixture is sufficiently diluted with either oxygen or fuel gas. Operators of oxyflame cutting, welding, or flame-hardening often experience such flashbacks. In 1951, Bernard Lewis of Cambridge and Gunther von Elbe of Berlin published Combustion, Flames and Explosions of Gases, which covered the subject in detail. Their work was carried out at the Physical Sciences Department of the US Bureau of Mines in Pittsburgh. They concluded that one should learn the distinction between a combustion wave and a detonation wave. A combustion wave propagates by heat transfer and diffusion, but a detonation wave is a shock wave sustained by the energy of compression of the explosive mixture in the wave. The former wave has a velocity slower than that of sound, whereas that of the latter is supersonic. The limits of detonability of an explosive mixture are found by diluting the explosive mixture with an excess of either constituent or with an inert gas (nitrogen). Increasing the dilution of the detonating mixture, a composition is reached at which a stable detonating mixture is no longer possible. Such a mixture, nevertheless, still propagates a combustion wave. The length that a combustion wave travels before a detonation wave is established is influenced by many factors as originally noted by LeChatelier in 1900. Table 12.1 gives wave velocities of various mixtures at room temperature (288 K), at 101.325 kPa, and at a speed of sound at 344 m=s. It can be noticed from Table 12.1 that acetylene and hydrogen are high on the detonation list, but methane especially when diluted with atmospheric air is at the bottom. The latter fact should encourage proponents of maritime transportation of LNG (Table 12.2 through Table 12.4).

12.1 SAFETY OVERVIEW In the 1950s when the second generation air separation plants were built with a capacity of 100–500 t=d, a series of accidents occurred ranging from small fires to large and even fatal explosions. This precipitated the formation of safety committees, safety seminars, and safety research projects. Later on, the American Society for Testing and Materials (ASTM) inaugurated a special committee G4 to carry out special studies to research the effects of oxygen under high pressure on various materials.


TABLE 12.1 Detonation Wave Velocities of Various Gases at 288 K and 101.325 kPa (Velocity of Sound Is 244 m=s) 2H2 þ O2 C2H2 þ 1.5O2 iC4H10 þ 4O2 C3H8 þ 3O2 C2H2 þ 1.5O2 þ N2 C5H12 þ 8O2 C2H6 þ 3.5O2 C2H5OH þ 3O2 C3H8 þ 6O2 iC4H6 þ 8O2 C6H6 þ 7.5O2 CH4 þ 2O2 CH4 þ 1.5O2 þ 2.5N2

2821 m=s 2716 m=s 2613 m=s 2600 m=s 2414 m=s 2371 m=s 2363 m=s 2356 m=s 2280 m=s 2270 m=s 2206 m=s 2146 m=s 1880 m=s


As the third generation of plants is built with capacities up to 3000 t and even 4000 t of oxygen per day, there is still no readymade formula for 100% safety. The safety of an air separation plant and its ancillary equipment is still designed on an empirical basis, often with no finite quantitative rules for guidance. It is important, therefore, that the basic problems and requirements for safety should be reviewed in detail. Safety is a critically important factor in the specification, design, operation, and evaluation of a process cycle for a complete oxygen plant. This chapter reviews the general parameters of safety, as well as those recommended by safety committees, such as ASTM International G4, the Compressed Gas Association, and other international organizations, which are followed by the gas industry today. It is hoped that this chapter will help existing plant owners and operators to understand and appreciate safe operating procedures. It will also help prospective purchasers to consider safety along with investment and operating costs in their economic evaluation of various processes.

TABLE 12.2 Limits of Detonability of Some Mixtures Mixture H2–O2 H2–air CO–O2, moist CO–O2, well dried (CO þ H2)–O2 (CO þ H2)–air NH3–O2 C2H2–O2 C2H2–air C3H8–O2

Lower Limit % Fuel

Upper Limit % Fuel

15 18.3 38 — 17.2 19 25.4 3.5 4.2 3.2

90 59 90 83 91 59 75 92 50 37



TABLE 12.3 Limits of Inflammability of Some Gases and Vapors with Air (Limits are at 288 K and 9 101.325 kPa) Compound Methane, CH4 Ethane, C2H6 Propane, C3H8 Butane, C4H10 Isobutane, iC4H10 Pentane, C5H12 Hexane, C6H14 Ethylene, C2H4 Propylene, C3H6 Acetylene, C2H2 Benzene, C6H6 Methyl alcohol, CH4O Ethyl alcohol, C2H6O Isopropyl alcohol, C3H8 Acetaldehyde, C2H4 Acetone, C3H6

Lower Limit (%)

Upper Limit (%)

5.0 3.0 2.12 1.86 1.80 1.40 1.18 2.75 2.00 2.50 1.40 6.72 3.28 2.02 3.97 2.55

15 12.50 9.35 8.41 8.44 7.80 7.40 28.60 11.10 80.00 7.10 36.50 18.95 11.80 57.00 12.80


Fortunately, incidents such as fires and explosions are not too common in the oxygen industry. This is primarily due to the diligence and conscientiousness of the industry itself. But the main problem, namely atmospheric pollution is still with us and is getting worse every day. Moreover, the influx of recently graduated engineers and new operators lacking the necessary hands-on background does not help the situation. Though present designers of oxygen plants have already learned a great deal about handling contaminants entering the plants, and are constantly researching new materials and safer designs, nevertheless, they still lack a quantitative solutions or an overall scientific formula to resolve problems of safety. Third generation oxygen plants are built with a production capacity of 2000–4000 t=d. Consequently, it is mandatory to increase efforts in designing and operating these megaplants with maximum safety to avoid potential disasters. TABLE 12.4 Limits of Inflammability of Some Gases with Oxygen Compound Hydrogen, H2 Carbon monoxide, CO Methane, CH4 Ethane, C2H6 Ethylene, C2H4 Propylene, C3H6 Cyclopropane, C3H6 Ammonia, NH3

Lower Limit (%)

Upper Limit (%)

4.65 15.50 5.40 4.10 2.90 2.10 2.45 13.5

93.9 95.0 59.2 50.5 79.9 52.8 63.1 79.0

Source: From Lewis, B. and von Elbe, G. in Combustion, Flame, and Explosion of Gases, Academic Press, New York, 1961. With permission. ß 2006 by Taylor & Francis Group, LLC.

12.2 CHEMISTRY OF IGNITION, COMBUSTION, AND EXPLOSION One of the main difficulties in trying to deduce what happened, or find reasons for, any fire or explosion involving an oxygen plant, is that in the presence of an excess high-purity oxygen, the combustion reaction is generally complete without any telltale residue or evidence of the combustible substance. Furthermore, when we try to simulate possible reactions in a chemical laboratory, we get involved with microchemistry whose results can only be accepted with certain reservations. It is a known fact that in any combustion reaction three elements are necessary: a combustible, oxygen, and a source of ignition energy (Figure 12.1). In an oxygen plant, oxygen is already available mostly in its pure form in very large quantities, especially in the main condenser–vaporizer. As for the combustible, its presence and concentration inside an oxygen-producing unit has been the main problem since the day that Carl von Linde built the first industrial unit.

12.2.1 SOURCE OF COMBUSTIBLES In the early days of the oxygen-producing industry, plants were always placed close to the acetylene generation units for operating convenience, and the presence of acetylene was well known as the prime cause of explosions. Nevertheless, in the 1950s there occurred a series of explosions in some comparatively large plants (for that period), with a capacity of 100 t=d of oxygen with no acetylene generation near the site. The immediate reason could have been that acetylene is always present in a polluted industrial area, but other hydrocarbons became suspect also.1 Subsequent tests did prove that not only acetylene, but all hydrocarbons found in the atmosphere of an oxygen plant should be considered as potentially dangerous, and especially those with a low solubility in liquid oxygen. The presence of oxides of nitrogen in the atmosphere and their influence on hydrocarbons should not be disregarded. It was already known at the time that in coke oven gas plants, nitric oxide (NO) can polymerize

Heat input and intensity quantity friction catalyst

rgy

ne

ne

Fu

el

itio

Ign

Composition mass geometry terminalconductivity oxideformation

Oxygen Total pressure purity dynamics mass

FIGURE 12.1 Ignition and fire triangle (modified Slusser–Miller triangle). (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


unsaturated heavy hydrocarbons to form a compound with highly explosive properties. Even nitrous oxide (N2O) can form cocrystals with acetylene, and the formation of this compound resulted in a reboiler explosion of a large plant.2 Nitrogen dioxide (NO2) can also dimerize to dinitrogen tetroxide (N2O4), which has explosive characteristics. Since 1950, greater attention has been paid to the analysis of all contaminants entering an oxygen plant, and the results have helped the industry to pinpoint specific factors and conditions, giving rise to an abnormal energy release. One must consider the fact, however, that contaminants, even in highly polluted industrial atmospheres, are found in terms of a few vppm or even less than 1 vppm. It is a dangerous assumption that such small quantities can be ignored or eliminated by a few safeguards. In large plants in the range of 1000 t=d, the accumulation of 1 vppm of acetylene in the atmosphere can reach up to 4.8 kg=d in the main condenser–vaporizer if no special precautions are undertaken. In the era of Revex cycles for prepurification and the use of rich liquid filters, the accumulation was serious. A test carried out on a 100 t=d oxygen unit operating for a period of 1 year with Revex and equipped with rich liquid adsorbers, the following results were registered (Table 12.7).3 Experience thus shows that in the operation of a plant with a Revex process cycle there is a gradual accumulation of organic material in the oxygen plant. It tends to collect on the walls of the reversing heat exchangers, piping, distillation columns, in the rich liquid and in the liquid oxygen. The liquid oxygen main condenser–vaporizer is the most critical locality as it receives all the liquid oxygen, and contains all the contaminants not removed by specifically designed prepurification procedures. A combustible gas may concentrate in a liquid oxygen mixture by .

.

Surpassing the solubility limit of the combustible in liquid oxygen, and the settling of the precipitate to produce a small volume of a highly concentrated mixture Surpassing the upper limit for homogeneous solution, which is usually above 0.5%

However, even contaminants such as methane, ethane, and ethylene that have high solubility limits have served as combustible sources for the release of energy through the drying up of oxygen splashes, or their occlusion in already existing precipitated material.

12.2.2 IGNITION ENERGY Dangerous mixtures of combustibles in oxygen-rich fluids do not explode on their own. They require ignition energy. This energy may be supplied as mechanical or thermal shock, friction, static electricity, heat or the introduction of trace amounts of a third reactant, which has the capability of undergoing rapid reaction with the fuel source. If one examines case histories of oxygen plant explosions, one finds recurring pattern. A great many of the incidents have taken place shortly after, or sometime after a plant stoppage and restart for reasons of deriming, an operational upset, a compressor stoppage, an electrical failure, etc. This indicated that the liquid oxygen level in the main condenser–vaporizer had dropped, and entrained contaminants were above their solubility limits. The result was subsequent precipitation and possible formation of occlusions or cocrystals. Then during the restart there was enough thermal or mechanical shock to initiate the explosion. Another possibility was an atmospheric inversion. One specific case involved a 100 t=d oxygen plant next to a nitric acid production plant. Just before the explosion it was observed that there was a dark red gas, probably a mixture of nitrogen dioxide (NO2), a red-brown gas, and dinitrogen tetroxide (N2O4), a colorless but dangerous gas, which was carried by


inversion into the air intake of the air separation plant. Fortunately, the explosion took place in the auxiliary vaporizer, which had been designed to remove the bottoms of the main condenser, and dispose them outside of the cold box. The auxiliary vaporizer had been placed at ground level near the outer casing of the cold box, far from the main condenser and the distillation columns. The original idea had its provenance from the early days of the industry when incidents from acetylene entering into oxygen units were quite common. This was the first accident experienced by the industry involving materials other than acetylene, which opened the door to serious investigations on safety (Figure 12.2 and Figure 12.3).

Crude argon

Crude argon column

Lowpressure column

Rich liquid filters

Vaporizer– condenser Auxiliary vaporizer

Liquid nitrogen Oxygen Purge

Highpressure column

Air from heat exchangers Rich liquid

Nitrogen reflux cooler

Rich liquid cooler

Nitrogen to heat exchanger

FIGURE 12.2 Two-column system with auxiliary vaporizer. (Courtesy of F.G. Kerry, Inc. With permission.)


Safety


Booster

Expander

Waste nitrogen

GOX WN2

P

Reg N2 Upper column

Primary exchangers Waste heater P

LOX filter

Prepurifier unit

LOX pump

Cooler and chiller

Air filter

Main condenser

P

Subcooler

LIN to storage

Lower column

P

Air compressor

Reflux

LIN

Air

Secondary condenser

Rich liquid

Motor LOX storage

FIGURE 12.3 Process schematic of air separation plant. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

415

Atmospheric inversion was also suspected later and more serious explosion involving a 3000 t=d oxygen plant located in an oil refinery and depot in the Far East. The incident took place at night during a dense fog. The culprit in this case was probably propane escaping from the refinery’s propane storage tanks nearby. From the description of the results of the explosion, the location of the ignition was in all probability in the main condenser–vaporizer of the oxygen unit.

12.3 CRITICAL AREAS IN AN AIR SEPARATION PLANT 12.3.1 GENERAL DESCRIPTION If one examines a standard low-pressure air separation process cycle (see Figure 12.3), one can better understand the behavior of contaminants. Air enters the plant after compression to 5–10 bar and after prepurification is cooled to about 103 K by countercurrent heat exchange with the outgoing products. Refrigeration is supplied by the expansion of the compressed air in gaseous or liquid form, after going through a series of throttling valves. Refrigeration losses due to heat gain are made up by expanding a small percentage of either air or nitrogen in an expansion turbine using the standard Claude cycle. The incoming air is condensed and partially rectified in a high-pressure column at 5–6 bar producing a liquid containing 35%–40% oxygen, and is therefore called rich liquid. In some designs, air is withdrawn from a few trays above the bottom of this column and sent to the expansion machine to balance refrigeration lost from any heat gain in the overall cycle. The temperature of the rich liquid is approximately 103 K. After being subcooled countercurrent with the outgoing nitrogen from the upper column, it is sent through an expansion valve into the middle of the upper column. As the rich liquid descends the upper (low-pressure) column, the nitrogen and part of the argon are more volatile, vaporize and ascend the column, whereas the oxygen is less volatile, descends and settles in the sump of the main condenser–vaporizer at a temperature of about 93 K. One can understand, therefore, that any dangerous contaminant that enters with the main airstream and into the high-pressure column will also enter into the main condenser along with the liquid oxygen. Its concentration will increase by almost 100% in the rich liquid and fivefold in the liquid oxygen in the main condenser. In other words, if 1 vppm of acetylene enters with the main air feed, the resultant liquid oxygen in the main condenser will contain 5 vppm. At the temperature of liquid oxygen in the sump of the upper column (93 K), acetylene is near its limit of solubility assuming that it is evenly distributed (which is never the case). The freezing point of acetylene is around 182 K. It can easily attach itself to the surfaces of cold equipment, i.e., exchangers, distillation columns, internals of the main condenser–vaporizer, etc. Hence, designers have always limited the acetylene content of the air feed to less than 0.1 vppm. If it ever reaches 1 vppm in the main condenser one has to consider taking emergency action or even shutting down the entire plant. As already noted, however, acetylene is not the only dangerous contaminant, and the elimination of other contaminants requires instituting different procedures. A review of the various critical areas of an air separation cycle wherein contaminants may accumulate dangerously will prove to be prudent.

12.4 PURIFICATION SYSTEMS 12.4.1 ADSORPTION SYSTEMS4,5 As explained in detail in Chapter 5, the application of pressure–temperature swing adsorption (PTSA) has made life easier for the process designer of air separation plants. This system was


able to remove all the acetylene as well as carbon dioxide, and the majority of other dangerous hydrocarbons. Nevertheless, it still permitted the escape of all methane and ethane but still retained at least 10% of propane.6 As it was not too long ago that propane was found to be a serious factor in plant safety, most conservative designers have employed a LOX filter trap at the recirculating outlet of the main condenser–vaporizer to remove any traces of frozen hydrocarbons that may have slipped through the adsorption system. This prevents their accumulation in liquid oxygen in the main condenser. The design of a 100% foolproof adsorption system has not yet been achieved by any major company. Due to the aggressive competitive marketing, designers are still trying to find the right balance between the quantity of adsorbent for removing hydrocarbons and water from the process air. Though not dangerous in itself water does hinder adsorption and regeneration of the adsorbent. As shown in Chapter 5 (see Section 5.3.2), the velocity of the air feed through the adsorbent bed should be less than 20 cm=s and close to 10 cm=s. Studies of more than 30 applications indicated that the minimum quantity of molecular sieve (type 13X) for the removal of all contaminants should be 50 kg per 1000 N m3=h of process airflow per hour of adsorption. The weight of activated alumina for removal of water upstream of the main adsorbent for hydrocarbons should be around 25% minimum of the main adsorbent. In areas where the atmosphere is heavily contaminated with propane the quantity of the molecular sieve should be increased. However, this has not yet been quantified. The mass of carbon steel to be used for fabrication of the adsorber vessels can be an important factor in the regeneration procedure as it draws an important quantity of regeneration heat, and may affect the overall adsorption capacity of the adsorbent. In some highpressure designs, which may involve very thick walls (say 50 mm and over), the use of a layer of insulation on the exposed internal walls of the vessel should be considered.

12.4.2 REVERSING HEAT EXCHANGERS: REVEX Though the use of Revex for prepurification of air separation plants has fallen to a small percentage, the system is still used and must be reviewed. Without question the contaminants can be removed from the process air during the cooling stage by condensation into the walls of the front end heat exchange system. The amount removed will be proportional to the difference between the partial pressure of the contaminant and its vapor pressure under the existing conditions. Contaminants will tend to be removed down to their respective equilibrium vapor pressure at 103 K. Some accumulation of contaminants will occur, however, and be a potential danger if plant operating conditions are upset. For example, if the exchangers warm-up due to a failure in the cooling cycle the revolatilized contaminants will be carried into the stream liquefied. Once introduced into the liquid stream, they will be carried to the remainder of the plant unless other safeguards are introduced. Another danger point in the Revex exchanger system is the possibility of freezing an explosive mixture such as a mixture of dangerous hydrocarbons with NO on the walls of the exchangers. During a defrosting period this mixture can explode and fracture the exchanger. This has already happened in the field. If any defrosting is to be carried out on Revex plants it should be done at night shifts under controlled conditions.

12.4.3 NONREVERSING HEAT EXCHANGERS (PRIMARY HEAT EXCHANGERS) With prepurification employing adsorbents, the primary exchangers become nonreversing, but also require attention. Though not recommended by the Compressed Gas Association, some designers send the contaminated and discarded liquid from either the main condenser–vaporizer


or from the auxiliary vaporizer into the product oxygen passages of the warm section of the primary heat exchangers (PHX). There it can be vaporized and become part of the main product. This is a dangerous procedure and should be avoided.7,8 As noted, this contaminated liquid can be disposed or saved in two different ways. It can be piped out of the cold box and dumped into a disposal system operated with steam, or piped out as a liquid and injected into the main product oxygen pipeline. This latter option is acceptable as long as the quantity to be disposed is relatively small compared with the main product oxygen. Even in the latter case, however, there have been events where the disposable liquid to be vaporized has frozen the outlet of the injector thus failing in its objective.

12.4.4 DISTILLATION COLUMN

AND

MAIN CONDENSER

The rich liquid at the bottom of the high-pressure column consists of approximately 50% (gas equivalent) of the process air. It includes all the oxygen and argon contained in the air plus any remaining nitrogen, as well as all contaminants not eliminated either by the reversing heat exchangers or by the prepurification adsorbers. If either system is used, keep in mind that the concentration of all contaminants is multiplied by 500% when they reach the critical areas of the distillation columns and especially the main condenser. When oxygen is removed from the main condenser as a gas and no liquid purge is maintained, the concentration of the contaminants will tend to increase. If their solubility limit is surpassed they will accumulate and precipitate in the main condenser. Table 12.5 shows maximum vapor pressures of contaminants above liquid oxygen at 92 K and 1.2 bar. As noted methane, ethylene, and possibly ethane are the only contaminants that can be carried outside the plant by product oxygen from liquid oxygen boiling at 92 K

TABLE 12.5 Vapor Concentrations of Contaminants in Different Conditions 1

2

Contaminant

Vapor Concentration in Air at 92 K at 600 kPa (vppm)

3 Vapor Concentration Above Saturation LOX Solution at 92 K at 121.6 kPa (vppm)

Methane Ethane Propane n-Butane Isobutane n-Pentane Ethylene Propylene 1-Butene Isobutylene 1-Pentene Acetylene 1,3-Butadiene Nitrous oxide Carbon dioxide

79,000 29 0.12 3.7 104 2 103 Less than 2 104 145 0.22 7 104 7 104 Less than 2 104 0.71 Less than 2 104 0.66 0.1

1,200 (1%) 0.5 (1%) 5 103 (1%) Less than 5 104 Less than 1 103 Less than 2 104 20 (1%) 0.05 Less than 8 104 Less than 6 104 Less than 1 104 0.14 Less than 1 104 0.12 0.015

Source: From Sefton, V.B., Contaminants, Source Physical and Chemical Properties, Basis of Limits, Report presented at Compressed Gas Association, October 1962. With permission.


TABLE 12.6 Contaminant Solubilities in Liquid Oxygen and at Freezing Point Component Methane Ethane Propane n-Butane Isobutane n-Pentane Isopentane n-Pentane n-Hexane n-Decane Ethylene Propylene 1-Butene Isobutylene 1-Pentene 2-Pentene 1-Hexene 1-Octene

Solubility at 92 K (vppm) Miscible 250,000 60,000 700 2,500 50 350 7 1.5 30,000 10,000 1,500 200 300 60 15 5

Freezing Point (K) 91 90 85 135 114 143 112 253 173 243 104 88 88 133 108 133 133 171

Compound

Solubility at 92 K (vppm)

Freezing Point (K)

Acetylene 1,3-Butadiene Isoprene

6 25

192 160 153

Methanol Ethanol Acetaldehyde Acetone Benzene

16 20 0.5 2 5

75 156 150 178 268

Carbon dioxide Nitrous oxide Nitric oxide Nitrogen dioxide Ammonia Sulfur dioxide Ozone

4.5 180 6 15

217 182 109 262 195 200 80

Miscible


and 1.2 bar. Yet, there is clear evidence of plant explosions traced directly to ethylene and propane in the presence of other contaminants such as oxides of nitrogen. Table 12.6 indicates the solubilities of various contaminants in liquid oxygen.

12.4.5 AUXILIARY VAPORIZERS The surest way to eliminate all danger is to increase the liquid purge to 100% of the total production in liquid form directly outside the plant as carried out in total liquid-producing plants. An alternative method is to use an auxiliary vaporizer. This has been the practice since the 1930s (see Figure 12.2). Its use may be explained as follows. The quantity of hydrocarbons and other dangerous impurities is going to increase in the liquid oxygen until a level is reached at which the impurities escaping with the gaseous oxygen are equal to the total entering the low-pressure column. Their concentration in liquid oxygen in equilibrium with gaseous oxygen can be fairly high, and in some cases approach the solubility limits of certain hydrocarbons. Rather than allow the whole sump of the low-pressure column to be so contaminated, product oxygen is removed as a liquid and vaporized in the auxiliary vaporizer. The liquid entering the auxiliary vaporizer is at the concentration level equal to the concentration of impurities in gaseous oxygen. (Equal concentration means much lower content than at equilibrium concentration.) A small stream of unvaporized liquid high in contaminants is purged from the system. This small purge is constantly analyzed for acetylene and other hydrocarbons. The auxiliary vaporizer can be isolated and derimed to remove hydrocarbons without stopping the plant. Experience, however, has proven that auxiliary vaporizers do not always eliminate explosions. They merely reduce the damage done to the main operating equipment. The explosions usually take place in the auxiliary vaporizer


Pressurized LOX to primary BAXH

LOX from main condenser Process air To high-pressure column

Pump

Filter

Bypass LOX storage

FIGURE 12.4 Pressurized secondary condenser. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

because the quantity of material exploding is so much smaller. Thus, the secondary or auxiliary vaporizer is a useful piece of equipment to improve safety. In the 1970s, the auxiliary vaporizer was further designed to operate as a secondary and pressure vaporizer due to its location at the lowest elevation in the cold box. The hydraulic pressure of the descending liquid raised the pressure of the liquid, which when vaporized could become product gaseous oxygen available at a pressure above atmospheric. This saved energy at the first stage of the oxygen compressor. This idea was quickly put into practice in 1978 in a large plant with an oxygen capacity of 1200 t=d at a pressure of 55 bar. In this case, a small quantity of liquid oxygen, 1% of total production, was withdrawn from the bottom of the secondary condenser, analyzed, and sent to the silencer stack of the molecular sieve adsorbers for disposal (Figure 12.4). As the industry progressed, the same auxiliary vaporizer was also used to transfer refrigeration to the incoming air from the PHX. This lowered the temperature another 5 K before entering the high-pressure column. In turn, the liquid oxygen product was vaporized to the desired temperature needed at the entrance of the PHX. Another solution was to use the refrigeration capacity of the LOX in the secondary condenser–vaporizer to liquefy a portion of the main process air going into the high-pressure column, but at a higher level in the column than that of the rich liquid. In all cases, a small portion of the liquid was withdrawn for analysis and disposal (see Figure 12.2).

12.4.6 ANCILLARY EQUIPMENT 12.4.6.1

FOR

SAFETY

Rich Liquid Filters

For many years, the oxygen industry has used adsorbent filters usually consisting of silica gel, in powder form or pellets to remove contaminants from the rich liquid before it enters to accumulate in the main condenser–vaporizer. Unfortunately, experience has also shown that these filters sometimes do not contain sufficient adsorbent material, or are improperly regenerated, or even badly designed, and explosions have occurred in spite of their use. On several occasions, explosions occurred at the filters themselves during regeneration. The expanding use of prepurification adsorption systems for the removal of dangerous


contaminants, however, has eliminated the use of rich liquid filters. These were costly and a problem to the operators (see Figure 12.2). 12.4.6.2

LOX Guard Filter (Figure 12.3)

In the operation of a Revex process cycle, and apart from the use of rich liquid filters (in duplicate), the product oxygen still in the liquid phase requires the use of a LOX guard filter trap to remove any possible stray and dangerous contaminants remaining in the liquid. With the use of prepurification adsorbent systems, the use of rich liquid filters has been eliminated, but the use of the final LOX guard filter is still strongly recommended. A single filter of a large capacity is acceptable. This can be bypassed during its deriming phase, depending on the atmospheric contamination in the area.

12.4.7 LIQUID OXYGEN STORAGE TANKS As noted previously, liquid oxygen removed from an air separation plant and piped into a storage tank may not need any special treatment in regard to safety. It has to be kept in mind, however, that constant vaporization takes place in the tank, and if any dangerous contaminants are entrained with the liquid their concentration in the tank will keep rising and possibly to a dangerous level. It is prudent, therefore, to install a drain near the bottom of the liquid tank to check a liquid purge periodically to analyze its contamination level of the liquid.

12.4.8 SUMMARY These general conclusions are as follows: .

.

.

The efficiency of contaminant removal by a heat exchanger cannot be predicted solely on the basis of contaminant vapor pressure data. There is a considerable evidence that solid accumulations will occur even though the average concentration in the bulk liquid phase is well below the solubility limit. Rich liquid silica gel filters are very efficient in removing small quantities of acetylene, but only partially effective in removing C3–C4–C5 paraffins. Their efficiency, moreover, is further reduced by the presence of carbon dioxide and nitrous oxide. Neither of these filters is very effective for the removal of methane and ethane.

12.5 PARAMETERS FOR THE SAFE DESIGN OF A PROCESS CYCLE To design an oxygen plant for maximum safety the process engineer must start with certain objectives: .

.

.

.

Liquid oxygen in the main and secondary condensers must be free from all harmful contaminants. Any contaminants entering the main and secondary condensers must be considerably below their solubility limits. Air to be used in an expansion machine should be taken from well above the sump of the high-pressure column. If air is used for an expansion machine that comes directly from the reversing exchangers, it should be removed from the system after expansion and not allowed to enter the low-pressure column. The process cycle must be as simple as possible and easy to operate. This will avoid complicated operating procedures and plant upsets, which may increase the danger of contaminant accumulation in critical areas of operations.


.

Process equipment, especially brazed aluminum heat exchangers, should be so designed, specified, and inspected during fabrication as to be trouble free and give long onstream time. This will avoid frequent stoppages which, unless properly carried out, may cause conditions for an unexpected energy release.

12.6 GENERAL DESIGN PROCEDURES Obviously prime consideration should be given to the location of the oxygen plant. Heavily contaminated areas should be avoided if at all possible. Unfortunately, most large oxygen plants serve the needs of metallurgical and petrochemical industries, which means that the entire area is polluted with dust, hydrocarbons, oxides of nitrogen, carbon dioxide, and the ever present acetylene. Even if a specific area is not heavily polluted, it is very difficult to predict any increase in contamination over the life of the plant. The design engineer, therefore, should project and evaluate the necessary degree of safety. The design engineer should also include in his design all known and successful features, in a general manner. Even then, there has to be included as part of the safety equipment an in-line gas chromatograph to monitor constantly, precisely, and quickly the level of individual contaminants going through the plant at the critical zones as shown in Table 12.7.

12.6.1 FRONT END PREPURIFICATION Whereas both the PTSA and the PSA systems are highly recommended for safe designs they are in their development phase for maximum contaminant removal, and still in the ‘‘What did we do in the last project?’’ stage. Designers are still pushed by market pressures to hone their designs, possibly to a dangerous limit (see Chapter 5).

12.6.2 REVERSING HEAT EXCHANGERS Properly designed, reversing heat exchangers will not only remove carbon dioxide and water, but also the bulk of other contaminants as well. Nevertheless, it is also known that small quantities of carbon dioxide, water, and other contaminants do go past the front end exchange system and accumulate in the rest of the plant. One is never sure if such a breakthrough is the result of poorly designed exchangers, careless plant operation, a bad process, poor analytical procedures, or a combination of these. If the overall process is poorly designed the plant operation will be unstable, and upsets will be frequent. The reversing exchanger system should be designed not only on the basis of correct vapor pressure data, but

TABLE 12.7 Contaminants Held over 1 Year Contaminants Oxides of nitrogen (NO þ NO2) Nitrous oxide (N2O) Hydrocarbons (C3–C5) Acetylene Carbon dioxide

Walls of Revex

Main Condenser

4–60 g 30–400 g 10–130 g 2–5 g —

7g 2.1 g 63 g 0.02 g 3–10 g



also by taking into account the entire geometric distribution of the various exchanger components such as exchanger passages, piping, and valves. This will result in a system whereby the incoming air and the outgoing products can be balanced efficiently and easily with very little operator control. This system will not be prone to frequent upsets.

12.6.3 NONREVERSING HEAT EXCHANGERS Design and operation of the PHX of the nonreversing type should not be ignored from a safety point of view. The upper warm end of the exchangers should not be used to vaporize and dispose of contaminated LOX by using the gaseous oxygen passages unless pressure of the oxygen is a minimum of 3 barA.7 According to Compressed Gas Association Report, CGA P-8, 1989, Article 8.17.3, one should be careful in the design of the PHX to avoid the possibility of condensing process air at the cold end of the exchanger. Such liquefaction may lead to a hazardous situation. The resulting rich liquid may contain significant concentrations of C3’s and C4’s.8

12.6.4 HIGH-PRESSURE COLUMN The lower (high-pressure) column is very important because its design sets the process pattern for the rest of the plant. First of all, it receives all incoming process air at its bottom. As the vaporized air rises to the top, it is rectified to a high-purity nitrogen depending on the number of distillation trays. As soon as it reaches the top, it is condensed in countercurrent heat exchange with the product in the main condenser. Part of the liquid nitrogen is sent to the top of the low-pressure column to serve as reflux, and the remaining liquid flows down the highpressure column over the distillation trays, stripping the rising air of its oxygen and argon content. As the entrained oxygen and argon are condensed, they descend to the sump, helping to concentrate all the incoming contaminants. This liquid, rich in oxygen, is then subcooled and sent to the low-pressure column along with all entrained contaminants. The rich liquid, therefore, must be constantly analyzed for dangerous contaminants. It should be considered as the red light to warn the operator that may be going on in the main condenser or the secondary condenser if the latter is in use.

12.6.5 MAIN CONDENSER This piece of equipment, sometimes called condenser–vaporizer or condenser–reboiler, is the heart of an air separation process cycle. Its design determines both the process efficiency in terms of heat transfer between the contained liquid oxygen vaporized and the circulating nitrogen from the high-pressure column that is liquefied. Its design also holds the key to the safety of the plant. Its purpose is to concentrate the oxygen product in liquid form, and allow it to vaporize into a gaseous product at the desired purity after nitrogen and argon are driven off by reboiling the liquid. The degree of vaporization or reboiling of liquid oxygen is a function of the heat transfer between this liquid oxygen and the pure gaseous nitrogen reaching the top of the high-pressure column. As a result of its heat transfer incoming nitrogen condenses and liquid oxygen vaporizes. The closer the temperature difference the lower the process energy requirements and the higher the process efficiency. The large quantity of liquid oxygen present in the main condenser presents a potential hazard if dangerous contaminants are allowed to enter and accumulate beyond their solubility limits. It is also known from field experience that even though prepurification systems and ancillary liquid adsorbers have been used, some dangerous contaminants breakthrough, and enter the main condenser. In the design of main condensers, therefore, one must take into consideration the important parameters for maximum safety.


The most important design feature from a process point of view is to use the maximum amount of heat transfer surface within a fixed volume. Shell and tube condensers have been almost completely eliminated by designers because they have been considered unsafe and costly especially for very large plants. While standard brazed aluminum heat exchangers answer the requirements, they can present problems unless they are built with larger than normal passages. Small passages and poorly formed joints or small cavities may provide an area for adherence for contaminants. Maximum safety requires close supervision during fabrication, not only by the supplier, but also by the purchaser. To eliminate the possibility of any energy release one has to rely on additional safeguards: 1. Design the equipment so that the liquid oxygen is in steady motion over all surfaces of the main condenser–reboiler to avoid the formation of dry spots on which contaminants can precipitate and adhere. The contours of the reboiler should not contain any locality or cavity where liquid can cease flowing, and form a stagnant pool in which contaminants accumulate beyond their solubility limit. 2. Design a continuous liquid purge from the main condenser (anywhere from 0.5% to 1% of the total oxygen production), which can be purged from the cold box, vaporized, and possibly be sent to the product line as gas. In this way one can avoid the steady accumulation of contaminants even though some process efficiency may be lost. 3. All serious explosions can be pinpointed either at the main condenser or auxiliary and secondary condenser because they contain the maximum concentration of dangerous contaminants. Designers should employ a LOX guard filter trap to adsorb any errant contaminants, then re-circulate the liquid oxygen product back to the main condenser. This adsorption trap should be oversized to prevent any unusual peaks in contaminant concentration due to an upset in the prepurification system, or an atmospheric irregularity, i.e., an atmospheric inversion. 4. Recirculating the liquid oxygen may be carried out either by a mechanical cryogenic pump or a thermosyphon pump system. Most designers prefer the use of a mechanical cryogenic pump because the recirculation of the LOX is positive, constant, and well determined in advance. Rate of recirculation depends on many factors but should be very high, at least up to 150% or even higher, to make sure that all the internal surfaces are continuously well irrigated. If the plant is well designed from a safety standpoint there should be no appreciable accumulation of contaminants in the main condenser, and only one adsorber trap with a large capacity needs to be used with a bypass. If mechanical pumps are used, however, two pumps should be piped up and hot-wired at all times. The second pump will serve as a 100% standby. If a thermosyphon system is used, the recirculating rate can (and does) vary, because it is a function of temperature differential of several streams as well as fluid dynamics. Thus, there is the danger of liquid level variations in the main condenser with the possible development of dry spots. With the use of this system, the design on the main condenser is very critical. Laboratory tests6 have indicated that silica gel is not very effective for removal of hydrocarbons. Molsieve 5A should be considered especially for ethane and even for propane. It should be stressed, nevertheless, that liquid oxygen filter traps are to be used as a final safeguard on a well-designed safety system. They should not be considered as a catchall in a carelessly designed plant. In the design and regeneration requirements for these adsorbers, the same criteria apply as for the rich liquid filters except that the minimum quantity of silica gel to be used is approximately 10 kg per 1000 N m3=h of incoming process air.


12.7 LIMITS OF CONTAMINANTS AND ANALYSIS As a rule, it is very difficult to ascertain the degree of contamination in any given area. One has to contend with the degree of industrialization, types of industries, prevailing winds, and atmospheric variations during the change of seasons. Table 12.8 provides the design of the appropriate safety measures. Based on the moles of CO2 per mole of air or oxygen one can determine the total combustion of hydrocarbons. This shows an actual analysis of contaminants in a main condenser using a single LOX guard trap with a bypass. Having defined the limits of contaminants in the process air, one has to monitor the quantities passing through the system to determine the adequacy of the safety measures, and also whether proper operating procedures are followed. Periodic and well carried out analyses should be made and reviewed by a supervisor at least every 8 h (see Section 11.4.14 for full details, also Table 12.7).

12.7.1 ARGON AS

A

CONTAMINANT

9

In 1946, Jost of Germany indicated that argon, because of its greater atomic weight, reduced the limit of an explosive mixture. In the late 1950s, this was proven when oxygen was used for the partial oxidation of hydrocarbons to produce ammonia. With the oxygen plant producing 95% oxygen (the balance mostly argon), it was integrated with the nitrogen-scrubbing unit, and an explosion occurred in the latter unit. This problem was eliminated when the oxygen purity was increased to at least 98% or higher.

12.7.2 PROPANE

AS A

CONTAMINANT

In 1983, Reyhing6 suggested that propane may be a serious contaminant in air separation plants. Later, in 1996, Lassmann,7 in concert with a team selected by the CGA committee, stated in a serious study that propane may be enriched to a higher concentration due to its very low volatility at 91 K. At around 1 bar, its limit of solubility may be reached. At higher pressures, however, around 3 bar propane is carried off with the gaseous phase oxygen due to a more favorable equilibrium constant at higher pressures (see Figure 12.5). At a pressure above 3 bar, it may be considered as a safe process system. TABLE 12.8 Accumulation of Dangerous Contaminants in an Air Separation Unit Contaminants (vppm)

Acetylene Process air Main condenser Oxides of nitrogen Process air Main condenser Hydrocarbons Process air Main condenser

Abnormal (vppm)

Dangerous (vppm)

Critical

Take Action

0–0.1 0–0.05

0.1–0.5 0.05–0.1

1 1

0.0–0.1 0.0

0.1–0.3 0.0–0.05

0.3 0.1

0.2 0–15

2–10 15–150

Acceptable (vppm)



10 150

100,000 ppm Solubility

10,000 ppm

0.3 ppm C3H8 in LOX to evaporator without molecular sieve

1,000 ppm 100 ppm

0.03 ppm C3H8 in LOX to evaporator with molecular sieve

10 ppm 1 ppm 10

Pressure

51.3

100 bar

FIGURE 12.5 Enrichment of propane in boiling LOX. (Courtesy of BOC Process Plants, 2006. With permission.)

12.8 ROTATING MACHINES AND OTHER EQUIPMENT 12.8.1 EXPANSION MACHINES The high-pressure column may also supply air or an air–nitrogen mixture at column pressure for the production of energy to offset loss from heat gain in the process cycle. If nitrogen is used, it must be taken near the top of the high-pressure column. If mostly an air–nitrogen mixture is used, it must be taken at a point several trays above the sump of the column to make sure that it contains no dangerous contaminants concentrated in the sump. It is not a good practice to use air coming from reversing heat exchangers for expansion machines, and send it directly to the upper, low-pressure column. This should be avoided.

12.8.2 LIQUID OXYGEN RECIRCULATING PUMPS On one occasion an explosion occurred in a recirculating pump, which had just gone through periodic maintenance and cleaning. The casing was completely destroyed and the flexible metallic hose connection on the inlet side distorted. Only the bronze impeller was left without apparent damage. The pump had been dismantled for maintenance work and needed a new sealing ring. The pertinent parts had been taken to the maintenance shop where the old ring was removed and replaced with a new ring. The parts were then thoroughly degreased by immersion in trichloroethane. Afterward, the parts still immersed in the solvent were returned to the site. No attempt had been made to dry the parts before reassembly. As soon as the pump had been completely reassembled, it was allowed to dry in this fully assembled form when the operators went out for lunch. After 1.5 h the operators returned and started the pump. In less than 5 min the explosion occurred. In open air, the thin film of solvent would have evaporated in much less than 45 min. In this case, however, the wetted pump parts were in an enclosed housing with little or no ventilation. After 1 h=25 min there was sufficient vapor to initiate an explosion in a mixture with liquid oxygen under pressure (Figure 12.6 and Figure 12.7). This was not the first time that a fire or an explosion involving a solvent was reported. Experimental tests have indicated that a mixture of 67% oxygen and 33% 1,1,1-trichloroethane


FIGURE 12.6 On right, a re-circulating LOX pump connection damaged from an explosion, compared with new one on left. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

at atmospheric pressure and 293 K will react if a spark of only 1000 V (1 mm in length) is generated in the mixture. Though this solvent has no flash point or fire point in air, within certain limits, it is combustible in high oxygen concentrations. Thus, any equipment degreased with this product or any other solvent should be thoroughly purged and dried with clean dry oil-free nitrogen, before use for oxygen service.

12.8.3 LIQUID PURGE LINES Explosions have also occurred in cryogenic liquid purge lines. Such explosions have been pinpointed to (a) changes in the product piping system made by the local operators without marking revisions to the original P & I drawings, and without advising the original supplier and (b) changes suggested in the basic design by a new piping draftsman without first reviewing the possibility of accumulation of dangerous contaminants accumulating in the liquid to be purged (see Figure 12.8).

12.8.4 LIQUID OXYGEN DISPOSAL10 With the ever increasing capacity of oxygen units since 1980, the disposal of large quantities of cryogenic liquids is becoming a major problem. The dumping of contaminated liquids, especially liquid oxygen, cannot be tolerated in industrial areas or near residential zones. These disposable liquids are from start-up operations, off-specification products, highly contaminated oxygen, and purging before and following a normal shutdown. Whereas there are numerous designs available for the design engineer the two most common and the

FIGURE 12.7 Damage done to a 4 in LOX purge drain by an explosion. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


(a) Incorrect design

(b) Correct design

Incorrect design 9-1, permits contaminated purge liquid to remain upstream of the valve, and allowing vaporization of the liquid, thus increasing the contamination and possible explosion, which has happened.

FIGURE 12.8 (a) Incorrect design. (b) Correct design. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

safest to use involve (a) a simple steam-heated horizontal tank with a vertical tower and (b) a water cooling tower. .

.

The advantages of the low-quality steam-heated horizontal tank include the ability to dispose large quantities, and a relative small area for installation. Because stack height may be increased as required, it can be located close to other areas of operation without problem. On the other hand, if there is an interruption in the supply of steam, the condensate disposal outlet may freeze, resulting in a tank rupture. The field staff has to be trained to be ever watchful to check steam and condensate lines at regular intervals during disposal operation (see Figure 12.9). The main advantage of this design is a safe vaporization of a cryogenic liquid using the large volume of air produced by the cooling tower fans. This will not result in a heavy vapor cloud blanketing, and the area is normally produced by other methods. If the piping


15 cm 5 cm

6m

13 cm

Liquid inlet

60 m 20 cm

Steam inlet

Condensate outlet

8m

FIGURE 12.9 Steam-heated horizontal tank (liquid oxygen disposal design). (From Hugill, J.T., Safe Liquid Oxygen Disposal, Compressed Gas Association, April 23, 1969, pp. 32–36. With permission.)

and disposal ring are properly designed, the capacity for disposal is almost unlimited. On the other hand, there is the possibility of power failure or breakdown of the cooling fans. In either case the air separation plant would have to be shutdown. Moreover, there has to been an elevation differential between the plant and the cooling tower to transfer the liquids to the disposal area by gravity. Although this system is simple in principle, the many conditions involved probably make the low-quality steam-heated horizontal tank preferable.

12.9 SAFE PRACTICES The concentration of dangerous contaminants passing through the separation system has to be watched closely, systematically, and must be recorded. If concentrations reach the abnormal limit shown in the Table 12.7, the liquid oxygen purge from the main or the secondary condenser should be increased until the situation is rectified by checking the plant operating procedures or regenerating the adsorbers more frequently. Consideration may also be given to shutdown the plant.

12.9.1 ANALYTICAL EQUIPMENT The main problem in analyzing contaminants is the choice of analytical apparatus. It seems that every designer or project engineer and every plant superintendent have their own opinion of the subject, and their own preference for a certain type of apparatus (see Section 11.4.14 and Table 12.7). It is not the purpose of this text to describe the various equipments available on the market or their operational details. Suffice it to state that qualified equipment does


exist, which is capable to measure accurately and quickly various dangerous or troublesome contaminants such as acetylene, oxides of nitrogen, hydrocarbons, and carbon dioxide. A word of advice with regard to hydrocarbons. Hydrocarbons can be measured individually by an in-line gas chromatograph, or a total combustion analyzer in terms of carbon dioxide. Obviously the latter is less expensive and quicker to read, but has the disadvantage that one also needs a separate acetylene analyzer. Moreover, one never knows the distribution of methane, ethane, propane, and other more dangerous hydrocarbons in the sample stream. Methane is highly soluble in liquid oxygen. Therefore, it evaporates very readily and is considered relatively harmless. Another point to consider is that if a large quantity of methane is present in the atmosphere, the instrument will indicate a high value for carbon dioxide and the average operator may be inclined needlessly to shutdown the plant. On the other hand, an in-line gas chromatograph is a more sophisticated instrument, which can analyze all hydrocarbons including acetylene individually. It requires, however, a qualified person to recalibrate this instrument at least on a weekly basis, as well as costly reference gas mixtures. Only a trained person can read and evaluate the results. It should be stressed, however, that unless the apparatus is specified, to be ultrasensitive, an indicated quantity less than 50 vppb of any contaminant may be deemed questionable.

12.10 SUMMARY In this era of megaplants, one must devote utmost attention to the safe design features and operating procedures, which should be part of an oxygen plant. Whereas some features may seem to be redundant, their extra cost is very small compared with the overall investment for the entire project. In the opinion of most experienced engineers, therefore, the extra cost may be considered as inexpensive insurance. Apart from the danger to human life, the present high investment costs for oxygen plants make it imperative to be up-to-date on all safety procedures. Those who have the responsibility to evaluate air separation systems must be in a position to specify safety requirements properly. In reviewing a supplier’s proposal they should know what questions to ask, and insist on correct technical answers to their inquiries. Safety has always been and should remain an integral part of the design and operation of the plant.

12.11 SAFETY IN THE DESIGN OF DYNAMIC OXYGEN SYSTEMS Design considerations for pressure, flow, flow measurement, and control of pipelines for transmission and measurement and control of pipelines for transmission and distribution of oxygen are not different than for any other gas provided the following codes are followed: ANSI B.31.3 ‘‘Petroleum Refinery Piping’’ for distribution of oxygen generated within the area of use. ANSI B.31.8 ‘‘Gas Transmission and Distribution Systems’’ for gas transmission from a generating plant outside of the area of use. In the design of an oxygen pipeline, however, one must take into consideration certain fundamentals of safety, especially for pure and pressurized oxygen. Nevertheless, if one understands the basic principles of combustion of materials used in an atmosphere of pure oxygen, design becomes much more simple.

12.12 CAUSES OF COMBUSTION In an analysis of all fires that have occurred in oxygen pipelines since the beginning of their use in 1940, one will observe that the majority of them have taken place in four general areas:


1. Points of ignition have occurred at a change of direction of flow of oxygen. 2. Points of ignition occurred whenever metallic or nonmetallic particles were found. 3. Points of ignition were found whenever oxygen flow velocity was higher than the design flow, in combination with high pressure. 4. Electricity whether in static form or as induced parasitic currents caused ignition. One has to understand that oxygen itself is not combustible; it merely accelerates combustion. The higher the oxygen purity and pressure, the greater the energy release. As noted in this chapter, for combustion to occur one needs a combustible material and energy of ignition. These conditions dictate the parameters of choice for cleanliness, prudent selection of materials, reasonable flow velocity, and flow control to eliminate the first energy of ignition and prevent a chain reaction of combustion. On all new piping for oxygen transmission or distribution, one encounters high humidity, mill scale, slag and metallic particles leftover from welding. One also finds organic particles such as grease, oil, or paint. In conditions of high oxygen velocity, these particles with a low point of ignition can enter into a combustion reaction due to adiabatic compression caused by rapid or sudden opening of a valve. Particles can also be entrained by high flow velocity and impinge on a surface of a valve component, an orifice plate or poorly aligned pipe edge. Once started, the heat of initial combustion may supply sufficient energy to another material with a higher ignition temperature, thus setting off a chain reaction, which can result in a fire and possibly an ultimate explosion. This principle is shown in Figure 12.10 and Figure 12.11. Combustion does not occur below a specific temperature that is determined by the pressure of oxygen. The reaction requires that the heat released by oxidation is sufficient to maintain the resulting temperature above the inflammability level of the material in question. Therefore, the specific heat and the thermal conductivity of the material are important factors in the reaction. The principle of ignition involves two phases. The first phase is called the heating phase, where a source of energy is imparted to the combustible material or mixture. In the second phase, the energy wave passes through a series of nonsteady states that may lead to either extinction or propagation. The second phase depends on the quantity of energy provided by the heating phase.

Heat of reaction

Heat of reaction

Kinetic energy

Ignition combustion heat

Ignition combustion heat

I

II

III

Organic or plastic material with low point of ignition

Small metallic particle with high calorific conductivity

Large metallic mass

Ignition fire explosion

FIGURE 12.10 Initial combustion heat and resulting change reaction. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)


Flow

A—Alignment and welding well done

Flow

B—Poor welding = Point of ignition and corrosion

Flow

C—Poor alignment = Point of ignition and corrosion

D—Socket weld = Dangerous and may lead to corrosion

FIGURE 12.11 Proper and improper welding of piping. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

12.13 TEST PROCEDURES AND RESULTS AS EXPLAINED BY DE JESSEY11 Initially, the oxygen industry did not experience any incidents in the field. The reasons were simple enough. Gas flows were low, and the project engineers used relatively small diameter copper pipes and bronze valves. After World War II, however, the booming metallurgical, chemical, and petrochemical industries imposed the requirement for large quantities of oxygen gas at pressures of almost 40 bar. Project engineers, therefore, had to look for new materials capable of safe handling large flows at high pressures, but within limits of economy. New materials had to be acceptable for use with oxygen, with velocity controlled within limits of safety. Subsequent results instructed the industry that combustion generally occurred in flanges, pressure-reducing valves, block valves, and orifice plates but rarely in the actual piping. Results of the tests were as follows: 1. Abrasion: A carbon steel pipe 30 mm in diameter was bent into four coils having an overall diameter of 1 m. High-velocity oxygen was passed through the tube entraining small metallic and nonmetallic particles. Oxygen pressure was 40 barA. During the test, the high-velocity particles reached ignition and could be seen glowing as they exited, but the pipe did not ignite. 2. Impingement: A mild steel plate was bombarded by particles of various materials entrained by an oxygen jet stream flowing at a velocity close to sonic, but the plate did not ignite. 3. Vibration: Tests were made to study the effects of vibration on various plastic materials that are normally used for trims in control valves, in relation to the possibility of igniting


the main control valve. The results were inconclusive because the materials fused before could not ignite. 4. Adiabatic compression: The formula indicating enthalpy (J=kg) for a diatomic gas varies in relation to temperature and pressure. The variation is given in Table 12.9. At higher temperatures the coefficient 0.28 becomes lower because the specific heat is reduced. At a pressure of 60 bar and at a temperature of 701 K the coefficient becomes 0.26. Therefore, the temperature is below the point of ignition of the majority of the nonmetallic materials that are currently used for valves, gaskets, couplings, and measuring equipment. In the tests the best results were obtained using materials such as Teflon, Kel-F, and Viton. Perbunam and Neoprene also gave acceptable results up to a pressure of 40 bar. 5. Electricity: Electrical transmission lines can induce parasitic electrical currents in pipelines. Moreover, parts of valves are linked to electrical systems. In the tests, a small quantity of dust, which normally accumulates in oxygen filters, was placed on the seat (Kel-F) of a valve. A voltage of only 1.5 V was enough to turn the material red. In regard to static electricity, an oxygen compressor newly installed in 1978 and thoroughly inspected for cleanliness caught fire on start-up. On reviewing the design, it was observed that the supplier had not considered the possibility of static electricity built-up between rotor and casing. This was remedied for the next machine by connecting the two elements with a heavy braided copper connection. This same machine is still functioning reliably since 1980. Observations Although the previous tests were of tremendous interest the following comments are offered to avoid any erroneous conclusions: 1. In the case of abrasion, the tube used in the tests had a smooth inside wall. If the wall had been rough with any obstructions, a fire may have occurred. 2. In the case of impingement, the plate did not ignite but the surface was badly scarred. Only its thickness and heat dispersion prevented ignition. 3. In the case of vibration, the materials did not ignite, but the heat of fusion or oxidation could have ignited another material with a lower ignition point nearby. 4. In the case of adiabatic compression, the results were conclusive enough to recommend prudence in the selection of materials for use in oxygen pipelines. TABLE 12.9 Variation of Enthalpy in Relation to Temperature and Pressure of Diatomic Gases T2=T1 ¼ (p2=p1)0.28 and if T1 ¼ 293 K and p1 ¼ 1 bar, T2 ¼ 293(p2)0.28 p2 (bar)

T2 (K)

10 20 40 60

558 678 823 922

Source: From DeJessey, L., Safety in Oxygen Pipeline Systems, Compressed Gas Association, April 23, 1969, 37–45. With permission.


5. In the case of electricity, the tests were not conclusive enough. Additional studies have been recommended by the industry. As noted from laboratory tests and from field experience, it is not difficult to conclude that fires are not a single simple reaction of combustion, but a chain of reactions involving three or four elements. All that is required is an ignition energy, a small element either metallic or organic with a low ignition point, and some other metallic body with low thermal conductivity. If these parameters are clearly understood, one can specify the rules of safety to eliminate as much as possible conditions that lead to combustion or explosion.

12.14 THE FOLLOWING RECOMMENDATIONS ARE IN ORDER 12.14.1 FLOW VELOCITIES Flow velocities should be limited in accordance with those recommended by the Compressed Gas Association Inc., and published in its report CGA—G4.4 covering Industrial Practices for Gaseous Oxygen Transmission and Distribution Piping Systems. Originally, project engineers were limited to a flow velocity of 8 m=s. Now designers follow rules recommended by CGA, which relate flow velocity and pressure. When calculating diameters of piping, however, one has to keep in mind that there is always the possibility of future increase in consumption. Therefore, the diameter of the pipe should accordingly be increased not to surpass the critical flow velocity. Furthermore, flow velocities prescribed by CGA may not take into consideration the existence of valves and fittings. To make allowance for these items it is recommended that any suggested flow velocity should be reduced by 50%.

12.14.2 A VERY CAREFUL SELECTION

OF

MATERIALS

Ignition and fire can occur when the rate of energy input exceeds the rate of heat dissipated by the material’s calorific conductivity. It also depends on (a) ambient temperature, mechanical heat input, and the catalyst, if the latter is present, (b) the mechanical properties and thermal characteristics of the material, the surface conditions: the composition of the metal, if it is an alloy; the oxide formation in terms of adherence and stability; properties of the oxidizer in terms of pressure as well as partial pressure, and the dynamics of contact between the oxidizer and the material. Since 1980 a number of tests have been carried out especially under the aegis of ASTM International Committee G-4 to determine the compatibility of various metals and materials used in an atmosphere of high-purity high-pressure oxygen. Nevertheless, in order to appreciate the results of the tests it is necessary to fully understand certain basic principles. In general, carbon steel is still used for oxygen piping (Figure 12.12), flanges, fittings, etc., and rightly so. One has to be very careful with the use of stainless steel, even though the latter material has a higher ignition point than that of carbon steel. Stainless steel has a low calorific conductivity and may take longer to ignite but once it reaches its ignition point it burns with explosive force, because the formation of its oxide is highly exothermic. Copper and its alloys such as bronze and Monel are acceptable and highly recommended for use with oxygen because their calorific conductivity is very high. Heat from their preliminary reaction with oxygen is dissipated before reaching any critical temperature. These alloys begin to melt before reaching any ignition point. For this reason, such alloys are generally specified for items such as valves of any type, orifice plates, or even small diameter piping whenever one encounters oxygen with a high flow velocity, or a high pressure, say over 40 bar, for example, in bypass valves of oxygen compressors and piping having a ‘‘T’’ connection.


D

D—10 D To be used for underground pipelines with large diameters and pressures higher than 25 bar

C

C—5 D

B

B—3 D

A

Is the minimum to be used for lines underground Preferred for pipelines using high velocity and high pressure up to and higher than 10 bar

A—1.5 D Used for prefabricated piping when alignment and welding are perfect

FIGURE 12.12 Recommended curvature for oxygen piping. (Courtesy of F.G. Kerry, Inc., 2006. With permission.)

With valves, especially those of the pressure-reducing type, one has to be very careful in their selection about material. If the specification calls for the use of stainless steel, one has to make sure that the internals (trim) is either of copper, bronze, or Monel, preferably the latter.

12.15 CONSIDERATION OF DYNAMIC OXYGEN CONDITIONS 12.15.1 TESTS In 1978, the US Department of Energy requested the Metallurgical Section of the Southern Research Institute in Birmingham, Alabama, to carry out an intensive research program on the ‘‘dynamic burn factor’’ (DBF) of various metals and alloys in a dynamic oxygen environment.12 Pears et al.12 began by assuming that the DBF was related to the heat generated by oxidation, divided by the quantity of heat required to melt or evaporate the underlying metal. This assumption was thought to give a good indication of the tendency of the material to burn in a dynamic oxygen environment, but not necessarily in a static environment. The DBF at the melt point was defined by the following equation for a pure metal: (12:1) DBFmp ¼ Df(metal oxide)298 = D298mp þ Dfusion(mp) where DBFmp is the dynamic burn factor at the melt point, Df(metal oxide) 298 is the heat of formation of the metal oxide, 298 is the room temperature, D298mp is the heat required to raise the metal’s temperature from room temperature to its melting point, and Dfusion(mp) is the heat required to cause fusion. The measurement of the tendency for alloys to burn at the melt point can also be calculated by modifying the equation to account for various alloy elements present in the material: Z mp X PBFmp ¼ (DHfoxide formation(1) Ni )= cpalloy þ DHfalloy (12:2) rt

where PBF is the predicted burn factor, Ni is the mole fraction of component i, and cp is the specific heat. In this equation the heat of oxidation of the individual metal contents


is multiplied by the molar concentration of each component present in the alloy. The sensible heat in the base metal is that required to heat it from the ambient conditions to the melting point. A sample calculation is offered for 410 stainless steel, where the principal components are 12% chrome and 88% iron by weight, and the mole fractions are 13% and 87%, respectively. The most stable form of their oxides are Cr2O3 and Fe2O3. When the heat of oxidation of each of the species is multiplied by the mole fraction of each element present, about 95 kcal of heat will be evolved per mole of oxidized stainless steel. The PBF at the melt point, therefore, is given by the summation of the heats of formation of the most stable oxides, multiplied by the mole percentage of the elements present, divided by the heat dissipation ability of the base metal. In this specific case of 410 stainless steel the PBF is 5.39.

12.15.2 EXAMPLE

Elements Cr Fe

Most Stable

DH Mole

Weight

Mole%

Oxide

Fraction

12 88

13 87

Cr2O3 Fe2O3

17.55 77.69

PBFmp ¼ S (heat of formation of most stable oxide mole percent)=heat dissipated by base material PBF ¼ 95:24=(14 þ 3:67) ¼ 5:39 Table 12.10 provides the PBF for various materials. Those marked with an asterisk (*) were not included in the practical experiments.

12.15.3 NONFERROUS METALS There was a discontinuous increase in the DBF when going from ferrous base metals to some nonferrous metals such as titanium, lead, zinc, magnesium, aluminum, and tin. Their burn factors ranged from 13 to 44. These calculations implied that these metals or alloys containing a substantial amount of these metals should burn vigorously and rapidly. It was also assumed that this may be related to the formation of a vapor species at the burning interface, and the calculations were remade assuming a vapor phase combustion using the expression given in the following equation: PBFbp ¼

DHf DH298mp DHmp DHmpbp DHbp

(12:3)

where PBFbp is the burn factor at the boiling point, DHf is the heat of formation of the metal oxide, 298 is the room temperature, DH298mp is the heat required to raise the metal’s temperature from room temperature to its melting point, DHmp is the heat of fusion of the metal, DHmpbp is the heat required to raise the temperature from the melting point to the boiling point, and DHbp is the burn ratio of a metal at its boiling temperature. Using the results of the calculations from Equation 12.3, Table 12.11 shows the materials in order of increasing calculated burn factor after taking into consideration the heat required to vaporize the metal. The team working at the Southwest Research Institute suggested that tin babbitt burns on a liquid phase, whereas titanium, aluminum, zinc, and magnesium are


TABLE 12.10 Predictive Burn Factor of Various Metals and Alloys Material Silver Copper* 90:10 Cu=Ni CDA 938 CDA 314 Monel 400 Co* Monel K 500 Ni* CDA 828 4140 Ductile iron Cast iron 1025 Fe* 17.4 PH 410 SS CA 15 304 SS Ti* Pb* Zn* Lead babbitt Mg* Al Tin babbitt Tin*

PBF at Melting Point 0.40 2.00 2.39 2.83 2.57 3.02 3.50 3.64 3.70 4.49 5.10 5.10 5.10 5.10 5.10 5.32 5.39 5.39 5.39 13.1 18.6 19.3 20.6 22.4 29.0 42.6 44.8

*Materials not included in the experimental program. Source: From Pears, C.D. et al., Final Report to the Department of Energy on Structural Materials Evaluation for Oxygen Centrifugal Compressors, Project 3528 Report XXXVI, Southern Research Institute, Birmingham, AL, 1978. With permission.

definitely capable of vapor phase combustion. Experimental observations made during the test program confirmed these assumptions.

12.15.4 FURTHER STUDIES After the above study was reported to the Department of Energy in 1978, another study was sponsored by ASTM on the same subject and reported by Monroe et al.13 in 1983. In this latter report, the previous term of burn factor was renamed the (Burn Ratio)mp at the melting point of the metals in first study, and the (Burn Ratio)bp at the boiling point of the metals in the second study (see Table 12.12 and Table 12.13). In the classification of the various metals, however, the same ranking materialized, as in the previous study carried out at the Southern Research Institute (see Table 12.9). As noted, one may conclude that the melt- and vapor-burn ratios arrived at may be useful indicators of the combustibility of metals and their alloys, and in predicting their mode of combustion, i.e., in liquid or vapor phase.


TABLE 12.11 Calculated Burn Factor of Nonferrous Metals Material Tin babbitt Sn* Pb* Lead babbitt Ti* Al Zn* Mg*

PBF (Boiling Point) 0.78 0.80 0.90 1.00 1.70 2.20 2.40 3.60

*Materials not included in the experimental program. Source: From Pears, C.D. et al., Final Report to the Department of Energy on Structural Materials Evaluation for Oxygen Centrifugal Compressors, Project 3528 Report XXXVI, Southern Research Institute, Birmingham, AL, 1978. With permission.

In metallic materials prone to ignition and fire a chain reaction is usually started from either a hydrocarbon contaminant, a polymer particulate, or even a small metallic particle leftover from welding. Initial resistance to ignition may be brought about by a protective coating on the metal’s surface. According to Pears et al.,12 however, such a protective coating must be a stoichiometric line compound to be effective, such as Cr2O3 and Al2O3. On the other hand, nickel oxide and iron oxide are not protective coatings because they are not line compounds. Even so, chromium oxide cannot be used for temperatures above 1072 K, because it volatilizes. Even aluminum oxide, a good choice, can only form at high temperatures. Above 1272 K it cannot withstand thermal shock as it has a tendency to spall. A later study was carried out by Beeson et al.,14 editors of ASTM Committee G4. Beeson and coworkers attempted to apply laboratory findings to the overall design of oxygen systems as deemed necessary by the industrial gas industry.

TABLE 12.12 Burn Ratio at Melting Point for Several Metals Material Silver Copper Nickel Iron Titanium Lead Zinc Magnesium Aluminum Tin

(BR)mp 0.40 2.00 3.70 5.10 13.10 18.6 19.3 22.4 29.0 44.8

Source: From Pears, C.D. et al., Final Report to the Department of Energy on Structural Materials Evaluation for Oxygen Centrifugal Compressors, Project 3528 Report XXXVI, Southern Research Institute, Birmingham, AL, 1978. With permission.


TABLE 12.13 Burn Ratio at Boiling Point for Several Metals Material

(BR)mp

Silver Copper Nickel Iron Tin Lead Titanium Aluminum Zinc Magnesium

0.05 0.2 0.5 0.5 0.8 0.9 1.7 2.2 2.4 3.6

Source: From Pears, C.D. et al., Final Report to the Department of Energy on Structural Materials Evaluation for Oxygen Centrifugal Compressors, Project 3528 Report XXXVI, Southern Research Institute, Birmingham, AL, 1978. With permission.

12.15.5 NICKEL

AND ITS

ALLOYS

These materials are resistant to ignition and combustion, have high strength with significant low-temperature toughness. Tests have also shown that nickel 200 wire mesh does not support combustion and is suitable for application as a filter.

12.15.6 INCONEL ALLOYS Inconel alloys such as Inconel 718 have been used in high-pressure systems up to 690 bar, but it has been found only slightly less ignitable than stainless steels. Nickel–copper alloys have been found the least ignitable alloys, and are very much used as structural materials. Monel 400 and K-Monel have not been ignited in impact tests, although some fusing has been observed. These alloys have been found safe to use with oxygen even at pressures of 690 bar. However, they should not be used in sintered form or as wire mesh because they are flammable in oxygen. Hastelloy nickel-based alloys, such as C-22 and C-27, are more ignition-resistant than Inconel 718 or stainless steel.

12.15.7 STAINLESS STEELS Stainless steels are more resistant to ignition and combustion than either titanium or aluminum alloys and are used extensively in high-pressure systems in the industry. Resistance to ignition is about the same for all stainless steels with the possible exception of stainless steel 440C, which is slightly superior. As mentioned earlier, though stainless steel has a high ignition temperature, once ignited it burns with explosive force because of the high-energy release from its oxide formation.

12.15.8 COPPER

AND ITS

ALLOYS

Copper and its alloys are suitable for use in oxygen systems at all pressures. Bulk copper in its pure form can resist ignition by particle impact, and can be useful as impingement plates in pipeline ‘‘T’’ connections. Nonetheless, copper and some of its alloys in finely divided configuration, such as wire mesh are flammable in oxygen. Though copper is resistant to ignition and combustion, it also has a low-ductility oxide, which is not tenacious and can shed easily.


12.15.9 ALUMINUM BRONZE Aluminum bronze is not recommended for oxygen systems because of its very high ignitability and flammability. Sintered bronze, however, is less flammable than either sintered Monel 400 or stainless steel for filter material.

12.15.10 ALUMINUM

AND ITS

ALLOYS

Aluminum and its alloys have the distinct advantage of offering high thermal conductivity as well as high strength to weight ratio, and are used extensively in the cryogenic industry even for pressure vessels. On the other hand, their application for pipelines, valves, and other components should be carried out with caution or even avoided because they can easily ignite in high-pressure oxygen. They burn rapidly and have a high heat of combustion. Moreover, the protective oxide can be rubbed off easily pushing the material into the flammability range. Therefore, they should not be applied whenever frictional heat is a possibility.

12.15.11 COMPATIBILITY OF ALUMINUM AND ITS ALLOYS FOR STRUCTURED PACKINGS (STRUCTURED PACKING CONSISTS OF CORRUGATED STRIPS COILED AND USED AS DISTILLATION TRAYS) In 1990, Dunbobbin et al.15 issued a report after carrying out a series of tests on the oxygen compatibility of high surface areas on metals such as aluminum, brass, stainless steel, and copper. Selection of the metals for the test was based on propagation of combustion and not merely on potential for ignition. The tests indicated that copper alloy 110 resisted propagation of combustion in all test conditions. Brass with 63% copper and 37% zinc had a threshold of flammability lower than that for aluminum but higher than that for stainless steel. For structured packing, however, the adiabaticity and high surface area of brass may permit sufficient heating of the thin metal to vaporize the zinc within the brass. The resultant zinc vapors will burn and may release sufficient energy to melt the copper, in effect burning the packing. In regard to stainless steel, the minimum threshold was taken to be lower than 70% level, similar to brass and lower than for aluminum. Stainless steel propagated combustion above approximately 70% in oxygen vapor or liquid. Aluminum reacted explosively in a liquid oxidant with vapor concentrations exceeding 97%. According to Dunbobbin et al.15 current knowledge would rank the oxygen compatibility of the aforementioned metals in best to worst order as follows: copper, brass, stainless steel, and aluminum. They recommended that copper should be used for structured packing in lowpressure columns for air separation plants.

12.15.12 SUPPLEMENTARY TESTS 16

ON

ALUMINUM

In 1993, Zawierucha et al. carried out similar tests, but by using only aluminum packing and simulating normal operating conditions. In the first series of tests the packing material used was aluminum 3003 with 95.25 mm diameter 76.2 mm height. The foil thickness was 0.25 mm to make up the single circular sample. The data indicated no explosions or serious combustion propagation. In the second series of tests involving a multiple packing element, which was more typical to a real design, the aluminum used was also aluminum 3003. The foil had a nominal thickness of 0.20–0.25 mm, and the weight was 62.67–80.6 g. High-purity LOX (99.9%) was used, and irrigation was continuous as in the previous case involving a single foil. (Irrigation implies a continuous downflow of LOX as in a shower, but the sample does not have any part submerged in the LOX.)


Data from the second series of tests indicated a high potential for combustion arrest, if a promoted ignition event occurred even when strong energetic promoters were used. There was no evidence of severe combustion propagation or explosive tendencies. Therefore, to avoid a violent energy release (VER), as suggested by Dunbobbin et al.,15 it is prudent to avoid even partial immersion of aluminum packing in high purity LOX. One should use a stable continuous irrigation of the aluminum packing.

12.15.13 REPLICATION

OF

ALUMINUM TESTING15

In 1993, Barthelemy17 replicated the testing procedures of Dunbobbin et al.15 and found the same results thereby verifying the work. Nevertheless, Barthelemy also confirmed Zawierucha’s conclusions that by applying proper irrigation and avoiding contact between the structured packing and the liquid oxygen, one can avoid VER. Barthelemy also believed that the violent reactions produced in Dunbobbin’ work were not true detonations, but the result of burning molten aluminum at very high-temperature falling in LOX of ultrahigh purity.

12.15.14 MACHINES USED

IN THE

FABRICATION

OF

STRUCTURED (CORRUGATED) PACKING

Crimping machines to be used for fabrication of structured packing have to be specified and operated carefully to avoid high contamination from lubrication oil. Helpful information may be found in a report from Kirzinger et al.18 They conclude that any oil residue already on the sheets up to 250 mg=m2 remained attached to the packing surface, and there was little if any decrease of oil even when submerged in LIN for up to 4 h. Kehat and Ball in their report mentioned that a quantity of 1 or even up to 5 g=m2 can be tolerated, as at this magnitude no ignition took place in an oxygen atmosphere.

12.15.15 SUMMARY Because of the importance of aluminum used in low-temperature air separation, Werley et al.19 combined their efforts in a report to summarize its advantages, and highlight information available to evaluate it for oxygen service. The report states as follows: Aluminum is an extensively studied metal for use in oxygen service. Many features of its ignition, propagation of combustion, and damage potential are understood, but more remains to be learned. Among the features are: The autoignition temperature of aluminum is reported to be higher than most other metals; and it is near the oxide melting point of about 20458C (2318 K). In powder form spontaneous ignition can occur at much lower temperatures. Aluminum can be ignited by mechanical impact at lower energies than most other metals, although test results vary. Fresh metal exposure does not cause ignition in the absence of other effects, but aluminum has one of the lower friction ignition thresholds of all metals. Aluminum fines have been ignited by sparks; but bulk samples are difficult to ignite due to the high thermal conductivity of the metal. Aluminum burns quickly compared with other metals, and in some conditions can appear to detonate. Aluminum will react with the oxides of other metals, and with fluorinated organics. Aluminum is widely held to be a material that burns in the vapor phase, a mode of combustion more rapid than that involving the liquid phase. Thermodynamic arguments support vapor-phase burning for aluminum, but mixed-phase com bustion is also seen as possible.


Small quantities of inert diluents in oxygen significantly alter the threshold for combustion of aluminum, rendering it less flammable. Argon has a much more pronounced effect than nitrogen. Aluminum is volatile among metals. Its flash point can be estimated from the partial pressure corresponding to the lower flammability limit concentration of aluminum vapor. The effects of alloying elements on aluminum flammability have not been investigated, but they can be significant. Lithium, for example, appears to raise the threshold for aluminum alloy propagation of combustion.

12.15.16 IRON ALLOYS14 Iron alloys are not a good choice for oxygen systems because they can ignite easily. Alloy steels between 5% and 9% nickel are used in cryogenics, especially for flat-bottom storage tanks. As indicated in Section 6.3.5.2, 6.3.5.3 and especially 6.3.5.4.3 an iron alloy with 9% nickel has a greater resistance to impact than 8% nickel steel. If one considers anything lower than 5% nickel, he should be advised that around 3.5% nickel, the steel alloy may suffer from embrittlement at the very low temperatures involved.

12.15.17 NONMETALLIC MATERIALS14 Nonmetallic materials such as elastomeric or thermosetting polymers, plastics or composites, solid and liquid lubricants are used extensively in oxygen systems, and a review of these is advisable. Commonly employed compounds are fluorinated compounds such as Viton and Fluorel, which are used for O-rings and diaphragms because of their flexibility. They have a glass transition temperature Tg below room temperature and can be used up to 520 K. Silicon rubbers have a very low Tg, but are not recommended because of their poor resistance to ignition. The plastics most frequently employed are: . . . .

Polytetrafluoroethylene (PTFE) Fluorinated ethylene propylene (FEP) Polymonochlorotrifluoroethylene (PCTFE) Amorphous polymers such as polyamides (Vespel SP21)

PTFE is the one most commonly used because of its high resistance to ignition on impact, its high autoignition temperature, as well as low heat of combustion. However, it has the annoying tendency to creep, and is often replaced by other plastics that are slightly less compatible with oxygen systems. In choosing a plastic reinforced with a nonplastic such as glass, one should be aware that such a composite has a lower ignition resistance.

12.15.18 LUBRICANTS14 Liquid lubricants used for oxygen systems are mainly fluorinated or halogenated chlorotrifluoroethylene (CTFE) fluids such as halocarbon thickened with higher molecular weight CTFEs. Greases are derived from perfluoroalkylethers that are thickened with PTFE or FEP, which are short-chain polymers. CTFE fluids thickened with silica have been found to allow moisture to penetrate the oil film resulting in serious corrosion. It should be avoided.

12.15.19 CAUTION Great care must be exercised in the selection of nonmetallic materials because of their lower resistance to ignition and the small size of particles in lubricant dispersions. Sources of ignition


involving them are not always obvious. For example, organic polymers are particularly susceptible to ignition in conditions of adiabatic compression in oxygen-rich environments when caused by a sudden valve actuation. Many incidents have been traced to such phenomena. Oxygen systems, especially those under high pressure, must be pressurized slowly.

REFERENCES 1. Kerry, F.G. 1956. Safe design and operation of low temperature air separation plants. Chem Eng Prog 52 (11): 44. 2. Rotzler, R.W., S.A. Glass, W.E. Gordon, and W.R. Heslop. 1960. Oxygen plant reboiler explosion. Chem Eng Prog 56 (6): 70. 3. Sefton, V.B. 1961. Contaminants, Source Physical and Chemical Properties, Basis of Limits. Presented at Compressed Gas Association, New York (October 5, 1962) Safety Symposium, January 1961. 4. Kerry, F.G. 1991. Front-ends for air separation plant—The cold facts. Chem Eng Prog (August 1991): 48–54. 5. Grenier, M., J.Y. Lehman, P. Petit, and D. Eyre. 1984. Hydrocarbon control in air separation units, Cryogenic processes and equipment, Book No. G00283. New York: American Society of Mechanical Engineers. 6. Reyhing, J. 1983. Removing hydrocarbons from the process air of air separation plants using molecular sieve adsorbers, Linde Reports on Science and Technology, No. 36, p. 14. 7. Lassmann, E.P. 1996. Enrichment of hydrocarbons in pressurized liquid oxygen evaporators. Linde Reports on Science and Technology, No. 57. 8. Safe Use of Brazed Aluminum Heat Exchangers for Producing Pressurized Oxygen, Compressed Gas Association, Bulletin CGA G-4.9, 1996. 9. Jost, W. 1946. Explosion and combustion processes in gases, 299. New York: McGraw-Hill. 10. Hugill, F.T. 1969. Safe Liquid Oxygen Disposal, Air Separation Plant Safety Symposium, Compressed Gas Association, April 23, 1969, pp. 32–36. 11. de Jessey, L. 1969. Safety Aspects Regarding Mass Distribution Methods by Pipeline, Compressed Gas Association, Safety Symposium Air Separation Plants, pp. 37–45. 12. Pears, C.D., R. Monroe, J.E. Wren III, and C.E. Bates. 1978. Final Report to the Department of Energy on Structural Materials Evaluation for Oxygen Centrifugal Compressors, Project 3528, Report XXXVI, Southern Research Institute, Birmingham, AL, 1978. 13. Monroe, R.W., C.E. Bates, and C.D. Pears. 1983. Metal combustion in high-pressure flowing oxygen. In Flammability and sensitivity of materials in enriched atmospheres, ASTM STP 812, B.L. Werley, ed., 126–149. Philadelphia: American Society for Testing and Materials. 14. Beeson, H.D., W.F. Stewart, and S.S. Woods. Safe Use of Oxygen and Oxygen Systems, Sponsored by ASTM, Committee G-4. 15. Dunbobbin, B.R., J.G. Hansel, and R.L. Werley. 1991. Oxygen compatibility off high surface area materials. In Flammability and sensitivity of materials in oxygen enriched atmospheres, fifth volume, 338–353ASTM STP 1111. Philadelphia: American Society for Testing and Materials. 16. Zawierucha, R., J.F. Million, S.L. Cooper, K. McLLroy, and J.R. Martin. 1993. Compatibility of aluminium packing with oxygen environments under simulated operating conditions. In Flammability and sensitivity of materials in oxygen enriched atmospheres, sixth volume, 255–275, ASTM STP 1197. Philadelphia: American Society for Testing and Materials. 17. Barthelemy, H.M. 1993. Compatibility of aluminium packings with oxygen-test results under simulated operating conditions. In Flammability and sensitivity of materials in oxygen enriched atmospheres, sixth volume, 276–290, ASTM STP 1197. Philadelphia: American Society for Testing and Materials. 18. Kirzinger, A., K. Baur, and E. Lassmann. 1993. The behavior of oil films on structured packing under cryogenic conditions. In Flammability and sensitivity of materials in oxygen enriched atmospheres, sixth volume, 291–299, ASTM STP 1197. Philadelphia: American Society for Testing and Materials.


19. Werley, B.L., H. Barthelemy, R. Gates, J.W. Slusser, K.B. Wilson, and R. Zawierucha. 1993. A critical review of flammability data for aluminum. In Flammability and sensitivity of materials in oxygen enriched atmospheres, sixth volume, 300–345, ASTM STP 1197. Philadelphia: American Society for Testing and Materials.

FOR FURTHER READING Lewis, B.W., and G. von Elbe. 1987. Combustion, flames and explosions of gases. New York: Academic Press. Safe use of aluminum structured packing for oxygen distillation, CGA G-4.8-1993, Compressed Gas Association Inc.


13

Cleaning for Oxygen Systems

13.1 OVERVIEW The question of ‘‘How clean is clean?’’ has been around for years, but it cannot be answered by any one engineer or any one set of conditions. After a thorough review of the available literature, specifications, and practices of various industrial gas companies as well as users, the following recommendation is offered to the project engineer: ‘‘Choose the correct cleaning procedure to fit the existing conditions, and specify in detail the necessary steps to be used by the contractor doing the work.’’ This chapter, however, should be used only for information or guidance. It should not be confused with, or take the place of any federal, state, or municipal specifications, or insurance requirements or national safety codes. Abnormal or unusual circumstances may require different or additional procedures. All engineers faced with the construction of a proposed oxygen system should do everything possible to achieve a high degree of cleanliness prior to the flow of oxygen through the system. Some years ago, during a session of the Compressed Gas Association on oxygen safety, a speaker berated the term ‘‘hospital clean,’’ so popular at the time because there were many circumstances that did not merit the procedure. Although many in the audience agreed with the speaker, there was one who stood up and defined the term in another way, using the difference between a professional golfer and a weekend amateur. The professional when faced with a 20 ft putt, takes an enormous time to study the situation in order to hole the ball with only one putt. He also knows that the odds of accomplishing it are high, but he also knows that even if he misses the putt, the ball will remain close enough to the hole for an easy tap in and a par. The weekend amateur, on the other hand, accepts the fact that he cannot make the putt, so he simply tries to get close, with the result that he will end up with a bogie or even a double bogie. In short, every effort should be made to achieve the highest possible degree of cleanliness, so that even if we do not obtain a ‘‘hospital clean’’ state, we can go to sleep knowing that the oxygen system is safe from any dangerous contaminants.

13.2 GENERAL CONSIDERATIONS Field cleaning and degreasing of piping and ancillary equipment for oxygen service should not be considered as a separate project isolated from other considerations involved in the overall planning of the entire project. Field cleaning of the piping must be programed to fit within the schedule of erection and construction of the pipeline in order to make sure that the cleaned pipe sections and the equipment are installed and erected immediately following cleaning. Unless a proper schedule is followed, cleaned piping and fittings may run the risk of being recontaminated by lying around with improper storage and handling, thereby defeating the very purpose of cleaning objectives.


13.3 CLEANING REQUIREMENTS FOR OXYGEN SYSTEMS As it is well known, the main initiator of a combustion reaction in an oxygen distribution system is ‘‘ignition energy.’’ As a rule, it is provided by a foreign material or particles that are not part of the designed system, but are found in the system either because of improper cleaning, careless construction, or incorrect testing of the oxygen distribution or transmission system. The people who write specifications for pipeline distribution or transmission systems very often specify material and equipment going into an oxygen system with the words ‘‘clean for oxygen service,’’ and leave the responsibility for the standard of cleanliness up to the supplier of the material or to the construction contractor. Unfortunately, they are very often unaware of what oxygen service truly means. The cleaning standards for an oxygen system should be as follows: (a) free from all moisture, and (b) free of all particulate matter such as rust, metallic scale, shop dirt, earth, filings, weld spatter, chips, powdered metal oxides, excelsior, packing material, labels, or any other foreign matter; free of all organic materials such as oil, crayon marks, paint, ink, hydrocarbons, greases, thread lubricants, varnishes, etc. All the above mentioned foreign matter including subsequent rust resulting from the left over moisture will cause ignition with oxygen under pressure, creating a fire or even an explosion. It is not enough, therefore, to merely specify ‘‘clean for oxygen service.’’ It is also necessary to set quantitative standards, which should define tolerances and limits for inspection permitted in the inspection methods.

13.3.1 INSPECTION STANDARDS

FOR

FIXED SURFACES

The following standards should apply to piping, fittings, and vessels, etc.: 1. Visual examination of the surface under a strong white light should indicate no moisture, slag, scale, organic material, foreign matter, or corrosion products. 2. Particulate matter should not be larger than 1000 mm and should amount to fewer than 100 particles (500–1000 mm=m2). 3. Black light (ultraviolet 3600–3800 AU) examination should indicate no hydrocarbon fluorescence. 4. A wipe test should show no appreciable discoloration and no evidence of oily residue. 5. If solvents are used after the distribution system is installed, the system is considered clean only when the final solvent flowing through the pipe leaves a maximum residue on evaporation of less than 100 ppm by weight in comparison with a fresh new liquid. 6. If chemical reagents such as acids or alkalis are used on a pipeline already installed, then the spent liquid of the last pass must undergo a test to ensure that it contains less than 100 ppm by weight of residue in comparison to a fresh liquid. 7. If sand blasting is used directly on an already installed pipeline with no further treatment, there is no ready quantitative standard of cleanliness. The contractor, therefore, must present a test or guarantee of cleanliness acceptable to the user and equivalent to the standards stated above.

13.3.2 INSPECTION STANDARDS

FOR

MOVABLE PARTS

The following standards should apply to valves, pumps, compressors, etc.: 1. Visual examination of the direct surface under a strong white light should indicate no moisture, slag, scale, organic material, or other foreign matter, and no corrosion products.


2. Particulate contamination should be limited to that suggested in Article 11.4.2.2 and 11.4.3.7, ASTM Bulletin G 93–96 of Committee G-4.1 3. Black light examination should indicate no hydrocarbon fluorescence. Any isolated particles of lint must be removed with dry oil-free nitrogen. 4. A wipe test should show no appreciable discoloration and no evidence of oily residue.

13.3.3 CLEANING PROCEDURES Cleaning procedures are only as effective as the quality of the workmanship and the degree of inspection. These aspects, therefore, must be given as serious consideration as the procedure itself. All materials and equipment purchased ‘‘cleaned for oxygen service’’ should be inspected at the fabricator before packing to assure that initial cleaning and packing are in accordance with the purchaser’s standards, and then reinspected when received at the field assuring that no contamination occurred in transit. If there is even the slightest doubt as to the cleanliness of an item the latter must be dismantled and re-cleaned. Following any cleaning procedures and prior to any field work, all ‘‘cleaned’’ items must be protected by clean coverings like plastic material and stored in clean areas. Their handling must be closely supervised to ensure that the clean surfaces are not contaminated by grease, dirt, dust, or any other contaminant including workmen’s gloves and even bare hands, which always contain natural skin-oils.

13.4 EQUIPMENT OTHER THAN PIPING Ancillary equipment (valves, fittings, orifice meters, instruments, etc.) is manufactured, assembled, tested, and packed outside the customer’s premises. It is not enough, therefore, to order these items as ‘‘cleaned for oxygen service.’’ The supplier should submit to the purchaser not only a certificate of the materials to be used for oxygen service, but also the assembly, testing, cleaning, inspection, and packing procedures used at his plant before shipment. Furthermore, the purchaser should have the right to impose his own specifications for the various procedures, if he is not satisfied with those outlined by the supplier. Once the final procedures are agreed to, the purchaser should either send his own inspector to the supplier’s shop to ensure that the established procedures are followed or wait until the equipment is received, and then unpack, disassemble, re-clean, reassemble, and repack the items under his own inspection procedures.

13.4.1 CLEANING PROCEDURES: GENERAL 1. All disassembled parts should be given a preliminary cleaning to remove chips, shop dirt, oil, etc. 2. Valve bodies should be given a hydrostatic test with clean potable water. 3. All metallic parts should be immersed in a suitable solvent such as methylene chloride perchloroethylene or acetone for a minimum of 2 h. 4. All nonmetallic parts should be wiped with a clean lint-free cloth slightly moistened with a suitable solvent referred to previously. 5. Once cleaned, all parts should be thoroughly dried with dry oil-free nitrogen, or dry oilfree air, and examined with a black light as outlined under ‘‘Testing and Inspection Procedures’’ to meet the necessary standards. 6. Metallic component parts should be reassembled and again immersed in a fresh solution of the solvent and rechecked under ‘‘black light.’’


7. Finally, the cleaned assembly should again be dried thoroughly using dry oil-free nitrogen or dry oil-free air. 8. Valves or other equipment should be assembled dry or with a lubricant that has been proven acceptable for oxygen service.

13.5 CLEANING PROCEDURES FOR CARBON STEEL PIPING There is no single standard specification or procedure for cleaning piping or equipment for oxygen service because of the variety of factors involved in a project. These factors include, but are not limited to, the following: . . . .

. . .

Geographical locality. Local weather conditions due to seasons or latitudes. Condition of the piping when received at the site. Local working conditions at the site with regard to available handling equipment and skilled labor. Quality of supervision and inspection. Availability and choice of commercial cleaning agents. Finally, the magnitude of the project in terms of pipe diameters, length of the system, and number of branch lines may help decide whether the piping should be totally pre-cleaned before or after construction, or a combination of both.

13.5.1 DEFINITION

AND

RECOGNITION OF CONTAMINANTS

The following information will be found useful in choosing the right method and procedure for cleaning carbon steel piping: Tight mill-scale: It is gray in color and cannot be removed manually with a putty knife. It can be removed by pickling or wire brushing. Loose mill-scale: It has a visible lamination, usually rust colored, and can be removed manually with a putty knife. It can also be removed by impact tools, pickling, or sand blasting. Tight heat furnace scale: It cannot be removed by peening. Use pickling or wire brushing. Loose heat furnace scale: It is easily removed when rapped. Use pickling or sand blasting. Dust or earth: This is readily visible. Use an alkali solution or detergent, then wash thoroughly and dry. Rust: This is a homogeneous coating of adhesive reddish ferric oxide that is difficult to remove. Use pickling or sand blasting. Slight rust: A homogeneous, very light bluish coat of powdery ferric oxide. It can be removed by acid pickling and either sand or shot blasting. Corrosion products: Products other than rust resulting from chemical attack on the base metal. They include pitted rust, and built up rust, which can be removed by acid pickling, and sand or shot blasting. Hydrocarbons: Grease, oil, wax, crayon marks, paint, ink, varnishes, thread lubricants, protective coatings, and general organic material. They can be removed by a caustic wash, and sand or shot blasting. Metal particles: This includes any metallic particulate matter such as weld spatter, residue from shot blasting, reaming, machining, or grinding machines. It can be removed by sand blasting, wire brushing, or acid pickling.


Moisture: If water has to be used for washing or rinsing purposes, immediately afterward it should be removed completely, and the cleaned pipe surfaces should be dried completely using dry oil-free nitrogen or dry oil-free air. Fluid oil-films2: According to the ASTM Report G 93–96, Article 10.1.4, more attention has been paid to fluid oil-films than solid particulates, because oil films are more easily ignited, and have a tendency to migrate as well as vaporize. When an oil film is removed completely, precious few solid particulates will remain behind, but these are generally inert.

13.6 CLEANING PROCEDURES AVAILABLE 13.6.1 BLAST CLEANING: GENERAL After a review of the various contaminants and their removal previously enumerated, it becomes obvious that the logical method is to use hard silica sand or shot blasting as the basic cleanser, and possibly the only one. In actual fact, sand blasting is still the most recommended basic step before any other cleaning agents are considered or used. If sand blasting can be carried out under strict supervision, and the cleaning of the surface can get down to bare metal, and if the cleaning is followed immediately by erection, then no further cleaning procedures are necessary: 1. The interior of all carbon steel piping and fittings should be sand-blasted using hard silica sand, or grit. Blasting should be done evenly and carefully to remove all contaminants. 2. Special machines are available with radial or rotating nozzles that are held in position at the center axis of the pipe by adjustable templates and drawn manually through the length of the pipe. 3. Before sand blasting is started, the piping and fittings should be already prepared to include cut-outs for nozzles, branch lines, and prefabricated into sections that can be readily handled in the field. 4. Once sand-blasted, no further work should be done on a pipe or fittings using an oxyflame cutting torch. A mechanical drill or saw should be used. 5. During the construction period, sand blasting should also be carried out on spool pieces that take the place of valves, orifice plates, or meters. After blasting, the piping and fittings should be blown out using dry oil-free nitrogen or dry oil-free air, and should have their ends sealed or plugged before construction begins.

13.6.2 SAND BLASTING

IN

PLACE (SANDJET)3

Aside from sand blasting individual pieces of piping, fittings, and accessories prior to construction, it is possible to sand-blast long lengths of single diameter piping in place, especially where few side branch lines exist. This procedure is comparatively quick, but must be carried out by an experienced group that specializes in this method of cleaning. This method involves the use of a controlled flow of hard silica sand or grit that is propelled through the length of the pipe by a propellant gas stream, usually nitrogen, under pressure and at high velocity. The sequence is continued until the entire length of the piping is scoured and cleaned down to the bare metal base. This method is used just prior to making the closing welds of long lengths of piping. The branch lines, if any, are treated as individual sections, and are cleaned separately.


If the distribution system has already been totally and completely installed, then the piping must be disassembled or cut in long straight sections and rewelded after cleaning. Under these conditions all valving, flow meters, flow nozzles, and orifice plates must not be erected during the cleaning operations, and should be removed and replaced by spool pieces. As noted, branch lines, if connected, must be disconnected or cut close enough to the nozzle, which, after the cleaning operations, should be inspected and cleaned manually if necessary. 13.6.2.1

Equipment

The equipment involves a tank containing the sand or grit; a truck containing the propellant gas, usually nitrogen in liquid form and subsequently vaporized, and the controlling manifold. 13.6.2.2

Procedure

1. If there is a significant amount of moisture in the piping, a quantity of clay propelled by the gas under pressure is sent through the piping. 2. After the pipe section is dried by the clay, a controlled quantity of abrasive material is aspirated into the piping by the propellant gas. The quantity of abrasive material injected into the pipe depends on the diameter, length, and the interior condition of the piping. 3. This cleaning procedure is a batch-type process, and should be repeated until the desired standard of cleanliness is achieved. 4. If any mill varnish exists in any of the piping, it should be removed with an appropriate solvent before the sand blasting begins. 5. Minimum velocity of the blow should be at least double the maximum oxygen velocity to be encountered in actual service. The duration of the blow should be sufficient to pass a volume 10 times that of the piping. 6. All nozzles should be inspected and washed with a solvent, and any inner surfaces not having a gray–white color should be cleaned with a stiff wire brush. 7. Finally, the pipeline should be blown out with dry oil-free nitrogen to remove all residual grit, dust, and other loose particles. On long transmission systems a brush type pig-dust ball may be used for a final clean-up. 8. All components removed or separated from the distribution or transmission system should be reinstalled, observing all the prescribed precautions necessary for oxygen service. 9. The completed pipeline should be filled with dry oil-free nitrogen, and maintained under pressure until the distribution system is placed in actual oxygen service. 13.6.2.3

Inspection and Control

1. Although the foregoing cleaning method is simple, quick, and straightforward, its control depends on the expertise and close supervision of the operating team. There is no direct way of examining the interior surface of the long pipe section. However, one may use small sample sections of the steel plate, tack-welded at the inlet and outlet ends of the pipe section, and at some intermediate points where spool pieces may be located. 2. These samples may be covered with contaminants such as oil, scale, varnish, weld slag, etc. After sand blasting, these test samples can be removed and studied for cleanliness. 3. Wipings may be taken at the sample points and examined under a black light for fluorescence.

13.6.3 SECONDARY CLEANING PROCEDURES Unless sand blasting is carried out under ideal and closely controlled conditions, and if the piping as well as the fittings are reasonably clean and free from mill-scale, rust, mud, etc., and if the piping after being sand-blasted can be handled and erected without any further contamination, there would be no necessity for any other cleansing treatment. Unfortunately, this is not always the case in actual field conditions where piping and fittings have to be handled by people as well as machines under adverse conditions of dust, rain, mud, and oil (the latter from hoisting equipment and pipefitters’ gloves.) Furthermore, a proper cleaning and construction program may not always be possible, especially if deliveries of piping and equipment do not follow a predetermined schedule. This necessitates further cleaning using chemical means, which may be termed secondary cleaning procedures. Small-diameter piping less than 100 mm that cannot be cleaned by sand blasting or other mechanical means must undergo cleaning by chemical agents. Chemical cleaning may also apply to distribution systems that have a great number of branch lines, which means a great deal of handling and many fittings. Depending on the magnitude of the project, the length and the diameter of the piping distribution system, the time interval between sand blasting and erection, and finally, the availability of specialized outside services for cleaning, one has two options: (a) to clean the piping in sections just prior to erection; or (b) to erect the pipe right after sand blasting, then further clean the pipe in place. The latter method is more convenient of the two, but cannot always be applied.

13.6.4 PRE-CLEANING

BEFORE

ERECTION (WITH CLEANING REAGENTS)

This procedure generally involves transmission piping of large diameters (over 250 mm), long runs underground, and important branch lines. This type of system is difficult to clean in place unless unusual procedures are taken. This also applies to small piping that cannot be cleaned internally by sand blasting. Before this treatment is started, the following preparations must be made: 1. All piping must be cut in suitable lengths—about 6 m. 2. All perforating and bevelling requiring welding and cutting out should be carried out in advance. 3. As much prefabrication as possible should be carried out in advance, and within limits of handling during the treatment. 4. Sand-blast the interior surfaces of the piping and fittings including spool pieces until the clean gray–white surfaces of the carbon steel base metal can be seen. If this objective is difficult to obtain, the following procedure should be taken. A bath should be prepared in the form of a tank whose size can handle piping and fittings already prepared, and a reasonable number at a time: 1. The first treatment should involve an alkaline solution such as sodium hydroxide (NaOH) or sodium carbonate (Na2CO3) for degreasing purposes. The temperature of the bath should be around 808C, and the treatment held for around 40 min. The time and temperature will vary according to the cleanliness of the piping and the commercial grade of the reagent used. 2. Rinse with clean potable water until a neutral pH value is obtained. 3. Immerse in a hydrochloric acid solution (HCl) or equivalent, using a 15% solution held for 2 h. Here again, the time and temperature depend on the commercial grade of the


4. 5. 6. 7. 8.

9.

10. 11. 12.

reagent used. This treatment is for the removal of rust, scale, iron oxide particles, corrosion products, etc. Rinse in clean potable water until a neutral pH value is obtained. Many commercial reagents, acid or alkaline, contain inhibitors to prevent or slow further rusting. The use of inhibitors is strongly recommended. Wash with ferrophosphate solution for 30 min with a bath temperature of around 808C. Passivate by using a sodium dichromate solution (Na2Cr2O72H2O) at 408C for 5 min. No further rinsing with water is necessary. If the commercial reagents do contain inhibitors or if these are added by the contractor, it would be preferable to use a weak alkaline solution (20%–25% of that used in step 1) with an inhibitor as a last operation after step 4 with no further treatment. This operation will slow down rusting and more importantly, passivate, and stop corrosion from any acid remaining from steps 3 and 4. Visual inspection under a strong white light should reveal no foreign particulates fewer than 10 (between 0.5 and 1.0 mm) per square meter. Under an ultraviolet lamp there should be no fluorescence from oil, grease, or organic matter. For small diameter piping, use a wipe test by rubbing the interior walls of the pipe with specially treated laboratory filter paper, and examine the latter under an ultraviolet light. See point 2 in Section 13.10.2. If not used immediately, piping, fittings, and accessories should have their exposed ends sealed hermetically, as much as possible, using plastic covers or plugs. During erection and construction, the clean piping must be handled by operators with clean gloves. Use hoists or cranes that must not drop any oil or grease. Finally, the entire oxygen distribution system, fully erected and with spool pieces still in place, should be completely dried using dry oil-free nitrogen or dry oil-free air.

13.6.4.1

Precautions

Special precautions should be taken during the cleaning procedure to protect operators against direct contact with the various acid and alkaline solutions. Proper clothing and gloves should be worn, and there should be strict supervision during cleaning operations. Precautions should be taken in the handling and mixing of the chemical reagents, and in the disposal of spent solutions.

13.6.5 CLEANING

AFTER

ERECTION: GENERAL

If at all possible, it is more convenient and more economical to do the secondary cleaning after erection, especially for internal distribution systems. First sand-blast the piping, fittings, and spool pieces, then proceed with the erection and construction of the oxygen distribution system. This alternative, however, depends entirely on the availability of outside contractors specializing in the cleaning of pipeline distribution systems. If this alternative is decided upon, the owner of the distribution system has several options depending on the local conditions and general economics.

13.6.6 ALTERNATIVE A—WITH SOLVENTS In 19874, an international agreement was achieved in Montreal, Quebec, called the Montreal Protocol, which called for a sharp reduction in the worldwide production of atmospheric ozone-depleting substances. This was followed up with the US EPA Clean Air Act of 1990, which also called for the phase-out of the production of these substances. Needless to


say, the Class I ozone depleters such as 1.1.1-trichloroethane, which has been widely used as an oxygen system cleaning solvent, have been phased out. The reader is recommended to get in touch with the US Significant New Alternatives Policy (SNAP) Program for alternatives acceptable for cleaning oxygen systems. Information may be obtained from SNAP coordinator, Stratospheric Protection Division 62021 US Environmental Protection Agency: 1. Proceed with preparation of the piping and fittings followed by primary cleaning with sand blasting as previously described. 2. To facilitate the inspection of piping components such as elbows, tees, and fittings with mill painted surfaces, it is best to carry out sand or grit blasting prior to prefabrication of piping sections in order to visually examine the internals following the operation. 3. Following sand or grit blasting, all internal surfaces should have a uniform gray–white color. All residue of grit should be removed by blowing with clean, dry, oil-free nitrogen or air. Surfaces should be inspected to make sure that no grit has penetrated into the surface. 4. If not used immediately, the piping, fittings, and accessories should have their exposed ends sealed hermetically as much as possible using plastic covers or plugs. 5. During the erection and construction, the clean piping and accessories must be handled by operators with clean gloves and by using hoists or cranes that do not drip any oil or grease. 6. The first pass of all welds should be carried out using a tungsten-arc inert gas (TIG) procedure and a small argon purge. After the root weld is complete, it should be washed thoroughly with a solvent and dried. Care should be taken that the piping remains clean during prefabrication and construction. If it is exposed to the atmosphere, dirt, water, or other contaminants, further grit blasting prior to cleaning should be carried out. 7. Once the distribution system or section is ready for cleaning, the cleaning contractor should use a cleaning solvent that is environmentally acceptable. One has to keep in mind that chlorinated solvents are being replaced by aqueous or semiaqueous detergents or emulsions often in conjunction with deionized water as part of the process. If possible, these solvents should include a corrosion inhibitor. 8. Prior to cleaning, the piping system should be blown by dry oil-free nitrogen or air to make sure that there is no water or moisture in the system. Certain metals are subject to corrosion once the inhibitor in the solvent is destroyed by moisture. 9. The solvent should be circulated through all piping sections, branch lines, and instrument connections. Flow rates should be sufficient to assure full flow in all lines. For large systems where the pumping rate would be excessive, the operation can be intermittent, provided that the minimum pumping time is 2 min. This should be continued until the entire piping section or distribution system is thoroughly flushed and cleaned. 10. The piping system may be considered clean when 1 L of a representative sample of the spent solvent is collected and analyzed for impurities. The analysis of the spent solvent should reveal less than 100 ppm (by weight) of total impurities. If the analysis reveals an impurity content of more than 100 ppm (by weight), the washing should be continued until the aforementioned standard of cleanliness is achieved. See Section 13.10. 11. Once the system is clean, the solvent should be drained completely, and the piping system thoroughly blown out and dried with hot dry oil-free nitrogen or air, preferably


the former. Chlorinated solvents are flammable in oxygen service and are very toxic. Purging should continue until all exits indicate no odor. 12. Methylene chloride and perchloroethylene are not ozone depleting, but highly toxic.Trichloroethylene is suspected of being a carcinogen. These solvents also react with aluminum and its alloys to produce hydrochloric acid vapor, which is toxic and corrosive. Moreover, these solvents should only be used for small materials, and the cleaning should be carried out in a well-ventilated room. For very small items, acetone may be used. 13. Safety procedures in handling these solvents should be in accordance with the manufacturer’s recommendations. During the cleaning operations, the inhibitor content of the solvent should be checked in accordance with the manufacturer’s standards.

13.6.7 ALTERNATIVE B—CLEANING AGENTS This alternative is equivalent to alternative A, but in the absence of chlorinated reagents, one may use chemical reagents both alkaline and acid. There are many such commercial products available on the market. Commercial cleaning agents suitable for cleaning piping systems are made by Oakite Products, Inc. (USA) and the Nalco Chemical Company and others. These chemical solutions are pumped through the distribution system in the same manner as the chlorinated solvents, but the procedure is different depending on the circumstances. The cleaning procedure to be carried out in this alternative B is as follows: 1. The first six steps including welding are exactly the same as in alternative A. Then contractor should pump the alkali solution first for degreasing purposes. 2. Rinse with clean potable water until a neutral pH value is obtained. 3. Follow with an acid solution to dissolve iron oxide, corrosion products, or particulates. 4. Rinse with clean potable water until a neutral pH value is obtained. 5. Passivate with a dilute alkali solution containing an inhibitor to retard rusting. 6. Depending on the type and strength of the final alkali solution, rinsing with potable water may or may not be necessary. Instructions should be obtained from the manufacturer of the chemicals. 7. During the operation, the solutions should be examined for contamination. The piping system or section may be considered clean when the representative samples or the spent reagents are collected and analyzed for impurities. As indicated previously, analysis of the spent solutions should reveal less than 100 ppm (by weight) of total impurities. If analysis reveals an impurity content of more than 100 ppm (by weight), washing should be continued until the aforementioned standard of cleanliness is achieved. 8. Once the system or section is clean, the chemicals should be removed and the entire system purged with dry nitrogen until the system is dry, and reaches a dew point of 408C. Dew points should be taken at every outlet until the stated dew point is reached. 9. Safety procedures in handling the chemical solutions should follow those indicated in the previous methods.

13.6.8 ALTERNATIVE C—WITH MOVABLE PISTONS This alternative is equivalent to alternative B. One may use the chemical reagents mentioned previously in conjunction with movable pistons or ‘‘pigs’’ as they are better known. The chemical solutions are placed in the piping system and sealed between the pistons. Then the whole assembly is pushed back and forth through the system by means of water pressure or


nitrogen gas under pressure, depending on the individual circumstances. The pigs are made of various materials, but polyethylene seems to be preferred. Care should be observed when using polyethylene pigs that no particles are caught and left in the weld joints or fittings, because polyethylene has a very low acceptability limit with oxygen.

13.6.9 CLEANING STAINLESS STEEL AND NONFERROUS METALS SUCH AS COPPER, ASSOCIATED FITTINGS, PARTS, AND FABRICATIONS N.B.: Do not use any cleaning agents containing hydrochloric acid or hydrochlorofluorocarbon agents (HCFCs). Presently there are many water-based cleaning agents available on the market such as Oakite (Oakite Products, Inc.), Blue Gold Industrial Cleaner (Modern Chemical Co.), and others. These products have been replacing CFCs and HCFCs that are either highly toxic or ozone depleting and are therefore not acceptable: 1. Prepare a solution of Oakite #77, which is an alkali material with a chelating action, for the removal of grease, smut, as well as light rust and paint. The solution should have a concentration of 45–60 g=L of water maintained at a temperature of 808C–908C. Allow work to soak in the solution for a period of time required for the complete removal of soils and contaminants. 2. Upon attaining the required cleanliness of parts or fabrications, remove from cleaning solution and immediately immerse in a potable water rinse at ambient temperature using violent air agitation. 3. If necessary, the following method should be used to remove stains, residue from initial cleaning, or similar discoloration. Immerse the work in a 25% by volume solution of Oakite #31 (an acidic detergent material) maintained at a temperature of 708C–808C for a period of time required for complete removal of all stains and discoloration. 4. Following the removal of stains or discolorations, the work should be thoroughly rinsed in a potable water rinse at ambient temperature using violent air agitation.

13.6.10 CLEANING ALUMINUM PIPING, FITTINGS, PARTS,

AND

FABRICATIONS

1. Prepare a solution of Oakite Aluminum Cleaner #164 at 45–75 g=L of water. Heat the solution to a temperature of 808C–908C. Allow parts to soak in the solution for the period of time required for the complete removal of soils and contaminants. Maintain mild agitation of solution by means of compressed air introduced into solution. 2. Upon attaining required cleanliness of parts or fabrications, remove from cleaning solution and immerse in potable water rinse at ambient temperature using violent air agitation. When rinsing the tubing, piping, and similarly fabricated parts, an internal pressure rinse is recommended, i.e., a high-velocity hose stream. 3. If necessary, the following method should be used to remove oxides and residue from initial cleaning. Immerse work in a solution of Oakite #34, 75 g=L at ambient temperatures (but not less than 158C) for a period of 5 min. 4. Following the above process, the work should be thoroughly rinsed in a potable water rinse at ambient temperature using violent air agitation.

13.6.11 ALTERNATE METHODS OF CLEANING STAINLESS STEEL PIPE, ALUMINUM PIPE, COPPER TUBING, AND THEIR FITTINGS For small pieces of equipment and material, such as tubing, valves (suitably dismantled), orifice plates, that may be slightly soiled due to handling and may require a light cleaning, a suitable solvent or an acceptable alternate substitute may be used:


1. Small tubing, fittings, etc., may be quickly immersed in a solvent=cleaner or swabbed until all traces of dirt and contamination are removed. 2. Products recommended for this purpose in order of preference are: . . . .

Blue Gold Industrial Cleaner. Completely safe and reusable. Acetone Methylene chloride (highly toxic) Perchloroethylene (highly toxic)

3. It should be kept in mind that the chlorinated solvents can be harmful to the metals as well as to personnel, and should be used sparingly. Furthermore, once cleaned, the material should be dried very quickly and thoroughly. Proper ventilation should always be used throughout the operation.

13.6.12 OXYGEN COMPRESSORS Generally, all oxygen compressors are cleaned thoroughly at the factory before packing and shipping. Unfortunately, they often arrive at the site and are erected long before start-up occurs. It is good practice, therefore, to dismantle and re-clean all parts in contact with oxygen prior to start-up: 1. The cleaning procedures and reagents to use are identical to those recommended in the previous chapters for piping and equipment. 2. If large pieces are to be cleaned by immersion, care should be taken not to expose them to the atmosphere before the passivation step is carried out. This procedure can be carried out by using a large tank with two compartments. In this manner, the piece or pieces to be cleaned are continuously submerged while the reagents are being prepared and fed from the separate compartment: . . .

Step 1—pickling Step 2—water flush Step 3—passivation

3. For small pieces that have hidden, sharp corners or welds such as impellers, it is strongly recommended to use a solvent that has no hydrogen radicals or use the Blue Gold Industrial Cleaner. Solvents should be used sparingly if possible, followed by thorough drying. 4. Normally, compressor pieces such as diaphragms have small threaded holes for the insertion of eyelet screws for handling purposes. Care should be taken to dry these small holes very thoroughly after the piece is cleaned. 5. If at all possible, the cleaning should be carried out by a cleaning contractor who has the necessary expertise and equipment to do a professional job. 6. In one specific instance, lint had to be removed from a huge steel-casting that was part of the compressor base. This was done expeditiously by quickly passing an oxy-acetylene torch about 30–40 cm over the surface of the casting until all lint was removed.

13.7 CLEANING AGENTS Although there is a wide variety of cleaning agents and solvents being used for cleaning oxygen pipe lines, they all follow a similar classification:


1. Cleaning agents having an alkaline base for the removal of grease, oil, and organic matter. 2. Cleaning agents having an acid base for the removal of scale, rust, and metallic particulate matter. 3. Solvents are used for the removal of grease, oil, and other organic material. 4. Detergents are used generally for the removal of earth, dirt, etc. These detergents as a rule leave a considerable residue, and if used on equipment or piping a very thorough rinsing with water should follow. 5. Inhibitors may be used alone or included in the cleaning agents to slow down rusting during the construction period. 6. Water-soluble cleaning agents such as Oakites are removed by flushing with large quantities of hot, clean water. The equipment is then dried by purging with oil-free air or nitrogen, preferably heated. When all of the outlets are warm to the touch, the equipment is visually inspected to be sure that all moisture has been removed. 7. Solvent cleaners must evaporate and be dried from all surfaces after cleaning. Again, warm, dry oil-free air or nitrogen is purged through the equipment until all outlets are warm to the touch and no trace of solvent odor can be detected at any outlet. A halogen sniffer can be used for this purpose.

13.8 PREPARATION OF CLEANING AGENTS Preparation and use of the various cleaning agents to be specified should follow a close consultation between the owner of the distribution system or contractor doing the operation and the manufacturer of the cleaning products to be used. Nevertheless, some typical preparations and procedures are offered below.

13.8.1 CAUSTIC OR ALKALINE SOLUTIONS 1. Caustic soda (NaOH): 220 g=L of water. Heated to 888C–1488C and held for 30 min 2. Sodium carbonate (Na2CO3): 150 g=L of water, plus 5 L of detergent (SAC). Heated to 808C and held for 10 min 3. Oakite #77 (Oakite Products): 45–60 g=L of water. Heated to 808C–908C and held for 30–60 min 4. Nalclean No 8900 (Nalco Chemical): 0.78–1.25 L per 10 L of water. Heated to 808C–858C 5. Blue Gold Industrial Cleaner (Modern Chemical Company): used in varying concentrations and times depending on the application

13.8.2 ACID SOLUTIONS 1. Hydrochloric acid (HCl): 15% solution, i.e., 500 L (30% HCl) plus 500 L of water plus 1 L of inhibitor. Heated to 808C and held for 30 min; or, 68 L of HCl (208 Baume´) plus 364 L of water. Heated to 888C–1488C and held for 30 min. 2. Oakite # 31, which is a phosphoric acid (Oakite Products): 5%–50% by volume depending on the extent of cleaning; 5%–10% recommended. Heated to 708C–808C and held for 30 min. 3. Nalclean 66 (Nalco Chemical): 18–50 kg per 1000 L of water. Ambient temperature is acceptable. Cleaning time 4–12 h depending on contamination. Phosphoric acid can be used for all metals. It can remove oxides, light rust, light soils, and welding fluxes.


13.8.3 AGENTS FOR STAINLESS STEEL, COPPER, AND ALUMINUM 1. Oakite stripper additive—neutral Application: Additive to any regular alkaline Oakite solution. Applicable for metals initially determined suitable for the Oakite solution to which it is added. Purpose: Reduces cleaning time considerably, and improves the cleaning ability of an Oakite solution. Used to remove acrylic and epoxy lacquers. Mixing proportions: Add 2%–10% by volume of Oakite stripper additive to any alkaline Oakite solution. Operating temperature: Do not heat in excess of 828C. Precautions: Never add stripper additive to an acid oakite. An active reaction will occur. Do not allow contact with skin or eyes. 2. Oakite #77—Strong alkaline Application: All metals, except aluminum. Purpose: Remove heavy deposits of soils and oils. Remove light deposits of rust and scale. Mixing proportions: 45–60 g=L of water. Operating temperature: 708C–828C. Precautions: Do not use at temperatures below 608C or when scum forms on solution. Do not allow contact with skin and eyes. 3. Oakite #31—Contains acid Application: All metals except aluminum. Purpose: Remove oxides, rust, oils and soils. Mixing proportions: 5% by volume Oakite #31 in water. Operating temperatures: 708C–828C. Precautions: Do not allow contact with skin or eyes. 4. Oakite #161—Mild alkaline powder material. Inhibits attack on aluminum. Contains antifoam agents Application: Cleaning aluminum by pressure spray or circulation method. Purpose: Remove all types of soils other than stains and oxides. Mixing proportions: 22–30 g=L of water. Equipment: Material safe on all metals. Ordinary steel solution holding tanks for spray equipment are suitable. 5. Oakite aluminum cleaner #164—Mild powder material. Inhibits against attack on aluminum. Contains anti-foaming agents Application: Cleaning aluminum. Purpose: To remove all types of soils other than stains and oxides. Mixing proportions: 45–60 g=L of water. Operating temperature: 828C–938C. Equipment: Safe on all metals. Ordinary steel tanks are suitable. 6. Oakite #34—Acid power–Type material Application: For treating aluminum, brass, bronze, and copper.


Purpose: Remove oxides and brighten metal surfaces by immersion or circulation. Mixing proportions: 45 g=L of water by spray or circulation, 75 g=L by tank immersion.

13.8.4 SOLVENTS The solvents used for cleaning equipment for oxygen service are generally of the chlorinated hydrocarbon type. Most are not compatible for oxygen service as explained in Section 13.6.6, and are limited to those indicated in Section 13.6.6, step 12. As noted previously, acetone may be used for small items. It is important to note that whenever these solvents are used, the piping or equipment should be purged with dry nitrogen until all traces of the specific solvent odor are eliminated. A halogen leak detector may be used with chlorinated solvents to determine when a system is completely purged. If the odor of the solvent is detected in the effluent purge gas, the system requires additional purging. Precautions: Work shall be carried out only in a well-ventilated room.

13.8.5 AQUEOUS OR SEMIAQUEOUS AGENTS Blue Gold Industrial Cleaner (Modern Chemical) is a basic product with a pH of 13 in its concentrated form (equivalent to lye) and 11 in a 1% solution equivalent to NH4OH. It is a suitable replacement cleaning agent for chlorinated hydrocarbons and has been field tested by Pratt & Whitney, Martin Marietta, and all the industrial gas companies. It has no harmful ingredients, is completely water-soluble, biodegradable, nontoxic, nonflammable, and noncorrosive. Although it is highly recommended as a secondary cleaner for pipelines and ancillary equipment, sand blasting is preferred for primary cleaning, especially if heavy contamination and scale are present.

13.9 DRYING GASES After undergoing a cleaning agent process, the oxygen piping system has to be dried. For this purpose, the choice is limited to either oil-free dry air or oil-free dry nitrogen with a preference for the latter: 1. The gases should have a dew point of 408C at atmospheric pressure. 2. Nitrogen or air should be compressed to the required pressure by dry oil-free compressors that use Teflon rings or labyrinth pistons (the latter made by Burckhardt, Switzerland). Obviously, if centrifugal compressors are available for this purpose they should be used. 3. The preferred drying gas is nitrogen that can be transported as a liquid, and vaporized to the desired pressure. With this procedure one is assured of dryness and complete freedom from oil. 4. It is hazardous to use compressed air from standard oil-lubricated compressors regardless of how good the oil removal system may be. Oil vapor adsorbers (even alumina) eventually become saturated and unless they are changed often, oil vapor will break through and contaminate the distribution system. 5. Compressed air for sand or grit blasting should also be completely oil-free and dry.

13.10 TESTING AND INSPECTION PROCEDURES5 As noted previously, cleaning procedures are only as effective as the quality of the workmanship and degree of inspection. ‘‘How clean is clean?’’ has always been a subject of debate,


because testing for cleanliness in the past has been more qualitative than quantitative, and has depended on individual perceptions rather than scientific procedures. Nevertheless, ASTM Committee G-4 has made an excellent attempt to quantify testing in a scientific manner5. For guidance purposes, the following procedures may be found useful in order to achieve the inspection standards prescribed.

13.10.1 INDIRECT INSPECTION 1. Visual examination of the cleaned surface under a strong white light should indicate no particulate matter such as slag, scale, rust, corrosion products, sand, etc. within the stated inspection standards. 2. For oil, grease, or other organic matter, use a fluorescent lamp with 3600–3800 AU black light. If any fluorescence shows up as a smear or film, the item should be re-cleaned. 3. For internal surfaces that cannot be examined as above, use a specially treated laboratory type filter and lint-free paper to wipe the surface. Examine it for fluorescence under black light. Do not use a cloth for this purpose as it may be stained and will fluoresce under a black light. 4. A water-break test is sometimes recommended for equipment such as valves whose surface can be placed horizontally. Wet the surface with clean potable water. If the water film remains unbroken for at least 5 s, the surface is considered clean.

13.10.2 AFTER USING SOLVENTS

OR

CHEMICALS

Under these conditions, the inspection is done on the solvent or chemical that has been used for cleaning and not on the part or item itself: 1. Chemical analysis: One liter of a representative sample of the spent solvent or chemical should be collected and analyzed for impurities. If analysis indicates that the impurity content is less than 100 mg (100 ppm) by weight, then the part is considered clean. Otherwise washing should continue. 2. Ultraviolet test: Collect some spent cleaning liquid in a clean eye dropper, measure 10 drops of solvent into a sheet of clean filter paper and allow it to evaporate. Examine under black light in darkness. The following indicators are helpful: . .

Fluorescence not indicated—impurity level less than 50 ppm Fluorescence definitely indicated—impurity level more than 100 ppm

3. Translucence test: Collect 1 L or more of a representative sample of the spent liquid in a glass beaker, bottle, or jar and hold it up against the light next to an equal amount of fresh new solvent or chemical. If there is an apparent difference in light transmission between the spent and the new solvent when viewed simultaneously, the contamination in the spent solvent or chemical exceeds 100 ppm, and the cleaning should continue. 4. Gravimetric solvent test (suggested in Article 10.3 by ASTM Committee G-4 in Recommended Practices G 93–98 for Surface-Film Cleanliness Limits)6: Evaporate a 1 L representative sample of the spent solvent or chemical almost to dryness, then transfer the remainder to an empty beaker that has already been weighed accurately. Heat the contents until the beaker is dry. Cool the beaker and reweigh. The weight difference per square meter due to the residue can be classified as follows:


General Conservative Target Level A Level B Level C Level D Level E Level F

Less than 11 g=m2 Less than 33 g=m2 Less than 66 g=m2 Less than 220 g=m2 Less than 550 g=m2 Specified by user or supplier of components in question

Extracted with permission from STP 1197—Flammability and Sensitivity of Materials in Oxygen Enriched Atmospheres, Copyright ASTM International, West Conshohocken, PA.

5. Alternatively, to find the particulate population in place of the film residue. The same solvent extraction may be used as in the previous test (Step 4), except that instead of complete vaporization, the solvent is passed through a laboratory filter paper to trap the particulates remaining in the solution, which is then counted. According to the ASTM G-4 Committee, the various levels of acceptance are as follows: Article 11.4.3.7, G 93–987 Level 175

300

500

Particles Allowed

Size Range (mm=100 mL)

0 1 5 20 5 0 5 20 No limit 25 0 5 20 100 No limit 100

Greater than 175 Between 100 and 175 Between 50 and 100 Less than 50 Fibers Higher than 300 Between 175 and 300 Between 100 and 175 Less than 100 Fibers Greater than 500 Between 300 and 500 Between 175 and 300 Between 100 and 175 Less than 100 Fibers

Extracted with permission from STP 1197—Flammability and Sensitivity of Materials in Oxygen Enriched Atmospheres, Copyright ASTM International, West Conshohocken, PA.

13.11 ANCILLARY TOOLS AND EQUIPMENT Aside from the various chemical agents and solvents, cleaning involves other tools and equipment as described below.

13.11.1 WIRE BRUSHES Wire brushes are indispensable for touching-up purposes on rust marks or particulate matter left over after cleaning or in areas that cannot be otherwise reached for proper cleaning;


otherwise either mechanical (the rotary type) or manual brushes can be used depending on the circumstances. It is recommended that the wire bristles be made of either stainless steel or brass alloy preferably the latter. This is to prevent any particulate matter from the bristle remaining in the equipment and acting as an ignition agent.

13.11.2 IMMERSION TANKS Immersion tanks for chemical solutions should be fabricated from ordinary carbon steel, big enough to handle the equipment and piping without overcrowding and external splashing of the chemical solutions. Two tanks are recommended and used in sequence with the same chemical solution. Tank A should be used first as a primary cleaner. Then the more or less cleaned equipment and piping should be immersed in tank B for the final cleaning. As soon as tank A becomes heavily contaminated and requires a fresh solution, tank B should be used for the primary cleaning, and tank A refilled with fresh solution for the final clean-up and vice versa.

13.11.3 PROTECTIVE CLOTHING Chemical-resistant clothes, aprons, and gloves are a must for the people who carry out chemical cleaning operations. The work itself must be done outside or in a well-ventilated area. Once cleaning is complete, the equipment and piping must be handled with clean gloves.

13.12 LABOR FORCE Needless to say, the quality and the conscientiousness of the personnel carrying out the cleaning operations must be of the highest caliber.

13.12.1 PERSONNEL Workmen shall be carefully chosen to fulfill the exacting requirements of cleaning procedures for transmission and distribution pipelines for oxygen systems. They should be specially trained for this type of work and taught to work as a team. They should be given lectures on the need for high standards of cleanliness, properties, and dangers of the various chemicals, and finally, what to do in case of accidents.

13.12.2 SUPERVISION The supervisor for this work has to be of the highest class and a completely responsible person. The supervisor should be not only experienced in the standards of acceptance but also familiar beforehand with the various testing and inspection procedures.

13.12.3 INSPECTION The owners of the transmission or distribution pipeline should have their own independent inspector on the job whose responsibility is to see that the cleaning procedures selected are executed properly, and that the prescribed standards for cleanliness are maintained.

13.12.4 CLEANING CONTRACTORS Unless impossible or unavailable, it is very advantageous and economical to engage outside cleaning contractors who specialize, and have had previous experience, in cleaning pipelines for oxygen service. The owners do not have to spend time and money to recruit and train their


own labor force. Outside cleaning contractors should be selected on the basis of quality of workmanship, and only after a thorough examination of their credentials for having done similar work. It is dangerous and false economy to select a cleaning contractor solely on the basis of price alone.

13.13 PROTECTION AND STORAGE Unless used immediately, all equipment that has been cleaned and inspected should be protected against contamination during storage and shipment.

13.13.1 PIPING If the cleaning schedule is ahead of construction, the cleaned piping must be protected and handled with great care because it is usually stored in the open, and is susceptible to rain, dust, and atmospheric pollution. The ends of all cleaned piping should be covered with a strong, heavy polyethylene sheet, sealed with a waterproof tape, or plugged with a tight plastic plug. Care should also be taken in the handling of the cleaned and sealed pipes so as not to pierce the plastic coverings or knock out the protective plugs. If this happens during storage, consideration should be given to re-cleaning the pipe in question. All pipes should be capped or sealed at the ends, and if possible have a small quantity of dry nitrogen flow continuously through a small aperture under a slight positive pressure so as to maintain proper dryness.

13.13.2 SMALL EQUIPMENT Small items such as valves, fittings, orifice plates, meters, instrumentation, and other ancillary equipment should be enclosed and sealed in strong polyethylene sacks with a thickness of at least 6 mils: 1. If the sack contains ferrous material (except stainless steel) then it is wise to enclose a small quantity of silica gel (Mobil Sova Beads) in the sack to reduce the moisture content and keep the items dry. The silica gel should itself be enclosed in a porous bag so as not to contaminate the cleaned item; the quantity of silica gel added is about 1 kg=m3 (1 g=L). 2. Dates of cleaning and inspection should be noted on the tag, and enclosed in the sack. Small equipment and spare parts that may have to be stored for an indefinite period should be housed in a special room in a weather-proof building under clean (though not necessarily ‘‘hospital clean’’) conditions. Handling should be controlled and supervised by a responsible person of the stores department. In fact, in some companies, overall supervision is carried out by the superintendent of the oxygen plant or by the engineer in charge of the project.

13.13.3 LARGE EQUIPMENT Cleaned equipment that is too large to be placed in individual sacks may be wrapped in heavy polyethylene sheets after all openings, manholes, nozzles, nipples, and inspection ports are properly sealed using plastic coverings or metal plugs, to assure that the cleaned interior will not be contaminated. If a polyethylene sheet is used, it should be secured tightly with a waterproof tape. In case of open flanges, covers should be used and their edges sealed tightly with a waterproof tape.


Inside the ferrous equipment, small porous bags of a desiccant such as silica gel or Mobil Sova bead may be placed in strategic locations to reduce moisture and prevent rusting. Such desiccant bags should be placed in prominent places, however, to attract attention and not be forgotten, when the equipment is unwrapped and ready for service.

REFERENCES 1. 2. 3. 4.

ASTM Committee G-4, Bulletin G 93–96, Article 11.4.2.2, p. 13. ASTM Committee G-4, Bulletin G 93–96, Article 10.1.4, p. 11. Union Carbide Industrial Services Company. Cleaning Equipment for Oxygen Service, Bulletin G-4.1, 1996, Compressed Gas Association, Article 3.2.5, pp. 7–8. 5. ASTM Committee G-4, Bulletin G 93–96, Article 11 complete, pp. 12–14. 6. ASTM Committee G-4, Bulletin G 93–98, Article 10.3, p. 11. 7. ASTM Committee G-4. Bulletin G 93–98, Article 11.4.3.7, p. 13.


14

Economics

14.1 GENERAL OVERVIEW Because engineering is an industrially applied science, it will be wise for the promising engineers to become acquainted with some business terminology that they will face throughout their career. The following terms listed alphabetically will be of use: Amortization: A plan to pay off a financial obligation, usually a capital asset, according to some prearranged program. Annuity: A series of equal sums of money payable to a beneficiary at regular intervals of time. Balance sheet: A presentation of the financial values of a company at a given time that equalizes the company’s assets to liabilities plus net worth. Book value: The original cost of an asset minus the accumulated depreciation. Capital: The monetary resources involved in entering or sustaining a project. Capital recovery: The act of replacement of the original cost of an asset, or project, plus interest. Cash flow: The real dollars passing into and out of the treasury of a commercial venture. Cost of capital: The expense, usually stated as a percentage, involved in borrowing capital. Critical path of a project: The longest chain of activities through a network delineating a project. Declining balance depreciation: Calculating depreciation in which the annual charge is a fixed percentage of the depreciated book value at the beginning of the year to which the depreciation applies. Depreciation: The loss of value because of obsolescence, or attrition according to some plan. Discounted cash flow: The investment analysis that compares the present worth of projected receipts or disbursements, which will occur at some designated future. Guarantee: A promise in writing by the guarantor that his process or equipment shall meet all the technical specifications outlined in the signed agreement or contract. Inflation: A prevailing increase in price of goods resulting in a decline of purchasing power. Internal rate of return: The rate of return that takes into account of only what a project earns, and disregards earnings on cash flows reinvested in other projects. Net worth: The current worth of investments in a company. Payback period: The period of time over which an asset, machine, facility, or project will produce sufficient net revenue to recover its initial investment costs. PERT: The acronym for Program Evaluation and Review Technique, which is a probabilistic project network scheduling method. Rate of return: The interest rate earned by an investment, project, or asset. Salvage value: The cost to be recovered from a used asset, machine, or property when sold or scrapped.


Sinking fund: The fund accumulated by periodic deposits and reserved exclusively for a specific purpose, i.e., retiring of a financial obligation or the replacement of a machine. Straight-line depreciation: The write-off spread uniformly over the estimated life of the asset in terms of time periods or units of production. Time value of money: The expected interest rate that capital should or will earn. Warranty: It outlines the penalty either in monetary value, or in terms of work and equipment, so as to make good his promised guarantees. Working capital: The capital necessary to get involved in a project or to sustain operations.

14.2 HISTORICAL BACKGROUND At the beginning of the twentieth century the production of industrial gases, especially oxygen, acetylene, and propane, was in the hands of a very few companies even on a worldwide basis. The use of these gases was limited to welding, flame cutting of steel shapes, cutting up steel scrap, in repair shops and in steel foundries for cutting risers. Production units for oxygen were measured in cubic meters per hour of air treated and a unit treating 1000 m3 of air per hour with an oxygen production of 200 m3=h was considered large. The product was supplied in high-pressure cylinders and at a pressure of 150 bar. Because of the mystique and the possible inherently high dangerous nature of the industry, competition for customers and pricing was, to say the least low, if nonexistent. Costs involved more than the cost of utilities, especially electric power, investment, operating personnel, and general overhead. The cost of delivering the gaseous product in heavy steel cylinders including an ample inventory of the same was a major cost factor in the sale of the product. As the demand for welding and oxy-flame cutting increased in industry so did the demand for a constant and reliable supply of these gases, especially oxygen. This placed a heavy load on North American oxygen producers during World War II, when the government practically took over steel production in terms of priorities. It was then that the industry began to think in terms of storing and delivering oxygen in liquid form, as well as by means of direct gas pipelines, which became the cheaper form of supply. These new methods of delivering products proved attractive for users with a large consumption. This development as well as the appearance of the cryogenic liquid pump for filling gaseous cylinders up to 150 bar facilitated the entrance of the gas industry into the modern era of industrial gases.

14.3 POST–WORLD WAR II DEVELOPMENT At the end of the war there was a sudden rise in the petrochemical industry as well as in the production of steel, both of which demanded almost an insatiable supply of oxygen, but at a very low price. This set of conditions pushed the gas industry in the design of large plants, and in the use of materials other than copper and lead soldering, which until then, was the staple procedure for air separation plant fabrication. Moreover, during the war large engineering companies, which formerly had experience in the design and fabrication of chemical and petrochemical plants, involving new distillation techniques, now entered into the competition for large air separation units. In fact, during the early 1950s one such company with very little experience in cryogenics built two air separation units, each with a design capacity of 1000 t of oxygen product with a purity of 95% for the production of chemicals by partial oxidation of natural gas in southern Texas. When the plant went into operation, the capacity of the oxygen unit fell short by 10%. Whereas the main


objective of cracking natural gas to produce chemicals did not succeed, it proved that with the proper engineering staff, possessing only a little experience in the field of cryogenics, and with the proper choice of materials, any group could design and build large gas separation plants successfully. The design of low-temperature separation and purification of gases was no longer considered a mysterious black box, as had been previously promoted by the few companies already in the field. It was now open to all and sundry who had the availability of venture capital and the process engineers who were ready to meet any challenge. Strangely enough, the old and conservative designers of gas separation plants did not stagnate. They met the challenge head-on. At the end of the war, a 300 t=d oxygen plant was built for International Nickel using Everdure metal and modern welding techniques. At the time, this plant was classified as a giant. Today, it may be classified as a small packaged unit, as already a 4000 mt=d of oxygen unit consisting of two parallel half-sized units has been operating for some years in Europe. Moreover, equipment for a 4000 mt=d oxygen unit came under fabrication in Europe in 2003 for a South African venue.

14.4 ECONOMIC OVERVIEW Because the first low-temperature separation plants were designed and handcrafted in Europe where power costs were high and labor costs were very low, a great deal of attention was originally paid to designing large heat transfer surfaces, low-temperature differentials, and very low-pressure drops. Capital investment was of secondary consideration because no one questioned, or could question, the selling price of the products. When low-temperature gas separation technology was transferred to North America at the end of World War II, the design philosophy changed completely. This was because fabrication costs were high, whereas energy costs were comparatively low, at least in the United States and in Canada. Investment, therefore, became a crucial factor and plant designs were simplified using standard equipment and material already available in the North American market. Plants were shop-prefabricated, and even prepackaged modules became a common item, at least for small units. (What is now considered a small unit is becoming increasingly debatable.) Now, a 1808 turn has been made, as the industry finds itself caught in a vise between increasing energy costs and inflationary equipment costs, especially in compressors that are a highly important cost factor in gas separation.

14.5 ENERGY COSTS A standard low-pressure oxygen plant requires that air be compressed to approximately 5 to 7 barA to produce complete separation between oxygen and nitrogen. Then, of course, the products have to be further compressed to their required delivery pressure. The air compressor is the biggest consumer of energy. In designing the oxygen plant, therefore, one must bear in mind the following factors which are explained in the following sections.

14.5.1 OXYGEN PURITY The difference in power between 95% and 99.5% may be as much as 10%. In fact, between 98% and 99.5% the energy cost differential may be as much as 6.5%. Is a purity of 99.5% necessary, or not? These differentials may seem small, but when one is talking of compressors of 30,000 hp, even 5% represents an expenditure of US $500,000 or higher over a 5 year payout.


110 109

Energy requirements (%)

108 107 106 105 104 103 102 101 100 99 98 93

94

95

97 98 96 Oxygen purity (%)

99

100

FIGURE 14.1 Percentage variation of energy consumption to purity of oxygen. (Courtesy of F.G. Kerry, Inc., 2004. With permission.)

Originally, an oxygen purity of 99.5% or even slightly higher was required because the oxygen was used primarily for the precision oxyacetylene cutting of steel shapes wherein a drop in oxygen purity from 99.5% to 99.2% resulted in a lower accuracy and production efficiency. A more recent steel production operation has specified an extremely high oxygen purity reducing the nitrogen content in the steel to almost zero. If the air separation plant is used for metallurgical or petrochemical purposes not requiring an oxygen purity of more than 95%–98%, it should be designed specifically for that purity. If necessary, a small fraction of the product oxygen may be further purified in a special side arm distillation column to arrive at a product purity of 99.5%, which can then be distributed throughout the plant for maintenance purposes or for carbon steel shape-cutting (Figure 14.1). It should also be pointed out to users of a lower purity oxygen product that if the product is to be further compressed to a pressure of say 30 bar or higher, a minimum purity of 98% may prove more costly in terms of compression energy, as it will increase the quantity of process air in the main air compressor, due to added compression of the unnecessary nitrogen compressed along with the oxygen product (Figure 14.2).

14.5.2 LIQUID OXYGEN PRODUCTION Should a very large plant be built in an isolated area with no possible product backup from a nearby merchant plant, a decision must be made whether the plant will stand alone as a gaseous unit or should produce enough liquid for a 5 d storage in order to continue production operations during a downtime period, due either to a planned or unplanned maintenance. In the latter case, there should be enough process air compression to increase the refrigeration capacity of the plant to operate at the specified gaseous capacity, as well as deliver enough liquid product to maintain the liquid level of the storage tank at full capacity and the liquid pumps at a cryogenic temperature at all times. In this case, two pumps should


1.00 Total

Relative energy

0.95

0.70 ration

Sepa

0.65

O2 comp ressio

0.30

n to 500 p

si

99.5% 0.25

95

96

98 97 Oxygen purity (%)

99

100

FIGURE 14.2 Effect of oxygen purity on actual energy requirements. (Courtesy of F.G. Kerry, Inc., 2004. With permission.)

be supplied, of equal capacity (100%), one to operate and the other as a standby. Periodically, each pump should be tested for full operation. It should also be kept in mind that to produce liquid oxygen from a gaseous air separation unit there will be a loss in gaseous production of at least 3 to 1. In other words, for every tonne of liquid oxygen recovered approximately 3 t of gaseous product is lost because of the loss in the refrigeration potential of the extracted liquid. Admittedly, this is not an efficient way to produce a liquid product, but can be a useful one. Another more efficient way, but with a much higher investment, is to add a separate standby liquefier, which could go into operation every time when liquid production is necessary (see Chapter 6).

14.5.3 PURE NITROGEN RECOVERY (PURITY

AT

99.9995%)

Most people have the erroneous idea that an oxygen plant automatically produces highpurity nitrogen gratis. Nothing could be further from the truth. If high-purity nitrogen product is required, the following facts have to be kept in mind. For every 1000 moles of air entering the plant, about 195 moles of oxygen are recovered and about 600 moles of waste nitrogen are needed for the sublimation of CO2 and H2O if either reversing heat exchangers or regenerators are used. For prepurification of process air using adsorption, approximately 750 moles of waste nitrogen are used for the regeneration of the adsorption unit and for the evaporative water-chilling unit (see Chapter 5). In any case, there is very little left for using


waste nitrogen for the expansion machine. One can use air or a combination of oxygen and nitrogen from the high-pressure column. The best ways to produce high-purity nitrogen as well as product oxygen from the same unit are to increase the quantity of process air, increase the number of passages in the primary heat exchangers, add more trays in the distillation columns, and increase the pressure in the process air feed. This means of course, extra energy and a higher investment, the amount of each depending on the quantity of pure nitrogen required and the range of purity. Another possibility is to replace the evaporative water chiller with a mechanical refrigeration unit.

14.5.4 ARGON AND RARE GAS RECOVERY People considering the purchase of very large air separation plants should first study the market for the above-mentioned gases before preparing the technical specifications for the type and size of the plant. Unless the owner and the operator of the plant have some form of guarantee for the future sale of these gases, especially argon, the owner should avoid the extra cost even though it is small in comparison with the overall investment. On the other hand, the owner should not turn a deaf ear to the possible sale before studying the market carefully. In one instance, during the early project stage of a plant with a total capacity of 3000 t=d oxygen, the owner refused even to consider the recovery of argon. Then, during a national scarcity of the product, which occurred during the construction period, the owner found himself signing a contract with a ready and eager purchaser for the full recovery of the argon product. This situation forced the costly reopening of the cold box, replacing the upper low-pressure distillation column completely as well as replacing and adding to the interconnecting cryogenic piping. The argon recovery connections could have been carried out originally at a very small fraction of the final cost.

14.5.5 PREPURIFICATION

OF

AIR

In areas where the cost of fuels is high it is advisable to review the possibility of using the pressure swing adsorption (PSA) system for the regeneration of the adsorbent in place of the PTSA system as described in Chapter 5. The former system will save the fuel costs required to raise the temperature to the degree necessary to regenerate the molsieve adsorbent.

14.6 INVESTMENT COSTS IN GENERAL In 1949, a representative of a well-known engineering company stated in a national engineering society meeting that his company was ready to build a standard low-pressure oxygen plant with a capacity of 500 t=d of oxygen, and all the trimmings for a price of US $2,000,000. This was quite shocking even for those days. A figure of US $3,000,000 might have been more correct. Recently, an air separation plant with a capacity of 300 t=d was constructed with all the trimmings including LOX and LIN storage for approximately US $20,000,000, even though modern concepts for prefabrication techniques were used, including the entire cold box fabricated and delivered in one packaged module (unheard of in 1949). As we have entered the era of mega plants with a capacity of as high as 4000 t=d, the size–cost escalation factor no longer follows an exponential curve. It now follows a less steep curve, which tends to become linear as one approaches and surpasses the 2000 t=d capacity (Figure 14.3 and Figure 14.4).


FIGURE 14.3 Computerized isometric of cold box internals. (Courtesy of F.G. Kerry, Inc., 2004. With permission.)

FIGURE 14.4 Computerized isometric of entire plant. (Courtesy of F.G. Kerry, Inc., 2004. With permission.)


The reasons may not be easy to explain but are plain to see. The primary exchangers are made of brazed aluminum fabricated under vacuum and carefully controlled during fabrication, but their cores have a limit in size. Therefore, for possible savings in costs one has to consider a multiple exchanger system with the usual piping manifolds and prefabricated in large modules for any possible saving in costs. For extra safety, the prepurification system must also employ an adsorption unit and should include preferably a radial design to achieve saving in both investment and process energy. Large centrifugal compressors and even expansion turbines for air separation plants in the range of 1000–4000 t=d are found in a specialty class, and hence are more expensive. Moreover, the number of fabricators of such units has decreased over the recent years, and production as well as delivery available has slowed. Of course, pricing has increased considerably. To offset the inevitable increase in price, it is possible to purchase an entire machine along with its motor drive and ancillary piping and valving completely prepackaged, in one module ready for delivery to a site, even overseas. This can save considerable cost in erection man-hour savings, can also be effected in instrumentation, and in the supply of electricals both of which consume a high number of specialty erection man-hours if carried out at the field. This phase of field operations is always the last to be completed and was generally the cause for delays in the startup. To overcome this problem, it is possible to prefabricate the entire instrumentation complex into one compact package including control room and facilities complete with all the instruments, wiring, controls, etc. converging into one juncture, so that there is only one connection to be made from the outside (see Chapter 11). Personnel facilities can also be included in the module. The same can be carried out for the entire substation for the electricals. All the motor control centers, switch boxes, relays, and wiring can be hooked up in a large cabinet similar in size to a normal transportable field office. This whole package can be transported to the site and placed on top of a concrete foundation or concrete piling to be connected directly externally by outside cables to the main transformer and to other field electricals.

14.6.1 Approximate Allocation of Investment Costs (Table 14.1) Moreover, Figure 14.5 through Figure 14.7 show a typical histogram of workforce loading for a large air separation plant with an oxygen capacity of over 2000 t=d. As one can observe from the above analysis, the actual value of the entire cryogenic system for air separation including air prepurification, plus all instrumentation along with a complete DCS system, has a total value of only 22% of the total final cost. The mechanical equipment including compressors, pumps, and electricals can reach a value of 31%, which outweighs the entire field costs. For this reason, a great deal of attention has to be given during the selection and purchasing of the mechanical and electrical equipment. Even if the process department designs a highly efficient cryogenic process cycle, it still does not completely counterbalance a careless selection of mechanical and electrical equipment. Nevertheless, if the process engineer can save energy no matter how small, it could show up as a lower operating cost, which would justify a higher cost of compression equipment. Yet, in some projects, for instance, it was found that the use of cryogenic liquid pumps was more economical than the use of product oxygen compressors, despite the former’s higher energy consumption. The present high purchase and erection cost of the gaseous oxygen compressors has almost made their selection uneconomical in spite of lower energy usage.

14.6.2 CONTINGENCIES Contingency is an educated allowance of funds to lower financial risk to the project. It may be from improper definition of the scope of supply by purchaser, or from any lingering doubts from the process and purchasing departments of the engineering company. It should not be


TABLE 14.1 Approximate Allocation of Investment Costs

(a) (b) (c)

(d)

(e)

Air separation unit plus ancillary equipment, i.e., cold box; prepurification unit; instrumentation and DCA unit; plus delivery to site Machinery such as main process air compressor; product compression; pumps; electricals, plus delivery to site Field work Foundations Buildings (if necessary) Construction of equipment and instrumentation Construction of mechanicals and piping Construction equipment (Cranes, etc.) Painting Construction support personnel Construction management Start-up costs Total equipment plus field work General field engineering Contingency Head office engineering þ fee þ consultants Total

22% 31%

4.2% 1.5% 5.6% 11.2% 1.4% 0.5% 1.0% 1.0% 0.5%

26.9% 79.6% 0.4% 5.0% 15.0% 100.0%


considered a source of cash reserve to bolster poorly defined areas. A reputable engineering company can lower the contingency factor by encouraging its engineers to develop a close relationship with its probable suppliers, give all bidders detailed information to arrive at a fair evaluation of their bids and avoid as much as possible continual revisions of design thereby increasing the cost of the project.

14.7 OPERATING COSTS In the past, it was safe to assume that these were made of about 49% energy costs and an equal amount for depreciation. Labor costs accounted for the remaining small amount. Any charge for return for capital used for investment was generally ignored. Now, however, the entire picture has changed. There are very few corporations with ready cash for venture capital, regardless of interest rates. The crucial factor in the construction of very large air separation plants is the financial charges. Ten percent per annum for the repayment of capital, plus the interest rate, plus any other fees, can quickly add up to an imposing figure in the total unit costs of operation. It is absolutely vital, therefore, to reduce capital expenditures by carefully determining the requirements of the oxygen plant, by logical planning in the field, and by strict supervision during the project stage. During the project stage, care should be exercised especially in the selection of the compression machinery. The units to be selected should have a good history of reliability in operation as well as ease of maintenance. Avoid the first of this kind, because it can lead to trouble. It is also recommended that unless the owner and the operator of a large oxygen plant have their own efficient maintenance department, it would be wiser to contract the work outside. No matter how big and well automated the oxygen plant may be, one still needs a minimum of nine people plus a supervisor. One has to have two men per shift, plus a swing



80 76

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Aug. Aug. Aug. Aug. Aug. Sep. Sep. Sep. Sep. 1 8 15 22 29 19 5 26 12

Oct. 3

Oct. 10

Oct. 17

Oct. 24

FIGURE 14.5 Workforce histogram—all trades. (ß Air Liquide, all rights reserved, 2006. With permission.)

Oct. 31

0

Nov. Nov. Nov. Nov. Dec. Dec. Dec. Dec. 7 14 21 28 5 12 19 26


Pipefitters 64 56 48 40 32 24 16 8 0

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4

3 0

0

0

3

1 July 4 1994

4

Aug. 1

2

1 Sep. 5

1

1

1

FIGURE 14.6 Workforce loading histogram—by craft. (ß Air Liquide, all rights reserved, 2006. With permission.)

Dec. 5

0


Ironworkers 16 14 12 10 8 6 4 2 0

15

10 7 3

3 July 4 1994

4 2

1

1 Aug. 1

5

6

5

4

6

1

1 Sep. 5

Oct. 3

FIGURE 14.7 Workforce loading histogram—by craft. (ß Air Liquide, all rights reserved, 2006. With permission.)

Nov. 7

Dec. 5

FIGURE 14.8 Automatic welding of stainless steel fractionation tower. (ß Air Liquide, all rights reserved, 2006. With permission.)

shift, plus one extra day operator for general purposes. As the air separation plant becomes larger and larger, however, the labor cost factor becomes relatively less important. Energy is no longer cheap, and may become even more expensive in the future. With larger oxygen plants to be considered, it may be wise to start thinking once again of steam (turbine) power for either the main air compressor or even product compression. Steam is too valuable an energy form to be used solely for heating purposes. In this connection, it should be possible to consider the use of low calorific coals or residual refinery fuels, which may have limited usefulness for other energy purposes, at least for the present. Some 20 years ago, this concept was put into practice in South America with the use of heavy refinery residuals to generate high-pressure steam to operate not only on ammonia plant but also a urea unit with a capacity of 1500 t=d, but also on an air separation unit with a capacity of 1200 t=d. The entire compression machinery used high-pressure steam in two pressure stages. Presently, there are more than 20 cogeneration plants worldwide. Figure 14.8 shows the relation of power to investment costs over a period of years.

14.8 MAINTENANCE There is always a difference of opinion between the plant superintendent and the head office management as to when the plant should stop to carry out routine preventive maintenance, therefore, reduce production, or even stop production altogether. Some industries have schedules that stop the entire plant every 2 to 3 years in order to carry out complete inspection and maintenance; other companies operate with duplicate machinery and accessories in order to operate at reduced capacities during maintenance periods and finally, some unfortunate plant owners operate their plants at full capacity until breakdown of important machinery. Even preventive maintenance is a very important operating cost on a yearly basis. Actual field experience has shown that including the cost of material and labor it amounts to


anywhere from 3% to 4% of the total annual operating cost. There are several options in carrying out preventive maintenance: first option is to carry out maintenance on a fixed schedule based on past experience; the other option is based on a computerized program, which indicates the time schedule of each individual machinery or material that requires maintenance. Some companies have found that the latter option is a better choice. This, however, depends on the type of industry, the importance of the individual equipment, and the cost of any downtime if duplication is not available. As a help to any plant superintendent, Table 14.2 gives the expected life of the equipment and the material involved in a standard air separation plant.

14.9 MARKETING HISTORY OF INDUSTRIAL GASES As noted, industrial gases were originally manufactured and supplied in cylinders by a very few companies. In the post–World War II era, very large consumers, who were used to own and operate their own industrial plants, insisted that they preferred to purchase industrial gas generators, whether they are oxygen units or whatever. Needless to say, in place of the original two or three suppliers worldwide, suddenly there appeared as many as eight at any one time, ready to supply and guarantee the operation of such units. The keen competition to design and sell these plants reached the level of an oriental rug bazaar. The victor of the sale had the misfortune of meeting the warrantees for equipment and operation imposed by purchasing agents who had honed their experience in large chemical and petrochemical corporations. Needless to say, the margin for error in pricing was infinitesimal both in process design and in cost estimation, so that important sums were lost in many projects. As one manager of projects of a large American supplier was heard to say philosophically, ‘‘We don’t consider it as money lost, but as an investment in the future.’’ Nevertheless, such a policy could not be maintained forever. After some years, when heads of corporations came from the accounting department, with an MBA degree, and not from the engineering division, the picture changed drastically. It was decided to return to the original and more interesting philosophy of the gas industry of selling gaseous products continuously on a long-term contract, rather than sell a gas plant to the user. In the past 15 years, large suppliers of industrial gases have sometimes refused to bid on the sale of a gas generating plant, except in exceptional cases, and have proposed instead to supply products to the user at a price that could be more attractive than the prospective user could possibly incur by operating his own plant, and this was over a long contractual period of time, say 15 years. In fact, in many projects both seller and user have found combined interests in the final products and in financial charges, so that the proposed plant may produce and supply product gases, steam, or electricity for the benefit of all parties.

14.10 CHALLENGING MARKET CONDITIONS One of the major factors that influences the investment costs of a gas separation and purification plant is the never-ending struggle between the commercial department desiring a low investment, and the process department that may be unsure of the energy required to operate the plant. Selecting the optimum refrigeration cycle to be used in a proposed process design is a problem that can only be solved by trial and error. In the past, the process was


TABLE 14.2 Expected Life of Air Separation Plant Equipment Expected Operational Life Air filter Structure Filters Air compressor Casings Bull gear and shaft Impellers and shaft Bearings and seals Motors Coolers Switch gears and motor control centers Direct contact aftercoolers Pumps Evaporative water chillers Refrigeration machines Prepurification unit, molecular sieve adsorber Vessels and piping Switch valves Adsorbent (zeolite) Activated alumina Cryogenic cold box Casing and structures Outdoor panels Insulation Reversing exchangers Reversing valves Nonreversing exchangers Cryogenic control valves Cryogenic piping Expansion turbines Oxygen compressors

Electrical switch gear Interconnecting field Piping and valves Gas storage tanks Cryogenic liquid vessels

30 years 1=2 to 1 year (depending on local conditions) 30 years (inspect internally every 5 years) 15–20 years (inspect every 5 years) 10–20 years (inspect every 5 years for erosion or cracks) Substitute every 5 years 15 years (clean and inspect every 3 years) 20–25 years (inspect every 2–3 years) 20–25 years (inspect every 2–3 years) 20–25 years (internal inspection every 2–3 years) 5–10 years (inspect every 2–3 years) (Same as direct water chillers) 10–15 years (inspect every 2–3 years) 20–25 years (inspect every 2–3 years) 10–25 years (inspect valve seats every 2–3 years) 10 years (with activated alumina upstream, but inspect every 5 years) 5–10 years (inspect every 3–5 years) 20–30 years (inspect every year for corrosion or leaks) 10–20 years (inspect annually for corrosion or painting) 10–15 years (inspect during routine maintenance) Inspect every 5–10 years for corrosion or leaks 15–20 years (maintenance every 2–3 years) 20–25 years (check for leakage under pressure every 5–10 years) 20–25 years (check every 2–3 years) 20–25 years (check every 2–3 years) 15–20 years (inspect every 5 years for erosion or hairline cracks) (See air compressors) (For centrifugal inspect every 3–5 years) (For reciprocating inspect every 2–3 years) 20–30 years 20–30 years (pressure test with ultrasonic equipment every 15–20 years) 20–30 years (after 20 years inspect for erosion and ductility) 30 years (every 5 years check for leaks)


carried out by one or more trained mathematicians working for a month or more. With the advent of the computer, however, and the application of simulation techniques, it can be solved in a matter of days. In other words, any contingency to be added for any possible process uncertainty can be reduced to an acceptable technical and financial limit.


In regard to the selection of the low-temperature process equipment such as heat exchangers, process piping, and valves, the estimating and the purchasing department should keep up to date information on the variety and designs available, costs and deliveries, keeping in constant touch with the suppliers. As to selection of machinery such as compressors, pumps, refrigeration units and their ancillary interconnecting piping that are an important factor in the overall costing of the plant, the designer’s machinery specialist should also keep in constant touch with the suppliers on the latest information in design, costs, and delivery. The same holds true for the latest information for instrumentation and controls, whose development is as rapid as the most recent publication of any engineering magazine. If the engineering company does not have an instrumentation specialist, it should have an independent consultant who should have up to date knowledge on the subject. Finally, as every one knows, field costs are high and rising due to inflation, which is constantly with us. Because of this situation every effort should be made to fabricate, or purchase complete shop-fabricated modules in order to simplify and accelerate field labor. There are two caveats to keep in mind: firstly, interconnecting piping whether it be process or utility is expensive to purchase, deliver and install in the field, therefore, every effort should be carried out to keep their lengths as short as possible, and prefabricate as much as possible; secondly, try to avoid the welding of aluminum sheets or piping in the field especially if the proposed plant is to be located in an isolated location faraway from any industrial area. Good aluminum welders are not easily available, and a contractor may have to initiate and operate a costly welding school. The use of computerized isometrics for piping and even the entire project will reduce the cost of material and field man–hour costs (see Figure 14.3 and Figure 14.4). The choice of material such as aluminum versus stainless steel for vessels has been debatable since 1950. It depends on the market price of each at the time of the project, the wages, and the availability of welders for aluminum and the quality of the performance after results of the x-rays are studied. On the other hand, stainless steel permits the use of smaller thicknesses and automatic welding, therefore, lower overall costs (see Figure 14.8). Two examples may be significant. In one case, the aluminum welders left the isolated construction site as soon as they were certified because of poor local living conditions. In another case, a European contractor found himself in South America with no welders of any value, either for mild steel or aluminum, and wired his home office for help. Within a week, 14 European tourists suddenly appeared in South America, and by sheer coincidence, they were all qualified aluminum welders. In a complex comprised of a partial oxidation unit, a 1200 mt=d oxygen plant and a 1200 mt=d ammonia=urea plant, carbon steel interconnecting piping was designed on the short side, with just enough freedom to justify a minimum acceptable maintenance operation. When the contractor was questioned on the design, especially when there was more than ample usable space around the plant, the reply was very simple and to the point: ‘‘piping is expensive.’’

14.11 INVESTING IN A PROJECT As noted in the beginning of this chapter, engineering is the industrial application of science. In fact, many engineering schools offer degrees in applied science. If one goes a step further into reality he can say that engineering is the commercial application of science. The study of engineering includes the design and construction of industrial projects involving large sums


of money, and which brings the aspiring engineer closer to the business neighborhood of accounting and finance. Commercial corporations that deal with industrial projects have to raise funds to finance promising projects, and for this reason it is important that an engineer should become very familiar with the following financial axioms.

14.11.1 RAISING INVESTMENT CAPITAL Raising funds for capital investment may include private funds up to say 10%. The private funds may still be issued for stock, whether preferred or common, or they may be treated as a loan to the company. The preferred stock up to 45% and the balance of 45% in common stock plus private funds will add up to 100%. The payments to investors will be made in dividends unless their contributions are treated as loans, not investments. The stockholders will recoup their original investment only if the stock price remains at or rises above original price. The average rate of interest payments to various stockholders is known as the average cost of capital (ACC), and the corporation’s cost of debt is equal to its borrowing rate less income tax rate. Before a decision is made on any project, therefore, management must study the return on investment (ROI) in relation to ACC. If the ROI is greater than ACC then the project is considered profitable and receives a green light. If ROI is lower, then the project should be rejected.

14.11.2 PRESENT VALUE

OF AN INVESTMENT

Discounted cash flow return on investment (DCFROI) is widely used in evaluating a project. The term explains that a dollar earned today has a greater purchasing power than a dollar to be earned at some future time. The depreciation of future cash receipts can be discounted over time, so that the annual cash flow is on an equal basis. The discount factor used is Discount factor ¼ 1=(1 þ r)n

(14:1)

where r is the interest rate and n is the year that interest rate is compounded. Each annual cash flow is calculated for its present discounted value: Cn ¼ Cn [1=(1 þ r)n ]

(14:2)

DCFROI, therefore, is developed by the following series: Net present value (NPV): NPV ¼ C0 þ C1 =(1 þ r)1 þ C2 =(1 þ r)2 þ þ Cn =(1 þ r)n

(14:3)

The summation of the above series is as follows: n X

[(1 þ r)n 1]=r(1 þ r)n

(14:4)

n¼1

The above equation represents the NPV of an annuity, which can be transposed to NPV ¼ C[(1 þ r)n 1]=r(1 þ r)n


(14:5)

If the above equation is adjusted to equal zero, the interest rate (IRR) produces a future cash flow equal to the total investment, and the break-even point is computed as follows: I =C ¼ [(1 þ r)n 1]=r(1 þ r)n

(14:6)

where I is the investment, C is cash flow, r is the rate of return (interest rate is considered as the internal rate of return [IRR]), and n is the years of project life. The term I=C is generally referred to as the payback period. Inflation does not have to be included in a DCF analysis. Its effect is negligible. It is also very important, especially in these days of ever increasing cost of energy, to use the NPV equation to evaluate the present value of energy consumption over the life of the project. Figure 14.9 and Figure 14.10 show a series

Net present value of electrical energy 01/09/91 Basis: 85% discount factor for inflation 98% on stream time or 8585 h/year Duration : 5; 10; 15; 20 years Cost of power $20; $20; $25; 30; $35; $40; $45; $50 per MW h

50 5 years

10 years

15 years

45

Cost of power in US$ per MW h

20 years 40

35

30

25

20

15 500

1000

1500

2000

2500

3000

3500

4000

Net present value in US

FIGURE 14.9 Relationship between cost of power and net present value of project (15–50 US$ per MW h). (Courtesy of F.G. Kerry, Inc., 1991. With permission.)


Cost of power in US$ per MW/h

Net present value of electrical energy Basic: 8.5% discount factor for inflation 98% on stream time 8585 h/year Duration 5, 10, 15, and 20 years Power cost 50, 60, 70, 80 and 90 $/MW

15 years

10 years

5 years 90

20 years

80 70 60 50 40 2000

2500

3000

3500

4000

4500

5000

5500

6000

6500

7000

Net present value in US

FIGURE 14.10 Relationship between cost of power and net present value of project (40–90 US$ per MW h). (Courtesy of F.G. Kerry, Inc., 1991. With permission.)

of charts indicating the relationship between NPV of energy at various unit costs and life of the project.

14.11.3 CAVEAT

ON THE

USE OF DCFROI

Though the use of DCFROI methodology is precise, one must keep in mind that it is also based on the accuracy of estimates; but these estimates can sometimes border on wishful thinking. It also assumes that the periodic cash flow funds can be invested at the same rate as that computed for the rate of return (ROI). Nevertheless, according to Horwitz, a former supervisor of economic evaluations and feasibility studies at H.K. Ferguson Co. in Cleveland, OH, ‘‘the DCF analysis is a very useful tool to compute the economic validity of a project, but it should not preempt sound judgment.’’

14.12 ENVOI The industrial gas industry deserves credit, first of all for its self-policing safety methods through its international organizations, which were founded from day one of the industry. The industry is also a close follower of the concept of coefficiency, which was established by the Committee on Sustainable Development at the 1992 Earth Summit in Rio de Janeiro, Brazil. This concept involves the supply of goods or services wherein environmental concerns are just as significant as financial considerations. Industrial gases have been used to reduce pollution and change the face of planet Earth. Oxygen has been used to treat and purify water and air, nitrogen to reduce electronic industry


waste as well as recycle solvents, hydrogen to provide cleaner fuel through hydrogen fuel cells. Glass manufacturers are controlling emissions resulting from the fusion of glass in hightemperature furnaces with special burners using oxygen of minimum purity of 98% to reduce the formation of NOx. Noncryogenic separation of industrial gases has been developed by reducing the consumption of energy. The development of a new bleaching process involving a mixture of oxygen, ozone, and hydrogen peroxide has considerably reduced the quantity of pollution effluents released into the environment from the paper industry. In Europe, oxygen is used to incinerate waste at high temperatures, and the combustion gases to drive turbines generating electricity as well as steam for municipalities. Oxygen is also used for coal gasification, and for the production of low sulfur-free diesel. As it has been stated in the centennial issue of a well-known international gas company: The molecular chemistry of gases is well known and it seems unlikely that any new discoveries will cause Lavoisier, the father of modern chemistry, to turn in his grave. On the other hand, there are always new applications to discover, or to develop some of which will require phenomenal quantities of well known industrial gases.


Appendix

A.1 A.2 A.3 A.4 A.5 A.6

A.7

A.8 A.9 A.10 A.11 A.12 A.13 A.14 A.15 A.16 A.17 A.18 A.19 A.20 A.21

Directory of base SI terms and derivatives. (From GOA report CED 78-128, Oct 20, 1978. With permission.) Engineering Values of Some Important Constants. Normal Volumetric Analysis of Atmospheric Air. (From Product and Reference Book, D00–143, 2003–2004. With permission of Leybold Vacuum Company.) Values of Product C* of Mean Free Path l. (From Product and Reference Book, D00– 142, 2003–2004. With permission of Leybold Vacuum Company.) Altitude and Atmospheric Pressures. (From Compressed Air and Gas Data, 33–154, 155. With permission of Ingersol-Rand Company Limited.) Weight of Water per Cubic Foot of Air at Various Temperatures as a Percent of Saturation (based on atmospheric pressure of 14.7 lbs ab [101.33kPa abs]). (From Compressed Air and Gas Data, 33–119. With permission of Ingersol-Rand Company Limited.) Noise Intensity Levels. (From Jewett, J.W. and Serway R.A., Physics for Scientists and Engineers (with PhysicsNow and InfoTrac) 6th edition, copyright 2004. Reprinted with permission of Brooks=Cole, a division of Thomas Learning.) Vapor-pressure (boiling point) curves of common refrigerants. (From King, G.R., Modern Refridgeration Practice, McGraw Hill Company, 53, 1971. With permission.) Vapor pressure of water vapor at varying temperatures. (From King, G.R, Modern Refrigeration Practice, McGraw Hill Company, 31, 1971. Vapor-pressure curves between triple points and critical points. (With kind permission of Linde BOC Process Plants.) Conversion Table for Oxygen, Nitrogen, and Argon. (With kind permission of Linde BOC Process Plants) Molar Heat Capacity of Some Gases at 300 K. (From Data Book, Air Liquide Process and Construction. With permission.) Conversion Table for Vacuum Pressures. (From Product and Reference Book, D00– 142, 2003–2004. With permission of Leybold Vacuum Company.) Scale of Various Acidities. (From Eshbach, O.W., Handbook of Engineering Fundamentals, 2nd edition, 11–09, 1952. With permission of John Wiley &Sons.) Boiling points, Latent Heat, and Specific Heat of Various Gases. (From Kerry, F.G., Inc. With permission.) Heating Value and Density of Various Fuels. (From Kerry, F.G., Inc. With permission.) Value Analysis of Equipment. (From Kerry, F.G., Inc. With permission.) Effect of oxygen purity on actual energy requirements. (From Kerry, F.G., Inc. With permission.) Flammability Characteristics of Various Gases and Liquids. Relation between boiling points and pressure for oxygen, nitrogen, and argon. (From Kerry, F.G., Inc. With permission.) Friction factors for clean commercial steel pipes. (With kind permission of the Crane Company.)


A.1 Derived units with special names

Base units

(kg·m/s2)

Newton Kilogram

kg Mass

N

Pascal (N/m2) Pa Pressure stress

Gray (J/kg) Gy Absorbed dose

Force

m2

M2

Area m/s Velocity

Length

J Becquerel Bq

m/s2 Second

s

Coulomb

Farad

(A·s)

mol

Amount of substance Ampere A Electric current Degree Kelvin Celsius K

(K)

Henry

Radian rad Plane angle Steradian sr Solid angle


(C/V)

Weber Wb (V·s)

Magnetic Lumen flux density Luminous flux

W Power, heat flow rate Volt

(W/A)

Tesla (Wb/nm2) T Magnetic flux density

(cd·sr)

lm

(J/s)

V Potential, electromotive force

(Wb/A)

H Inductance

8C

Thermodynamic Celsius temperature temperature t 8C = Tk 273.15 Candela cd Luminous intensity suplementary units

Watt

Frequency

F Electric Capacitance charge Siemens S (1/Ω) ohm (V/A) Ω Resistance C

(N·m)

Energy, work, quantity of heat

Hertz Hz (1/s)

(1/s)

Activity (of ionizing radiation source)

Acceleration

Time Mole

Joule

m

Meter

Lux

(Im/m2)

lx

Illuminance

M2

A.2 Designation (Alphabetically)

Symbol

Value and Unit

Atomic mass unit Avogadro constant

mu NA

1.66051027 kg 6.02251023 mol1

Boltzmann constant

k

Electron rest mass Elementary charge Molar gas constant

me e R

1.38051023 J K1 mbarl 13.8051023 K 9.10911031 kg 1.60211019 As 8.314 Jmol1 K1 mbarl ¼ 83:14 molk

Molar volume of the ideal gas Standard acceleration of free fall Planck constant

Vo gn h

Stefan–Boltzmann constant

s

22.414 m3 kmol1 22.4141 mol1 9.8066 ms2 6.62561034 J s W 5.669108 2 4 m K

Speed of light in vacuum Standard reference density of a gas

e me c ln

As kg 2.9979108 m s1 kg m3

Standard reference pressure Standard reference temperature

pn Tn

101.325 Pa ¼ 1013 mbar Tn ¼ 273.15 K, J ¼ 08C

Specific electron charge


Remarks

Number of particles per mol, formerly: Loschmidt number

R ¼ NA k DIN 1343; formerly: molar volume at 08C and 1013 mbar

Also: unit conductance, radiation constant

1.75881011

Density at q ¼ 08C and pn ¼ 1013 mbar DIN 1343 (Nov. 1975) DIN 1343 (Nov. 1975)

A.3 Component

Mol Fraction

Partial Pressure (mbar)

Nitrogen 0.78084 Oxygen 0.20946 Argon 0.00934 Carbon dioxide 0.0350 106 Neon 1.818 105 Krypton 1.14 106 Xenon 8.6 108 H2 deuteride 3.12 104 Deuterium 1.56 104 Helium 3 1.0 107 Hydrogen 5 105 Others Methane 2.0 104 N2O 5.0 105 Caveat In industrial areas carbon dioxide may reach as high metallurgical areas check for H2S and SO2.


792 212 9.47 0.31 1.9 102 1.1 103 9.0 105

5.3 103 5.0 104 2.0 103 5.0 104 as 450 vppm. In

A.4 Abbreviation H2 He Ne Ar Kr Xe Hg O2 N2 HCl CO2 H2O NH3 C2H5OH Cl2 Air


Gas

C * 5 lp (cm mbar)

Hydrogen Helium Neon Argon Krypton Xenon Mercury Oxygen Nitrogen Hydrochloric acid Carbon dioxide Water vapor Ammonia Ethanol Chlorine Air

12.00 103 18.00 103 12.30 103 6.40 103 4.80 103 3.60 103 3.05 103 6.50 103 6.10 103 4.35 103 3.95 103 3.95 103 4.60 103 2.10 103 3.05 103 6.67 103


A.5 Altitude Above Sea Level Feetb

Miles

5,000 4,500 4,000 3,500 3,000 2,500 2,000 1,500 1,000 500 0 500 1,000 1,500 2,000 2,500 3,000

— — — — — — — — — — — — — — — — —

Temperaturea Metersb 1526 1373 1220 1068 915 763 610 458 305 153 0 153 305 458 610 763 915

8F 77 75 73 71 70 68 66 64 63 61 59 57 55 54 52 50 48

Barometerb

8C 25 24 23 22 21 20 19 18 17 16 15 14 13 12 11 10 9

Atmospheric Pressure

Inches Hg Abs.

mm Hg Abs.

psia

kg= cm2 Abs.

kPa A

35.58 35.00 34.42 33.84 33.27 32.70 32.14 31.58 31.02 30.47 29.92 29.38 28.86 28.33 27.82 27.32 26.82

903.7 889.0 874.3 859.5 845.1 830.6 816.4 802.1 787.9 773.9 760.0 746.3 733.0 719.6 706.6 693.9 681.2

17.48 17.19 16.90 16.62 16.34 16.06 15.78 15.51 15.23 14.96 14.696 14.43 14.16 13.91 13.66 13.41 13.17

1.229 1.209 1.188 1.169 1.149 1.129 1.109 1.091 1.071 1.052 1.0333 1.015 0.956 0.978 0.960 0.943 0.926

120.5 118.5 116.5 114.6 112.7 110.7 108.8 106.9 105.0 103.1 101.33 99.49 97.63 95.91 94.19 92.46 90.81


3,500 4,000 4,500 5,000 6,000 7,000 8,000 9,000 10,000 15,000 20,000 25,000 30,000 35,000 40,000 45,000 50,000 55,000 60,000 70,000 80,000 90,000 100,000

— — — 0.95 1.1 1.3 1.5 1.7 1.9 2.8 3.8 4.7 5.7 6.6 7.6 8.5 9.5 10.4 11.4 13.3 15.2 17.1 18.9

1068 1220 1373 1526 1831 2136 2441 2746 3050 4577 6102 7628 9153 10,679 12,204 13,730 15,255 16,781 18,306 21,357 24,408 27,459 30,510

47 45 43 41 38 34 31 27 23 6 12 30 48 66 70 70 70 70 70 67 62 57 51

8 7 6 5 3 1 1 3 5 14 24 34 44 57 57 57 57 57 55 52 59 46

26.33 25.84 25.37 24.90 23.99 23.10 22.23 21.39 20.58 16.89 13.76 11.12 8.903 7.060 5.558 4.375 3.444 2.712 2.135 1.325 c 8.2731 5.2001 3.2901

668.8 656.3 644.4 632.5 609.3 586.7 564.6 543.3 522.7 429.0 349.5 282.4 226.1 179.3 141.2 111.1 87.5 68.9 54.2 33.7 21.0 13.2 8.36

12.93 12.69 12.46 12.23 11.78 11.34 10.91 10.50 10.10 8.29 6.76 5.46 4.37 3.47 2.73 2.15 1.69 1.33 1.05 0.651 0.406 0.255 0.162

0.909 0.892 0.876 0.860 0.828 0.797 0.767 0.738 0.710 0.583 0.475 0.384 0.307 0.244 0.192 0.151 0.119 0.0935 0.0738 0.0458 0.0285 0.0179 0.0114

89.15 87.49 85.91 84.33 81.22 78.19 75.22 72.40 69.64 57.16 46.61 37.65 30.13 23.93 18.82 14.82 11.65 9.17 7.24 4.49 2.80 1.76 1.12


A.6 Percentage of Saturation Temperature (8F)

10

20

30

40

50

60

70

80

90

100

Temperature (8C)

10 0 10 20 30 32 35 40 45 50 55 60 62 64 66 68

Grains 0.028 0.048 0.078 0.124 0.194 0.211 0.237 0.285 0.341 0.408 0.485 0.574 0.614 0.656 0.701 0.748

Grains 0.057 0.096 0.155 0.247 0.387 0.422 0.473 0.570 0.683 0.815 0.970 1.149 1.228 1.313 1.402 1.496

Grains 0.086 0.144 0.233 0.370 0.580 0.634 0.710 0.855 0.024 1.223 1.455 1.724 1.843 1.969 2.103 2.244

Grains 0.114 0.192 0.310 0.494 0.774 0.845 0.946 1.140 1.366 1.630 1.940 2.298 2.457 2.625 2.804 2.992

Grains 0.142 0.240 0.388 0.618 0.968 1.056 1.183 1.424 1.707 2.038 2.424 2.872 3.071 3.282 3.504 3.740

Grains 0.171 0.289 0.466 0.741 1.161 1.268 1.420 1.709 2.048 2.446 2.909 3.447 3.685 3.938 4.205 4.488

Grains 0.200 0.337 0.543 0.864 1.354 1.479 1.656 1.994 2.390 2.853 3.394 4.022 4.299 4.594 4.906 5.236

Grains 0.228 0.385 0.621 0.988 1.548 1.690 1.893 2.279 2.731 3.261 3.879 4.596 4.914 5.250 5.607 5.984

Grains 0.256 0.433 0.698 1.112 1.742 1.902 2.129 2.564 3.073 3.668 4.364 5.170 5.528 5.907 6.308 6.732

Grains 0.285 0.481 0.776 1.235 1.935 2.113 2.366 2.849 3.414 4.076 4.849 5.745 6.142 6.563 7.009 7.480

23.3 17.8 11.7 6.7 1.1 0 1.7 4.4 7.2 10.0 12.8 15.6 16.7 17.8 18.9 20.0


70 72 74 76 78 80 82 84 86 88 90 92 94 96 98 100

0.798 0.851 0.907 0.966 1.028 1.093 1.163 1.236 1.313 1.394 1.479 1.569 1.663 1.763 1.867 1.977

1.596 1.702 1.813 1.931 2.055 2.187 2.325 2.471 2.625 2.787 2.958 3.138 3.327 3.525 3.734 3.953

2.394 2.552 2.720 2.896 3.083 3.280 3.488 3.707 3.938 4.181 4.437 4.707 4.990 5.288 5.601 5.930

3.192 3.403 3.626 3.862 4.111 4.374 4.650 4.942 5.251 5.575 5.916 6.276 6.654 7.050 7.468 7.906

3.990 4.254 4.533 4.828 5.138 5.467 5.813 6.178 6.564 6.968 7.395 7.844 8.317 8.813 9.336 9.883

4.788 5.105 5.440 5.793 6.166 6.560 6.976 7.414 7.877 8.362 8.874 9.413 9.980 10.576 11.203 11.860

Condensed from a circular of the US Weather Bureau. 1 Grain ¼ 1,7000 lb; grains 0.0648 ¼ grams.

5.586 5.956 6.346 6.758 7.194 7.654 8.138 8.649 9.189 9.756 10.353 10.982 11.644 12.338 13.070 13.836

6.384 6.806 7.253 7.724 8.222 8.747 9.301 9.885 10.502 11.150 11.832 12.551 13.307 14.101 14.937 15.813

7.182 7.657 8.159 8.690 9.249 9.841 10.463 11.120 11.814 12.543 13.311 14.120 14.971 15.863 16.804 17.789

7.980 8.508 9.066 9.655 10.277 10.934 11.625 12.326 13.137 13.997 14.780 15.639 16.624 17.676 18.661 19.766

21.1 22.2 23.3 24.4 25.6 26.7 27.8 28.9 30.0 31.1 32.2 33.3 34.4 35.6 36.7 37.8

A.7 Comparing Noise Levels Using Decibels (b dB) In physics, a decibel is a unit expressing the intensity of a sound wave equal to 20 times the logarithm (base 10) of the ratio of pressure produced by the sound wave to a reference pressure, usually 0.0002 microbar, or 1 microbar (mbar). b ¼ 10 log (I=Io ) where Io ¼ reference intensity (threshold of hearing) ¼ 1012 W=m2 I ¼ intensity in W=m2 at level b and is measured in decibels (dB)

Source of Sound Nearby jet airplane Jackhammer, machine gun Siren, rock concert Subway, power mower Busy traffic Vacuum cleaner Normal conversation Mosquito buzzing Whisper Rustling leaves Threshold of hearing


b (dB) 150 130 120 100 80 70 50 40 30 10 0

A.8 Critical 1000 800 600

Critical

400

( id e

200

fri g

t7

44

io x nd o 13 rb nt a r Ca e frig Re

)

nia mo Am

t ran ige r f e (R

er f ri g Re

7) 71

500 ant

So

lid

100 80 60 40

Re

n era

20

10 8

psig

6 3

02 t5 2 n a 2 er nt rig era f e 12 R frig nt e a r R ge fri e R

4

11

r

ge

ri ef

t an

nt

R

g

fri

Re

11

a er

13

t1

ge

fri

2

Re

n ra

t7

18

)

1 0.8

rig

er

an

0.6

W

at er

(R

ef

0.4

0.2

0.1 –100

–80

–60


–40

–20

0 20 40 Temperature (8F)

60

80

100

120

1.0

2.0

50

0.9

1.8

45

0.8

1.6

40

1.4

35

1.2

30

1.0

25

0.8

20

0.6

15

0.4

10

0.2

5

Absolute pressure (in.Hg)

Absolute pressure (psi)

0.7

0.6

0.5

0.4 0.3

0.2 Vapor pressure of water at 608F is 0.52 in.Hg or 0.26 psia

0.1

20

30


40

50

60 70 80 Temperature (F)

90

100

Absolute pressure (mm·Hg)

A.9

A.10 xxxxxxxx 8 7 6 5

1 103 9 8 7 6 5

Kr Kr Kr

4

Kr

Kr

Kr

Kr

Kr

ne

3

4

le ety

Ac 2

n

go 1

Xe

C 2l

2

F 2

e an

nt

H

ex

Pe

CC

C

C

L

3

Tr

LF

e

ne

en

Cr

yle Eth

Eth

Tr Tr

SO

rz

an

Tr

Se

10 9 8 0 10 7 6 5 4

x

xx

xx

x xx xx xx xx xxx xx

xx

) (VA Air

Tr

F

2

Tr

Tr

TR Tr

2

−1

Tr Tr

3

To lle ne

1 9 8 7 6 5 4

Tr

3

2

L CC

Tr

100 9 1 8 10 7 6 5 4

3

N

yle op ne Pr pa o Pr

Kr

2

1 9 8 7 6 5 4

H2

ne

2

3

e

2

LF 3

CC

S H2

e

CC

4

an

N

1

Me tha No ne

Tr

3

3

2 CO

G

Ar

101 9 8 2 7 10 6 5 4

Kr

n

IQ r(L Ai

2

o pt

y

)

3

1 9 8 7 6 5

Kr

Kr

Tr

2

Tr

Tr

10−2 1

− 1 10

60

70

80

90 100 110 120 130 140 150 160 170 180 190 200 210 220 230 240 250 260 270 280 290 300 K

−340 −320 −300 −280 −260 −240 −220 −200 −180 −160 −140 −120 −100 −80 −60 −40 −20


0

20

40

60

80 ˚F


A.11

1 Metric ton (t)

1 Short ton (st)

1 scf gas at 608 F, 1 atm

1 scf gas at 708F, 1 atm

1 m3 gas at 08C, 1 atm

1 gallon (US) liquid

1 liter (dm3) liquid

O2 N2 Ar O2 N2 Ar O2 N2 Ar O2 N2 Ar O2 N2 Ar O2 N2 Ar O2 N2 Ar

t

st

1

1.102311

0.907185

1

3.8275105 3.3494105 4.7784105 3.7552105 3.2862105 4.6881105 1.4286103 1.2502103 1.7836103 4.320103 3.056103 5.276103 1.141103 0.8074103 1.394103

4.2191105 3.6921105 5.2672105 4.1394105 3.6224105 5.1678105 1.5748103 1.3781103 1.9660103 4.762103 3.369103 5.816103 1.258103 0.8900103 1.537103

scf Gas @ 608F

scf Gas @ 708F

m3 Gas @ 08C

Gall (US) Liquid

Liter Liquid

26127 29856 20928 23702 27085 18985

26630 30430 21330 24158 27606 19351

700.0 799.9 560.7 635.0 725.6 508.6

1

1.019243

2.67911102

0.981120

1

2.62853102

37.32579

38.04405

1

231.5 327.2 189.5 210.0 296.8 171.9 8.861 103 10.96 103 9.056 103 8.693 103 10.75 103 8.885 103 0.3307 0.4091 0.3380

876.3 1238.6 717.4 795.0 1123.6 650.8 3.354 102 4.148 102 3.428 102 3.291 102 4.070 102 3.363 102 1.252 1.548 1.280

112.9 91.25 110.4 29.81 24.11 29.17

115.0 93.00 112.5 30.39 24.57 29.73

3.024 2.445 2.958 0.7988 0.6458 0.7815

1

3.785412

0.264172

1

Note: 1 atm ¼ 14.696 psia. Liquid quantities measured at 1 atm at the respective boiling points.

A.12 Molar Heat Capacity (cal=mol K) Gas Monatomic Gases He AV Diatomic Gases H2 N2 O2 CO Cl2 Polyatomic Gases CO2 SO2 H2O CH4

CP

CV

CP 2 CV

g 5 CP=CV

4.97 4.97

2.98 2.98

1.99 1.99

1.67 1.67

6.87 6.95 7.03 7.01 8.29

4.88 4.96 5.04 5.02 6.15

1.99 1.99 1.99 1.99 2.14

1.41 1.40 1.40 1.40 1.35

8.83 9.65 8.46 8.49

6.80 7.50 6.46 6.48

2.03 2.15 2.00 2.01

1.30 1.29 1.30 1.31

Note: All values obtained at 300 K.



A.13 torr mtorr bar mbar mbar Pa MPa

103 106 1.33 1.33103 1.33106 1.33105 1.33101

102 105 1.33101 1.33102 1.33105 1.33104 1.33102

101 104 1.33 102 1.33 101 1.33 104 1.33 103 1.33 103

1 103 1.33 103 1.33 1.33 103 1.33 102 1.33 104

101 102 1.33 104 1.33 101 1.33 102 1.33 101 1.33 105

102 101 1.33 105 1.33 102 1.33 101 1.33 1.33 106

103 1 1.33 106 1.33 103 1.33 1.33 101 1.33 107

104 101 1.33 107 1.33 104 1.33 101 1.33 102 1.33 108

105 102 1.33 108 1.33 105 1.33 102 1.33 103 1.33 109

106 103 1.33 109 1.33 106 1.33 103 1.33 104 1.33 1010

A.14 Cooling Water Acidity Influence of cooling water acidity on corrosion which may affect the cooling water tower, heat exchangers piping etc. Water dissociates in hydrogen and hydroxyl ions: H2 O ¼ (Hþ ) þ (HO) or quantitatively or numerically

(A:1)

[(Hþ )(HO) ]=H2 O

(A:2)

(Hþ )(OH ) ¼ KION ¼ 1014

(A:3)

Equation A.3 for dissociation of water (Hþ)(OH) ¼ KION ¼ 1014 expresses the independence of hydrogen and hydroxyl (ion) concentrations in aqueous solutions or log (Hþ ) log (OH ) ¼ 14 And log (Hþ) ¼ pH ¼ measure of acidity. The lower the pH, the higher the acidity (lower the alkalinity). At pH ¼ 7 the solutions are equal or neutral (pure distilled water). Lowest recorded rainfall pH (Wheeling, WV, Fall, 1978)

Battery acid

Lemon juice Unpolluted rain Lye

Distilled water

Vinegar

Ammonia

Acid rain

0

1

2

3


4

5

6

7

8

9

10

11

12

13

14

A.15 Boiling Point @ 1=0133 bar

Latent Heat (MJ=kmol)

Specific Heat Capacity (kJ=kmol 3 K) 0.083 0.92 1.79 5.55 5.71 6.05 6.55 6.83 9.04

Helium Hydrogen Neon Nitrogen Air Carbon monoxide Argon Oxygen Krypton

He H2 Ne N2 — CO Ar O2 Kr

4.23 20.4 27.1 77.3 80.9 81.6 87.3 90.2 119.8

268.92 252.8 246.1 195.9 192.3 191.6 185.9 183.0 153.4

Xenon Methane Ethylene Ethane Propylene Propane iso-Butylene n-Butane

Xe CH4 C2H4 C2H6 C3H6 C3H6 C3H10 n-C4H10

165.0 111.7 169.5 184.6 225.5 231.1 266.3 272.7

108.2 161.5 103.7 88.6 47.7 42.1 6.9 0.5

Ammonia Freon 12 Freon 22

NH3 CCl2F2 CHClF2

239.8 243.37 232.4

Water

H2O

373.15


17.9 18.9 37.3 57.5 — 60.8 42.0 54.0 44.8

12.63 8.18 14.7 14.8 18.5 18.8 22.5 22.4

44.3 55.1 68.0 73.0 103.0 98.0 — 133.0

33.4 29.78 40.78

23.3

76.1

100.00

40.67

76.0

A.16 Fuel

kcal=Nm3

kg=Nm3

Hydrogen Methane Ethane Propane Butane Ethylene Propylene Butylene Bezene Toluene Xylene Coke oven gas Blast furnace Fuel oil

2582.4 8588.88 15432.2 22438.3 29547.8 14228.5 20933.6 27055.7 33919.3 39977.0 46428.2 4749.0 865.0 9528=kg

0.0900 0.7188 1.3605 2.0257 2.7045 1.2640 1.9203 2.5089 3.5157 4.1203 4.7469 0.5182 1.3040 —


A.17 Concept

Air filter Air compressor Air cooler Air separation unit Argon separation unit Oxygen compressor Liquid related equipment C ¼ Complexity I ¼ Importance P ¼ Weight (value) V ¼ Cost Score ¼ I (C þ P þ V)


C

I

P

V

Score

Total (%)

1 4 2 5 3 5 1

2 5 4 5 3 5 1

1 5 1 5 1 4 1

1 5 1 5 2 4 2

6 70 16 75 12 64 4

2.42 28.36 6.48 30.36 4.86 25.91 1.62

A.18

1.00

Total

Relative energy

0.95

0.70 ation

Separ

0.65

O2 compre ss

0.30

ion to 500

psi

99.5% 0.25 95

96

97

98

Oxgen purity (%)


99

100

A.19

Product

Boiling Point Ka

Flash Point K (8C)b

Auto-ignition Point K (8C)c

Acetaldehyde Acetone Acetylene Allyl chloride Ammonia Benzene 1,3-Butadiene Butane Butyl acetate 1-Butene 2-Butene N-butyl formate Carbon disulfide Carbon monoxide Cyclopropane 1,1-Dichloroethylene 2,2-Dimethylbutane Ethane Ethanol Ethyl acetate Ethyl bromide Ethyl chloride Ethylene Ethylene oxide Ethyl ether Ethyl glycol Ethyl formate Ethylglycol acetate Gasoline Gasoline (octane 60) Gasoline (octane 92) Gasoline (octane 100) Heptane Hexane Hydrogen Isobutane Isopropanol Isopropyl ether Methane Methanol Methylethylketone Ethylisobutylketone Methyl acetate Methylamine Methyl butene Methyl chloride Methyl formate N-butanol Pentane

294.26 329.26 189.60 318.15 239.82 353.15 268.71 237.59 399.26 266.48 274.26 380.37 319.26 83.15 240.15 310.37 322.59 184.60 351.48 350.37 311.48 285.37 169.50 286.71 307.59 407.15 327.59 429.15 — — — — 371.48 342.15 20.40 261.48 355.15 341.15 109.26 339.26 353.15 390.15 330.37 266.15 303.93 249.26 305.15 391.15 309.26

275.37(2.22) 255.37(17.78) — 241.48( 31.67) — 262.04(11.11) — — 299.26(26.11) — — 290.93(17.78) 243.15(30) — — 252.59(20.56) 225.37(47.78) — 224.82(48.33) 270.93(2.22) — 223.15(50.00) —

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