Handbook of Lead-Free Solder Technology for Microelectronic Assemblies

Handbook of lead-Free Solder Technology for Microelectronic AssembIies edited by Karl J. Puttlitz ISM Corporate Staff...

Author: Karl J. Puttlitz | Kathleen A. Stalter

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Handbook of lead-Free Solder Technology for Microelectronic AssembIies

edited by

Karl J. Puttlitz ISM Corporate Staff Corporate Environmental Affairs ISM Corporation East Fishkill, New York, U.S.A.

Kathleen A. Stalter ISM Technology Group ISM Corporation Hopewell Junction, New York, U.S.A.

a% MARCEL

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Although great care has been taken to provide accurate and current information, neither the author(s) nor the publisher, nor anyone else associated with this publication, shall be liable for any loss, damage, or liability directly or indirectly caused or alleged to be caused by this book. The material contained herein is not intended to provide specific advice or recommendations for any specific situation. 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 A catalog record for this book is available from the Library of Congress. ISBN: 0-8247-4870-0 This book is printed on acid-free paper. Headquarters Marcel Dekker, Inc. 270 Madison Avenue, New York, NY 10016, U.S.A. tel: 212-696-9000; fax: 212-685-4540 Distribution and Customer Service Marcel Dekker, Inc. Cimarron Road, Monticello, New York 12701, U.S.A. tel: 800-228-1160; fax: 845-796-1772 Eastern Hemisphere Distribution Marcel Dekker AG Hutgasse 4, Postfach 812, CH-4001 Basel, Switzerland tel: 41-61-260-6300; fax: 41-61-260-6333 World Wide Web http://www.dekker.com The publisher offers discounts on this book when ordered in bulk quantities. For more information, write to Special Sales/Professional Marketing at the headquarters address above. Copyright n 2004 by Marcel Dekker, Inc. All Rights Reserved. Neither this book nor any part may be reproduced or transmitted in any form or by any means, electronic or mechanical, including photocopying, microfilming, and recording, or by any information storage and retrieval system, without permission in writing from the publisher. Current printing (last digit): 10 9 8 7 6 5 4 3 2 1 PRINTED IN THE UNITED STATES OF AMERICA

MECHANICAL ENGINEERING A Series of Textbooks and Reference Books

Founding Editor

L. L. Faulkner Columbus Division, Battelle Memorial Institute and Department of Mechanical Engineering The Ohio State University Columbus, Ohio

1. Spring Designer's Handbook, Harold Carlson 2. Computer-AidedGraphics and Design, Daniel L. Ryan 3. Lubrication Fundamentals,J. George Wills 4. Solar Engineering for Domestic Buildings, William A. Himmelman 5. Applied Engineering Mechanics: Statics and Dynamics, G. Boothroydand C. Poli 6. Centrifugal Pump Clinic, lgor J. Karassik 7. Computer-AidedKinetics for Machine Design, Daniel L. Ryan 8. Plastics Products Design Handbook, Part A: Materials and Components; Part 6: Processes and Design for Processes, edited by Edward Miller 9. Turbomachinery: Basic 'Theory and Applications, Earl Logan, Jr. 10. Vibrations of Shells and Plates, Werner Soedel 1 1. Flat and Corrugated Diaphragm Design Handbook, Mario Di Giovanni 12. Practical Stress Analysis in Engineering Design, Alexander Blake 13. An lntroduction to the Design and Behavior of Bolted Joints, John H. Bickford 14. Optimal Engineering Design: Principles and Applications, James N. Siddall 15. Spring Manufacturing Handbook, Harold Carlson 16. hdustrial Noise Control: Fundamentalsand Applications, edited by Lewis H. Bell 1 7. Gears and Their Vibration: A Basic Approach to Understanding Gear Noise, J. Derek Smith 18. Chains for Power Transmission and Material Handling: Design and Applications Handbook, American Chain Association 19. Corrosion and Corrosion Protection Handbook, edited by Philip A. Schweitzer 20. Gear Drive Systems: Design and Application, Peter Lynwander 21 . Controlling In-Plant Airborne Contaminants: Systems Design and Calculations,John D. Constance 22. CAD/CAM Systems Planning and Implementation,Charles S. Knox 23. Probabilistic Engineering Design: Principles and Applications, James N. Siddall 24. Traction Drives: Selection and Application, Frederick W. Heilich 111 and Eugene E. Shube 25. Finite Element Methods: An Introduction, Ronald L. Huston and Chris E. Passerello 26. Mechanical Fastening of Plastics: An Engineering Handbook, Brayton Lincoln, Kenneth J. Gomes, and James F. Braden 27. Lubrication in Practice: Second Edition, edited by W. S. Robertson 28. Principles of Automated Drafting, Daniel L. Ryan 29. Practical Seal Design, edited by Leonard J. Martini 30. Engineering Documentationfor CAD/CAM Applications, Charles S. Knox 31 . Design Dimensioning with Computer Graphics Applications, Jerome C. Lange 32. Mechanism Analysis: Simplified Graphical and Analytical Techniques, Lyndon 0. Barton 33. CAD/CAM Systems: Justification, Implementation, Productivity Measurement, Edward J. Preston, George W. Crawford, and Mark E. Coticchia 34. Steam Plant Calculations Manual, V . Ganapathy 35. Design Assurance for Engineers and Managers, John A. Burgess

36. Heat Transfer Fluids and Systems for Process and Energy Applications, Jasbir Singh 37. Potential Flows: Computer Graphic Solutions, Robert H. Kirchhoff 38. Computer-Aided Graphics and Design: Second Edition, Daniel L. Ryan 39. Electronically Controlled Proportional Valves: Selection and Application, Michael J. Tonyan, edited by Tobi Goldoftas 40. Pressure Gauge Handbook, AMETEK, U.S. Gauge Division, edited by Philip W. Harland 41. Fabric Filtration for Combustion Sources: Fundamentals and Basic Technology, R. P. Donovan 42. Design of Mechanical Joints, Alexander Blake 43. CAD/CAM Dictionary, Edward J. Preston, George W. Crawford, and Mark E. Coticchia 44. Machinery Adhesives for Locking, Retaining, and Sealing, Girard S. Haviland 45. Couplings and Joints: Design, Selection, and Application, Jon R. Mancuso 46. Shaft Alignment Handbook, John Piotrowski 47. BASIC Programs for Steam Plant Engineers: Boilers, Combustion, Fluid Flow, and Heat Transfer,V. Ganapathy 48. Solving Mechanical Design Problems with Computer Graphics, Jerome C. Lange 49. Plastics Gearing: Selection and Application, Clifford E. Adams 50. Clutches and Brakes: Design and Selection, William C. Orthwein 51. Transducers in Mechanical and Electronic Design, Harry L. Trietley 52. Metallurgical Applications of Shock-Wave and High-Strain-Rate Phenomena, edited by Lawrence E. Murr, Karl P. Staudhammer, and Marc A. Meyers 53. Magnesium Products Design, Robert S. Busk 54. How to Integrate CAD/CAM Systems: Management and Technology, William D. Engelke 55. Cam Design and Manufacture: Second Edition; with cam design software for the ISM PC and compatibles, disk included, Preben W. Jensen 56. Solid-state A C Motor Controls: Selection and Application, Sylvester Campbell 57. Fundamentals of Robotics, David D. Ardayfio 58. Belt Selection and Application for Engineers, edited by Wallace D. Erickson 59. Developing Three-DimensionalCAD Software with the ISM PC, C. Stan Wei 60. Organizing Data for ClM Applications, Charles S. Knox, with contributions by Thomas C. Boos, Ross S. Culverhouse, and Paul F. Muchnicki 61. Computer-Aided Simulation in Railway Dynamics, by Rao V. Dukkipati and Joseph R. Amyot 62. Fiber-Reinforced Composites:Materials, Manufacturing, and Design, P. K. Mallick 63. Photoelectric Sensors and Controls: Selection and Application, Scott M. Juds 64. Finite Element Analysis with Personal Computers, Edward R. Champion, Jr., and J. Michael Ensminger 65. Ultrasonics: Fundamentals, Technology, Applications: Second Edition, Revised and Expanded, Dale Ensminger 66. Applied Finite Element Modeling: Practical Problem Solving for Engineers, Jeffrey M. Steele 67. Measurement and lnstrurnentation in Engineering: Principles and Basic Laboratory Experiments, Francis S. Tse and Ivan E. Morse 68. Centrifugal Pump Clinic: Second Edition, Revised and Expanded, lgor J. Karassik 69. Practical Stress Analysis in Engineering Design: Second Edition, Revised and Expanded, Alexander Blake 70. An Introduction to the Design and Behavior of Bolted Joints: Second Edition, Revised and Expanded, John H. Bickford 71. High Vacuum Technology: A Practical Guide, Marsbed H. Hablanian 72. Pressure Sensors: Selection and Application, Duane Tandeske 73. Zinc Handbook: Properties, Processing, and Use in Design, Frank Porter 74. Thermal Fatigue of Metals, Andrzej Weronski and Tadeusz Hejwowski 75. Classical and Modem Mechanisms for Engineers and Inventors, Preben W. Jensen 76. Handbook of Electronic Package Design, edited by Michael Pecht 77. Shock-Wave and High-Strain-Rate Phenomena in Materials, edited by Marc A. Meyers, Lawrence E. Murr, and Karl P. Staudhammer 78. Industrial Refrigeration: Principles, Design and Applications, P. C. Koelet

79. Applied Combustion, Eugene L. Keating 80. Engine Oils and Automotive Lubrication, edited by Wilfried J. Bartz 8 1. Mechanism Analysis: Simplified and Graphical Techniques, Second Edition, Revised and Expanded, Lyndon 0. Barton 82. Fundamental Fluid Mechanics for the Practicing Engineer, James W. Murdock 83. Fiber-Reinforced Composites: Materials, Manufacturing, and Design, Second Edition, Revised and Expanded, P. K. Mallick 84. Numerical Methods for Engineering Applications, Edward R. Champion, Jr. 85. Turbomachinery: Basic Theory and Applications, Second Edition, Revised and Expanded, Earl Logan, Jr. 86. Vibrations of Shells and Plates: Second Edition, Revised and Expanded, Werner Soedel 87. Steam Plant Calculations Manual: Second Edition, Revised and Expanded, V. Ganapathy 88. Industrial Noise Control: Fundamentals and Applications, Second Edition, Revised and Expanded, Lewis H. Bell and Douglas H. Bell 89. Finite Elements: Their Design and Performance, Richard H. MacNeal 90. Mechanical Properties of Polymers and Composites: Second Edition, Revised and Expanded, Lawrence E. Nielsen and Robert F. Landel 91. Mechanical Wear Prediction and Prevention, Raymond G. Bayer 92. Mechanical Power Transmission Components, edited by David W. South and Jon R. Mancuso 93. Handbook of Turbomachinery,edited by Earl Logan, Jr. 94. Engineering Documentation Control Practices and Procedures, Ray E. Monahan 95. Refractory Linings: Thermomechanical Design and Applications, Charles A. Schacht 96. Geometric Dimensioning and Tolerancing: Applications and Techniques for Use in Design, Manufacturing, and Inspection, James D. Meadows 97. An Introduction to the Design and Behavior of Bolted Joints: Third Edition, Revised and Expanded, John H. Bickford 98. Shaft Alignment Handbook: Second Edition, Revised and Expanded, John Piotrowski 99. Computer-Aided Design of Polymer-Matrix Composite Structures, edited by S. V. Hoa 100. Friction Science and Technology, Peter J. Blau 101. Introduction to Plastics and Composites: Mechanical Properties and Engineering Applications, Edward MiIler 102. Practical Fracture Mechanics in Design, Alexander Blake 103. Pump Characteristics and Applications, Michael W. Volk 104. Optical Principles and Technology for Engineers, James E. Stewart 105. Optimizing the Shape of Mechanical Elements and Structures, A. A. Seireg and Jorge Rodriguez 106. Kinematics and Dynamics of Machinery, Vladimir Stejskal and Michael ValaSek 107. Shaft Seals for Dynamic Applications, Les Horve 108. Reliability-Based Mechanical Design, edited by Thomas A. Cruse 109. Mechanical Fastening, Joining, and Assembly, James A. Speck 110. Turbomachinery Fluid Dynamics and Heat Transfer, edited by Chunill Hah I 1I.High-Vacuum Technology: A Practical Guide, Second Edition, Revised and Expanded, Marsbed H. Hablanian 112. Geometric Dimensioning and Tolerancing: Workbook and Answerbook, James D. Meadows I13. Handbook of Materials Selection for Engineering Applications, edited by G. T. Murray 114. Handbook of Thermoplastic Piping System Design, Thomas Sixsmith and Reinhard Hanselka 115. Practical Guide to Finite Elements: A Solid Mechanics Approach, Steven M. Lepi 116. Applied Computational Fluid Dynamics, edited by Vijay K. Garg 117. Fluid Sealing Technology,Heinz K. Muller and Bernard S. Nau 118. Friction and Lubrication in Mechanical Design, A. A. Seireg 119. Influence Functions and Matrices, Yuri A. Melnikov 120. Mechanical Analysis of Electronic Packaging Systems, Stephen A. McKeown

121. Couplings and Joints: Design, Selection, and Application, Second Edition, Revised and Expanded, Jon R. Mancuso 122. Thermodynamics:Processes and Applications, Earl Logan, Jr. 123. Gear Noise and Vibration,J. Derek Smith 124. Practical Fluid Mechanics for Engineering Applications, John J. Bloomer 125. Handbook of Hydraulic fluid Technology,edited by George E. Totten 126. Heat Exchanger Design Handbook,T. Kuppan 127. Designing for Product Sound Quality, Richard H. Lyon 128. Probability Applications in Mechanical Design, Franklin E. Fisher and Joy R. Fisher 129. Nickel Alloys, edited by Ulrich Heubner 130. Rotating Machinery Vibration: Problem Analysis and Troubleshooting, Maurice L. Adams, Jr. 131. Formulas for Dynamic Analysis, Ronald Huston and C. Q. Liu 132. Handbook of Machinery Dynamics, Lynn L. Faulkner and Earl Logan, Jr. 133. Rapid Prototyping Technology: Selection and Application, Ken Cooper 134. Reciprocating Machinery Dynamics: Design and Analysis, Abdulla S. Rangwala 135. Maintenance Excellence: Optimizing Equipment Life-Cycle Decisions, edited by John D. Campbell and Andrew K. S. Jardine 136. Practical Guide to Industrial Boiler Systems, Ralph L. Vandagriff 137. Lubrication Fundamentals: Second Edition, Revised and Expanded, D. M. Pirro and A. A. Wessol 138. Mechanical Life Cycle Handbook: Good Environmental Design and Manufacturing, edited by Mahendra S. Hundal 139. Micromachining of Engineering Materials, edited by Joseph McGeough 140. Control Strategies for Dynamic Systems: Design and Implementation, John H. Lumkes, Jr. 141. Practical Guide to Pressure Vessel Manufacturing, Sunil Pullarcot 142. Nondestructive Evaluation: Theory, Techniques, and Applications, edited by Peter J. Shull 143. Diesel Engine Engineering: Dynamics, Design, and Control, Andrei Makartchouk 144. Handbook of Machine Tool Analysis, loan D. Marinescu, Constantin Ispas, and Dan Boboc 145. Implementing Concurrent Engineering in Small Companies, Susan Carlson Skalak 146. Practical Guide to the Packaging of Electronics: Thermal and Mechanical Design and Analysis, Ali Jamnia 147. Bearing Design in Machinery: Engineering Tribology and Lubrication, Avraham Harnoy 148. Mechanical Reliability Improvement: Probability and Statistics for Experi-mental Testing, R. E. Little 149. Industrial Boilers and Heat Recovery Steam Generators: Design, Applications, and Calculations, V. Ganapathy 150. The CAD Guidebook: A Basic Manual for Understanding and Improving Computer-Aided Design, Stephen J. Schoonmaker 15 1. Industrial Noise Control and Acoustics, Randall F. Barron 152. Mechanical Properties of Engineering Materials, Wole Soboyejo 153. Reliability Verification, Testing, and Analysis in Engineering Design, Gary S. Wasserman 154. Fundamental Mechanics of Fluids: Third Edition, I . G. Currie 155. Intermediate Heat Transfer, Kau-Fui Vincent Wong 15 6. H VAC Water Chillers and Cooling Towers: Fundamentals, Application, and Operations, Herbert W. Stanford Ill 157. Gear Noise and Vibration: Second Edition, Revised and Expanded, J. Derek Smith

158. Handbook of Turbomachinery: Second Edition, Revised and Expanded, Earl Logan, Jr., and Ramendra Roy 15 9. Piping and Pipeline Engineering: Design, Construction, Maintenance, Integrity, and Repair, George A. Antaki 160. Turbomachinery: Design and Theory, Rama S. R. Gorla and Aijaz Ahmed Khan 161. Target Costing: Market-Driven Product Design, M. Bradford Clifton, Henry M. B. Bird, Robert E. Albano, and Wesley P. Townsend 162. Fluidized Bed Combustion, Simeon N. 0 ka 163. Theory of Dimensioning: An Introduction to Parameterizing Geometric Models, Vijay Srinivasan 164. Handbook of Mechanical Alloy Design, George E. Totten, Lin Xie, and Kiyoshi Funatani 165. Structural Analysis of Polymeric Composite Materials, Mark E. Tuttle 166. Modeling and Simulation for Material Selection and Mechanical Design, George E. Totten, Lin Xie, and Kiyoshi Funatani 167. Handbook of Pneumatic Conveying Engineering, David Mills, Mark G. Jones, and Vijay K. Agarwal 168. Clutches and Brakes: Design and Selection, Second Edition, William C . Orthwein 169. Fundamentals of Fluid Film Lubrication: Second Edition, Revised and Expanded, Bernard J. Hamrock, Steven R. Schmid, and Bo 0. Jacobson 170. Handbook of Lead-Free Solder Technology for Microelectronic Assemblies, edited by Karl J. Puttlitz and Kathleen A. Stalter

Additional Volumes in Preparation

Mechanical Wear Fundamentals and Testing: Second Edition, Revised and Expanded, Raymond G. Bayer Engineering Design for Wear: Second Edition, Revised and Expanded, Raymond G . Bayer Progressing Cavity Pumps, Do wnhole Pumps, and Mudmotors, Lev Nelik

Mechanical Engineering Sofhyare Spring Design with an IBM PC, Al Dietrich Mechanical Design Failure Analysis: With Failure Analysis System Software for the IBM PC, David G. Ullman

Preface

The inspiration and need for this book arose from our task at IBM of implementing leadfree/reduced products to be in conformity with legislation proposed by the European Union (i.e. the RoHS Directive), and from the growing worldwide concern about the toxic agents utilized in electronics manufacture and products, particularly lead. The lead (Pb)-free implementation issue is of paramount interest to the electronics industry due to the magnitude of the undertaking and the many unanswered questions that must be addressed. It is an area that has been much discussed in trade press publications, at workshops, and it has been the subject of portions or entire international conferences. Given the importance and level of activity worldwide, it is both timely and appropriate to assemble the information surrounding the key aspects of the lead-free issue, particularly as they relate to electronic assemblies, from among the most knowledgeable practitioners and researchers in the world. Virtually every major electronics component or product manufacturer is involved in activities in response to both legislative and market pressures to eliminate toxic materials, particularly lead, from their products. The purpose of this book, therefore, is to serve as a handbook/primer for those who are contemplating or have already started the process of conversion, from standard Sn-Pb solder microelectronic assemblies to Pb-free. The book serves as a single source that covers all key aspects associated with this issue, including health, total environmental impact, legislation, market economics, and the materials, process, and reliability challenges; not just a narrow focus on a particular aspect. Accordingly, the book also serves as a valuable resource for those actively engaged in procuring, manufacturing, and utilizing lead-free products in all regions comprising the world market. The audience for this book includes almost everyone either peripherally or directly involved in bringing microelectronics products to the marketplace, and users as well. This book would certainly be a valuable resource for individuals engaged in producing and marketing low-end products such as cameras, camcorders, cell phones, pagers, games, lapand desktop computers, and other consumer products. Some of these products have been targeted for conversion to lead elimination/reduction and are already available in the marketplace. Typically, these products tend to be high-volume items, destined for disposal in landfills. Although high-end product disposal at end of life is likely to be by recycling, or refurbishing for the secondary market, individuals involved with high-end products will also find the book very useful. While this handbook thoroughly discusses the technological aspects surrounding lead-free implementation, it also provides a detailed understanding of lead-related health issues, tradeoffs and implications in changing over to Pb-free/reduced assemblies; and global regulations, consequences, dispensations, and exemptions with iii

iv

PREFACE

regard to utilizing lead. Since the book contains material that relates to a wide range of electronics products from low- to high-end applications, the book will be useful to a very large audience. The book will serve as a primer for newcomers in the field, as it does not assume prior knowledge of the subjects discussed and provides the necessary background information. It will be a valuable resource tool for those already in the field. Specifically, the book is intended for:

Scientists and engineers engaged in research, development and design, assembly, manufacturing, applications, quality, reliability, equipment engineers and testing Management, to gain a thorough understanding of the technology challenges, risks, and trade-offs Marketing, sales, intellectual-property-oriented individuals, and others peripherally involved, who require a working knowledge or a general understanding of the issues Students, for a thorough understanding of the forces that come into play in shaping business decisions such as ‘‘green’’ manufacturing, especially in electronics; materials technology fundamentals and issues, and potential solutions surrounding Pb-free solders This book is written in a manner that assumes the reader has some knowledge or facility with science or engineering-based principles, but does not assume prior knowledge of the material discussed here. In addition to serving as a handbook, it can be used for instructional purposes, with each chapter being a self-contained unit. That is, those key concepts or principles necessary to understand the material introduced in a chapter are briefly explained in the early sections of a chapter, or at the point of discussion. Also, brief segments may be introduced as restatements or summaries of concepts discussed in some detail in another chapter. This is done as a convenience to the reader, to avoid interruption in seeking information in other chapters. However, this book is well cross-referenced, providing the reader the opportunity for a more thorough explanation and understanding concerning a particular area of interest if desired. The introduction in Chapter 1 provides an overview of why the lead-free matter has become a worldwide issue, some concerns in introducing lead-free substitutes, and perspectives for a successful introduction. The book is divided into three parts, each addressing a major aspect related to the elimination of lead in electronic assemblies: the driving forces; the impact on manufacturing; and reliability-related factors. In Part I, ‘‘Background and Driving Forces,’’ Chapters 2 through 5 provide an understanding of the key issues associated with the worldwide lead-free movement, and the forces that are shaping and driving the movement. For example, Chapter 2 provides a detailed discussion of the toxic effects of lead on humans, how lead enters the body and its effect on various organ systems. Chapter 3 addresses the potential environmental impacts of substituting various materials utilized in microelectronic assemblies. In addition to health and environmental issues, there are also significant legislative drivers, both national and regional, discussed in Chapter 4 that manufacturers and users alike must be aware of because of the constraints and requirements they impose. The marketplace provides a significant driving force for lead-free products (Chapter 5), which is steadily increasing in importance due to worldwide identification with, and sentiment for, the removal of toxic materials in the manufacture of products. This driver may, in fact, ultimately require manufacturers to convert their products to lead-free long before legislation requires it.

PREFACE

v

In Part II, ‘‘Materials and Manufacturing Considerations,’’ Chapters 6 through 18 provide detailed discussions of the changes and impacts associated with various materials utilized, as well as manufacturing practices associated with the fabrication of microelectronic assemblies. Chapter 6 describes the key features of the Pb-Sn system, which the industry unanimously adopted worldwide and successfully utilized for the past 40 years, and serves as a reference or benchmark for comparing replacement systems. Chapter 7 defines the nature and mechanical properties that a solder material must collectively possess to meet the requirements imposed by solder joints in microelectronic assemblies at both the die and package levels. Chapters 8 and 9 describe the melting characteristics, microstructural features, morphology and other characteristics that determine the properties of interest for various binary, ternary and higher-order alloy systems considered as potential candidates to replace the Pb-Sn system. Although Pb-Sn solders possess many characteristics desirable for solder joints, they are not ideal, exhibiting low shear strength and creep resistance, particularly at elevatedtemperature application conditions. Chapter 10 describes several dispersion-strengthening methods for Pb-free solders to help overcome these traditional shortcomings of solder materials. Chapter 11 addresses the fundamentals governing wetting and spreading of liquids on solid surfaces and how they relate to lead-free solders on metallized terminal pads to assure sound connections. The mechanism and role of fluxing agents are also discussed. As a follow-on, Chapter 12 describes the metallurgical compatibility, processing advantages/disadvantages and reliability aspects associated with various types of lead-free board and component finishes in combination with several popular lead-free solder candidates under consideration. Chapter 13 discusses the method of formation, morphology, stability, and properties of intermetallic compounds (IMC) that form at solder/metallized surface interfaces that play an integral part in the quality and reliability of solder joints. Chapter 14 discusses the assembly differences that arise when utilizing lead-free solders as compared to eutectic Sn-Pb, and the impacts of those differences. Of particular interest are situations that necessitate and benefit from inert atmospheres (Chapter 15), which is a significant departure from standard eutectic Sn-Pb practice. Chapter 16 discusses some logistical and other impacts in changing over to a lead-free soldering assembly operation. The manufacturing experiences of various organizations that successfully introduced lead-free technology to their product lines are also reviewed. Numerous studies have been conducted by consortia around the world to determine which solder systems or alloys exhibited the most promise as suitable replacements for eutectic Sn-Pb solder. The major studies and their findings are reviewed in Chapter 17. It is anticipated that non-solder alternatives (Chapter 18) will play an increasing role in providing suitable, cost-effective interconnection solutions for some product segments, particularly in the area of consumer electronics. In Part III, ‘‘Reliability Considerations,’’ Chapters 19 through 22 describe a variety of reliability issues of concern in solders utilized in electronic assemblies, but particularly so for high Sn-based solders typical of lead-free solder alloys. Chapter 19 discusses general reliability concepts in relation to solder alloy composition, phasing, and morphology. Reliability aspects reported in the literature in relation to the use of lead-free alloys for flipchip and various package-level interconnections are also reviewed. Chapter 20 describes the factors associated with thermomigration and electromigration, which are of particular increased interest due to higher electrical power levels in flip-chip solder joints, and the introduction of lead-free solders. Tin-whisker formation and the potential for creating electrical shorts in the field is an issue mostly confined to Sn-rich finishes on components and boards. Chapter 21 describes how and why they form, and addresses the issue from the perspective that most lead-free replacement alloys are Sn-rich, and what the studies in this area have revealed. Finally, Chapter 22 provides a comprehensive explanation of various phe-

vi

PREFACE

nomena associated with Sn-based systems that can have significant reliability implications, among them, soft-error problems due to the presence of radioactive species, catastrophic degradation as a result of an allotrophic phase transformation, and structure-related issues. We are most grateful to our parents for their unending devotion, sacrifices, and love. They instilled the desire to seek the truth and enlightenment to better understand ‘‘how’’ and ‘‘why,’’ and to freely share our talents and knowledge with others in helping to give back what was given to us. We are also thankful to our spouses, Dianne Puttlitz and Mark Stalter, being ever mindful of their understanding and patience over the many months the manuscript was in preparation, and without whose support and encouragement the completion of this book would have been impossible. Finally, we are very appreciative of the effort put forth and information provided by the many contributors to this book who so generously shared their knowledge and expertise, and devoted so much of their precious free time to make this book a reality. Karl J. Puttlitz Kathleen A. Stalter

Contents

Preface

iii

PART I: BACKGROUND AND DRIVING FORCES 1. Overview of Lead-Free Solder Issues Including Selection Karl J. Puttlitz 2. Health and Environmental Effects of Lead and Other Commonly Used Elements in Microelectronics Edwin B. Smith III 3. Environmental Impact of Lead and Alternatives in Electronics Laura J. Turbini and Dennis Bernier 4. Environmental Stewardship with Regional Perspectives and Drives of the Lead-Free Issue Susanna Pelzel, Minna Juuti, and Yukio Sugimoto 5. Market, Product, and Corporate Policy Trends E. Jan Vardaman

1

49 83

115 149

PART II: MATERIALS AND MANUFACTURING CONSIDERATIONS 6. The Metallurgical Aspects, Properties, and Applications of Solders from the Lead–Tin System Paul T. Vianco

167

7. Physical Basis for Mechanical Properties of Solders Morris E. Fine

211

8. Sn-Ag and Sn-Ag-X Solders and Properties Karl J. Puttlitz

239

9. Bi-Sn, Sn-Sb, Sn-Cu, Sn-Zn, and Sn-In Solder-Based Systems and Their Properties Sung K. Kang vii

281

viii

CONTENTS

10. High-Temperature Lead-Free Solders with Dispersoids K. N. Subramanian, Sunglak Choi, and Fu Guo

301

11. Solder Wetting and Spreading Timothy J. Singler, Stephan J. Meschter, and James Spalik

331

12. Lead-Free Finishes for Printed Circuit Boards and Components Rob Schetty

431

13. Formation of Intermetallic Compounds at Pb-Sn/Metal and LeadFree/Metal Interfaces in Solder Joints Eric J. Cotts, Richard Chromik, Peter Borgesen, Robert Kinyanjui, and Anis Zribi

465

14. Electronics Assembly and the Impact of Lead-Free Materials Anthony A. Primavera

495

15. Use of Inert Atmospheres in Lead-Free Soldering Martin Theriault, Jason Uner, and Armin Rahn

569

16. Pb-Free Component Conversion and Some Manufacturing Experiences Kathleen A. Stalter, Tom Baggio, and Kenichiro Suetsugu

591

17. Major International Lead (Pb)-Free Solder Studies Carol A. Handwerker, Frank W. Gayle, Erik E. de Kluizenaar, and Katsuaki Suganuma

665

18. Electrically Conductive Adhesives–A Lead-Free Alternative Daoquiang Lu and C. P. Wong

729

PART III:

RELIABILITY CONSIDERATIONS

19. Reliability Aspects of Lead-Free Solders in Electronic Assemblies Puligandla Viswanadham, Steven O. Dunford, and Jorma Kivilahti

769

20. The Physics and Materials Science of Electromigration and Thermomigration in Solders James R. Lloyd, King-Ning Tu, and Jasvir Jaspal

827

21. The Structure and Kinetics of Tin-Whisker Formation and Growth on High Tin Content Finishes W. J. Choi, George Galyon, King-Ning Tu, and T. Y. Lee

851

22. Degradation Phenomena Michael J. Sullivan and Steve Kilpatrick Index Biographies

915

979 1015

Contributors

Tom Baggio

Panasonic Factory Automation, Tokyo, Japan

Dennis Bernier

Kester Northrop Grummen, Des Plaines, Illinois, U.S.A. Universal Instruments Corporation, Cleveland, Ohio, U.S.A.

Peter Borgesen, Ph.D.

Samsung Electronics, Seoul, South Korea

Sunglak Choi, Ph.D. W. J. Choi, Ph.D.

Intel Corporation, Santa Clara, California, U.S.A.

Richard Chromic, Ph.D. York, U.S.A.

State University of New York at Binghamton, Binghamton, New

Eric J. Cotts, Ph.D. York, U.S.A. Steven O. Dunford

State University of New York at Binghamton, Binghamton, New

Nokia Research Center, Irving, Texas, U.S.A.

Morris E. Fine, Ph.D.

Northwestern University, Evanston, Illinois, U.S.A.

George Galyon, Ph.D. New York, U.S.A.

Systems Development Group, IBM Corporation, Poughkeepsie,

Frank W. Gayle, Sc.D. Maryland, U.S.A.

National Institutes of Standards and Technology, Gaithersburg,

Fu Guo, Ph.D.

Michigan State University, East Lansing, Michigan, U.S.A.

Carol A. Handwerker, Sc.D. burg, Maryland, U.S.A. Jasvir Jaspal Minna Juuti

National Institute of Standards and Technology, Gaithers-

IBM Corporation, Hopewell Junction, New York, U.S.A. IBM Environmental Affairs, Sweden ix

x

CONTRIBUTORS

Sung K. Kang, Ph.D. T.J. Watson Research Center, IBM Corporation, Yorktown Heights, New York, U.S.A. Steve Kilpatrick, Ph.D. Physical Scientist, Sensors and Electron Devices Directorate, Micro-Devices Branch, U.S. Army Research Labs, Adelphi, Maryland, U.S.A. Robert Kinyanjui U.S.A.

State University of New York at Binghamton, Binghamton, New York,

Jorma Kivilahti, Sc.D.

Helsinki University of Technology, Helsinki, Finland

Erik E. de Kluizenaar, Ph.D. T. Y. Lee, Ph.D.

Philips CFT, The Netherlands

Hanbat National University, Daejeon, Korea (South)

James R. Lloyd, Ph.D. T.J. Watson Research Center, IBM Corporation, Yorktown Heights, New York, U.S.A. Daoquiang Lu, Ph.D.

Georgia Institute of Technology, Atlanta, Georgia, U.S.A.

Stephan J. Meschter, Ph.D. U.S.A.

BAE Systems Platform Solutions, Johnson City New York,

Studio Legale Paggi, Venice, Italy

Susanna Pelzel, J.D.

Anthony A. Primavera, Ph.D. U.S.A. Karl J. Puttlitz, Ph.D. New York, U.S.A. Armin Rahn, Ph.D. Rob Schetty

Universal Instruments Corporation, Cleveland, Ohio,

Corporate Environmental Affairs, IBM Corporation, East Fishkill,

Consultant, St. Catherines, Ontario, Canada

Advanced Technology Division, Technic Inc., Plainview, New York, U.S.A.

Timothy J. Singler, Ph.D. New York, U.S.A. Edwin B. Smith III

State University of New York at Binghamton, Binghamton,

Hart InterCivic, Lafayette, Colorado, U.S.A.

James Spalik, Ph.D. IBM Corporation, Endicott, New York, U.S.A. Kathleen A. Stalter U.S.A.

Technology Group, IBM Corporation, East Fishkill, New York,

K. N. Subramanian, Ph.D. Department of Chemical Engineering and Material Science, Michigan State University, East Lansing, Michigan, U.S.A. Kenichiro Suetsugu, Ph.D.

Matsushita Electric Industrial Co., Ltd., Tokyo, Japan

CONTRIBUTORS

xi

Katsuaki Suganuma, D.Eng. Yukio Sugimoto

IBM Corporation, Tokyo, Japan

Michael J. Sullivan, Ph.D. Martin Theriault

Consultant, Red Hook, New York, U.S.A.

Air Liquide America Corporation, Houston, Texas, U.S.A.

King-Ning Tu, Ph.D. U.S.A.

University of California at Los Angeles, Los Angeles, California,

Laura J. Turbini, Ph.D. Jason Uner

Osaka University, Osaka, Japan

University of Toronto, Toronto, Ontario, Canada, U.S.A.

Air Liquide Corporation, Houston, Texas, U.S.A.

E. Jan Vardaman

TechSearch International, Inc. Austin, Texas, U.S.A.

Paul T. Vianco, Ph.D.

Sandia National Laboratories, Albequerque, New Mexico, U.S.A.

Puligandla Viswanadham, Ph.D.

Nokia Mobile Phones, Inc., Irving, Texas, U.S.A.

C. P. Wong, Ph.D. School of Material Science and Engineering, Georgia Institute of Technology, Atlanta, Georgia, U.S.A. Anis Zribi, Ph.D. U.S.A.

State University of New York at Binghamton, Binghamton, New York,

1 Overview of Lead-Free Solder Issues Including Selection Karl J. Puttlitz IBM Corporation, East Fishkill, New York, U.S.A.

I.

INTRODUCTION

The use of lead (Pb) has been widely accepted in the electronics industry (utilized in a variety of applications for more than five decades), but most notably the use of eutectic tin–lead (Sn–Pb) solders to attach discrete components to printed circuit boards (PCBs). These attachments typically also serve as the electrical interconnection between the attached components and the PCB. Lead-based solders are also used as a coating or finish on metal terminations on PCBs, and on the peripheral leads of components.

A.

Lead Utilization and Consumption

Approximately 5 million tons of Pb are consumed annually on a global basis, of which storage batteries represent the largest single application (about 81%) by far. Lead oxides (for glass, ceramic products, chemicals, etc.) and ammunition together account for about 10%, and solders for electronic assemblies account for only about 0.5%, as noted in Table 1 [1]. Although storage batteries constitute a large percentage of consumption due to an alleged near-100% recycling practice, batteries have not been targeted for a ban because they are believed not to contribute to pollution or contamination, and hence do not pose a problem [2]. The largest use of Pb in electronics is in the manufacture of cathode ray tubes as a component of the glass that protects TV and computer monitor viewers from undue exposure to x-radiation. Only about one tenth of the total (0.05%) is currently used in solder alloys [3]. There are laws in the United States that control the use of lead in gasoline and plumbing solder, and banned its use in consumer paint since 1978 [4].

B.

Efforts to Restrict Lead Use

However, due to an increased awareness of the potential health hazards associated with the toxicity of lead in humans, which can lead to disorders of the nervous and reproductive systems, and also affect neurological and physical development, actions have been taken to eliminate or reduce the use of Pb in a variety of products. Among these are solders used for plumbing, and additives in paint and gasoline since the late 1970s. Banning or restricting the use of Pb did not become an issue until the early 1990s in the United States, but the proposed legislation to ban Pb was never enacted into law due, in large part, to pressures brought about by the electronics

1

2

PUTTLITZ

TABLE 1

Lead Consumption by Product

Product Storage batteries Oxides used for glass, ceramic products, pigments, and chemicals Ammunition Sheet lead Cable covering Casting metals Brass and bronze billets and ingots Pipes, traps, and other extruded products Solder (excluding electronic solder) Electronic solder Miscellaneous

Consumption (%) 80.81 4.78 4.69 1.79 1.40 1.13 0.72 0.72 0.70 0.49 2.77

Source: Ref. 1.

industry. However, there has been a growing worldwide movement to eliminate the use of lead in products due mainly to health-related concerns. This will now be mandated through European Union (EU) legislation banning the use of lead in electronics products with some exemptions starting July 1, 2006. Environmental issues, particularly lead elimination, had not received much attention in the electronics industry until the late 1990s when the initial Waste from Electrical and Electronic Equipment (WEEE) Directive was drafted by the EU. It became clear that responsible environmental practices can have a positive impact on business. Failure to embrace prescribed environmental practices would result in a degraded competitive posture and an erosion in market share. Cleaning procedures that were, at one time, an integral part of PCB assembly today have mostly converted to no-clean processes. This shift, like many others, benefits the environment while also contributing to the bottom line [5].

C.

This Handbook

This handbook discusses all the key aspects of the lead-free issue as it relates the implementation of lead-free soldering technology in the electronics industry. Among those areas discussed in detail are the background issues that give rise to the driving forces for lead-free implementation; the factors related to selecting alternative material systems such as flux, printed circuit board, board/component lead finishes, and lead-free solder alloys suitable for both reflow and wave soldering; and also manufacturing and reliability-related considerations.

II.

DRIVING FORCES FOR LEAD-FREE SOLDERS

A.

Health-Related Aspects

Candidate solder materials to replace eutectic Sn–Pb solders must satisfy toxicity regulations and limits set by the government. U.S. occupational exposure limits for selected materials are listed in Table 2 [6]. These are generally the same as for other industrialized countries (e.g., UK limits are the same for inorganic Pb) [6]. Despite the long-term utilization of Pb, it is now well recognized that Pb can have some very serious ill effects on humans. Depending on the aspects, Pb poisoning is sometimes classified as alimentary, neuromotor, or encephalic. Lead poisoning is the result of prolonged exposure to Pb or its compounds.

LEAD-FREE SOLDER ISSUES INCLUDING SELECTION TABLE 2

3

Occupational Exposure Limits of Selected Materials

Material Zinc oxide fume Inorganic tin Copper dust Antimony and compounds Copper fume Organic tin Indium Silver dust and fume Silver, soluble compounds Inorganic lead Bismuth Gallium

U.S. Occupational Safety and Health Administration, permissible exposure lights (mg/m3) 5 2 1 0.5 0.1 0.1 0.1 0.1 0.01 0.05 Not assigned Not assigned

Source: Ref. 6.

1. Contaminated Drinking Water Historians have speculated that Pb poisoning played a role in the collapse of the Roman Empire because of the widespread use of Pb for drinking water pipelines [7]. However, this source of Pb was recognized early on as a source for ingesting Pb, so the use of Pb pipes and lead-containing solders for plumbing of drinking-water lines has been prohibited by federal legislation in the United States (Section 1417, Amendment to the Safe Drinking Water Act, June 19, 1986). a. Leachate from Landfills. Based on similar concerns, the electronics industry is quickly moving ahead with lead-free assemblies and components to avoid potentially Pb-contaminated drinking water. This time, however, the source is alleged to be Pb leached from landfills containing electronic wastes (lead from component finishes, printed circuit board assemblies, etc.) of consumer products such as TVs, PCs, telephones, camcorders, handheld games, refrigerators, microwaves, etc. It is contended that the leachate would find its way into the groundwater system with the potential of contaminating municipal water supplies, as illustrated in Fig. 1. b. Lead Concentration Limits. In Japan, the Pb elution environmental standard in landfills in 0.3 mg/L. In toxic materials detection tests conducted by the Japanese Environmental Agency, it was determined that the amount of lead leaching from the pulverized remains of TV cathode ray tubes, PCBs from PCs, and pachinko machines far exceeded the environmental standard [8]. In the United States, the regulatory limit for Pb in drinking water is 0.015 mg/L (EPA 40 CFR 141). In utilizing the Toxicity Characteristics Leaching Procedure (TCLP) per EPA 40 CFR 261, the concentration limit is 5 mg/L [2]. It has been demonstrated that the concentration of lead leached from solders can be several hundred times higher than the acceptable limit [9].

B.

Legislation

There are several forces driving the movement toward the elimination of Pb. An initial driving force is the fear of legislation, the most important of which is the EU’s proposal (RoHS Directive) to ban lead in all electronics products, including information technology (IT) products such as monitors, printers, and computers. The basis for the legislation is the allegation that Pb could be leached from electronic wastes discarded in landfills, and that the leachate will find its way into the groundwater system, ultimately contaminating the drinking water supplies of municipalities.

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PUTTLITZ

FIG. 1 An illustration depicting the sequence of events that results in the ingestion of lead leached from electronic wastes in landfills. (From Ref. 69. Courtesy of Panasonic Corporation.)

This legislation, which will have a major impact on the electronics industry, is a precautionary measure because there are no known instances of water supplies having been contaminated in this manner. Nevertheless, there is a growing concern about electronic wastes and their Pb content because they are increasing at a rate three times greater than other solid wastes [10].

C.

Market Advantages

Japan has taken aggressive steps in instituting recycling laws to reduce the volume of wastes to be disposed. Because Japan is a small, highly populated country, it does not wish to utilize its precious space for landfills. Accordingly, it requires many products at end-of-life to meet prescribed specifications for disposition, including the removal of lead-bearing parts. To help reduce disassembly costs at end-of-life, Japanese original equipment manufacturers (OEMs) took the approach of introducing lead-free technology into their assemblies, particularly the solders used to reflow attach components to printed circuit boards, and lead-free components and board finishes. This strategy put the Japanese OEMs at the forefront of lead-free implementation. They have made lead-free products available in several niche markets, such as laptop and desktop computers, and some handheld consumer electronic products. Many observers of the lead-free solder issue feel that the market-driven aspect is more important than legal mandates because the market can force a transition to lead-free solders much more quickly as it can provide a substantial marketing advantage in some geographies, particularly for consumer products. Japanese electronics manufacturers have taken the lead in ‘‘green’’ marketing and manufacturing. Consider, for example, a lead-free minidisc player made by Panasonic, bearing a green leaf environmentally safe symbol that was released in October 1998. The product gained a significant increase in market share, moving from 4.7% of the minidisc player market to 15% in a very short time. Much of this success was attributed to the product’s green status. In Japan, products bearing a green sticker are typically not more expensive than competitors’ lead-bearing products—in some instances even less expensive. It appears that there are tradeoffs that can be made to work around cost issues [11]. Although not as pervasive as in Japan, there have been some early lead-free activities in both Europe and the United States as well. For example, the Ford Motor Company launched a large

LEAD-FREE SOLDER ISSUES INCLUDING SELECTION

5

green marketing campaign, having declared targets of 2002 for lead-free electronic assemblies, and totally lead-free vehicles by 2004 [12]. Green marketing is rapidly being viewed as a powerful and effective marketing tool. Many respected companies attribute some of their success to having taken a proactive environmental approach. Corporate executives are becoming increasingly aware that benefiting the environment can lead to bottom-line benefits. Among the benefits are a reduction in expenses and the creation of new sources of revenue [13,14].

D.

Mechanical Properties

There are technology-based reasons for eliminating Pb–Sn solders. The trend for continued densification in electronic packages is driving the need for even tighter pitches, which, correspondingly, requires finer-line circuitry. Because this trend results in solder joints with smaller cross sections, it also requires solders whose inherent mechanical properties are better suited to withstand the increasingly severe mechanical demands compared to eutectic Sn–Pb solder. It has been reported that many lead-free candidate solders exhibit significantly better strength and fatigue life properties than eutectic Sn–Pb [6].

E.

International Standards Organization (ISO) 14001 Certification

The international community is becoming ever more environmentally conscious, causing manufacturers and their materials and equipment supplies to develop products and processes that reduce the impact on the environment. The two primary factors are: the growing support for and embracement of the worldwide lead-free movement, and ISO 14001 certification. The certification is a standard for environmental management published by the ISO in 1996. It emphasizes establishing management systems that ensure environmentally sound business practices as opposed to focusing on specific performance targets. ISO 14001, whose attributes and features are listed in Table 3, can also be tailored to individual companies owing to its generic structure. The intent is to integrate the environmental management system (EMS) with companywide business management practices and decisions [15]. The certification supports techniques that are environmentally friendly. Electronics manufacturers are turning to vendors that can provide products that meet ‘‘green’’ standards. For example, the implementation of ISO 14000 prevents volatiles from flux to escape into the atmosphere, a problem aggravated by lead-free solders due to the higher reflow temperatures of furnaces. The reduced flux volume that results in enhanced throughput and efficiency also serves to preserve the internal cleanliness of reflow furnaces by minimizing interruptions for maintenance purposes. Some flux separation and collection systems can remove 95% of flux residues from an oven interior, and extend maintenance for 30–90 days in high-volume operations [16].

TABLE 3

Attributes and Features of Environmental Standard ISO 14001

Standard Sponsor Purpose of standard Emphasis Structure Intent Advantage

Description International Standards Organization Provide an environmental management system Environmentally sound business practices Can be tailored to individual organizations, big or small Integration of environmental management system with company-wide business practices and decisions Benefits the environment, and users are increasingly turning to products that meet ‘‘green’’ standards

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PUTTLITZ

TABLE 4

Some Useful Standards Related to Lead-Free Technology

Standard Association of Connecting Electronics Industries, IPC-J0Std-006 ICP-J-Std-002, 003 MIL-M-38510 JEDEC J-Std-620A

F.

Subject Detailed list of both tin–lead and lead-free solders Board and component lead solderability Plating thickness for good component lead solderability Plastic package moisture sensitivity, heat tolerance

Standards/Practices/Information

Because lead-free technology entered the scene relatively recently, practices are still being established and few standards exist. Some useful standards that do exist are listed in Table 4.

III.

WORLDWIDE ACTIVITIES/PRACTICES

A.

Early Initiatives

Some of the efforts to ban lead were initiated in the U.S. Congress. In 1991, the U.S. Senate proposed the Reid Bill (S.391, Lead Exposure Reduction Act), which proposed restrictions on lead solders, limiting the lead content to 0.1% or less. Subsequently, the bill was modified (S.729, 1993) to propose a ban on all lead-bearing alloys including electronic solders, and a tax of US$1.69 per kilogram on primary lead and US$0.83 per kilogram on secondary lead used in the industry. Although Pb solders were ultimately removed from the bill due to intense lobbying by the U.S. electronics industry, it serves to provide global awareness that Sn–Pb solders may, at some point, be restricted. As a result, many companies worldwide instituted investigations addressing the feasibility of a variety of Pb-free solders [2]. Environmental consciousness was firmly in place in industries, including the electronics industry, which responded to several challenges, most notably the ban on chlorofluorocarbon (CFC) cleaning agents. Prior to 1992, flux residues from soldering operations were removed mainly by industry choice. After the Montreal Protocol and subsequent Amsterdam Regulations that banned manufacturers from using ozone-depleting chemicals effective January 1, 1996, the industry responded by implementing suitable solutions [17].

B.

Europe

In 1994, Denmark, Sweden, Norway, Finland, and Iceland signed a pact to phase out Pb, with Swedish manufacturers under a voluntary ban effective 2000. There are producer responsibility laws in effect in Holland, Switzerland, and Norway. In 1997, the Swedish government identified Pb as an element it intended to eliminate from products over the following 10 years. The Sweden Environmental Quality Objectives direct that new products, including batteries, introduced in Sweden should be mostly Pb-free by 2020 [18]. In 1991, the UK, in collaboration with GEC, BNR-Europe, Multicore Solders, and ITRI, and supported by the Department of Trade and Industry (DTI), undertook a project to make an initial assessment of alternative solders for electronic assemblies. That was followed by a European collaboration in 1994 to address issues associated with the implementation of Pb-free soldering technologies as did several large companies in the UK. 1. WEEE Directive The EU introduced a draft directive (law) called the WEEE, which called for a ban on Pb in all electronics, except automotives. The scope of the ban is far-reaching, banning the selling or


7

import of electronics containing lead, normally as part of interconnections that attach components to printed wiring boards (PWBs). But Pb can also be present as part of finishes on cards/ boards, peripheral leads of components, or materials used to fabricate components and peripherals (e.g., monitors). 2. Take Back/Recycling In parallel with initial investigations of lead-free alternatives that commenced in the early 1990s, there were also recycling laws proposed, mostly in Asian countries, that addressed the related issue of an ever-increasing volume of end-of-life electronic wastes to be disposed in landfills. Products must be designed for disassembly to remove parts containing environmentally objectionable materials, including Pb. Various product ‘‘take-back’’ schemes involving electronic wastes have been devised whereby manufacturers of these products are mainly responsible for the collection of wastes from designated locations, businesses, and households when owners decide to discard electrical and electronics products. So-called ‘‘producer responsibility’’ laws for electronics and electrical waste equipment are anticipated to be widespread in Europe. Such laws have already been in effect for some time in Holland and Switzerland, and are also in effect in Norway (1999) and Sweden (2000). Threshold limits for recycling specified materials are typically defined [2].

C.

Asia Pacific

There is a wide disparity in the approach to environmental matters among Asian countries, from virtually no regulations to strict legislation diverting discarded electronic products from landfills. In those cases where regulation is lax, it is most often due to a corresponding lack of infrastructure. For example, in 1995, Thailand had a staff of 43 to oversee 50,000 industrial plants. A reliance on self-regulation of the private sector is not uncommon in these areas where the trend is to adopt environmental standards such as ISO 14000, the environmental impact assessment required by the World Bank and the Asian Development Bank. Although ‘‘command-andcontrol’’ regulations will likely increase in the United States, Europe, Japan, and South Asian countries, other Asian countries will more likely use market-driven regulation and self-regulation because they have experience with these methods and lack necessary resources [19]. 1. Recycling Due to the very significant involvement on the part of the electronics industry, there is a high degree of interest and activity in the Pb-free issue in Japan, Singapore, South Korea, and Taiwan. For example, there has been a recycling law in effect in Taiwan since March 1998 requiring that all electronics products undergo proper waste processing at a product’s end-of-life. Japan enacted the ‘‘Electronics Appliance Recycling’’ law in June 1998 requiring appliance manufacturers by 2001 to collect and recycle their own products. Japan appears to be leaning toward high-technology disassembly coupled with material recovery (i.e., recycling) with many private-operated and government-operated recycling plants coming on-line. An example of these is the highly automated plant that went into operation in April 1998 that has the capacity to service a population center of about 1 million people, equivalent to about 150,000 units (TVs, washing machines, etc.) [19]. So as the EU is focused on a Pb ban, the Japanese Ministry of International Trade and Industry (MITI) is approaching the issue quite differently through socalled ‘‘take-back’’ legislation (i.e., recycling). In the European Directive, the manufacturer or importer is responsible for the cost of recycling [20]. 2. Lead Reduction/Elimination Whereas the official Japanese approach to the ‘‘Pb in landfills’’ issue is recycling, Japanese OEMs, and material and component suppliers, apparently motivated by the benefits of reducing disassembly costs at end-of-life by eliminating lead-containing parts and the specter of significant potential market opportunities, are moving in the Pb reduction/elimination direction. Early in 1998, the Japanese Electronics Industry Development Association (JEIDA) and the Japanese

8

PUTTLITZ

Institute of Electronics Packaging presented a Pb-free roadmap. Japanese OEMs announced their own plans to drastically reduce Pb levels in their products by 2002, and some even earlier [21]. The targets of several Japanese OEMs who were among the first to promote and adapt leadfree manufacturing to their products are listed in Table 5.

D.

North America

Since the initial attempts in congress to pass lead-free legislation in relation to electronic products, there has been little activity in the United States until the late 1990s, when it was clear that the Japanese were well positioned to exclude products from Japan that did not meet standards they might require owing to their early lead-free activities. Furthermore, because some Japanese lead-free products were being offered in some niche markets, it demonstrated the viability and manufacturability of lead-free technology, thus justifying the EU-proposed legislation to ban the use of Pb in electronics products. 1. Low Activity Level Early On Significant levels of lead-free activities were carried out over a wide range of consumer and industrial electronic products in the early 1990s. These initiatives were gradually abandoned in many areas due to the lack of legislative pressures to pursue alternatives to lead-containing products and materials, particularly solders for PWB assemblies. A notable exception was the automobile industry, where companies such as Delco, Ford, and Motorola continued to investigate and qualify lead-free products. Some of these activities were motivated by the desire to achieve higher strength and greater fatigue resistance characteristic of most lead-free alloys compared to eutectic Sn–Pb at higher service temperatures. Nortel Networks was one of the few North American companies to utilize lead-free solders in consumer electronics (Meridian phones). Texas Instruments (TI) provided Ni/Pd finishes for its component leads early on, replacing Sn–Pb coatings. TI also replaced the use of high-Pb solder for chip attachments to components with Ag-filled epoxies. These joints, however, are not suitable for high-power applications [19]. 2. Drivers Motivate Activity Aside from these few exceptions, it was both market-driven concerns and legislative concerns that served to revive interest in the lead-free issue among manufacturers in the United States. Although industry associations initially lobbied against the issue, they eventually supported it since the movement gained support worldwide. In fact, National Electronics Manufacturers Initiative (NEMI) initiated a very ambitious and aggressive lead-free effort that sought to select a suitable lead-free alloy, define a process window, produce populated cards/boards, and demonstrate the reliability of the solder joints of a broad variety of components through standard stress testing.

TABLE 5 Products

Several Early Manufacturers to Initiate Lead Reduction/Elimination in a Variety of

Company Hitachi

Action

Year

50% reduction from 1997 levels 1999 Eliminate Pb interconnects April 2001 Toshiba Elimination in mobile phones 2001 Sony All Pb eliminated, except high density electronics 2001 Matsushita (Panasonic) All Pb from products 2001 Been producing 40,000 minidisc players per month Pb-free October 1998 NEC 50% reduction from 1997 levels 2002 NTT No Pb or Cd in newly purchased equipment Now Source: Ref. 21.


9

An IPC statement probably reflects the opinion of most U.S. manufacturers, ‘‘. . . lead in electronics is not perceived as a health issue, but government and commercial drivers will push for its adoption anyway. Thus IPC will facilitate activities to enable it to happen’’ [22].

IV.

LEAD-FREE SOLDERS AND SELECTION CONSIDERATIONS

A.

Solder Alloys

1. Brief History The use of solders extends back to about 4000 BC and, as indicated in the timeline of Fig. 2, served as an important factor through the Bronze and Iron Ages up to the current so-called ‘‘Silicon Age.’’ The oldest archaeological evidence relates to gold-based hard solders since gold is impervious to corrosion, so an ample number of objects have been discovered. The earliest of these date back to about 4000 BC when gold-based hard solders were developed by artisans in Mesopotamia [24]. Both the Celts and the Gauls performed high-quality soft soldering using the same solder alloy. The use of Sn–Pb solders in ancient cultures is much more obscure, but by 350 BC, the Greeks sealed bronze-based water pumps and air pumps with Sn–Pb alloys, and the Romans used Sn–Pb solders extensively in the construction of aqueducts [25] and to close the seams of lead water pipes. They used an alloy, ‘‘terrarium,’’ consisting of one part Sn and two parts Pb, still used today for that purpose. There are references to lead–tin soldering going back to biblical days, where its use to solder water pipes is mentioned in the Old Testament [26]. There were few changes to soldering practice during the Iron Age and Middle Age; however, its use was greatly expanded during the Industrial Revolution, such as the use of soft solders for plumbing (i.e., copper tubing). Early in the 20th century, soldering was introduced to the electronics industry as a reliable method of connecting copper wires. More recently, the solder joint took on the dual function of serving both a mechanical and electrical connection, challenging the strength, creep, and fatigue resistance of Sn–Pb solders [23]. In electronics, soldering continues to be the predominant means to attach devices to printed circuit boards or microelectronic chip carrier packages.

FIG. 2 Solder utilization time line. (From Ref. 23.)

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PUTTLITZ

2. Definition and Classification Soldering may be defined as the joining of metals by utilizing another metal, which has a melting point or range below about 450jC (1000jF) and is lower than the metals to be joined. This differs from brazing, in which the added metal generally has a melting range above 450jC (1000jF), but also below the metals to be joined [27]. Fig. 3 illustrates the melting ranges of some typical solder and braze materials. Solders are typically classified as either soft or hard. Some confusion often results from this convention because some hard solders are really braze materials if the melting temperature criterion is applied. Soft solders typically consist of alloys containing lead and tin, but also often contain indium (In), bismuth (Bi), antimony (Sb), or silver (Ag). In practice, most soft alloys melt at temperatures lower than 450jC, usually between 180jC and 300jC. High-tin solders, typical of lead-free solders, tend to be stiffer, harder, and less ductile compared to high-lead solders. Hard solders often contain metals such as Au, Zn, Al, and Si. 3. Traditional Electronics-Based Solders Lead-bearing solders had become the largest group of alloys utilized in electronics products; among these were eutectic (63Sn–37Pb) and near-eutectic (60Sn–40Pb) solders used extensively for PCB assemblies. The melting point or range of these alloys is sufficiently low to maintain the integrity of both components and laminate materials, such as FR-4, typically used for PCBs. There are some specialized applications of several high-Pb alloys (Table 6) utilized as flip-chip solder bumps (e.g., 97Pb–Sn, 95Pb–Sn), and as high-melt balls or columns (90Pb–10Sn) for ball

FIG. 3 Melting ranges for fusible alloys. (From Ref. 27.)

LEAD-FREE SOLDER ISSUES INCLUDING SELECTION TABLE 6

11

Some Solders Utilized for Electronic Assemblies

Application

Solder alloys

Flip-chip solder bumps BGA solder bumps

Soldering to thick film metallizations (e.g., Au–Pd, Au–Pt, Ag–Pd) Soldering to thin films (e.g., Cu, Ni, Au) Low-temperature operations (e.g., attaching temperature-sensitive components, last step in a temperature hierarchy)

Pb–3Sn, Pb–5Sn, Pb–50In, eutectic Sn–Cu Dual-solder bumps: Pb–10Sn (ball) and eutectic Sn–37Pb (ball and card attachments) Single-solder bumps; eutectic Sn–37Pb, eutectic Sn–Pb–2Ag, Sn–Ag–Cu Pb–10Sn, Pb–50In Eutectic Sn–37Pb, Au–20Sn Eutectic Sn–Bi

and column grid array components attached to cards or boards with a low-melt solder (eutectic Sn–Pb) [28]. In all Pb–Sn alloys, Pb is not a metallurgically active species (i.e., does not enter into any reactions with bonding pads or surface metallizations), so that there are no interfacial intermetallic compounds (IMCs) containing Pb, nor any such compounds dispersed within the bulk solder. The presence of lead, however, affects the solid-state growth kinetics of Sn–Cu intermetallic compounds. These factors are all discussed in detail in Chap. 6. The presence of Pb also helps to prevent the white tin (h)-to-gray tin (a) allotropic phase transformation from occurring in high-Sn solders upon cooling below 13jC. The loss of structural integrity that accompanies this transformation during cold-temperature service is referred to as tin pest [29]. The details of this phenomenon are discussed in Chap. 22. There are several non Pb–Sn solders utilized in electronics for specialized applications. One of these, 50Pb–50In, is utilized for hybrid microcircuit applications involving soldering to precious metal (Au–Pd, Au–Pt), thick film, and thin film circuits [30]. One advantage of this solder is that its Au thieving (dissolution rate) is much slower than Sn–Pb solders. a. Role of Lead in Lead–Tin Solders. Lead plays a very important and unusual role in Sn– Pb solder alloys in general, as noted in Table 7, and in solder joints of electronic assemblies in particular. As noted earlier, Pb does not participate in the formation of any intermetallic compounds, neither at the terminal ends of solder joints nor distributed within the bulk solder. Pb is effectively a ‘‘neutral diluent’’ for Sn. It reduces the melt temperature of alloys with Sn, and improves the mechanical characteristics of Sn as a solder joint material (i.e., reduces stiffness, increases plasticity, etc.) without significantly affecting solder chemistry (i.e., interactions between Sn and terminal metallizations, namely Cu and Ni) [31].

TABLE 7

Role of Lead (Pb) in Tin–Lead Solders

Feature Melting point Surface tension Activity Cost

Comments Reduces the melting point of Sn Reduces surface tension of molten Sn–Pb solders, providing their characteristic good wetting properties Does not participate in reactions with circuit metallizations (e.g., Cu, Ni) Intermetallic compounds formed by reaction with Sn Trade off for Sn content reduces alloy cost

12 B.

PUTTLITZ Lead-Free Definition

There is no universally accepted lead-free definition. In fact, even the EU has not provided (and is not required to provide) a definition until 1 year after the RoHS Directive goes into effect. However, various industry organizations have adopted a 0.1% maximum residual lead content. This value appears somewhat arbitrary (i.e., not based on some health-related threshold value).

C.

Selection Criteria for Suitable Lead-Free Solder Candidates

Lead-free solders must conform to a broad range of properties and characteristics to be considered as suitable candidates to replace eutectic Sn–Pb solders for use in electronic assemblies [31]. Lead-free solder candidates must also be compatible with product packaging trends of everincreasing densification to meet the demands for more functionality and portability. Decreases in package size correspondingly result in smaller solder joint interconnections. While undergoing size reductions, solder joints of high-density, high-performance electronic packages are expected to withstand increasing mechanical, electrical, and thermal burdens, which can lead to reduced reliability and even failure during manufacture or in the field [32]. There are two basic requirements that all solder interconnection alloys must satisfy. First, they must wet and bond to metallic terminals or features on components and substrates. This implies that the solder must be capable of forming intermetallic compounds with Cu, Ni, and other metals routinely utilized in microelectronic packages. Second, a solder must have a melting point that is sufficiently low to be reflowed as a paste when surface mount reflow attaching components to printed circuit boards, yet high enough to avoid any ill effects due to operating temperatures in the field. The most important characteristics that must be considered in selecting suitable lead-free solder candidates are listed in Table 8. Many of the lead-free solder systems of interest consist of some minor additions of third or fourth elements to binary alloy systems to achieve a desired benefit such as enhanced wettability or reliability, etc. 1. Toxicity Clearly, the toxic metallic elements cadmium (Cd), thallium (Th), and mercury (Hg) can be eliminated immediately. It should be noted that some geographies (e.g., Japan, Scandinavia) regard Sb as toxic [33]. Although the metal itself is generally not thought to be toxic, the trioxide compound—not typically formed at soldering temperatures but can be formed at incineration temperatures—is considered toxic. Because incineration is widely practiced in Europe and Japan to dispose wastes as opposed to using landfills, alloys containing Sb are not likely candidates for widespread application.

TABLE 8

Criteria and Desired Attributes for Candidate Lead-Free Solders

Criteria and desired attributes . . . . . . .

Not toxic Available in sufficient quantities to meet current and future requirements Exhibit sufficient electrical and thermal conductivity Possess adequate mechanical properties: strength, toughness, fatigue and creep resistance Compatible with typical terminal metallizations (e.g., Cu, Ni, Ag, Au, Sn, etc.) Affordable (i.e., economically viable) Have acceptable melt and process temperatures Not too different from eutectic Sn–Pb Avoid deleterious temperature effects on cards, boards, and components Low dressing characteristics


13

2. Supply and Abundance The number of viable elements to replace Pb is significantly reduced when factoring in the condition of sufficient abundance and availability. Availability refers to the quantity that is produced and therefore in most cases is available to manufacture products, whereas abundance refers to world reserves in the Earth’s crust. It should be noted that politics and a region’s stability can play an important role (i.e., a particular element can be both sufficiently abundant and be in substantial supply, but not readily available due to manipulation by cartels or restrictions imposed by governments). Reserve implies that, in addition to its state of abundance, the metal can be economically extracted by utilizing current technologies at the posted market price. Availability is an important consideration, as can be deduced from Table 9, which provides a comparison for those elements most likely to be substituted for Pb. Consider, for example, the case of Bi. If the entire world’s supply of Bi were utilized for microelectronics applications, only about 7800 tons of eutectic Sn–Bi could be produced, or only a maximum of 8% Bi could be added to Sn to produce a weight of solders equivalent to the use of eutectic Sn–Pb in electronics, about 60,000 tons. Thus, an alloy that contains a significant amount of scarce elements is not a viable replacement for eutectic Sn–Pb solder [6]. 3. Cost Because lead is one of the least expensive elements on Earth, any substitute will have some impact on cost, particularly so because it is likely that the substitutes will be from the group of elements Sn, Ag, Cu, and In. If either In or Ag is utilized, their contents must be kept low because both are relatively expensive when compared to lead, as noted in Table 10. Besides idium’s high cost (typically in the range of US$200.00 per kilogram), it also has a supply problem. Only about 200 tons of indium are produced globally each year [36]. a. Metal Density Factor. There is a significant difference on a cost-per-pound basis between the elements that comprise most of the leading lead-free alloy candidates and the baseline eutectic Sn–Pb solder. Solders are purchased on the basis of weight, not volume. For example, solder paste is typically purchased by the kilogram, and solder wire or bar is purchased by the pound. The candidate substitute metals are much less dense compared to Pb (Table 11). Similarly, Table 12 shows that the density of the most popular Pb-free candidate alloys is much less than eutectic Sn–Pb solders. For example, the popular lead-free alloy, Sn–3Ag–0.7Cu, is nearly 20% less dense. The implication is that 1 kg of eutectic Sn–Pb yields about 1000 jars of solder paste,

TABLE 9 Cost and World Reserves of Metals Utilized in Microelectronic Assemblies Metal Antimony (Sb) Bismuth (Bi) Copper (Cu) Gold (Au) Indium Lead (Pb) Nickel (Ni) Palladium (Pd) Platinum (Pt) Silver (Ag) Tin (Sn) Zinc (Zn) Source: Ref. 34.

World reserves (thousand metric tons)

Price (US$/lb)

3,200 260 650,000 77 6 130,000 150,000 79,000 — 420 12,000 440,000

0.68 3.50 0.89 4480 85.45 0.44 3.91 5700 8602 76.54 3.70 0.65

14

PUTTLITZ

TABLE 10 Relative Cost of Some Candidate Element Replacements for Lead in Electronic Assemblies Replacement elements for lead Lead (Pb) reference Zinc (Zn) Antimony (Sb) Copper (Cu) Tin (Sn) Bismuth (Bi) Indium (In) Silver (Ag)

Relative cost 1 1.3 2.2 2.5 6.4 7.1 194 212

Source: Ref. 35.

whereas it only requires about 0.8 kg of the Pb-free solder to yield 1000 jars of solder paste [37]. Thus, the increased cost for materials is substantially offset by the differential density factor because a solder joint requires essentially the same solder volume independent of whether it consists of lead-bearing or lead-free solders. b. Metal Processing Factor. The processing costs associated with fabricated products such as solder pastes often dominate over the material cost. Therefore, there is little difference between eutectic Sn–Pb and lead-free solder alloy paste, as is shown in Table 13. The table also shows, however, that the cost of products that do not have significant process costs associated with them, such as a bar stock utilized for wave soldering or a wire for repair operations, is dictated by the cost of raw materials [38]. 4. Melt Temperature and Range Both gallium (Ga) and mercury (Hg) can be eliminated from consideration as base materials just based on their low melting points. The melting point of the elements Ag, Cu, Ni, and Au are too high to serve as base materials and some are also too expensive, only leaving In and Sn. Indium is too expensive to use in quantities required of a base alloy. That leaves Sn, whose solder alloy melt points are typically between about 120jC and 280jC, providing a baseline for lead-free solder alternatives to eutectic Sn–Pb. Eutectic alloys with either Cu or Ag have melt points close to pure Sn, so only a few percent of these elements can be added to Sn before the melting point level is considered unacceptable [31].

TABLE 11 Relative Cost of Candidate Elements to Replace Lead in Electronic Assemblies Based on Volume Element Zinc (Zn) Lead (Pb) Antimony (Sb) Copper (Cu) Tin (Sn) Bismuth (Bi) Silver (Ag) Indium (In) Source: Ref. 37.

Metal cost per pound (US$, approximately)

Density at 25jC(lb/in3)

Cost per standard volume (US$/in3)

Cost factor relative to Pb

0.50 0.45 0.80 0.65 3.50 3.40 84.20 125.00

0.258 0.416 0.239 0.324 0.354 0.354 0.379 0.264

0.129 0.184 0.191 0.211 1.20 1.20 31.91 33.00

0.70 1 1.04 1.15 6.52 6.52 173.40 173.30


15

TABLE 12 Relative Cost of Some Candidate Lead-Free Alloys to Replace Eutectic Sn–Pb Solders in Electronic Assemblies Based on Their Densities

Alloy 62 Sn–37 Pb 42 Sn–58 Bi 77.2 Sn–20 In–2.8 Ag 91 Sn–9 Zn 91.8 Sn–3.4 Ag–4.8 Bi 90 Sn–7.5 Bi–20.5 Ag–Cu 96.2 Sn–2.5 Ag–0.8 Cu–0.5 Sb 96.3 Sn–3 Ag–0.7 Cu 95 Sn–3.5 Ag–1.5 In 93.5 Sn–3.5 Ag–3 Bi 96.5 Sn–3.5 Ag 99.3 Sn–0.7 Cu 95 Sn–5 Sb

Melting range (jC)

Metal cost (US$/lb)

Density at 25jC (lb/in3)

Metal cost (US$/in3)

Percent increase cost/in3 compared to 63Sn–37Pb

183 139 179–189 199 208–215 186–212

2.37 3.44 30.06 3.23 6.24 5.09

0.318 0.316 0.267 0.263 0.272 0.273

0.75 1.09 8.02 0.85 1.7 1.39

0 45 970 13 125 85

213–219

5.48

0.267

1.46

95

217–218 218 216–220 221 227 232–240

5.90 8.15 5.92 6.32 3.48 3.37

0.268 0.268 0.269 0.368 0.264 0.263

1.58 2.18 1.59 2.33 0.92 0.88

110 190 110 125 23 17

Source: Ref. 41.

To be compatible with standard surface mount and wave soldering processes, and to avoid PCB and component damage require a solder alloy melting point not higher than about 230jC, whereas operating requirements may be as broad as a 55jC to 155jC. This implies a minimum melting point of at least 175jC [31]. The compositions of some lead-free solder alloys and their corresponding melt temperatures are listed in Tables 14–Tables 16, covering melt temperatures between about 180jC and 200jC. Some higher-temperature compositions are listed in Table 17, but are only of utility for niche applications owing to such issues as cost (Au–Sn) and health concerns (Sb-containing alloys).

TABLE 13

Relative Cost of Some Lead-Free Solders in Bar Stock and Solder Paste Form

Solder alloy

Bar relative cost (US$/kg)

Paste relative cost (US$/kg)

Sn–63 Pb (baseline) 1 Sn–3.5 Ag 2.9 Sn–3 Ag–2 Bi 2.17 Sn–2.6 Ag–0.8 Cu–0.5 Sb 2.06 Sn–3.4 Ag–4.8 Bi 2.26 Sn–3.5 Ag–0.5 Cu–1 Zn 2.27 Sn–4.6 Ag–1.7 Cu 2.25 Sn–3.2 Ag–0.7 Cu 2.21 Sn–3.5 Ag–1.3 Cu 2.28 Relative cost of selected metals Pb=1, Zn=1.7, Cu=3, Sb=3.9, Bi=8.6, Sn=11, Ag=260, Au=15,000 Source: Ref. 38.

1 1.07 1.06 1.05 1.06 1.06 1.08 1.06 1.06

16

PUTTLITZ

TABLE 14 Some Lead-Free Solder Alloys with a Melting Point or Liquidus Temperature Below 150jC Composition Bi–33 In Sn–52 In Sn–50 In Sn–58 Bi Sn–57 Bi–0.1 Ag

Melting range (jC) 109.(E) 118.(E) 118.–125 138.(E) 138.–140

E=eutectic composition.

For lower-temperature applications, 42Sn–58Bi (eutectic, mp=139jC), near-eutectic compositions, and Sn–Bi with Cu, In, Au, Sb, or Zn additions are possibilities, but a special flux may be required because a standard flux may not be sufficiently active at lower temperatures [39]. a. Minimize Pasty Range. A solder with a discrete melting point is preferred (i.e., one with a eutectic composition). However, if this is not possible, the melting range should be no more than 10jC, and less than 5jC is desirable. Alloys with a so-called large pasty or mushy range are known to be susceptible to hot cracking during solidification, causing solders to lift away from annular metallization surrounding the tops of plated through-holes (PTHs), a condition referred to as fillet lifting [40]. Accordingly, eutectic or near-eutectic composition alloys are desired to prevent this condition. 5. Wettability Solder wettability is the capability of an alloy in the molten state to spread over the solid surface of another metal, usually metallization on a terminal pad. It is the precursor of ‘‘solderability,’’ which describes the solder’s ability to form a solder joint on a PCB. Joint formation includes such aspects as fillet formation between a vertical surface and a horizontal surface, and capillary flow of the solder as required to fill holes or gaps. a. Effect of Elemental Additions on Surface Tension of Tin. All high-Sn solder alloys exhibit a higher surface tension and are more chemically aggressive (i.e., greater tendency to form intermetallic compounds with termination pad metallizations) compared to eutectic Sn–Pb solder [41]. Lead has the important effect of lowering the surface tension of Sn–Pb solders. For example, the surface tension of eutectic Sn–Pb at 280jC is about 470 dyn/cm compared with 550 dyn/cm for pure Sn at its melting point of 232jC [42]. Minimizing surface tension improves wetting and solderability. None of the elemental additions made to Sn in forming lead-free alloys duplicates the effects of Pb in terms of lowering surface tension, thus is likely the reason that the

TABLE 15 Some Lead-Free Solder Alloys with a Melting Point or Liquidus Temperature Between 150jC and 200jC Composition Sn–20 Bi–10 In Sn–20 In–2.8 Ag Sn–8 Zn–3 Bi Sn–9 Zn E=eutectic composition.

Melting range (jC) 143.–193 175.–186 189.–199 198.5 (E)

LEAD-FREE SOLDER ISSUES INCLUDING SELECTION TABLE 16 Some Lead-Free Solder Alloys with a Melting Point or Liquidus Temperature Between 200jC and 230jC Composition

Melting range

Sn–3.4 Sn–7.5 Sn–7.5 Sn–3.5

Ag–4.8 Bi Bi–2.0 Ag Bi–2.0 Ag–0.5 Cu Ag–3.0 Bi

Sn–3.5 Sn–2.0 Sn–2.0 Sn–3.4

Ag–5.0 Ag–3.0 Ag–4.0 Ag–4.8

Sn–3.5 Sn–3.0 Sn–3.5 Sn–3.5 Sn–3.2 Sn–4.7 Sn–2.5 Sn–3.8 Sn–2.0

Ag–0.7 Cu Ag–0.7 Cu Sb–2.0 Bi–1.5 Cu Ag–1.5 In Ag–0.5 Cu Ag–1.7 Cu Ag–0.8 Cu–0.5 Sb Ag–0.7 Cu Ag–0.8 Cu–0.5 Sb

Bi–0.7 Cu Bi–0.75 Cu Bi–0.5 Cu–0.1 Ge Bi

Sn–2.0 Ag–0.75 Cu Sn–4.0 Ag–0.5 Cu Sn–4.0 Ag–1.0 Cu Sn–3.5 Ag–1 Zn Sn–3.5 Ag Sn–2.0 Ag Sn–0.7 Cu Sn–2.0 Cu–0.8 Sb–0.2 Ag Sn–5 Sb

205.–210 207.–212 186.–212 206.–213a 216.–220a 210.–215 210.–215 216. 208.–216a 200.–216a 217.(E) 217.–218 218.(E) 218.(E) 217.–218 216.–218 213.–218 217.–219a 213.–219a 216.–222 217.–219 217.–219 217.–219 218.–221 221.(E) 221.–226 227.(E) 226.–228 232.–240a 236.–240a

E=eutectic composition. a Different values reported in the literature.

TABLE 17 Some Lead-Free Solder Alloys with a Melting Point or Liquidus Temperature Above 230jC Composition Sn–5 Sb Au–20 Sn Sn–25 Ag–10 Sb

E=eutectic composition.

Melting range (jC)

Characteristics

232.–240 280.(E) 365.(E)

Poor wetting Very expensive Low ductility Unsuitable thermomechanical fatigue resistance for die attach

17

18 TABLE 18

PUTTLITZ Effect of Temperature on the Surface Tension of Near-Eutectic (60Sn–40Pb) Solder

Conditions (jC) 211 260 260, with rosin (R) flux on Cu 260, with rosin mildly activated (RMA) flux on Cu

Surface tension (dyn/cm)

Reference

489.9 483.3 390 F 40 382 F 8

Schwaneke et al. [43] Schwaneke et al. [43] Vianco et al. [42] Vianco et al. [42]

wettability of lead-free alloys is typically worse when compared to studies that use eutectic Sn–Pb as the benchmark. b. Effect of Temperature on the Surface Tension of Pb–Sn Solder. Surface tension has not been found to be a strong function of temperature over the range of temperatures associated with typical solder assembly operations, as noted in Table 18. The wettability and solderability of lead-free solder candidates must be of sufficient adequacy to form solder joints whose reliability is not degraded relative to eutectic Sn–Pb joints due to wettability-related issues. 6. Minimum Complexity in Materials Fabrication It is significantly more difficult to manufacture powder for a multielement (ternary or higherorder) alloy than for a binary alloy. Consider, for example, that the powder for a Sn–Pb eutectic solder paste is prepared by melting ingots of pure Sn and Pb whose weight ratio is 63% and 37% in a pot, respectively. The molten solder alloy is atomized by a special process to form powder that is sieved to be in compliance with specified particle size limits. However, ternary, quaternary, and higher-order alloys require tight control and accuracy of composition. Manufacture will likely require an additional precomposing step. For example, to prepare an alloy consisting of three elements of weight percentages 96.3, 3.0, and 0.7, respectively (i.e., for a Sn–Ag–Cu alloy) demands a high degree of precision, which will result in additional cost [37]. 7. Minimum Metallurgical Complexity It is generally agreed upon that none of the lead-free binary systems investigated is likely to be a consensus replacement for eutectic Sn–Pb in assembly reflow applications, although this may be the case for wave soldering. Viable candidates typically consist of a small quantity of a third or fourth element added to a lead-free binary alloy compared to the relatively uncomplicated eutectic Sn–Pb solder.

TABLE 19

Required Mechanical Property Enhancements Beyond Eutectic Sn–Pb

Properties Shear strength

Fatigue resistance

Elevated temperature compatibility

Remarks .

Solder joints typically experience shear loading –Often due to coefficient of thermal expansion (CTE) mismatches . Increased number of consumer electronics products –Particularly communication, handheld devices –Experiences many on/off cycles . Increased applications involving harsh conditions (e.g., vehicle engine compartment) –Needs improved creep resistance and microstructural stability


19

a. Enhanced Properties Required. There are an increasing number of applications that require significantly enhanced mechanical properties, as listed in Table 19. For example, the operating temperature range for consumer and telecommunications electronics products normally can span the range from about 55jC to 100jC. In addition, there are critical applications such as avionics and automotive where under-the-hood, engine-mounted electronics typically experience many thermal cycles, which in some cases span the range from 55jC to 165jC and higher. Accordingly, there are now numerous demanding applications that require solder characteristics to exceed those of eutectic Sn–Pb [44–48], independent of the need to define a lead-free solder, making it necessary to mutually satisfy both requirements. b. Effect of Elemental Additions. Both Cu and Ag have a low solubility in Sn, forming intermetallic compounds instead, so there is little opportunity for solid solution strengthening. The hypereutectic alloys contain primary IMCs. The elements Zn, Bi, and In all exhibit some solid solubility in Sn, and so can contribute to strengthening the Sn matrix. They must be introduced as higher-order additions because the binary systems with Sn (e.g., Sn–In and Sn–Bi) have melting points too low for mainstream electronic assembly. Whereas Zn additions can cause several concerns (poor corrosion properties), the stability of ZnO necessitates the use of highly active fluxes, and pastes containing fine Zn particles can be expected to be unstable due to interactions between the powder and the medium, resulting in unacceptably short shelf-lives [33]. 8.

Adequate Properties for Operational Requirements a. Eutectic Sn–Pb Solder Was Satisfactory. For most applications, the processing and reliability characteristics of Sn–Pb-based solders have served the electronics industry well for more than 40 years, providing a comfort factor that will not be easy to emulate, although a

TABLE 20 Summary of the Base Metal Selection for Lead-Free Solders in Consideration of Key Elimination Factors Elimination factor Unstable

Too toxic

Abundance

MP too high/cost

Remainder

a b c

Metal

Symbol

Melting point (jC)

Cesium Rubidium Potassium Sodium Lithium Magnesium Mercury Thallium Cadmium Lead Gallium Indium Bismuth Copper Nickel Silver Gold Tina Zincb Antimonyc

Cs Rb K Na Li Mg Hg Tl Cd Pb Ga In Bi Cu Ni Ag Au Sn Zn Sb

29 39 64 98 179 651 39 304 321 327 29.8 157 271 1083 1453 960 1063 630 232 419

Tin (Sn) is the best choice, taking all key aspects into consideration. Zinc (Zn) is very active chemically, causes corrosion and oxidation issues. Antimony (Sb) has some components known to be very toxic.

20

PUTTLITZ

significant number of studies on lead-free alloys, some of which are discussed in Chaps. 8 and 9, have demonstrated more than adequate reliability characteristics [36].

D.

Major Lead-Free Candidate Solder Alloys

Although there has been a significant worldwide effort to identify a drop-in replacement for eutectic Sn–Pb solder, none has been found. Based on elemental selection criteria and considerations such as cost, abundance, toxicity, etc., discussed earlier and summarized in Table 20, the industry has focused on several lead-free solder candidates that appear suitable for the assembly of most electronics products in consideration of the various factors noted in Section C Many of these alloys offer a distinct advantage over eutectic Sn–Pb, such as high joint strength, enhanced fatigue resistance, improved high-temperature life, and harder solder joints [36]; all are noted as desired and sought-after characteristics in Table 19. Both tin and indium would serve as good base materials in that both readily wet the metallizations used for terminations in electronic assemblies and form IMCs with those metals to achieve good bonding. However, there is both a cost and a supply issue associated with In, thus leaving Sn. Another advantage is the existence of substantial experience and database in the use of tin solders compiled over a long history. Tin typically cannot be utilized elementally as a solder owing to its high melting point (232jC). Elemental additions are therefore necessary to reduce the melting point of Sn. Typically, Ag, Al, Bi, Ga, and In are used for that purpose. Other elemental additions such as Ni and Sb, which do not lower the melting point of Sn, are added to enhance the properties of the Sn matrix. The TABLE 21

Major Lead-Free Solder System Candidates to Replace Eutectic Sn–Pb

Alloy system Binary Sn–Ag

Su–Cu

Sn–Bi

Sn–Zn

Ternary Sn–Ag–Cu

Sn–Ag–Bi

Other additions Ag, Al, Bi, Ga, In, Cu Ni, Sb

Remarks Fairly high melt temperature (221jC) Alternative for wave soldering, but cost an issue Some reliability concerns related to Ag3Sn platelet growth Fairly high melt temperature (227jC) Wave solder candidate Only moderate wetting, but sufficient for most applications Not applicable as a general assembly solder due to cost, low melt temperature (138jC), and availability Good for low-temperature end of a solder hierarchy Cheap Lower melt temperature than most Pb-free solders Zn is highly active, presents potential corrosion and process concerns Leading candidate system for reflow soldering Lower melt point than Sn–Ag, Sn–Cu binary alloys Adequate wetting, mechanical properties Much reduced Cu scavenger characteristics compared to Sn–Ag Even lower melt temperature than Su–Ag–Cu Best fatigue characteristics among most popular Pb-free alloys Poses some reliability and end-of-life reclamation concerns Reduce melting temperature of Sn-based solders, and some enhance mechanical properties Do not reduce the melt temperature of Sn-based solders, but added to enhance properties


21

most popular Pb-free alloy candidates are listed in Table 21 and are discussed in the following sections. 1. Based on Sn–Ag System The most promising alloys identified are primarily based on the Sn–Ag system, with high tin contents, typically greater than 90%, and with a small addition of either Cu or Bi to create a ternary alloy. Sometimes a very minor addition, less than 1%, of a fourth element is made to enhance a particular desired characteristic, such as strength or reliability. In most cases, the binary eutectic (Sn–3.5Ag) is not adequate for several reasons, which include a high melting point (221jC) that would necessitate subjecting electronic assemblies to high temperatures during manufacture, and exhibits the worst wetting characteristics among high-Sn solder alloys [2]. However, there is a substantial experience base in the use of eutectic Sn–Ag for a variety of niche applications over the years. 2. Bismuth Addition to the Sn–Ag System Several alloys have been created through the addition of 5% Bi or less to lower the melting point of eutectic Sn–Ag. A typical example is Sn–3.4Ag–4.8Bi, which melts between 208jC and 215jC. In addition to lowering the melting point, Bi also improves the wettability of eutectic Sn–Ag. In fact, Sn–Ag–Bi alloys within the V5% Bi range generally exhibit the best solderability among the group of the most popular lead-free solder alloys [38]. The addition of Bi greatly increases the fatigue resistance of eutectic Sn–Ag solder. A fourth element such as Ge or Cu is sometimes added to improve alloy strength characteristics. However, the addition of Bi can also have adverse effects, as it can result in the brittleness of Sn–Ag–Bi alloys, and the presence of Pb on components and board finishes can form a lowmelting (96jC) eutectic, which in turn can result in poor fatigue resistance. In addition, fillet lifting has been observed at through-hole joints as wave-soldered joints cool down. Alloys with a solidification range, such as Sn–Ag–Cu alloys, are susceptible to this condition where the solder separates from the IMC layer. Although favored, even alloys without a so-called pasty range, such as eutectic alloys, are not necessarily immune to fillet lifting, as discussed in Chap. 17. Bismuth also has the tendency to easily oxidize, so fluxes that are active at lower temperatures should be used. 3. Copper Addition to Sn–Ag System The addition of minor amounts of Cu, typically in the range of 0.7%, lowers the melting point of eutectic Sn–Ag. The melting point of Sn–Ag–Cu alloys is slightly higher than Sn–Ag–Bi alloys, but is less costly. The ternary eutectic Sn–Ag–Cu alloy melts at 217jC. The presence of Cu in the solder greatly reduces the Cu dissolution rate from terminal pads with which the solder is in contact. In addition, the yields and reliability of eutectic Sn–Ag–Cu or near-eutectic alloys have been found to be at least comparable to eutectic Sn–Pb. These alloys are also reported to exhibit better solderability and reliability compared to eutectic Sn–Ag or Sn–Cu binary alloys [2]. 4. An Assembly Approach An approach which many in the electronic assembly industry believe is viable in utilizing lead-free solders consists not of a single solder, as in the past with eutectic Sn–Pb, but the choice of a small number of solders depending on the type of soldering operation. An alloy from either the Sn–Ag– Cu or Sn–Ag–Bi system is suitable for solder reflow operations (e.g., the attachment of components to PWBs by reflowing solder paste). The use of Sn–0.7Cu (mp=227jC) has been considered as a candidate for wave soldering applications because it is the least expensive of the lead-free alloy candidates of interest. Material cost is a key consideration in wave soldering because of the need to periodically change the entire contents of the bath due to contamination buildup. The alloy has soldering qualities equivalent to eutectic Sn–Pb. Although the eutectic Sn– Cu alloy also possesses the poorest mechanical properties among the lead-free solder candidates [2], it is sufficient for most applications. The Sn–Ag–Cu and Sn–Ag alloys can be utilized for wave soldering, and so, too, for rework operations.

Assignees

Harris Engelhard Willard Motorola IBM IBM IBM IBM IBM IBM Indium Co. Indium Co. Motorola Touchston Ford Lucent Lucent Lucent Seelig Seelig US Army

4758407 4778733 4806309 5229070 5328860 5344607 5393188 5411703 5368814 5414303 5455004 5580520 5410184 5435955 5429689 5538686 5589433 5698160 5352407 5405577 5520572

87 – 93 92 – 99 90 – 95 90 78 90 93 94 42 70 – 90 82 – 90 71 – 92 92 – 97 79 – 97 80 86 40 – 60 59 – 82 93 – 98 90 – 99 86 – 97

Sn (wt.%)

0.2t– 0.5 2t– 11 1.5t– 3.5 0.5t– 3.5 0.3t– 4.5

40t– 60

0t– 5

0t– 4 0t.5

2t– 4

3

2

0.1t– 0.5 0.05t– 3 0.1t– 0.5

Ag

0t– 1 4t– 15

1t– 4 5. 10 2 2 2 58 2t– 10 1t– 5

Bi (wt.%)

Partial List of U.S. Patents on Lead-Free Solders

US patent number

TABLE 22

0t– 9.3

5t– 15 5t

2 8t– 20 3t– 6 4t– 26

5t 10t 8

In

0.2t– 2 0.2t– 2

1 3

3t– 5

4t– 6

Sb

0.2t– 2 0.1t– 3 0.5t

3t– 8 3t– 15

1t 1t

3t– 5 0.7t– 6

Cu

16t– 30

9

4t– 6

Zn

Interim

Set (1)

Nl (1)

Others

Sn Sb Cu Sn Cu Ag Sn Sb Bi Sn Bi In Sn Bi In Ag Sn In Bi Sn Ag Bi Sb Sn Sb Bi Cu Bi Sn In Sn In Bi Sn In Ag Bi Sn In Ag Sn Cu Sn Cu Ag Sn Bi In Sn Zn Ag Sn Bi Ag Sn Zn Ag Sn Ag Sb Cu Sn Ag Sb Cu Sn Ag Bi In Cu

Major composition

22 PUTTLITZ

Iowa St. U. Mitsui Toyota IBM Lucent Ford

Ford Johnson Samsung KIMM Ford IBM Delco H-Tech Mitsui Hitachi Matsushita Alpha Fry Murata Fuji Elec

5527628 5658258 5733501 5730932 5762866 5755898

5833921 5837191 5843371 5851482 5863493 5874043 5938862 5985212 5993736 5942185 6077477 6086687 6139979 6156132

89 90 65 – 95 80 76 – 96 37 – 57 48 – 58 43 – 58 95 77 – 89 80 91 – 97 70 – 74 84 – 90 >75 91 – 95 72 – 87 81 – 91 >90 92 – 96 40 – 70 3t– 6 0t– 0.25 0t– 5

30t– 58

2t– 4

2t– 5 6.5t– 7.5 2.5t– 3.5

1t– 2 0t.6 3t– 4

3.5t– 7.7 1t– 4 0.8t– 5 3 1t– 6

2t– 3 10t– 23 5t– 10 0t–9.25

6t– 14 1t– 20

0.5t– 1.5 0.1t– 9.5 12 0.2t– 6 37t– 57 40t– 60 38t– 52

0.1t– 1.0 0t– 9.25

12t– 24 7t– 11 Dt-6

2t– 5 0.1t– 3

3t– 4 0.1t– 9.5 5 0.2t– 6 6t– 10 2t– 5 2

0t– 9.25 3t– 5 0t– 5

5t– 15 0.75t– 2

0.1t– 10

0.1t– 2 0t– 9.25 0.7t– 2.0 0t– 1

0.5t– 1.5 0.1t– 9.5

0t– 3

1t– 4 0t.6

1t– 4

Ge

0t– 0.2

0.5t– 2 3t– 5

0t.01

0.2t– 8

Tl (0.2) Nl (0.5) Nl (0.2)

Ga (100,000 lb annually . As of January 2001, reporting threshold lowered to 100 lb for each facility emitting Pb or Pb compounds July 1, relative to previous calendar year . Studies to consider use of Pb in certain alloys: e.g., some stainless steels, brasses, and bronzes

ENVIRONMENTAL STEWARDSHIP

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PBT chemicals indicates that even small quantities of lead and lead compounds are of concern, and that low reporting thresholds under EPCRA, Section 313, may be appropriate. Accordingly, the January 2001 rule complements the October 1999 PBT final rule by lowering reporting thresholds for toxic chemicals in general for lead and lead compounds. 2. Lower Reporting Thresholds Under the original EPCRA, Section 313 reporting thresholds, covered facilities were not required to report lead and lead compounds emissions to air, water, and land unless they manufactured or processed more than 25,000 lb annually or used more than 100,000 lb annually. However, the January 2001 rule dramatically lowered the reporting threshold to 100 lb or more annually for each facility emitting lead and lead compounds. The new TRI reporting requirements went into effect starting 2001. 3. PBT-based Modifications In addition, the new rule includes the following modifications to reporting under EPCRA, Section 313 for lead and lead compounds, based on their status as PBT chemicals: (1) eliminating the de minimis exemption for lead and lead compounds. This exemption was introduced in 1998 for facilities newly covered by EPCRA, Section 313, regarding low concentrations of toxic chemicals in mixtures that are imported, processed, or otherwise manufactured as impurities; (2) excluding lead and lead compounds from the eligibility for the alternate threshold of 1 million lb and Form A; (3) eliminating range reporting, as it could misrepresent data accuracy; (4) reporting of all releases or other waste management quantities greater than 1/10 of a pound of lead and lead compounds. At present, the EPA allows facilities to report whole numbers and to round releases of 0.5 lb or less to zero when reporting on EPCRA, Section 313, listed chemicals; (5) exempting reporting obligations for lead in certain alloys: a final decision on lower reporting thresholds for lead contained in stainless steel, brass, bronze alloys will be deferred until completion of ongoing scientific studies.

VI.

ELECTRONIC WASTE MANAGEMENT: UNITED STATES

A.

Federal Standards for Lead in Landfills

The Resource Conservation and Recovery Act (RCRA) [35] of 1976, as amended in 1984, gave the U.S. EPA the authority to regulate hazardous waste at all stages, including its generation, transportation, treatment, storage, and disposal. To this end, EPA issued Regulations which set forth criteria for identifying the characteristics of hazardous waste, its listing, and requirements to which it must comply with. Waste can be classified as hazardous either because it is listed as hazardous by EPA Regulations, or because it exhibits certain characteristics. The ‘‘toxicity’’ characteristic applies to a list of 40 substances, including lead. If a waste leachate derived from a test called the Toxicity Characteristic Leaching Procedure (TCLP) contains any listed substances at or above the defined regulatory levels, the waste is considered hazardous. Chapter 2 provides a more complete discussion of the TCLP and other tests utilized to determine toxicity levels. Hazardous waste must be treated in accordance with EPA standards before being placed in a landfill. Discarded printed wiring board and cathode ray tubes usually qualify as hazardous because they commonly exhibit the toxicity characteristic for lead. That is, the test extract contains more than 5.0 mg Pb/L.

B.

State Recycling Programs

At the state level, the management of end-of-life electronics, especially computers and cathode ray tubes, is an emerging environmental issue on which a number of initiatives have been taken.

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These states, at the time of this writing, include California, Connecticut, Florida, Massachusetts, Minnesota, New Jersey, New York, South Carolina, and Wisconsin. In general, state programs tend to promote the recycling of end-of-life electronics, and to reduce disposal by landfill or incineration. Household and commercial users are encouraged to bring end-of-life electronic equipment back to designated collection points, free of charge. Recycling facilities determine its further destination (e.g., whether reused, upgraded, donated to nonprofit organizations or schools, recycled, etc.) A brief overview of state-level activities on WEEE recycling in the United States is discussed in the following sections. 1. California A pilot project was initiated in San Francisco to collect discarded computers. Residents were afforded the opportunity to drop off end-of-life computers at a number of electronics stores and recycling facilities. The collected units were dispositioned for reuse, upgrading, donation to charities, or recycling. 2. Connecticut The Connecticut Department of Environmental Protection is encouraging a system of regional and municipal drop-off sites for recycling of electronics. The project is financed by individual towns. In addition, the agency is considering introducing a general permit for the collection of certain recyclable equipment, including consumer electronics. 3. Florida and Massachusetts Both Florida and Massachusetts focused on strategies aimed to encourage recycling of cathode ray tubes (CRTs). Although the shared goal is to encourage recycling and decrease the number of CRTs placed in landfills or incinerated, the regulatory framework chosen by the two states is different. The Florida Department of Environmental Protection (DEP) emphasizes that CTRs are hazardous waste under the Resource Conservation and Recovery Act (RCRA) when discarded into landfills or waste incinerators. Under the DEP’s interpretation of RCRA hazardous waste regulations [36], CTRs are considered as materials, rather than wastes, when reused as a substitute for commercial purposes, i.e., when utilized to manufacture new CRTs, or as a fluxing agent at a secondary lead smelter. Only CTRs disposed in landfills or incinerated are regulated as hazardous waste under RCRA. The state of Massachusetts has exempted unbroken CRTs from hazardous waste requirements under RCRA. Disposed of broken or crushed CTRs are subject to full hazardous waste requirements. However, intact CRTs that can still be reused and those which will be crushed and recycled (i.e., as spent material being reclaimed), will be treated as nonhazardous waste under solid waste regulations. In addition, the state of Massachusetts has enacted a rule that prohibits the disposal of CRTs in landfills and municipal waste-combustion facilities by both households and commercial users, as of April 1, 2000. If the screen is intact, a CTR can be safely recycled, but lead is released in the environment when the tube is crushed or recycled. The EPA supports the Massachusetts landfill ban for CTRs, which goes beyond federal rules. 4. Minnesota Minnesota’s Office of Environmental Assistance is developing proposed legislation for electronics product stewardship, and is collaborating with Sony Electronics, Panasonic, and the American Plastics Council to evaluate collection strategies and recycling market opportunities. 5. New Jersey The U.S. EPA has funded a grant for an educational pilot project coordinated by the New Jersey Solid Waste Policy Group at Rutgers University. Under this program, students produced a video addressed to children between ages 3 and 6 on electronics recycling.


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6. New York Rockland County launched a program to collect discarded electronics from both households and business. This program is funded through a grant from the New York Department of Environment. 7. South Carolina South Carolina’s legislature considered a Bill establishing a statewide electronic equipment recycling program. The program would be funded by a tax on the sale of electronics products containing cathode ray tubes. However, the Bill did not survive the 1999 session. 8. Wisconsin The Department of Natural Resources (DNR) has instituted an end-of-life computer collection program. As a consequence, three cities in Wisconsin have taken initiatives for the collection of discarded computers.

VII. A.

ELECTRONIC WASTE MANAGEMENT: JAPAN Recycle Focus

Although much of the initial activity surrounding the elimination of lead in processes and products was in Japan, there are no laws banning its use. The emphasis and governmental policies have been on recycling and reuse, with both government and private sector plants coming on board for those purposes. It is anticipated that the ‘‘Law for Recycling of Specified Kinds of Home Appliances,’’ enacted in 1998, will be expanded to include discarded PCs. Japan, being a relatively small country, must take the appropriate steps to efficiently utilize available real estate. Projections have indicated that the amount of end-of-life disposal exceeded available landfills, and would place a significant burden on available space in the future. Among the corrective measures put in place was an ambitious recycling and reuse strategy to significantly reduce the quantity of material to be placed in landfills. Japanese industry responded to this objective by putting voluntary product roadmaps in place to start reducing the use of lead in their products in the late 1990s. (See Chapter 1)

B.

Lead-Free Drivers

The migration toward a voluntary ban on lead was motivated by several factors: to reduce or eliminate costs associated with the treatment of end-of-life products that contained Pb to be in conformity with requirements to deposit materials into landfills. At the same time, Japanese companies were receiving the benefit of presenting a positive image of their organizations and products to consumers by taking a leadership role to eliminate lead, an issue that has been receiving worldwide attention and support. Japanese companies were among the first to market lead-reduced products, typically utilizing a lead-free solder to attach components whose peripheral leads (e.g., Cu, Alloy 42, etc.) were finished with a thin, plated, Pb-free coating to Pb-free features on a printed circuit board (PCB). The Japanese Pb-free approach was further encouraged and enabled through the Law on Promotion of Green Purchasing that went into effect April 2001.

C.

Key Program Elements

There are several key aspects of Japan’s environmental consciousness program that are briefly discussed in the following sections, particularly as it relates to the electronics industry. 1. PC Green Label A ‘‘PC Green Label’’ is intended to indicate that a PC manufacturer’s overall efforts and activities in introducing computer products are environmentally sound. PC manufacturers must disclose

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TABLE 11 Substance Group List of Prohibition-ranked Substances Level 1 [38] Legal limitsa

No.

Substance group name

1

PCB (substances that contain three or more chlorine atoms), PCT, and their alternatives

2

Asbestos

0.1 wt.%

3

Specific organic tin compounds

0.1 wt.%

4

Aliphatic halogenated hydrocarbons (Specific CFCs, halon, chlorine-based organic compounds) Specific bromine-based flame retardants Penlachlorophenols

5 6 7

Azo compounds, Aromatic amines

8

Polychlorinated naphthalene (Substances that contain three or more chlorine atoms) Fluoroacetic acids

9

11

Organic phosphorous compounds Hexachlorobenzene

12

Benzene

13 14

Yellow phosphorus 2,4,6-tri-tertbutylphenol

10

Applicable laws and regulations Law Concerning Examination and Regulation of Manufacture and Handling of Chemical Substances, EU Hazardous Substances Regulation, ChemG EU Hazardous Substances Regulation, ChemG, Labor Safety, and Hygiene Law EU Hazardous Substances Regulation, ChemG, Law Concerning Examination, and Regulation of Manufacture and Handling of Chemical Substances Ozone Layer Protection Law, Montreal Protocol, ChemG, EU Hazardous Substances Regulation German Dioxin Ordinance

0.01 wt.% 0.1 wt.%

0.1 wt.%

EU hazardous Substances Regulation, ChemG German Consumer Goods Ordinance, ChemG, EU Hazardous Substances Regulation, Law Concerning Examination, and Regulation of Manufacture and Handling of Chemical Substances, Labor Safety, and Hygiene Law Law Concerning Examination and Regulation of Manufacture and Handling of Chemical Substances Poisonous and Deleterious Substances Control Law Poisonous and Deleterious Substances Control Law Law Concerning Examination and Regulation of Manufacture and Handling of Chemical Substances EU Hazardous Substances Regulation ChemG, Labor Safety, and Hygiene Law Labor Safety and Hygiene Law Law Concerning Examination and Regulation of Manufacture and Handling of Chemical Substances

Major use Insulation oil, heat medium

Heat insulation materials Coating materials, antiseptic agents

Cleaning agents, refrigerants, foaming agents

Resin-based flame retardants Insecticides, insect repellents Dyes

Antiseptic agents, insecticides, lubricants

Hygiene materials, agrichemicals Agrochemicals Insecticides, antifoulants

Solvents, cleaning agents Matches Antioxidants


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TABLE 11 Continued No.

Substance group name

15

4-Nitrodiphenyl and its salts Octamethyl pyrophosphoramide Bis-(chloromethyl) ether Aluminum phosphide

16 17 18

Legal limitsa

Labor Safety and Hygiene Law

100 ppm (packaging materials)c

Poisonous and Deleterious Substances Control Law Labor Safety and Hygiene Law Poisonous and Deleterious Substances Control Law EU Hazardous Substances Regulation, EU Directive, ChemG Chemical Substance Control Laws (Netherlands, Denmark), Packaging Material Heavy Metal Regulation ChemG, EU Hazardous Substances Regulation, EU Directive, Proposition 65, Packaging Materials Heavy Metal Regulation EU Directive, Packaging Material Heavy Metal Regulation


EU Directive, Packaging Material Heavy Metal Regulation

19

Cadmium and its compounds (limited to inclusion in resins, coating materials, ink, plating,b and packaging)

75 ppm, 100 ppm (packaging materials)c

20

Lead compounds (limited to inclusion in coating materials and packaging materials) Hexavalent chromium compounds (limited to inclusion in packaging materials) Mercury compounds (limited to inclusion in packaging materials)


21

22

Applicable laws and regulations

Major use Synthetic intermediates Insecticides, insect repellents Insecticides, insect repellents Resins, packaging materials, coating materials ink

Coating materials, packaging materials

Packaging materials

Packaging materials

a Legal limits show the maximum contents designated by the EU Hazardous Substances Regulation and others. b Use for production of air planes, electrical contacts and glass materials are excluded. c Regulation on packaging materials: Requires to reduce total weight of heavy materials (lead, cadmium, mercury and hexavalent chromium) contained in packaging materials to 100 ppm or below.

environmental information helpful to consumers in selecting (or using) a computer. For example, chemical or other materials that can adversely impact health and/or the environment are not to be intentionally added to their products during manufacturing or recycling processes. Plastic frame components weighing more than 25 g, such as the PC enclosure, display, keyboard, and mouse, are not to contain lead (Pb), cadmium (Cd), or specific flame retardants (PBBs, PBDEs). Also, Pb, Cd, chromium (Cr), and mercury (Hg) are not to be added to inks used in owner’s manuals and shipping containers. The types and use of eco-labels are discussed in Chapter 5. 2. Guidelines for Standardization of Material Declaration The Guidelines for Standardization of Materials was developed by the Japan Green Procurement Survey Standardization Initiative (JGPSSI) to provide a common set of requirements for companies that implement green procurement surveys. The guidelines aid in reducing the burden on suppliers inundated with these types of surveys, and improve the value and accuracy of the responses provided. The guidelines were drafted to address both the ‘‘Chemical Substance Survey’’ and ‘‘Survey Response Format.’’ Nonparticipating companies are free to utilize the guidelines that are published in both English and Japanese. If other languages are required, the surveying companies that require them must create them.

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3. Company Policies, Standards, and Guidelines Many Japanese companies have an internal standard that addresses environmentally hazardous substances. Among these are Sony, Pioneer, Matsushita, Canon, Mitsubishi, and NEC, to name a few. The standard serves as a policy guideline for dealing with environmentally hazardous substances in products and processes, defines a lead-free implementation timeline, and defines lead-free procurement goals for materials, parts, and products. In addition, the standard serves to clarify abolition/reduction plans as they relate to materials and substances being removed voluntarily, such as Pb and Pb-containing substances, and those that are prohibited or banned. The standard sets monitoring checkpoints to assure the implementation of company goals are on schedule. 4. Materials Clarification A materials classification system has been defined by many Japanese electronics manufacturers that usually consist of three categories [37,38]. a. Banned or Prohibited Substances. These are substances for which there are no conditions that relate to use, purpose, content, or phaseout date. These substances are strictly prohibited from any intentional use, and whose use must be immediately halted if found to be present in a substance, Also, these substances are typically subject to a so-called nonuse warranty that guarantees no intentional use of the banned substances. This category was created based on the domestic and foreign laws and regulations as typified by those substances listed by Matsushita Electric Group In Table 11. There is also a subcategory group of substances that some Japanese companies ban but for which there are extenuating circumstances. If their use in confirmed, an alternative technology must be explored so these materials can be discontinued at some point. If, however, there are no viable alternatives, a supplier is required to notify the customer and must obtain an exemption (i.e., off-spec) [38]. b. Reduced Use or Restricted Substances. Japanese companies have placed restrictions on certain substances, some of which are used in volume quantities, when manufacturing their products, if they have been identified as having potential harmful effects on humans or the environment. Each company sets their own targets for phasing out these restricted materials, or significantly reducing their utilization. c. Managed Substances. There is yet a final class of materials whose use is not banned nor restricted, but subject to being managed and accounted for by properly channeling these materials through recycling and prescribed end-of-life processing procedures and operations. 5. Green Procurement Standards Many Japanese companies have committed to procuring materials and assemblies that have a minimal impact on the environment, and thus have aggressively pursued green procurement objectives. These policies have received a high priority among Japanese electronics manufacturers because environmentally related issues have gained universal attention. It therefore behooves corporate entities to take a leadership role in issues involving social responsibility.

TABLE 12

List of Supplier Actions and Activities for Favorable Selection Consideration [39]

Action/Activity . . . . . .

Established an environmental management system (EMS) Corporate philosophy and policy in place focused on global environmental preservation Organizations dedicated to the pursuit of environmental conservation Established a system to evaluate and manage product environmental assessment Have established environmental preservation educational programs for employees Address streamlining physical distribution for resource and energy conservation, and reduced exhaust emissions


145

However, green procurement goals cannot be achieved without the support, cooperation, and active participation of suppliers being full partners with manufacturer’s environmental programs. To foster supplier progress, manufacturers now include environmental considerations as an integral part of their supplier and materials selection criteria. a. Selection of Suppliers. Electronic manufacturers, like other manufacturers, traditionally select their suppliers on the basis of price, quality, on-time delivery, service, technical development capability, etc. Environmental aspects are now considered as well as part of the terms and conditions for doing business. Key among these conditions is the earliest possible implementation of environmental preservation activities, such as those listed in Table 12. Compliance with ISO 14001 is generally considered a precondition for selection. Additionally, suppliers are favored that aggressively pursue continuous improvement plans. Suppliers are expected to conduct periodic self-evaluations that measure their progress, and are expected to make these results available to manufacturers. b. Selection of Materials. In a manner similar to suppliers, Japanese manufacturers now include environmental impact considerations as part of their materials solution procedure. Part of the information suppliers provide to manufacturers includes the development of new materials whose environmental impact is less than those they replace [39]. D. Examples of the Impact The beneficial impact of converting the solder used on printed circuit board (PCB) assemblies is clear from data provided by several Japanese electronics manufacturers, who are among the first to convert their lines to lead (Pb)-free assembly. For example, Matsushita Electric (Panasonic) noted by the end of March 2003 that Pb-free solder was utilized in about 12,000 models manufactured in 22 domestic and 79 overseas facilities. Up to that point, Matsushita produced a total of about 70 million Pb-free solder PCB assemblies. In the prior year (2002), before completely switching over to lead-free, Matsushita utilized 2400 tons of solder for its products [40].

REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13. 14. 15.

64 Fed. Reg. Directive 76/769 (on the restrictions on the marketing and use of certain substances and preparations, Annex I). Directive 85/210 on the approximation of the laws of the Member States concerning the lead content of petrol. Alterio, J.M. ‘‘Billion PCs Add Up To Just Too Much Garbage,’’ Toxic Trade News, URL: http:// www.ban.org/../ban_news/billion_pcs.html. National Recycling Coalition. ‘‘Electronics Recycling Initiative,’’ URL http://www.nrc-recycle.org/ resources/electrnics/managing.htm. Minnesota Office of Environmental Assistance. ‘‘Product Stewardship,’’ URL: http://www.moea. state.mm.us/stewardship/policy2001-ps.pdf. U.S. Environmental Protection Agency. ‘‘Electronics Take Back Policies,’’ URL: http://www.epa. gov/jtr/netshare/elect.htm. Explanatory Memorandum to the European Commission’s proposals, page 11, refers to a Study of the Nordic Council of Ministers of 1995. PRRDy, ‘‘Product Stewardship,’’ URL: http://www.prrdy.com/hous-productstew.htm. National Pollution Prevention Center. ‘‘Sustainable Product Systems: Life Cycle Tools and Applications,’’ URL: http://www.umich.edu/fnppcpub/shortcourse97.html. Hendrickson, C.; Conway-Schempf, N.; Lave, L.; McMichael, F. ‘‘Introduction to Green Design,’’ Green Design Initiative, (Carnegie Mellon Univ.), URL: http://www.gdi.ce.cmu.edu/GreenDesign/ gdedintro.html. ‘‘Putting Used Materials Back to Work - A Strategy for the Long Term,’’ URL: http://www.web.net/ fjjackson/usedmat.html. ‘‘An Introduction to Product Takeback,’’ URL: http://www.ce.berkley.edu/fhorvath/NATO_AWR/ FILES/klausner.pdf. Excerpted from, Environmental Manager: Environmental Solutions that Make Good Business Sense, ‘‘Extended Producer Responsibility;’’ John Wiley and Sons, Aug. 1998; 10(1) URL: http://www. informinc.org/eprenvman.php. Klausner, M.; Hendrickson, C. ‘‘Product Takeback Systems Design, Green Design Initiative, URL: http://www.ce.cmu/GreenDesign/research/takeback.pdf.

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Michele Raymond. ‘‘EU Moves on Electronics Mandates,’’ URL:http://www.wasteage.com. Wood, D.; Bender, M.; Murry, M. ‘‘Computer TakeBack Campaign,’’ GrassRoots Recycling Network, URL: http://www.grrn.org/e-scrap/e-scrap_release_03-19-02.html. 18. Northwest Product Stewardship Council, ‘‘Electronics Policies and Legislation,’’ URL: http:// www.productstewardship.net/PoliciesElectronics.html. 19. Jackson, D. ‘‘Industries Agree on Computer Recycling,’’ Waste Age, May 1, 2002, URL: http:// www.wasteage.com/magazinearticle.asp?magazinearticleid=148084&magazineid=12&siteid= 27&releaseid=10166moc. 20. COM (2000) 347 final. 21. COM (2001) 316 final. 22. Directive 2000/53 on end-of life vehicles, O.J. L 269/00. 23. Opinion of the Scientific Committee on Toxicology, Ecotoxicology and the Environment, May 5, 2000. 24. Statutory Order No. 1012 of Nov. 13, 2000. 25. Council Regulation (EEC) No. 259/93 of 1 February 1993 on the supervision and control of shipments of waste within, into and out of the European Community, O.J.L 30/93. 26. The Ordinance on the Take-back of Refrigerating Equipment, Fed. Law Gazette No. 408/1992, as amended by Fed. Law Gazette No. 168/1995. 27. Statutory Order No. 1067 on Management of Waste from Electrical and Electronic Products, December, which went into effect December 1999. 28. Kreislaufwirtschaftsgesetz, in effect since 1996. 29. Waste Management Decree, February 5, 1997, No. 22 (as amended by Decree No. 389 of Nov. 8, 1997, No. 389), Article 44. 30. Waste Management Decree, Art. 44, Par. 4. 31. Ordinance on Collection and Recovery of End-of-Life Electronics, Mar. 16, 1998, in effect since July 1, 1999. 32. Regulations on Hazardous Waste, May 19, 1994. 33. Lead Exposure Reduction Act, S. 391, Oct. 8, 1991. 34. Section 6607 of PPA. 35. 42 U.S.C. s/s 6901 et seq. (1976). 36. 40 CFR 261.2 (e) (ii). 37. Pioneer Corporation. Environmental Hazardous Substance List, (Ver. 4.0), Feb. 2003. 38. Matsushita Electric Group. Chemical Substances Management Rank Guidelines, Ver. 2.1(For Products, Kadoma, Osaka, Japan, Dec. 2002. 39. Matsushita Electric Group. Green Procurement Standards, Ver. 2.0, Dec. 2002. 40. Matsushita Electric (Panasonic) news release, May 15, 2003.

16. 17.

SUGGESTED READING Kester Solder Lead Free Update, ‘‘A Concise Snapshot of the Lead-free Soldering Issue,’’ 2000. Electronics Industry Environmental Roadmap: ‘‘Activities in Europe and Japan, Microelectronics and Computer Technology Corporation’’ (MCC), 2000. National Physical Laboratory (NPL). Electronics Interconnection Lead-free News, UK, 2000. Lead-Free Legislation, UK–Europe, Brian Richards, 2000. Richards, B.P., et al. ‘‘An analysis of the current status of Lead-Free Soldering,’’ Apr. 1999. Richards, B.P.; Nimmo, K. ‘‘Lead-free Soldering—Update 2000,’’ Apr. 2000. Charles, H.K.; Sinnadurai, N. Microelectronics, Rising to the Environmental Challenge? Denmark plans a Generalized Ban on Lead. Environment Daily (ENDs), Mar. 19, 1999. Danish Bid to ban Lead ‘‘Unjustified’’. Environment Daily (ENDs), June 7, 2000. Auken defends proposed Danish Lead Ban. Environment Daily (ENDs), June 15, 2000. Swedish Environment Protection Agency. Swedish Environmental Quality Objectives, 2000. The Future Environment—our, Common Responsibility, Final Report of the Committee on Environmental Objectives; Stockholm, June 2000. U.S. Environmental Protection Agency, 40 CFR Part 372, Lead and Lead Compounds, Lowering of Reporting Thresholds; Community Right-to-know Toxic Chemical Release reporting Final Rule, Federal Register, Jan. 17, 2001. Cathode Ray Tubes banned from Disposal at Transfer Stations, Landfills, Incinerators. BNA Daily Environment, Apr. 6, 2000. EC to get the Lead Out, IPC says U.S. will too. U.S. Tech Interactive, Apr. 2, 1999. An assessment of the use of lead in electronic assembly, Part I. Circuit World, 1993; 19 (2). Price, J.L., Ed.; Reclaiming End-of-life Cathode Ray Tubes (CRTs) and Electronics: a Florida Update. Hazardous Materials Management Conference, Tucson, Arizona, Nov. 15, 1999. National Recycling Coalition, Electronics Recycling Initiatives, 2000. Florida’s Strategy for the Management of End-of-life Cathode Ray Tubes (CRTs), Computers and Other Electronic Equipment, Discussion Paper, Sep. 2, 1999. IPC Roadmap, a Guide for Assembly of Lead-free Electronics, draft June 2000.


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EACEM, European Association of Consumer Electronics Manufacturers. Position Paper on the Commission’s Proposed Directive on Waste Electrical and Electronic Equipment; Jan. 2001. American Electronics Association (AEA). Position Paper on Lead; Aug. 1999. Euro Coop’s Comments to the Commission’s Proposal for a Pair of Two Directives on Waste Electrical and Electronic Equipment, COM (2000) 347, Brussels, Feb. 14, 2001. APME’s Position Paper on the Proposal for a Directive on Waste Electrical and Electronic (E&E) Equipment, Sep. 18, 2000. EICTA Statement on the Proposal for a Directive of the European Parliament and the Council on Waste Electrical and Electronic Equipment, and the Proposal for a Directive on the Restriction of the Use of Certain Hazardous Substances in Electronic and Electrical Equipment, COM (2000) 347 final, Brussels, Sep. 22, 2000. AEA–EIA Joint Principles on Proposed European Waste and Substance Restrictions Directives, Nov. 2000. Orgalime’s Position Paper on Commission’s Proposals for Directives of the European Parliament and the Council on Waste Electrical and Electronic Equipment, and Restriction of the Use of Certain Hazardous Substances in Electronic and Electrical Equipment, Sep. 5, 2000. European Environmental Bureau (EEB)’s Comments on the Commission’s Proposal for a Directive on Waste from Electrical and Electronic Equipment (WEEE), and the Restriction of the Use of Certain Hazardous Substances in Electronic and Electrical Equipment (EEE), Sep. 2000. NEMA (Japanese National Electrical Manufacturers Association), Statement of Principles on EU Directives Regarding Waste in Electrical and Electronic Equipment and Substance Bans. AMCHAM (American Chamber of Commerce)’s Discussion Paper on the Commission’s Proposals for a Directive on Waste from Electrical and Electronic Equipment (WEEE), and a Directive the Restriction of the Use of Certain Hazardous Substances in Electronic and Electrical Equipment (EEE), Oct. 5, 2000.

5 Market, Product, and Corporate Policy Trends E. Jan Vardaman TechSearch International, Inc., Austin, Texas, U.S.A.

I.

INTRODUCTION

The United States is regarded as the first country to discuss regulations on the usage of lead after numerous studies on the toxicity of lead in 1989. Legislative attempts ensued; the lack of a drop-in replacement for electronics resulted in abandonment of the movement. In the late 1990s, Japan undertook a self-imposed restriction with many electronics manufacturers aiming to gain brand awareness by claiming ‘‘environmentally friendly operations.’’ In 1997, Japan’s technology research gained world recognition when Matsushita Electric introduced a consumer product using lead-free solder that gained market share over competitors. A variety of European Committee legislation has been proposed prohibiting the use of toxic substances containing lead by 2004, accelerating the movement for lead-free solder. Largely ignored in North America, Japan declared support for total elimination of lead and started work on legislation to ban lead. While timelines have been revised for European legislation limiting lead use, the date is proposed to be July 1, 2006 with various exemptions for industrial segments such as military and aerospace electronics [1]. The environmentally friendly manufacturing movement has demonstrated that it is not primarily driven by legislative activity, rather, by consumer emotions. The issue is being pulled by market forces, not pushed by technology. Given consumer demand, a great number of consumer products with eco-labels are expected to be introduced in the market over the next few years— especially in Europe and Japan. The green manufacturing wave will be concentrated in the segment of the electronics industry with products that have a short life such as mobile phones, personal computers, television sets, radios, etc. that typically are discarded in landfills. However, lead-free adoption is not limited to these products. Some automotive and computer products have adopted lead-free solder materials, and future telecommunications and additional computer products are anticipated on the horizon. While the exact definition of ‘‘lead-free’’ is debatable, there are numerous examples of products being manufactured and sold that are either completely lead-free or are manufactured with reduced amounts of lead. Many products are consumer or portable products such as consumer electronics; however, lead-free workstation servers and high-end mainframes have been introduced in Japan.

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150 II.

JAN VARDAMAN ENVIRONMENTALLY PREFERABLE PURCHASING TRENDS

Environmentally preferable purchasing insures that environmental considerations are included in purchasing decisions along with traditional factors such as product price and performance [2]. The U.S. Environmental Protection Agency (EPA) encourages that a product’s environmental impact be based on its life cycle, that is, from acquiring the raw material, through manufacture, packaging, distribution, use, and disposal. The impacts can include a number of factors, among them: energy and water efficiency, toxic material content, recycled material content, resources required, adverse effect on humans and animals, adverse effects on plants, water, air, and soil to cite a few. These types of assessments can be both costly and lengthy. So for practical reasons, it is the normal practice to address only the most important factors for their potential of impacting the environment during a product’s life.

A.

Public vs. Private Sector

Environmentally preferable purchasing has been practiced in the public sector for sometime. The government can utilize such tools as federal statutes and Presidential Executive Orders to help promote social and economic programs in achieving Congressional or Presidential policy objectives. That is, in addition to the normal considerations, government contracts would also include environmental considerations. An Executive Order, Federal Acquisition Recycling, and Waste Prevention (Oct. 20, 1993), required the EPA to define the environmentally preferable purchasing guide for federal agencies. Later (Sept. 1998), Executive Order 13101, Greening the Government through Waste Prevention, Recycling and Federal Acquisition, mandated that the federal agencies purchase ‘‘. . .products and services that have a lesser or reduced effect on human health and the environment when compared with competing products and services that serve the same purpose.’’ Many state and local governments practice similar incentives through the EPA’s Environmentally Preferable Purchasing (EPP) Program, in hopes that some of the experiences in the public sector will carry over to the private sector. Although public and private sector procurement share some similarities, they play distinctively different roles in the market place. Government agencies typically buy products designed and manufactured by the private sector, and as purchasers are concerned about price, performance, and availability. Unlike the government, companies are also suppliers and must take actions that make their products competitive, such as cost, time to market, performance, and other factors that affect sales and profit. Typically, these are not matters that play a role in government purchasing decisions.

B.

Move Towards Environmental Concerns

Private sector purchasing decisions typically do not take into account economic, social, or political factors; but this is changing, especially as it relates to the issue of the environment. Companies are rapidly becoming aware that it is good business to be environmentally conscious. Given the worldwide interest in the environmental movement, companies feel the need to respond by providing more environmentally friendly products and utilizing this as a method to differentiate their products from competitors.

C.

Cost Savings

Some companies have realized significant cost savings in implementing environmentally friendly processes and products. For example, the elimination of flux cleaning postsoldering operations of electronic circuit boards by the introduction of no-clean fluxes has resulted in enormous savings by the industry (see Chap. 11, Sec. 6). A broader perspective is often required to take advantage of the cost savings afforded by environmentally friendly approaches. That is, an assessment must be made on the basis of all aspects of production and sales, not just a side-by-side comparison of an individual aspect. As in the flux application case, there is only a modest cost advantage of a no-

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clean flux over a standard flux, which is lost in the more thorough preparation required to assure adequate solder wetting of parts to be soldered with no-clean flux (i.e., low resin content, thus lower activity) compared to a mildy activated rosin (RMA) flux. However, all of this is overshadowed by the significant cost savings in eliminating the flux clean step and the costs associated with treating the cleaning solutions or meeting the legal requirements for their proper disposal.

D.

Selection Guides

In order to supply products that are environmentally friendly requires procuring environmentally friendly parts to manufacture those products. There are organizations such as the Green Purchasing Network (GPN) in Japan which provides a list of vendors that supply environmentally preferable products and compares their environmental attributes through their guidebooks. The nonprofit organization, Green Seal, publishes monthly the ‘‘Choose Green Report,’’ which identifies green products, as well as a buying guide. The aforementioned guides help procurement organizations in meeting the challenge of determining products available in the market that are environmentally preferable. Many large companies generate lists of approved vendors, i.e., those that supply environmentally approved products. This serves to streamline the procurement process in only having to interact with acceptable sources.

E.

Procuring and Providing Environmentally Friendly Products

Depending on the product and the target customer audience, a product may be focused on just one environmental attribute or several attributes. Utilizing products made of recycled materials is a popular single attribute focus. For example, virtually all containers in which fast-food restaurants serve food prominently indicate they are constructed of recycled materials, as a means of impressing their customers that they are environmentally conscious and responsible corporate citizens. The focus may be on environmental attributes that are based on the company’s culture, what it deems important. For example, it is reported that a producer of high-quality skin and hair-care products avoids the use of animal-derived and petroleum-based chemicals whenever possible. Synthetic alternatives are utilized if performance requirements are satisfied since they are more environmentally preferable than those derived from natural sources.

F.

An Industry Consensus Required

In order to turn the entire electronics industry around from its more than 40 years of lead–tin solder assembly practices, a close working relationship is required with the supply chain, i.e., the vendors that provide materials, parts, and assembly services. The industry must come to some consensus concerning lead-free materials, particularly solders and component lead finishes. Vendors cannot be expected to provide multiple versions of individual part numbers. In addition, consensus will allow price reductions that result from economies of scale. As always, companies that desire products from vendors that lie outside the de facto standards (i.e., the industry consensus) must be prepared to bear the burden of additional costs, sourcing issues, and, in some cases, long lead times.

III.

USE OF ECO-LABELS FOR PRODUCT DIFFERENTIATION

Given the interest in environmental issues, many companies are taking serious steps to utilize environmental capability to differentiate their product in the marketplace as an integral part of their marketing strategy. This approach is often referred to as Design-for-Environment or DfE and others such as green design, life-cycle engineering, environmentally conscious design, etc. where the primary intent is to prevent the potential environmental impacts through prudent

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design choices rather than attempting to mitigate them at some point in their life cycle [3]. Manufacturers utilize the practice of eco-labels affixed to their products as a means of communicating to consumers that they have met an approved environmental standard as a method of differentiating their products in the market place.

A.

Types of Labels

Eco-labeling is a voluntary method of certifying environmental performance that is practiced around the world. It identifies a specific product based on life-cycle considerations. There are many different voluntary environmental performance labels with the International Organization for Standardization (ISO) having identified three general categories, whose characteristics are summarized in Table 1. Briefly, they can be described as [4]:

Type I—a voluntary, multiple-criteria-based, third-party program that awards a license authorizing the use of environmental labels indicating environmental preference of a product in a particular category based upon life-cycle considerations Type II—environmental self-declaration claims Type III—voluntary programs that provide quantified environmental data of a product relative to parameters set by a qualified third party based on a life-cycle assessment that has been verified by that or another qualified third party Understanding that there must be guiding standards and investigations by independent third parties to assure consumers that a labeled product does, in fact, meet certain prescribed criteria, there are both public and private organizations that assure credibility by serving as impartial

TABLE 1

Several Key Aspects of Eco-Label Categories Categories

Key aspects

I

II

Criteria Definition

Independent 3rd party

Manufacturers self-declarations

Scope

Specific products (e.g., computers, consumer products, etc.) Officially designated third party, (e.g., test laboratory) Environmental organization label or seal Japan: Eco-Mark

Single or multiple environmental attribute(s)

Certification

Designation

Examples

European Union: Eco-Flower Sweden: TCO Label Germany: Blue Angel

III Quantitative life cycle impact assessments of products—difficult to obtain meaningful data Specific products

Use of standardized formats defined by industry organizations Industry associations

Officially designated third parties

Japan Electronics Information Technology Assoc. (JEITA) has an eco-label for PCs

Not gained much acceptance to date


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third-party monitors. Several organizations engaged in eco-label activities such as providing guidance, information, setting standards, and granting certification are given in Table 2.

B.

Benefits

The ultimate gain and benefit of eco-label programs is to improve some aspect(s) of environmental quality. However, in doing so, there are also benefits derived by both producers and consumers. The eco-labels are typically well recognized. For example, the Nordic Swan label is recognized by over 90% of people in Sweden and thus is a very cost-efficient way of communicating that the products bearing that label are environmentally friendly, and the companies that produce them take responsibility for the environment. In addition, 77% of Swedes consider that if a product bears the Swan label, it is likely to be reliable and of good quality [5]. There is anecdotal information that suggests there is a widespread awareness and embracing of environmentally labeled products and a willingness to pay a premium for these products. Even early on (1988), 68% of German households were familiar with the Blue Angel label and its connection to environmental protection. Similarly, in a 1994 survey, 78% of households in Singapore recognized the Green Label and approximately a third considered it as part of their purchasing decision. In a 1993 poll, more than 50% of the population in Japan was familiar with the

TABLE 2

Several Organizations Engaged in Eco-Label Activities

Organization/address Environmental Management Market (http://www.environmarket.com/edp)

Global Ecolabelling Network c/o Terra Choice Environmental Services, Inc. 2781 Lancaster Road Ste. 400 Ottawa, ON., Canada K1B-1A7 Phone: 613-247-1900, 613-247-2228 ([email protected]; http://www.gen.gr.jp) Green Seal 1001 Connecticut Ave. NW Ste. 827 Washington, D.C. 20036 Phone: 202-872-6400 Fax: 202-872-4324 (http://www.greenseal.org) International Organization for Standardization 1, rue de Varembre’ Case Postale 56 Geneve 20, Switzerland CH-1211 Phone: 4122-7490-111, 4122-7333-430 ([email protected]; http://www.iso.ch) U.S. Federal Trace Commission

Source: Ref. 6.

Information/services Environmental product declaration —Product specific requirements —How to calculate and report environmental performance of particular products —GEN is an association of international eco-labeling organizations —Provides information on various organizations —Source for comprehensive news and information on eco-labeling —Independent certifying eco-label organization in the U.S. —Sets the standards and awards the Green Seal eco-label to qualifying products

—ISO Standards 14020 through 14025 established guidelines for both first- and third-party environmental labeling

—Govern environmental marketing claims —Set standards for the use of language on products

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EcoMark. One in five Canadians indicated someone in the household purchased a product with the Ecologo label in 1996 [6]. In addition to providing producers the potential for increasing their market share, public recognition and acceptance has the added advantage of driving producers to continuously search for new and better manufacturing practices to benefit the environment and provide enhanced environmentally sound products.

C.

Eco-Labeling Practice

1. Products Must Meet Standards Independent of whether the product is certified by a first or third party, it must be tested to demonstrate compliance with environmental standards. Most eco-labeling programs have clearly defined standards. Additional guidance in general eco-labeling standards and marketing is available from a variety of organizations, some of which are listed in Table 2 [7]. Prospective applicants must retain a test facility approved by the eco-labeling organization to confirm that the applicant’s product(s) to be covered by the license conforms with the criteria defined by the ecolabels organization. The findings of the tests are kept on file for the period that the license is in force. Licenses are required to maintain records that demonstrate that the eco-labels criteria are met on a continuing basis for all products covered. In addition, to further assure the validity that eco-labels criteria are met, eco-labeling organizations conduct periodic on-site inspections of licensees, subcontractors, importers, etc. without prior notice [8]. If an inspection or test reveals noncompliance, the licensee is required to take corrective action relative to the product, manufacturing process, or quality system. If the noncompliance persists, the license to use the eco-label is revoked. A revocation can be permanent or temporary until noncompliance issues have been appropriately remedied. Revocations are typically announced by eco-label-granting organizations [8]. 2. Term An eco-label license is typically valid for prescribed period, normally 3 years because it is anticipated that product environmental requirements will gradually increase. Companies must then reapply and an assessment must be conducted to determine if product(s) conforms to the new standards or requirements. If not, the license is withdrawn, thereby ensuring that companies pace with technological progress in providing increasingly environmentally friendly products [9]. 3. Fees Applications are typically responsible for an initial application fee and all costs associated with tests and inspections during the application phase [8]. The programs are usually financed in part

TABLE 3 Survey Ranking of Audio Equipment Attributes Rank 1–3 1–3 1–3 4–5 4–5 6–8 6–8 6–8

Item Price/cost of ownership Reliability/durability/build quality Design Sound quality Technical specification/feature Brand name Ease of use Output power

Source: Ref. 10.

MARKET, PRODUCT, AND CORPORATE POLICY TRENDS TABLE 4

155

Survey Ranking of Environmental Attributes

Rank

Item

1 2 3 4 5 6

Energy consumption Hazardous substances Material application Recyclability Life-cycle impact Packaging

Mentioned by (%) 100 60 53 47 23 17

(100) (53) (37) (23) (37) (30)

Source: Ref. 10.

through government subsidies and fees based on a product’s annual sales. For example, in the case of Sweden’s Nordic Swan, the assessment is 0.4% [9].

D.

Green Marketing

Producers such as automakers have not seen a strong correlation between customer’s words and actions. That is, there are claims that environmental performance is important in their purchasing decision, but postpurchase surveys place environmental concerns well below other factors such as a vehicle’s comfort, size, color, engine performance, etc. [2]. Studies have been conducted that if environmental characteristics are the sole focus, the marketing will be relatively unsuccessful. However, if other product benefits are linked to the environmental attributes, most consumers will give such products favorable consideration [10]. 1. Coupling Environmental with Other Parameters A study determined that customers typically do not make buying decisions on their own but take the advice of relatives, friends, and particularly the sales force who exercise a strong influence in approximately half the sales decisions irrespective of product price. The first step is to determine key parameters of interest for a particular product. Consider, for example, audio equipment. From a table (similar to Table 3), selected are the top three or four attributes (e.g., price, cost of ownership, reliability, quality) mentioned most frequently. These selections represent the best candidates to link with environmental attributes to enhance the marketing value of an audio product. Similarly, the environmental factors were ranked by customers as listed in Table 4. Based on the results, it was concluded that energy consumption, hazardous substances, materials application, and recyclability were the most important environmental aspects to couple with other product attributes to help differentiate them from competitor products and increase their marketability [11].

IV.

MATTERS IN JAPAN

Japan’s official policy for dealing with end-of-life electronics products is focused on recycling. Recycling plants in Japan began operation as the home appliance recycle law (Law for Recycling Specific Appliances for Home Use) took effect on April 1, 2001. However, many existed before the legislation was enacted [12]. This legislation provided Japanese manufacturers additional motivation to embrace lead-free technology because it eliminated or reduced disassembly costs associated with removing lead-containing components prior to being dispositioned. By adopting this approach, Japanese manufacturers were well positioned in the marketplace with some introducing products in niche applications that contained little or no lead quite early. Many Japanese companies indicate that it is not just a matter of appearing to be environmentally friendly, but rather a necessity for survival in the world market [13]. Lead-free materials have typically been utilized in consumer electronics such as pagers, televisions, laptop computers, minidisk players, etc. Almost all of these products are eco-labeled to attract environmentally

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conscious consumers. For example, Matsushita’s minidisc players sold both in and outside Japan under the Panasonic brand bare a green leaf logo, which states, ‘‘Produced for the Environment.’’

A.

Japanese Corporate Policies

The committee of the Japan Institute of Electronics Packaging (JIEP) defined Japan’s lead-free soldering roadmap on January 30, 1998. Following the roadmap introduction, several major Japanese consumer electronic manufacturers initiated their own road map and publicly announced accelerated plans to eliminate lead solder completely by 2001. Deadlines for all products have moved out, but many virtually lead-free products have been introduced. Companies such as Matsushita (Panasonic), Sony, Toshiba, and Hitachi pushed forward and met this aggressive timeline. Numerous Japanese companies have announced lead-free goals in concert with other environmental objectives. Japanese firms have been driven partly by the threat of pending European Union legislation and partly by the goal of introducing eco-labeled products to gain market share. Every major Japanese electronics manufacturer has a lead-free product plan and an annual environmental report. While many companies announced goals of making all their products lead-free by 2001, total lead reduction has not been achieved. Nevertheless, some products have been introduced and corporate policies for lead elimination remain in place. Hitachi’s plan was to make all its products lead-free by 2001. The company has been heavily involved in the examination of potential leadfree solders for use in Hitachi’s various product lines. Matsushita (Panasonic) planned to eliminate all lead from major electronics products by 2001. NEC started its lead-free solder task force in April 1998, and in May 1998, a lead-free solder promotion committee was established. With its corporate goal of lead reduction, all NEC divisions were researching lead-free solder for product applications. While the workstation division finds the development challenging because of the extensive use of ball grid array (BGA) components and flip chips that both use solder balls, other divisions are shipping products. By 2003, NEC’s pager and mobile phone division will purchase 50% less solder for use in its factories as compared with 1998 purchase levels. NEC is moving ahead with its lead-free plans and has proposed Sn–Ag–Cu as the industry standard because it can be used with existing components and equipment [14].

B.

Japanese Consortia

Active since the early 1990s, the Lead-Free Soldering Committee of the Japan Institute of Electronics Packaging (JIEP) includes more than 40 companies and three academic institutions. The group has focused its efforts on the evaluation of alternative alloy systems. A project similar to the National Center for Manufacturing Sciences, lead-free project was initiated in Japan (1998) under the New Energy and Industrial Technology Development Organization (NEDO). Oversight for NEDO falls under Japan’s Ministry of International Trade and Industry (MITI). The project goal was to establish a database on lead-free soldering, to select suitable lead-free solder candidates, and to establish the process for lead-free solders. The 2-year (April 1998 to March 2000) w350 million project was composed of members from major electronics, device, and solder companies, in addition to several universities. The database was established, solder compositions including Sn–Ag–Cu have been proposed, and lead-free soldering processes have been implemented. In January 1999, NEDO initiated the Research and Development Program for Lead-free Solder Standardization in an effort to draw together and standardize various lead-free studies and independent corporate projects. The Japanese Welding Association (JWA) was assigned to study the basic performance characteristics of leadfree solder. Numerous conference papers have been published detailing performance characteristics of various lead-free solders.

C.

Japanese Products

Many Japanese companies have introduced lead-free products and others are planned in the area of consumer electronics, computers, and automotive applications.


157

1. Hitachi Hitachi initiated production of several lead-free products in 1999, including a laptop computer and camcorder. In its FLORA220CX notebook computer, which started production in February 1999, Hitachi used both a Sn–2Ag–2Bi–0.5Cu alloy solder paste and a Sn–3.5Ag–0.75Cu board finish for an audio board. An Au–Sn or Ag–Sn alloy was used for soldering optical components, and a Sn–Ag–Bi alloy was used to attach the audio substrate portion of the notebook computer. In March 1999, Hitachi began mass production of a lead-free camcorder (VMH845L) that employed a Sn–3.2Ag–2.7Bi–0.5In solder paste and a Sn–3.5Ag–0.75Cu board finish. 2. Fujitsu Fujitsu developed two lead-free solders for use in the GS8900 global server series, introduced in October 1999 and also in the VPP5000 supercomputer. The flip chip solder bumps for the system are a Sn–Ag alloy (melting point of 220jC). The flip-chip-bumped ICs are mounted on a thin-film AlN multichip module. Fujitsu reports less concern with alpha particles, a general concern with lead (Pb)-based solders [15,16]. The CPUs are packaged in a pin grid array (PGA) and mounted on the server motherboard using a pin-in-paste method or a microconnector method. Fujitsu developed a Sn–Bi–Ag alloy with a melting point of approximately 130jC to mount components on the motherboard (Fig. 1). Components are mounted on a preflux-treated bare copper motherboard and then reflowed. Fujitsu has licensed its lead-free alloy to several solder manufacturers. The company is working to introduce additional lead-free products such as workstations, personal computer, hard disk drives, and mobile phones. 3. Matsushita Matsushita introduced its first lead-free solder minidisc player in October 1998, the MJ30. The minidisc player was assembled using surface mount technology and a Sn–Ag–Bi–Cu lead-free solder paste. For through-hole assemblies, the eutectic Sn–Cu alloy was selected. The company

FIG. 1 Components mounted on a Fujitsu System Board with a Sn–Bi–Ag alloy. (Courtesy of Fujitsu, Ltd.)

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reported that lead-free solder could be introduced into manufacturing without any major problems regarding material properties and product quality and reliability. The company introduced two other lead-free minidisc players in 1999, the MJ70 and MR1000, and retired the MJ30 model [17]. Between October 1998 and the end of 1999, Matsushita manufactured 1.6 million portable minidisc players and 1.8 million video recorders with lead-free solder [1]. The minidisc player uses a proprietary Sn–Ag–Bi–X alloy, also used in Matsushita’s phase change optical disk (PD) drives. Matsushita also began producing lead-free professional videotape recorders in August 1999, which used a Sn–Cu solder. The company also introduced leadfree headphone stereo sets manufactured in Taiwan. Additional product introductions included mobile phones, rice cookers, medical equipment, and personal computers. In its effort to develop a lead-free product, Matsushita spent 4 years evaluating approximately 50 lead-free solder alternatives and several component finishes including Sn–Pb, Sn–Bi, Pd–Ni, and Au–Ni [17]. Matsushita announced the availability of lead-free semiconductor products starting in April 2000. Matsushita utilized three solder alternatives, depending on package type, to achieve lead-free soldering for semiconductor products. By late 1999, Matsushita stated that it had started using its lead-free solder in its Panasert line of surface-mounting equipment. 4. NEC NEC introduced lead-free solder pagers and mobile phones in 1999. The company began research in 1997, testing Sn–Ag–Cu alloys, then in December 1998, selected a Sn–3.5Ag–0.75Cu material with a melting point of 220jC and an electroless Ni–Au board finish. In cases where largecomponent leads are soldered with a soldering iron or to repair SMT-mounted components, NEC uses Sn–Ag solder wire. The NEC lead-free mobile phone project started in December 1998, with trial production in September 1999. The same lead-free solder material and board finish combination was selected for the pager and used for the DoCoMo cellular phone produced for NTT. Test results for NEC’s mobile phones indicate that the use of lead-free solders have provided improved solder joint reliability. NEC’s personal computer division introduced a lead-free laptop computer in October 1999. NEC adopted lead-free solder materials in its portable notebook computers, the VersaPro NX series, models VA33H-BS and VA33H-BC. Mobile PC model numbers PC-LB33H12C, PCLB33H12D, PC-LB33H12CW, PC-LB33H12DW, PC-LB400J12C, and PC-LB400J12D are marketed as lead-free. The lead-free solder (Sn–Zn–Bi alloy) was used for both main and daughter boards, such as the card/bus interface boards. With a 3% Bi addition, the solder melting point is between 187jC and 197jC. The personal computer division selected Showa Denko’s solder paste and NEC Niigata manufacturers the boards. The solder volume requirement is onethird less than with previous notebook computers, and the board finish is a flash of nickel. An improved flux and nitrogen reflow is also used. NEC envisions several issues with the use of lead-free solders. There is concern of possibility of overheating small chips because the reflow temperatures are typically higher than with eutectic Sn–Pb solders. Because many semiconductor packages are purchased outside NEC, there is concern about the component finishes on these packages and the resulting brittle intermetallics that could be formed. 5. Others In April 1999, Sony announced that it would change the material composition of its lead-free solder from Sn–4Bi–2Ag–0.5Cu–0.1Ge to Sn–2.5Ag–1Bi–0.5Cu. Sony released a television manufactured with lead-free solder in 2000 and introduced the DCR TRV-30 digital camcorder using Sn–2.5Ag–1Bi–0.5Cu lead-free solder and a halogen-free PCB. Toshiba announced plans to adopt lead-free solder for all digital products, including PCs, televisions, DVD players, and digital cameras, by March 2003. Toshiba’s plans included eliminating lead entirely from its mobile phones by 2002. Toshiba developed a CDMA mobile phone that utilized a Sn–Ag alloy and a lead-free board finish based on an Alpha Metals silver electroless plating.


159

The Oki Group launched environmental protection programs in 1993 and set a goal of achieving lead-free solder for all Japanese products by the end of March 2004. Nagano Oki implemented lead-free solder in 20% of all new products by 2001 and set a corporate deadline to achieve 100% by 2004. Japanese automotive electronics makers have also introduced lead-free solutions. Toyota introduced lead-free wiring harnesses in all new cars starting in 1998. Wiring harnesses typically contain between 150 and 200 g of lead.

D.

Japanese Eco-Labels

Many Japanese companies are using eco-labels in product marketing efforts. These eco-labels range from leaves in various shades of green to images of trees or fields of flowers. Examples of eco-labels and eco-seals are shown in Fig. 2. While there has been no reported measure of marketing success with these labels, their use continues. In addition, many major Japanese companies such as Matsushita, NEC, and Sony publish full reports on their corporate environmental policies.

E.

Lead-Free Roadmaps

Japanese companies are ahead of the rest of the world in the introduction of lead-free products. The Japan Institute of Electronics Packaging’s Technology Roadmap for Electronics Packaging in 2010 documents the path toward lead-free product introduction. Lead-free component production is noted to have started in 2000 with rapid progress anticipated between 2003 and 2005. The use of lead-free solder is expected to become pervasive in Japan. While there are a number of products with lead-free solder today, many additional products are expected. The roadmap predicts that the use of lead-free solder will surpass conventional lead-based solders by 2003. While conventional lead-based solders will still be utilized until at least 2010, the volume is

FIG. 2 Examples of eco-labels and eco-seals: a) TerraChoice (Canada); b) Sony eco-info label; c) GreenPlus (Japan); d) KELA eco-label (Korea); e) NEC eco-logo; f) Ecoflower (European Union); g) Environmental Choice (New Zealand); h) Green Cross International; i) Panasonic Panaleaf1; j) Panasonic Panaleaf2; k) Green Seal (USA); l) TCO ’03 display label; m) EPA (USA); n) Nordic Swan; o) Blue Angel (Germany); p) Terra Choice (Canada); q) Eco Mark (Japan).

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FIG. 3 Lead-free replacement time line depicting the components shift to lead-free over time. (From Ref. [18].)

anticipated to sharply decline. The use of conductive polymers, as a substitute for solder, is expected to emerge particularly for consumer electronics applications [18]. Projections for solder use are shown in Fig. 3. While many companies lament that it will be difficult to move to totally lead-free manufacturing, most companies anticipate that lead-free manufacturing will soon become a reality (Table 5) [19].

TABLE 5

Japan Jisso Technology Roadmap

Application Desktop PC

Notebook PC

Handheld PC

PDA

Cellular phone

AV equipment

Source: Ref. 20.

Environmentally friendly characteristic Pb-free materials Percentage of Pb-free utilization (%) Pb-free materials Percentage of Pb-free utilization (%) Pb-free materials Percentage of Pb-free utilization (%) Pb-free materials Percentage of Pb-free utilization (%) Pb-free materials Percentage of Pb-free utilization (%) Pb-free materials Percentage of Pb-free utilization (%)

1999

2000

2005

2010

— 0

Pd, Ag 10

Pd, Ag 70

Pd, Ag 100

Sn–Ag —

Sn–Ag 50

Sn–Ag 50–100

Sn–Ag 100

Sn–Pb 0–20

Sn 50–100

Sn 100

Sn 100

— —

— 0

— 50

— 100

— —

Pd–Sn Application

Pd–Sn —

Pd–Sn —

— —

Sn–Ag–Cu Begin 1999

Sn–Ag–Cu 80

Sn–Ag–Cu 95

MARKET, PRODUCT, AND CORPORATE POLICY TRENDS V.

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MATTERS IN EUROPE

European mobile phone and consumer electronics makers have introduced lead-free products, as have some telecommunications and automotive electronics manufacturers. Many additional European companies have indicated that they are testing lead-free materials in their products and will eventually introduce them for mass production.

A.

European Corporate Policies

Mobile phone companies such as Ericsson and Nokia have made lead-free and halogen-free product plans part of their environmental policies. While some consumer electronics companies, such as Philips, favor product recycling as the best environmental choice, most European companies have lead-free policies driven by legislation initiatives.

B.

European Consortia

The European Community (EC) has many lead-free programs dating back to the early 1990s. The European Brite-Euram Consortia started work in April 1996 under the Improved Design Life and Environmentally Aware Manufacture of Electronic Assemblies by Lead-Free Soldering (IDEALS). The consortium recommended Sn–Ag–Cu as the general purpose solder. The BriteEuram IDEALS program included Siemens, Marconi, Philips, Cookson, and the University of Ireland. The project, which ran from 1996 to April 1999, investigated many potential alternative alloys. It was determined that Sn–Ag–X lead-free solders offer superior creep resistance at room temperature and 100jC compared to Sn–40Pb. Results of this work have been presented, in addition to factors to consider when developing and implementing lead-free alloys such as manufacturability, availability, and cost. The International Tin Research Institute, ITRI Ltd., based in Uxbridge, England, has been involved in solder research since 1932 and has conducted lead-free solder studies since the late 1980s. ITRI opened Soldertec, a lead-free soldering technology center supported by more than 50 industrial members in 1999. The center studies a variety of aspects related to the development of lead-free technology, including production and reliability concerns. The European Cooperative administrated by the Science and Technical research (COST) is studying three lead-free solder alternatives, including two for wave soldering applications recommend by the U.S. consortia, NEMI. Sweden’s IVF Industrial Research and Development Corporation is conducting reliability testing with support from member companies such as Volvo, ABB, and Ericsson. A collaborative effort between Infineon Technologies, Philips Semiconductors, and ST Microelectronics is focused on establishing an upper lead concentration limit for semiconductor packages and common approaches for solderability testing and assessment of moisture-sensitivity levels [20].

C.

European Products

Alcatel is currently producing a line of lead-free mobile telephones and boasts the first lead-free assembly line in France. A line of lead-free PBX systems for telecommunications is offered worldwide. The company also has the distinction of having the second lead-free production line in Europe after Marconi Communications. Alcatel evaluated two lead-free alloys and determined that Sn–3.8Ag–0.7Cu provided the best results. In addition to conducting lead-free solder alloy research, Alcatel evaluated four PCB finishes—OSP, silver, palladium, and NiAu. As with other manufacturers, Alcatel depends upon the termination choices suppliers make available. Many lead frame terminations contain lead, but some suppliers are providing components with leadfree finishes such as pure tin or Ni/Pd. Bosch is involved with lead-free activities including solder paste alternatives. The likely candidates are alloys within the Sn–Ag–Cu or Sn–Ag–Bi solder systems. Some automotive electronics products at Bosch are already lead-free. According to material suppliers, Bosch introduced conductive adhesives in its hybrid devices to a significant extent.

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Ericsson has produced a mobile phone that is both lead-free and uses a halogen-free PCB. Ericsson has a joint venture with Sony for mobile phone design and production that may more closely link lead-free product development in Europe and Japan. In April 1999, Marconi Communications initiated production of a telephone manufactured with lead-free solder, hailed as the first lead-free line in Europe. The telephone utilized a Sn–Ag– Cu alloy that was introduced to Marconi’s existing surface mount assembly line. The only modification required was a slight alteration to the convection reflow profile [21]. In collaboration with Multicore Solders, Nokia developed a lead-free solder alloy suitable for use in mobile telephones. The alternative Sn–3.8Ag–0.7Cu alloy utilizes existing manufacturing processes to replace lead-based alloys in the assembly of surface-mount PCBs. In production trials, 20 Nokia 2110 GSM digital cellular handsets were assembled using the lead-free solder. The alternative alloy was found to exhibit better solderability than the traditional Sn–Pb solder. Siemens initiated lead-free solder development activities in the mid-1990s. The first Siemens demonstrators using lead-free solder included industrial PCs, automotive, and telecommunications products. These products all have double-sided SMT boards with complete lead-free solutions. The smallest component is a 0402 device and the largest is a QFP with a 0.5-mm pitch. After analyzing 162 different alloys, Siemens selected three. The melting point for the Sn–3.8Ag– 0.7Cu alloy was too high; therefore Siemens added bismuth or another element to lower the melting point.

VI.

MATTERS IN NORTH AMERICA

In the absence of legislation requiring the reduction or elimination of lead, either pending or enacted, North American electronics companies have not introduced many products that employ lead-free solders and materials. However, with an eye on both market activities and legislative actions in Europe and Japan, many North American companies have begun lead-free development programs that include assembly evaluations and testing of commercially available and inhouse-produced lead-free solder materials in their products. While many companies are involved in investigations and development of lead-free electronics and manufacturing, other than components, few have immediate plans to introduce systems in North America. Many North American companies are concerned that lead-free components and assemblies will not meet the reliability or functionality requirements necessary for high-end equipment supplied to banks, air traffic control systems, web-based businesses, and other mission-critical applications. Accordingly, manufacturers of aerospace and military electronics have no plans to introduce lead-free solders. There are too many reliability concerns to utilize lead-free solder materials in high-reliability equipment related to the mechanical characteristics of the materials themselves and the effects of high temperatures to process them. The effect of new and modified intermetallic compound phases within solder joints and at the interfaces is yet an additional concern, and there are many more.

A.

North American Corporate Policies

Lead-free programs are being adopted by a number of companies, especially original equipment manufacturers (OEMS) involved in multinational activities. Companies such as Hewlett-Packard are expected to introduce lead-free consumer products before adopting more widespread application. Personal computer makers are addressing the introduction of lead-free systems consistent with market demand. Network communications companies such as Cisco, Lucent, and 3COM are also investigating lead-free manufacturing for future products, whose introduction will also be based on market demand. Personal communications companies such as Motorola have announced lead-free portable products. Automotive makers, who have the additional problem of dealing with harsh conditions for electronics mounted outside the firewall, have been investigating lead-free solders capable of high-temperature stability and mechanical shock. Additional companies such as Hewlett-Packard, IBM, and Sun Microsystems have substantial


163

lead-free programs. Numerous component suppliers such as AMD, National Semiconductor, Texas Instruments, and others offer components with lead-free finishes.

B.

North American Consortia

The National Center for Manufacturing Sciences (NCMS) backed by a consortium of North American industrial corporations, national laboratories, and academic institutions conducted an evaluation of alternatives to eutectic Sn–Pb solder between 1992 and 1996. Corporate participants included ATT-Lucent, Ford Motor Company, GM-Delco Electronics, GM-Hughes Aircraft, Hamilton Standard, National Institute for Standards and Technology (NIST), Rensselear Polytechnic Institute, Rockwell International, Sandia National Laboratories, Texas Instruments, and the U.S. Navy’s Electronics Manufacturing Productivity Facility. The goal of the project was to determine whether safe, cost-effective, nontoxic, reliable lead solder alternatives could be identified. With over 79 alloys examined, the study results indicated that while a number of possible lead-free solders were functionally possible, there was no universal drop-in replacement for lead-based solders in electronics applications. This study is discussed in detail in Chap. 17. The North American Electronics Manufacturing Initiative (NEMI) is a consortium of more than 50 U.S. electronics manufacturers, suppliers, associations, government agencies, and universities. A NEMI task force, the Lead-Free Assembly Project, was formed with the goal to enable North American companies to produce lead-free products by 2001 and achieve the total elimination of lead solder by 2004. The group recommended an alloy in the Sn–Ag–Cu system as an industry standard and utilized it in their study. The NEMI lead-free roadmap called for an increase in component tolerances to 260jC based on the initial belief that the same degree of superheating (i.e., the heating above the melting point or liquidus) would be required in the case of lead-free solders as is the practice for Pb–Sn solders. It is now understood that acceptable lead-free solder joints can be made at much lower temperatures (approximately 235–240jC), so this criterion was reduced. Even with this reduction, it is probable that most components will eventually be built to be more temperature- and moisture-resistant. However, this transition may take many years. Electronics assemblers implementing lead-free processes may be faced with serious issues during this transition, issues that may continue to linger well past the time when market and legislation forces have made leadfree assembly essential or required, respectively. The High-Density Packaging User Group (HD-PUG) is a nonprofit organization of users and suppliers in the telecommunications and computer industries. HD-PUG includes members from North America, Europe, and Japan. The group, which endorsed NEMI’s recommended alloy, is involved in lead-free issues among other environmental concerns such as halogen-free electronic materials.

C.

North American Products

Lucent Technologies introduced a power system that produces SMC-SMW series d.c.-to-d.c. converter power modules using its patented SnTech pure-tin electroplating process. Components for the power device are supplied by National Semiconductor and Fairchild with a SnTech lead finish. Motorola’s first lead-free product was a two-way radio. Other lead-free products include pagers and mobile phones. Nortel Networks began its lead-free research in 1991 as part of an industry consortium with ITRI Ltd., Multicore Solders, and GE-Marconi. The company selected the desktop telephone for its lead-free product redesign because it allowed the company to test the lead-free manufacturing process on an assortment of lead shapes [22]. Prototype phones were constructed on existing manufacturing lines in 1996 and intended to be operationally equivalent to those manufactured with traditional Sn–Pb solder. When introduced in 1997, it was the first desktop telephone produced entirely with a lead-free interconnection technology. A commercially available (Multi-

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core Canada, Inc.) Sn–Cu alloy, with a melting point of 227jC, was used on the PCB of the Meridian 9316 office telephone. Nortel also employed lead-free component leads and used an organic surface protective (OSP) film on the substrate. Several factors motivated Nortel to produce the lead-free telephone, including public concern over the effect of lead on the human body and the defeated U.S. Lead Exposure Reduction Act in 1994. In addition, efforts in Europe to ban lead-containing electronics waste initiated by the Netherlands served as a driving force to develop the telephone [23]. Other contributing factors to Nortel’s lead-free work included cost issues: lead-free products would incur lower handling and disposal costs, and product end-of-life liability and disposal costs would be avoided. Visteon participates in external lead-free programs and has internal activities to develop lead-free solder materials. Visteon’s first lead-free application is for a mass airflow sensor used in vehicle engine compartments. Production started in the fall of 2001 with Sn–Ag–Cu alloy bumps fabricated by stencil printing by business partner Hitachi. The first application is a single chip with eight bumps mounted on a ceramic carrier.

VII.

FUTURE TRENDS

While an increasing number of environmentally friendly products continue to be introduced in Japan and Europe, some North American companies have taken a ‘‘wait and see’’ approach while others are pacing themselves based on market demands and supply line availability to the lead-free issue. Many Japanese companies have adopted the use of lead-free solder and lead-free components. The use of ‘‘green labels’’ helps to promote lead-free product introduction using guidelines established by the Japan Electronics and Information Technology Industries Association (JEITA). The combination of proposed legislative action, particularly in Japan and Europe, and the market appeal of environmentally friendly products has driven many leading electronics makers to begin lead-free solder implementation in their products [24]. Ultimately, the move to lead-free products will be market-driven. While the issue of lead elimination in electronic assembly remains a major topic at international conferences, alternatives are still debated, and there is no universal consensus choice. Higher processing temperatures, component compatibility, and reliability remain major concerns for lead-free alternatives. The establishment of the Global Environmental Coordination Initiative (GECI), which combines the efforts of nine consortia, is a promising development in assessing the activities of an ever-growing number of organizations generating lead-free-related information.

REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. 10.

Asakura, H. Lead-free solder needed for the LSI industry to survive: worldwide trend is accelerating, Pb-free is a condition for participation in the LSI business. Nikkei Microdevices July 2001, 193, 56– 69. United States Environmental Protection Agency. Private Sector Pioneers. Environmentally Preferable Purchasing Program; EPA742-R-99-001, June 1999. Johnson E. Select Committee on Environment, Transport, and Regional Affairs Report; http:// www.parliament.the-stationary-office.co.uk/pa/cm199899/cmse/ect/cmenvtra/149, Dec. 1998. Global Ecolabeling Network, http://www.gen.gr.jp.eco.html. Why Label a Product with a Swan? http://www.svanen.nu/Eng/producer. United States Environmental Protection Agency. Overview of Labeling Programs Worldwide. Environmentally Preferable Purchasing; EPA742-R-98-009, Dec. 1998. GreenBiz.com, Eco-labeling, http://www.greenbiz.com/resources/marketing/tools.cfm?LinkAdvID= 7839. Nordic Ecolabeling Board. Regulations for Nordic Ecolabeling of Products; http://www.svanen.nu/ Regelverk/regelverk_eng.pdf, Dec. 2001. Get to Know the Nordic Swan, http://www.svanen.nu. Stevels A. Green Marketing of Consumer Electronics. Proc. Electronics Goes Green Conf.; Berlin, Germany, Sept. 2000; 539–542.

MARKET, PRODUCT, AND CORPORATE POLICY TRENDS 11. 12. 13. 14. 15. 16. 17. 18. 19. 20. 21. 22. 23. 24.

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Stevels A.; Agema, R.; Hoedemaker, E. Green Marketing of Consumer Electronics II. Proc. EcoDesign 2001; Tokyo, Japan, Dec. 2001. Ohkubo, H. Technology Topics: Tokyo Eco-Recycling Plant Begins Operations. Denshi Zairyo June 2001, 40 (6), 101–103. Asakura, H. Lead-free solder needed for the LSI industry to survive: worldwide trend is accelerating, Pb-free is a condition for participation in the LSI business. Nikkei Microdevices July 2001, 193, 56– 61. Asakura, H. NEC cautions about problems with Sn–Zn solders. Nikkei Microdevices September 2001, 195, 141–143. Higashi, O., et al. Lead-free Assembly Technology for High Reliability Server Computer, presented at the EIAJ Conference, Japan, December 1999. Ito, M. Fujitsu and Fujitsu lab developed low cost mounting technique with Pb-free solders. Nikkei Electronics December 13, 1999, 759, 188–189. Baggio, T.; Suetsugu, K. Guidelines for lead-free processing. Surf. Mount Technol. September 1999, 60. JIEP Technology Roadmap for Electronics Packaging in 2010, translated by Semiconductor Equipment and Materials International, 1999; 32–36. Japan Jisso Technology Roadmap, EIAJ, August 1999; 51. BradleyE. Environmental work goes on. Electron. News January 21, 2002. Grusd, A. The Future is Now: No Longer Creeping towards Lead-Free. Proceedings of IPC Works International Summit on Lead-Free Electronics Assemblies; p. S-03-11-1, Minneapolis, Minnesota, October 27, 1999. Trumble, B. Getting the lead-free telephone to market. IEEE Spectrum, May 1998; 58 pp. www.NortelNetworks.com-corporate-community-habitat-commsol-green.html., November 1999. Krishna, J., et al. Thermal and Bending Fatigue of PBGA Assemblies with Lead-Free Solder Pastes. International Symposium on Microelectronics Proceedings; October 9–11, 2001; 553–558.

6 The Metallurgical Aspects, Properties, and Applications of Solders from the Lead–Tin System Paul T. Vianco Sandia National Laboratories, Albuquerque, New Mexico, U.S.A.

I. INTRODUCTION A. History of Soft Solders Containing Lead and Tin The history of joining with the metals lead (Pb) and tin (Sn) as well as their alloys dates back several thousand years [1]. Unfortunately, archeological evidence of the use of Pb and Sn is lacking for largely two reasons. First, these elements are degraded by corrosion under long-term exposure to groundwater and soil. Therefore, actual artifacts are rare. Second, there is an absence of written records of Pb and Sn use because such practices were generally performed by slaves and the lower-class artisans. Therefore, istorians or upper-class individuals having the ability to read and write had very little interest in these materials and associated processes [1]. The first of these elements used for joining was Pb. Lead was used to join copper (Cu) by the Mesopotamians around 3000 BC because of its availability in the region as well as its low melting temperature (327jC) that was compatible with wood-fired stoves and kilns. Tin was exploited around 1900 BC by the Celtic and Gaul cultures of Northern Europe where ore deposits were plentiful. It was not long before alloys of Sn and Pb, that is, the soft solders* were used by these cultures to join Cu or Cu-alloy tools and cooking utensils. The Romans, whose empire extended into Northern Europe, learned of Sn and Pb–Sn alloys and subsequently used them to seal the Pb liners of their aqueducts. It was not until the Industrial Revolution of the 19th century did the use of soft solders extend beyond light tool and jewelry-making applications. The availability of portable heat sources that used combustible gases or electricity enabled soldering to be performed at remote sites. Soon, Pb–Sn soldering found use in the plumbing of water and gas conduit; the construction of heat exchangers; the sealing of food containers; and the assembly of light-duty tools as well as sheet metal construction [2].

* The joining together of metal objects with liquid or ‘‘filler’’ metals and alloys actually began well before the use of Pb and soft solders. The Au-based filler metals or hard solders were used to construct jewelry items by the Mesopotamians as early as 4000 BC. 167

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B. Electronic Applications The 20th century ushered in the era of electronics in which Pb–Sn solders had their most farreaching role. The first electrical connections in early radio and telecommunications equipment were mechanical attachments. The application of solder to those connections prevented them from loosening, thereby reducing static in signal transmission and reception. The low melting temperature eutectic and near-eutectic compositions, 63Sn–37Pb (wt.%), 50Sn–50Pb, and 60Sn– 40Pb were used for these applications. Whether through-hole solder joints or advanced surface mount technology (Fig. 1), soldering with Pb–Sn alloys has remained as the primary assembly methodology in the electronics industry today. The materials systems used in electronic devices, components, and subassemblies have been specifically developed for processing with one or more of the Pb–Sn alloys. The plastic molding compounds for component packaging as well as the organic polymers used in circuit board laminates have been engineered to accommodate eutectic Pb–Sn solder assembly processes. Similarly, paste formulations and flux technologies as well as furnaces, soldering irons, and cleaning processes have been optimized for Pb–Sn interconnections. The versatility of Pb–Sn solders is exemplified by their use in the low-cost, high-volume assembly processes associated with consumer electronics, as well as in low-volume, highreliability avionics and military electronics. Advancements in Pb–Sn soldering in each of these arenas required research and development activities that have elevated a ‘‘garage practice’’ to the level of a critical materials technology. An appreciation of the mechanical properties of a Pb–Sn solder, such as yield and fatigue strength, requires an understanding of the microstructure of the particular alloy. Interface interactions between the Pb–Sn solder and a base metal result in the formation of intermetallic compounds. The intermetallic compound composition as well as both liquid-state and solid-state growth kinetics are critical factors in the long-term reliability of solder interconnections.

FIG. 1 A typical surface mount soldering technology printed wiring assembly. (Courtesy of Sandia National Laboratories.)

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C. Attributes of Pb–Sn Solders It is also important to appreciate the role that Pb–Sn solders have had in the miniaturization and high-volume manufacturing of electronics assemblies. It is these two attributes that have resulted in both low-cost, consumer electronics (e.g., cell phones, personal computers, etc.) as well as highly functional, military products (‘‘smart’’ weapon systems) available today. The highly functional devices, which are the backbone of consumer or military electronics, can be rendered inoperative by the failure of a single solder joint, that solder joint being one of several hundred or thousand interconnections on a package or assembly. The technical attributes of Pb–Sn solders have been accompanied by availability and cost advantages of the Sn and Pb metals. The ores of Sn and Pb are abundant. In addition, the Sn and Pb metals are readily extracted from those ores. The fact that Pb and Sn ores are located in numerous countries around the world serves to stabilize their respective price structures. Finally, the ease of reclaiming these metals from scrap keeps their secondary market prices low as well. Despite the many attributes exhibited by Pb–Sn solders in electronics applications, these materials are being highly scrutinized, primarily because of the potential impact of their Pb content on human health and the environment. An important action taken to limit Pb exposure to the U.S. population was banning Pb–Sn solders from plumbing conduit for potable water [3]. Today, the electronics industry is addressing the technical and cost issues that would accompany the elimination of Pb–Sn solders from both commercial and military products.

II. PHYSICAL METALLURGY ASSOCIATED WITH THE LEAD–TIN SYSTEM A. Microstructures and the Binary Alloy Phase Diagram of the Pb–Sn System 1. The Phase Diagram The importance of understanding the physical metallurgy of the Pb–Sn alloy system stems from the microstructure–properties relationship. That is, the microstructure of a Pb–Sn alloy, or any material for that matter, determines its mechanical properties. In the case of Pb–Sn solders, this relationship is particularly critical because (1) the microstructure of these materials is sensitive to the cooling rate used in an assembly process, and (2) the microstructure can be readily altered during the course of the service life experienced by an interconnection. The microstructure of Pb–Sn solders can be described by the binary alloy phase diagram. The phase diagram is a two-dimensional construct that predicts phase development as a function of the temperature and material composition. An important caveat is that the phase diagram represents the material at equilibrium. Strictly speaking, equilibrium is achieved when the material has been cooled at an infinitely slow rate such that no further changes occur to the microstructure at the target temperature. Unfortunately, an equilibrium microstructure is often not realized in Pb–Sn solders because of the conditions that prevail in manufacturing processes and applications that utilize these materials. First, manufacturing processes result in cooling rates that are relatively fast. Therefore, the microstructure is far from equilibrium. Second, applications, even those at room temperature (25jC) are close to the melting temperatures of Pb–Sn alloys. As a result, there is sufficient thermal energy to support solid-state diffusion processes, which will, in turn, alter the microstructure of the solder and thus the performance of interconnections. 2. The Pb–Sn Binary Alloy Phase Diagram Shown in Fig. 2 is the binary alloy phase diagram for the Pb–Sn system [4]. This rendition is based upon weight percent quantities of Pb and Sn. A calculation of the corresponding atomic percent Sn can be performed with Eq. (1). The atomic percent of Pb is (100% Sn at.%). Atomic percent Sn ¼ ½A=ð57:3 þ 0:427AÞ 100%

ð1Þ

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FIG. 2 The Pb–Sn binary alloy equilibrium phase diagram. (From Ref. 4.)

In Eq. (1), A is the weight percent of Sn. The abscissa in Fig. 2 is the alloy composition, beginning with 100% Pb and progressing to 100% Sn. The ordinate axis is temperature (jC). The various curved and straight lines separate areas referred to as phase fields. The nomenclature for these lines and phase fields will be illustrated by the following examples. 3. Lead–Tin System: Example 1—15Pb–85Sn The Pb–Sn phase diagram is reproduced in Fig. 3 [5]. A line has been drawn, representing the hypereutectic composition, 15Pb–85Sn. A hypereutectic alloy composition has a Sn concentration that exceeds the eutectic value of 61.8 wt.%, that is, a Sn concentration in the range of 61.8 < Sn (wt.%) < 97.5. The equilibrium phases are described as a function of decreasing temperature along that line. The line begins in the liquid phase (L) field at point A; the corresponding temperature is 275jC. The 15Pb–85Sn solder remains entirely liquid as the temperature is decreased to point B. At point B, the alloy enters the two-phase field designated L + h, where L is the remaining liquid phase and h is the newly formed, Sn-rich solid phase; the latter is referred to as the proeutectic solid phase. The temperature at point B is referred to as the liquidus temperature, which, in this example, is 210jC. The composition of the liquid phase remains that of 15Pb–85Sn. The composition of the proeutectic solid phase is represented by point BV, which is 1.5Pb–98.5Sn as determined by the projection of BV down to the x-axis. A further temperature decrease causes two phenomena to take place. First, there is an increase in the quantity of solid phase and a decrease in the quantity of liquid phase. The fluidity of the alloy is reduced and the material takes on the consistency of a paste or mush (i.e., slush); hence, this two-phase region is often referred to as the pasty range or the mushy zone. Second, there is a change in the respective compositions of the liquid and solid phases. For example, when the temperature of the alloy drops to that of point C, 195jC, the composition of the newly formed


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FIG. 3 The Pb–Sn binary alloy phase diagram reproduced to show equilibrium phase evolution of the 15Pb–85Sn composition as a function of temperature. (From Ref. 5.)

solid phase is CV and that of the remaining liquid phase is CW. The composition of the solid phase tracks along the line M–DV; this line is referred to as the solidus line. The composition of the liquid phase tracks along the line M–DW; this line is referred to as the liquidus line. At any temperature within the two-phase field, the horizontal line CV–CW connecting the liquid and solid phase compositions is referred to as the tie line. It was noted that within the two-phase field, the alloy is comprised of a mixture of solid and liquid phases. At any particular temperature, the weight fraction of solid and liquid phases can be determined by the lever rule. The weight fraction of solid phase is calculated by dividing the length C–CW by the total tie line length, CV–CW. Likewise, the weight fraction of liquid phase is computed by dividing the length C–CV by the length, CV–CW. a. Reaction at the Eutectic Temperature. The temperature is decreased further to that of point D, 183jC. The alloy is still two phased, with the solid phase having the composition of point DV (2.5Pb–97.5Sn) and the liquid phase having the composition associated with point DW, that is, 38Pb–62Sn or, the more commonly cited composition, 37Pb–63Sn. The alloy is 65% solid phase and 35% liquid phase. When the temperature is decreased to slightly less than 183jC, the remaining liquid phase solidifies and the alloy enters the two-phase field (a + h). The microstructure is a mixture of two phases, a and h. The composition of the Pb-rich a phase is determined by point Dj (19Pb–81Sn). The composition of the Sn-rich h phase remains as that of DV (2.5Pb–97.5Sn). The a and h phases are each solid solutions of Pb and Sn. Below 183jC, the a or Pb-rich phase exhibits a greater solubility for Sn than is the solubility of Pb in the Sn-rich h phase. b. Cooling Below the Eutectic Temperature. Further cooling of the 15Pb–85Sn composition results in relatively small changes to the compositions of the a and h phases. At point E (125jC), the compositions of the a and h phases correspond to points EW and EV, respectively. At

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room temperature (25jC), the phase diagram indicates that a phase is 99Pb–1Sn and the h phase is effectively 100Sn. c. Microstructure Depends on Cooling Rate. An optical micrograph of the 15Pb–85Sn solder is shown in Fig. 4. This particular sample was fabricated at a cooling rate of 10jC/min. A homogeneous, cellular microstructure had developed. The light regions are the Sn-rich h phase and the dark areas outlining the Sn-rich phase are the Pb-rich a phase. Faster cooling rates can produce proeutectic Sn-rich dendrites dispersed within the microstructure [6]. The microstructure depends upon cooling rate. Extremely fast cooling rates may not allow the Pb and Sn atoms to rearrange themselves to allow precipitation of the solid proeutectic, Snrich phase (point BV). The alloy remains entirely liquid for several degrees below the liquidus temperature. This phenomenon is referred to as undercooling. The faster the cooling rate, the greater is the difference between the liquidus temperature and the temperature at which solidification initiates and thus, the greater the deviation of the microstructure away from equilibrium. The nonequilibrium microstructure is retained after solidification because solid-state, mass transport (diffusion) processes are not sufficiently fast to rearrange the elemental atoms into an equilibrium configuration. The characteristics of the nonequilibrium, Pb–Sn solder microstructure are threefold. First, the compositions of the phases are not accurately represented by the phase diagram boundary lines. Second, the relative quantities of the phases are not accurately described by the lever rule. Third, the spatial distribution of phases in the microstructure are sensitive to the cooling rate. For example, in terms of individual phase compositions, excessive Sn may be retained in the Pb-rich a phase, causing a supersaturation condition. In effect, the phase boundary line between the a single-phase field and the (a + h) two-phase field is shifted to the right. (A more detailed discussion of approximating a nonequilibrium microstructure from the equilibrium phase diagram can be found in Refs. 5 and 7.) A consequence to the solder microstructure caused by a supersaturated Pb-rich phase is the precipitation of Sn particles in the Pb-rich phase. The above discussion targeted the 15Pb–85Sn composition. A similar analysis would be applicable to the hypoeutectic compositions that have Sn contents less than the eutectic value of 61.8 wt.%, that is, Sn concentrations in the range of 19 < Sn (wt.%) < 61.8. In the latter case, a Pb-rich (a), proeutectic phase forms as hypoeutectic Pb–Sn compositions are cooled through the pasty range (L + a) as would occur, for example, between the points F and G for the composition 70Pb–30Sn in Fig. 3.

FIG. 4 Optical micrograph of the 15Pb–85Sn solder microstructure. The cooling rate used to produce this microstructure was 10jC/min. (Courtesy of Sandia National Laboratories.)


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4. Lead–Tin System: Example 2—38.1Pb–61.9Sn (Eutectic Composition) Referring to Fig. 3, the 38.1Pb–61.9Sn alloy is the eutectic composition. However, more commonly, the eutectic composition is identified as 37Pb–63Sn as it is done in this handbook. Solidification of the eutectic composition occurs at a single temperature as the alloy cools from the liquid phase (L) to the (a + h) two-phase solid. The pasty ranges, (L + a) and (L + h), are avoided during solidification. The as-cast (i.e., as-solidified) microstructure of the Pb–Sn eutectic composition depends strongly upon the cooling rate [6]. A lamellar microstructure develops, as is illustrated in Fig. 5. A slower cooling rate produces dispersed Pb-rich phase particles in a Sn-rich phase matrix. The micrograph shown in Fig. 6 very effectively illustrates the latter microstructure, albeit it was produced from a sample isothermally aged at 70jC for 400 days. In general, time-independent strength properties (i.e., the stress/strain curve) increase as the interlamellae or interparticle spacing decreases. However, time-dependent strength behavior (creep) exhibits the reverse dependence. Because typical applied stresses and service temperatures can activate both timeindependent and time-dependent deformation mechanisms in eutectic Pb–Sn solder, the impact of the microstructure on solder joint performance cannot always be easily predicted. 5. Lead–Tin System: Example 3—95Pb–5Sn The third and final alloy used to illustrate the features of the Pb–Sn binary system is the 95Pb–5Sn composition. The composition line is shown in Fig. 7 [5]. The unique property of this composition, which represents all alloys having a Pb content that exceeds 81 wt.%, is that it solidifies without a eutectic reaction. a. Solidification Without a Eutectic Reaction. At point Z (350jC), the alloy is entirely liquid. At the liquidus temperature designated by point Y (315jC), solidification begins as the alloy enters the two-phase pasty range (L + a). The solid phase has the composition represented by the point YV, 97.5Pb–2.5Sn. The alloy remains in the (L + a) until the temperature has dropped to just above point X, 307jC. Then, the solid phase has the composition 95Pb–5Sn and the liquid phase has a composition 92Pb–8Sn. When the temperature drops below 307jC, the alloy solidifies completely and enters the (Pb) solid solution field. For temperatures between points X (307jC) and W (95jC), the solder is a single-phase solid solution material having the 95Sn–5Pb composition.

FIG. 5 Optical micrograph showing the lamellar microstructure in the 37Pb–63Sn solder resulting from a cooling rate of 10jC/min. (Courtesy of Sandia National Laboratories.)

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FIG. 6 Optical micrograph of the 37Pb–63Sn solder microstructure following isothermal aging at 70jC for 400 days. This micrograph is for illustrative purposes only, as it represents the solder microstructure that would be observed with slower cooling rates. (Courtesy of Sandia National Laboratories.)

FIG. 7 The Pb–Sn binary alloy equilibrium phase diagram reproduced to shown the 95Pb–5Sn composition profile as a function of temperature. (From Ref. 5.)


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b. Solid-State Solubility Adjustment. Decreasing the temperature below point W (95jC) results in the alloy entering the solid, two-phase (a + h) field. Phase separation occurs as the single-phase, solid-solution microstructure separates into the two-phase microstructure via solidstate diffusion processes. The a phase has the same composition as the starting alloy, 95Pb–5Sn (W). The newly appearing h phase will have the composition represented by the projection of point WV on the x-axis, 2Pb–98Sn. The solder microstructure remains in the (a + h) field upon further cooling to room temperature. The high Pb, a phase composition follows the line between the (Pb) and (a + h) fields, while the h phase composition follows that of the line between the h phase (Sn) and the (a + h) fields. The quantity of phases present follows the lever arm rule. Shown in Fig. 8 is the microstructure of the 95Pb–5Sn solder resulting from a cooling rate of 10jC/min. This cooling rate is sufficiently slow such that the resulting microstructure is a reasonable representation of equilibrium. The Sn-rich h phase has precipitated along the grain boundaries of the more prevalent, Pb-rich a phase. Under a very rapid cooling rate, a single Pbrich phase may be observed that is supersaturated with Sn. Subsequent to solidification, the h phase would form as Sn precipitates in the Pb-rich a phase; the rate of Sn precipitation and thus, h phase formation, would depend upon time and service temperature.

B. Physical Metallurgy of Intermetallic Compound Formation with the Pb–Sn System 1. Intermetallic Compounds as Interface Reaction Products The basis of filler metal joining, of which soldering is one of the two primary techniques (the other being brazing), is that the low-melting-temperature filler metal bonds to the base (or substrate) metal. Adhesion is provided by a reaction at the mutual interface, involving some or all of the constituent elements from either the solder filler metal or the base metal. Those reactions can take place during the actual soldering process and may continue afterward when the (solidified) interconnection is placed into service. The interface reaction product formed between Pb–Sn solders and most base metals is commonly referred to as an intermetallic compound layer. Intermetallic compounds exhibit highly directional bonding similar to ceramic materials and, as such, typically have a well-defined stoichiometry, high melting temperatures, high strength, and very low ductility. Unfortunately, the binary alloy phase diagrams provide only an approximate indication of intermetallic com-

FIG. 8 Optical micrograph of the 95Pb–5Sn solder microstructure. The cooling rate was 10jC/min. (Courtesy of Sandia National Laboratories.)

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pound compositions that may develop at the interface. The uncertainty arises due to underlying assumptions of the phase diagram—the constituents are fully intermixed; the constituents are of finite quantities; and the system is at equilibrium—do not prevail in solder/base metal couples. In fact, some solder/base metal couples form metastable phases at their mutual interface during solid-state (aging) reactions [8–10]. The metastable phases can appear and subsequently disappear during the development of the intermetallic compound layer.*

C. Tin as the Reactive Constituent Of the two elements comprising Pb–Sn solders, only the Sn constituent reacts directly with the base metal. The most commonly used base metal is copper (Cu). Two intermetallic compound phases are observed: the DV or Cu6Sn5 phase and the q or Cu3Sn phase [11]. The Cu6Sn5 phase is favored when the high-Sn, Pb–Sn solders (including the 37Pb–63Sn eutectic) are joined to copper [12]. A nominal layer thickness of 0.5–1.5 Am is observed in as-soldered joints [13]. Exposing the Cu base metal to molten Pb–Sn solder for several minutes at temperatures as high as 292jC— clearly, conditions that are well in excess of any automated or manual soldering process—causes only a very slight increase of the intermetallic compound layer thickness [12]. When the Sn content of the solder is reduced to approximately 27 wt.%, the intermetallic compound layer composition favors formation of the Cu-rich, Cu3Sn stoichiometry. The Cu3Sn layer actually develops from an initial Cu6Sn5 sublayer.

D. Liquid Solder Interface Reactions In the presence of molten solder, the interface reactions include the simultaneous process of base metal dissolution or base metal erosion as well as intermetallic compound layer formation. The extent of dissolution depends upon the composition of the base metal as demonstrated in Fig. 9 for the case of molten 40Pb–60Sn solder and the base metals Au, Ag, Pd, Pt, Ni, and Cu [14]. The dissolution rate of Cu, specifically as a function of temperature for several molten Pb–Sn compositions is shown in Fig. 10 [15]. A lower Sn content reduces the extent of base metal dissolution. Dissolved base metal can potentially saturate the nearby molten Pb–Sn solder field. Upon solidification, the base metal, M, precipitates out as MxSny intermetallic compound particles located near to, and at, the interface. For example, dissolved Cu will precipitate as Cu6Sn5 particles, forming a zone of such particles adjacent to a contiguous intermetallic layer. This morphology is shown in Fig. 11. Optical micrographs were taken of the 37Pb–63Sn solder/Cu interface after the base metal had been exposed to molten solder (215jC) for 5, 60, 300, and 900 sec.

E. Solid-State Interface Reactions Intermetallic compound layer growth can continue after solidification of the joint. The Sn component reacts with the base metal. The Pb component is rejected, forming a Pb-rich layer between the intermetallic compound layer and the solder field (see Fig. 12). The reaction kinetics have been studied by isothermal aging experiments. The aging of 37Pb–63Sn/Cu couples results in the Cu6Sn5 layer composition adjoining the solder field [16–18]. The Cu3Sn phase appears as a sublayer at the Cu6Sn5/Cu interface after prolonged time periods and/or higher aging temperatures. The graph in Fig. 13 shows the effect of aging time and temperature on the total intermetallic compound layer thickness for the 37Pb–63Sn/Cu couples [16]. The Cu3Sn sublayer

* The binary alloy phase diagram can provide some insight into potential intermetallic compound compositions [5]. For example, only single-phase fields in the phase diagram can form at the solder/base metal interface. Dual- or higher-order phase fields would not appear, as they cannot develop the concentration gradients required to support diffusion; rather, the latter would exist solely as interfaces.


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FIG. 9 Dissolution rates of Au, Ag, Cu, Pd, Pt, Ni, and Cu in molten 40Pb–60Sn solder as a function of solder pot temperature. (From Ref. 14.)

FIG. 10 Dissolution rates of Cu wire as a function of molten solder temperature for various Pb–Sn compositions. (From Ref. 15.)

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FIG. 11 Optical micrographs of the interface region generated by the exposure of Cu base metal to molten 37Pb–63Sn solder (215jC) for (a) 5 sec, (b) 60 sec, (c) 300 sec, and (d) 900 sec. (Courtesy of Sandia National Laboratories.)

appeared only under aging temperatures of 135j and 170jC. The reaction kinetics of the 37Pb– 63Sn/Cu system are described by Eqs. (2) (3) through (4): Cu3 Sn : x ¼ xo þ 1:82 103 t0:58 expð56; 100=RTÞ Cu6 Sn5 : x ¼ xo þ 3:87 103 t0:31 expð40; 400=RTÞ

xo ¼ 0 m xo ¼ 1:26 106 m

Cu3 Sn þ Cu6 Sn5 : x ¼ xo þ 1:08 102 t0:36 expð45; 500=RTÞ xo ¼ 1:26 106 m

ð2Þ ð3Þ ð4Þ

where x is the layer thickness (m), xo is the initial layer thickness (m), t is time (sec), R is the universal gas constant (8.314 J/mol K), and T is the temperature (K).


179

FIG. 11 Continued.

The Cu3Sn intermetallic compound (sub)layer composition is favored under conditions in which Sn availability is limited. A reduced Sn supply occurs when there is only a very thin Pb–Sn solder layer (e.g., an electroplated Pb–Sn coating) present over the Cu base metal, or when the solder composition has a low Sn content. For example, use of high-Pb solders on Cu, such as the 95Pb–5Sn composition, favors the formation of the Cu3Sn composition [19].

F. Interface Reactions with Nickel and Iron Lead–tin solder joints are commonly made to nickel (Ni), the Ni being either a bulk material or a thin solderable finish. The reaction results in the formation of primarily Ni3Sn4 intermetallic compound. The Pb component of the solder has no explicit role in the interface reaction. The reactivity between the Sn component of the Pb–Sn solder and Ni is considerably slower compared to Cu for the liquid-state dissolution reaction (see Fig. 9) as well as the

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FIG. 12 Optical micrograph of the interface between 37Pb–63Sn solder and copper (Cu) showing the intermetallic compound layer that developed after solid-state aging at 135jC for 400 days. (Courtesy of Sandia National Laboratories.)

solid-state interface reactions [20,21]. The low reactivity between Pb–Sn solder and Ni allows the latter to be an effective thin-film barrier layer between the solder and other underlying base metal materials. A very low reactivity has also been observed between Pb–Sn solders and iron (Fe). The observed intermetallic compound layer has the composition FeSn2 [22]. Base metal dissolution data are not available for molten Pb–Sn solders on Fe. Solid-state intermetallic compound layer growth at the 37Pb–63Sn/Fe interface is negligible (< 0.3 Am) [23]. The aging of 37Pb–63Sn/Fe couples at 170jC for 100 days produced a layer thickness of 0.86 F 0.57 Am compared to 17 Am for 37Pb–63Sn/Cu couples exposed to the same conditions.

FIG. 13 Total intermetallic compound layer (Cu6Sn5 + Cu3Sn) thickness as a function of aging time and temperature for 37Pb–63Sn/Cu couples. (From Ref. 16.)

SOLDERS FROM THE LEAD–TIN SYSTEM G.

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Interface Reactions with Gold, Silver, and Palladium

An important interface reaction is that which occurs between Pb–Sn solders and gold (Au). Gold is rapidly dissolved into Pb–Sn solders (see Fig. 9). Fast solid-state interface reaction rates between Pb–Sn solder and Au can result in the rapid growth of an intermetallic compound layer; even aging at room temperature. Layer development can degrade solder joint strength, particularly because of the poor adhesion between the Au–Sn intermetallic compound layer and the remaining Au layer (Fig. 14). The intermetallic compound layer compositions that develop between Pb–Sn solder and Au as a consequence of solid-state reactions are AuSn2 and AuSn4 [24]. The AuSn2 appears after relatively short-term aging and is located between the Au-based metal and AuSn4 layer that adjoins the solder field (Fig. 15a). Further solid-state aging causes the AuSn2 layer to disappear in favor of the AuSn4 composition (Fig. 15b). The ternary phase diagram has been constructed for the Au–Pb–Sn system [25,26]. The reaction kinetics for solid-state growth of the total intermetallic compound layer (AuSn2 + AuSn4) in 37Pb–63Sn solder/Au couples is described by the following equation [24]: x ¼ xo þ 3:87 103 t0:31 expð40; 400=RTÞ

xo ¼ 1:26 106 m

ð5Þ

where x is the total layer thickness (m), xo is the initial layer thickness (m), t is aging time (sec), R is the universal gas constant (8.314 J/mol K), and T is the aging temperature (K). The Pb component of the solder, which does not participate in the intermetallic compound reaction, accumulates at the interface between the intermetallic compound layer and the solder field. Two other base metals utilized as finishes in electronics assemblies to form Pb–Sn solder joints are silver (Ag) and palladium (Pd). The predominant intermetallic compound compositions that develop for these base metals are Ag3Sn and PdSn4, respectively [22]. The relative dissolution rates of these base metals in 40Pb–60Sn solder are shown in Fig. 9. There have been no long-term aging studies performed on Pb–Sn/Ag or Pb–Sn/Pd couples from which to predict the solid-state intermetallic compound layer morphology or growth kinetics.

FIG. 14 Optical micrograph showing a crack in a 37Pb–63Sn solder joint made to a Au coating. The crack formed at the interface between the Au–Sn intermetallic compound and the remaining Au layer. (Courtesy of Sandia National Laboratories.)

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FIG. 15 SEM micrographs showing development of the intermetallic compound layer between 37Pb–63Sn and Au base metal following solid-state aging for 90 days at (a) 70jC and (b) 100jC. (Courtesy of Sandia National Laboratories.)

H. Intermetallic Compounds in the Bulk Solder The inadvertent presence of intermetallic compound particles in Pb–Sn solder is most often a result of excessive base metal dissolution into molten solder. In most cases, the base metal is a single element such as Cu, Ni, or Au. In other applications, the base metal may be an alloy of several elements, in which case, some of the elements may exhibit preferential dissolution into the solder [27]. Intermetallic compound particles (e.g., Cu6Sn5, AuSn4, Ni3Sn4, PdSn4, etc.) precipitate upon solidification, when the base metal concentration exceeds its solid solubility limit in either the Pb-rich or Sn-rich phases. The solubility limits are nearly zero for most Sn/base metal combinations as shown in Table 1 [28].


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TABLE 1 Solubility Limits of Common Base Metals in Tin Metal Au Bi Cu Ni Pb

Pd Pt

Temperature (jC)

Solubility limit (wt.%)

0–100 100 50 0–100 0–100 100 50 0 0–100 0–100

f0 10 4 f0 f0 1.5 0.7 f0 f0 f0

Intermetallic compound particles may not develop upon solidification when the solder is cooled at a rapid rate. The solder becomes supersaturated with the base metal element(s). However, the precipitation reaction may occur later on as a thermally activated process, during the service life of an interconnection. Finally, some contaminants simply do not react with Sn. For example, Bi will precipitate as simply Bi particles when it exceeds the solubility limit in Sn (generally 3–5 wt.%) or Pb (0 wt.%).

I.

Effects of Intermetallic Compound Formation on the Mechanical Properties of Pb–Sn Solder Joints

The impact of intermetallic compound layers on the mechanical properties of Pb–Sn solder joints can be described in general guidelines only. The absence of more specific rules is a result of the mechanical and material complexity of a solder joint and, more specifically, the high strain sensitivity of Pb–Sn solders. The mechanical complexity of a solder joint stems from the interactions between gap size, solder joint area, and applied load (tensile, shear, or mixed mode). A thinner gap causes a joint to appear stronger under tensile stress than would be calculated from the bulk strength properties of the solder. The higher joint strength arises from the plane strain constraint imposed on solder deformation by the joint area. In addition, as the intermetallic compound layer occupies a larger percentage of the thickness of smaller joints, its properties have a greater influence on those of the overall interconnection. The material complexity of solder joints arises from the presence of the base metal(s), the intermetallic compound layers, and the ‘‘bulk’’ solder [5]. Each one of those materials has different mechanical properties (e.g., yield strength, ductility, ultimate strength, etc.). In addition, Pb–Sn materials have strength properties that are very strain rate sensitive. The yield strength and ultimate tensile strength values increase dramatically with faster strain rates [29]. Also, there are two interfaces associated with each intermetallic compound layer—the base metal/intermetallic compound interface and the solder/intermetallic compound interface—which can exhibit significantly different adhesion properties. The incorporation of intermetallic compound particles into a solder can cause the strength of the solder to increase via the precipitation hardening phenomenon. The degree of increase depends upon the concentration of those particles. Concurrent with the strength increase is a loss of ductility. However, an excess of intermetallic compound particles will degrade the solder and, hence, solder joint mechanical properties. The Pb–Sn solders and specifically, the 37Pb–63Sn and 40Pb–60Sn alloys, are embrittled when the Au content exceeds approximately 4 to 5 wt.%. A high concentration of AuSn4 particles results in an easy path for crack propagation [30–32]. This point is illustrated in Fig. 16.

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FIG. 16 Optical micrograph showing crack propagation in a 37Pb–63Sn solder joint containing an excess of AuSn4 intermetallic compound particles. (Courtesy of Sandia National Laboratories.)

J. Guidelines In summary, the overall impact of the intermetallic compound layer on the strength of the solder joint is dependent upon the factors and effects listed in Table 2. Given the factors/effects in Table 2, the following guidelines have been developed: Failures will occur preferentially in the Pb–Sn solder under slow strain rate load conditions. As such, an intermetallic compound layer has minimal impact on failures that occur due to timedependent deformation (creep) and low-cycle fatigue (e.g., thermal mechanical fatigue). The failure path occurs in the solder, near to one of the intermetallic compound layers. Cracking in the solder is preempted when one of the interfaces associated with the intermetallic compound layer is inherently weak, such as in the case of the AuSn4/Au structure (Fig. 14).

TABLE 2 Factors that Affect the Role of the Intermetallic Compound Layer on the Strength of Solder Joints Factor Solder intrinsic strength Solder strain rate sensitivity Intermetallic compound material properties Gap size of the solder joint

Rate of applied load (stress)

Effect Solder strength equals or exceeds the strength of the interface structures. Faster strain rate increases the apparent strength of the solder [see (1) above]. Brittle intermetallic compound and/or poor adhesion with the solder and/or base metal. A thinner gap causes an increase in the apparent strength of the solder [see (1) above]. Strain rate sensitive solders will exhibit an apparent increase in strength at faster strain rates [see (1) above].


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As solder strength increases, there is an increased likelihood for failure to occur in an intermetallic compound layer. A study of Mode I failures in several solders joints containing the Cu–Sn intermetallic compound layer resulted in the following trend [33]: Intermetallic compound layer failure was preferred when the local displacement rate was greater than or equal to approximately 2 103 mm/sec and the apparent yield strength of the solder was in excess of 20 MPa. The likelihood of failure in the intermetallic compound layer increases with the inherent brittleness and/or thickness of the layer. For example, the Ni3Sn4 intermetallic compound tends to be more brittle than the Cu6Sn5 compound [34]. A thicker layer increases the likelihood that intrinsic flaws will initiate cracking in the brittle intermetallic compound material.

III. PB–SN SOLDERING PROCESSES A. Liquid Solder Wettability Wettability is the capacity of the molten solder to react with the base metal at their mutual interface and establish a bond there. The wettability of liquid solders is largely an intrinsic property between the specific solder composition and the base metal composition. On the other hand, spontaneous spreading is impacted by several parameters: base metal surface condition, wettability (solder/base metal interface reaction), and the flux. The synergistic roles of the molten solder, base metal, and flux on the combined properties of wetting and spreading, that is, the alloy’s solderability, can be described by Young’s equation: cSF cSL ¼ cLF cos hc

ð6Þ

where cSF is the solid (base metal)–flux interfacial tension; cSL is the solid (base metal)–liquid (molten solder) interfacial tension; cLF is the liquid (molten solder)–flux interfacial tension; and hc is the contact angle. Young’s equation represents the equilibrium balance of the interfacial tension parameters for both horizontal and vertical surfaces, as shown schematically in Fig. 17 [5]. The contact angle provides a generalized parameter for solderability, as it takes into account the effects of all of the interfacial tensions. As such, solderability is optimized when the contact angle, hc, is minimized. The contact angle is minimized by maximizing cSF; this is accomplished by the removal of contaminants and the oxide layer from the base metal surface. The value of hc can also be minimized by a small value of cSL. The value of cSL is determined by the solder–base metal interface reaction and, thus, can be altered by using a different base metal or placing a metal coating of alternative composition over the base metal surface. Variations among the Pb–Sn solder compositions will have little impact on cSL since the same Sn-based intermetallic compound will be formed at the interface. Finally, the contact angle can also be minimized by minimizing the

FIG. 17 Schematic diagrams of solder spreading on horizontal and vertical open surface configurations. The interfacial tensions balance one another, determining the contact angle as defined by Young’s equation. (From Ref. 5.)

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solder–flux interfacial tension, cLF. The value of cLF is dependent on the compositions of both the flux and the solder.

B. Solderability Performance of Pb–Sn Alloys The excellent solderability of Pb–Sn solders is a result of their good wetting and spreading properties [5]. Wettability is the capacity of a molten solder to react with the base metal at their mutual interface and establish a bond. In the Pb–Sn system, the interface reaction is typically between Sn and the base metal, in the form of an intermetallic compound layer. Spontaneous spreading of molten solder over a horizontal surface ends with the sessile drop profile. On the other hand, spontaneous spreading on a vertical surface results in a meniscus, which then solidifies into a solder fillet. Good spreadability allows molten Pb–Sn solders to cover a base metal surface without necessarily having to be placed over that entire surface. Lead–tin solders also exhibit good solderability in closed solder joint geometries such as gaps and holes. The spontaneous filling of through-hole solder joints is assisted by capillary flow. A contributing factor toward the excellent solderability of Pb–Sn alloys is the low surface tension of molten Pb–Sn solders, which is a direct result of the Pb constituent in the solder. The surface tension of molten Pb is 450 dynes/cm (at Tmelt = 326jC) and that of molten Sn is 550 dynes/cm (at Tmelt = 232jC) [35].* Listed in Table 3 are the solderability parameters, hc, cLF, and (cSF cSL) of the 40Pb–60Sn and 96.5Sn–3.5Ag solders for the same test conditions [37]. The values of these parameters were determined by the meniscometer wetting balance technique [38]. A lower contact angle is exhibited by the Pb–Sn solder in Table 3, when compared to corresponding 96.5Sn–3.5Ag solder data, a result that is directly attributed to the reduced value of cLF of the Pb–Sn alloy. Also included are solderability parameters for these solders on Au–Ni-plated Kovark.y The further reduction in the contact angle was a direct result of the reduction in cLF which, in turn, resulted from the dissolved Au preferentially locating at the solder–flux interface. The low surface tension (and low cLF interfacial tension) of Pb–Sn alloys also contributes to excellent capillary flow. Molten Pb–Sn solders can readily enter small clearances associated with through-hole solder interconnections and other configurations.z This behavior was documented in a series of capillary flow experiments carried out between parallel Cu plates, using the specimen shown in Fig. 18a; the superior meniscus (capillary) rise of Pb–Sn solders as a function of gap thickness, when compared to several Pb-free solders, is shown in Fig. 18b [5,39].

C.

Inadequate Wetting Conditions

Different conditions of Pb–Sn solderability are illustrated by the stereo photographs in Fig. 19. Adequate solderability results in a solder film that uniformly covers the base metal surface (Fig. 19a). Inadequate solderability can result in two defective conditions: nonwetting (Fig. 19b) or dewetting (Fig. 19c). The condition of nonwetting occurs when the solder simply does not wet and spread over the base metal surface. The solder ‘‘balls up,’’ leaving exposed base metal surface. The second defect, dewetting, is actually a two-step process. Initially, the relatively thick solder readily wets and spreads over a marginally solderable, base metal surface. The film thickness actually supports that initial wetting and spreading action. Then, as the solder film thins out

* Similarly, bismuth (Bi) has a relatively low surface tension, a property that is imparted into alloys with this constituent: 58Bi–42Sn and 91.84Sn–3.33Ag–4.83Bi, thereby improving the solderability of the high-Sn, hightemperature alloys [36]. y Kovar is a trademark of Carpenter Technologies, Inc., Reading, PA. z It is interesting to note that the often-cited equation for capillary rise shows a direct proportion between meniscus rise and interfacial tension, which is opposite to the observed trend. This discrepancy arises from an additional effect caused by the relatively high surface tensions of molten solders and the relatively small gap sizes of solder joints. The consequence is a constraint on the minimum contact angle that can be realized by the molten solder: a high cLF value causes a high hc value, which reduces solderability.


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TABLE 3 Solderability Parameters hc, cLF, and (cSF cSL) for 40Pb–60Sn and 96.5Sn–3.5Ag Solders on Cu- and Au–Ni-plated Kovar Base metal Cu Au–Ni plated Kovar Cu Au–Ni plated Kovar

Solder alloy (wt.%)

Temperature (jC)

Contact angle (jC)

cLF (dyne/cm)

(cSF cSL)

40Pb–60Sn 40Pb–60Sn

260 260

17 F 4 11 F 9

380 F 10 330 F 15

370 F 10 325 F 25

96.5Pb–3.5Sn 96.5Pb–3.5Sn

260 260

36 F 3 34 F 4

460 F 30 370 F 30

380 F 40 310 F 40

Source: Ref. 5.

FIG. 18 The parallel plate capillary rise test geometry (a) used to determine the meniscus rise, (b) as a function of gap size for the 37Pb–63Sn, 95Sn–5Sb, 96.5Sn–3.5Ag, 91.84Sn–3.33Ag–4.83Bi, and 95.5Sn–4.0Cu–0.5Ag solders. (From Ref. 5.)

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FIG. 19 Stereo photographs showing the cases of (a) wetting, (b) nonwetting, and (c) dewetting of a 40Pb–60Sn solder film on an Fe–Ni alloy. (Courtesy of Sandia National Laboratories.)

toward completion of the joint, it can no longer support coverage of a marginally solderable, base metal surface and retracts into the localized mounds of solder separated by areas of a thinly coated, base metal surface.

D.

Interfacial Reactions—Coatings on Base Metals

Some applications require the Pb–Sn solder to be used on a base metal that is not readily solderable. In order to provide a processable level of solderability for difficult-to-solder materials, metal coatings are applied to the base metal surface. 1. Protective Finishes One approach uses a single coating—the protective finish. As the terminology implies, the protective finish prevents the loss of base metal solderability due to rapid reoxidation prior to soldering. The role of the protective finish in a soldering process is illustrated in Fig. 20 [5]. The protective finish is initially wetted; therefore, it must be solderable. Then, the coating is dissolved into the molten solder. The solder proceeds to wet and spread over the underlying (pristine) base


189

FIG. 20 Schematic diagram showing the processes associated with molten solder spreading over a protective finish. (From Ref. 5.)

metal surface. The protective finish must be sufficiently thick to protect the base metal solderability, but not too thick so as to pose a significant contamination concern for the Pb–Sn solder joint. Common protective finishes include the Pb–Sn alloys (37Pb–63Sn or 95Pb–5Sn) as well as 100Sn. Some ‘‘Sn coatings’’ actually contain a small Pb addition to prevent the formation of tin whiskers, which are of concern with 100Sn electroplated layers. Lead–tin or Sn coatings along with recommended thicknesses are listed in Table 4 [40]. The benefit of fusing, that is, melting a Pb–Sn or Sn electroplated layer, is that the resolidified layer provides a more effective barrier against contamination and reoxidation. The preference for hot-solder dipped coatings is decreasing for microelectronics assemblies due to the variation in film thickness due to (molten solder) surface tension effects and the local geometry (flat surface vs. a corner). The lack of planarity is a concern for fine-pitch applications. However, an attribute of hot-solder dipped coatings is that they provide an on-the-spot assessment of the base metal solderability prior to final assembly. 2. Solderable Finishes In those cases in which it is impractical to sustain base metal solderability, a metallic coating is applied to the base metal. The molten solder wets and spreads over the coating and, ultimately, makes a bond to it. This coating is referred to as the solderable layer. Typically, a protective finish

TABLE 4 Commonly Used Pb–Sn and Sn Protective Finishes Composition Electroplated Pb–Sn (37Pb–63Sn or 40Pb–60Sn) Electroplated Sn (100Sn) Electroplated and fused Pb–Sn Electroplated and fused Sn Electroplated Pb–Sn and Sn Source: Ref. 40.

Thickness (Am) 7.6–23 7.6–13 2.5–13 2.5–13 >51

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FIG. 21 Schematic diagram depicting the processes associated with molten solder spreading over a solderable finish/protective finish combination. (From Ref. 5.)

is added on top of the solderable layer to protect the solderability of the latter after its deposition, but before the soldering process. The functional limitations placed on solderable finish/protective finish systems are illustrated in Fig. 21 [5]. As noted above, spreading of the solder front requires that the surface of the protective finish be solderable (free of contamination or oxidation). The protective finish is sacrificial as it is dissolved into the solder. Therefore, the protective finish must be sufficiently thick to protect the solderability of the underlying solderable layer, but not excessively thick to pose a significant contamination concern for the subsequent solder joint (e.g., Au in Pb–Sn solder). The solder then wets to the surface of the solderable layer. The requirements placed on the solderable layer are the following: it must be solderable; it must have adequate thickness so as not to be completely dissolved by molten solder during the initial assembly process or during subsequent process steps (including repair or rework cycles); and the layer must adhere to the base metal surface. Several solderable and protective finishes along with their thickness as specified by MILSTD 1276D are listed in Table 5 [5,40]. One of the most commonly used solderable-plusprotective layer combinations is Ni and Au. The likelihood of Au embrittling Pb–Sn solder joints is reduced by a hot-solder dipping operation that replaces the Au with a Pb–Sn finish [41,42]. Copper is a frequently used solderable coating; it is often coated with a Pb–Sn or Sn protective finish (electroplated, plated and fused, or hot-solder dipped). Other protective layers include

TABLE 5 Commonly Used Solderable Finish/Protective Finish Combinations Solderable layer (Am)

Protective layer (Am)

Ni (1.5–3.8)

Au (1.3–2.5)

Ag (3.8–8.9)

None or hot solder dipped None or hot solder dipped

Ni (1.5–3.8)

Ni–Pd (1.3–3.5) Source: Refs. 5 and 40.

None or hot solder dipped

Types of base metals Cu, Cu alloys, Zn, Al, Fe–Ni alloys, steels Ni and Ni alloys, stainless steels Ni and Ni alloys, Cu, Cu alloys, Al and Al alloys* Stainless steels


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electroplated Ag and Pd–Ni alloy. The solderabilities of both the Ag and Pd–Ni coatings can degrade due to tarnishing (formation of Ag2S) and oxidation (PdO), respectively.

E.

Flux Technology and Controlled Atmospheres

1. Function of the Flux The flux provides three functions in a soldering process: (1) removal of light oxide layers from the base metal (or coating) surface; (2) prevention of base metal reoxidation and excessive oxidation of the molten solder surface; and (3) reduction of the surface tension of the molten solder. Current flux technology and associated solder paste materials have been developed almost exclusively for eutectic, near-eutectic, and high-Pb, Pb–Sn solders. The coupling of flux technology to Pb–Sn solders grew in response to the need for high-volume, low-defect soldering processes for the electronics industry. Those processes include furnace reflow, wave soldering, and hand soldering. The three flux functions must be accomplished within narrowly defined process cycles. 2. Flux Categories Soldering processes associated with Pb–Sn solders have been performed with fluxes from each of three categories, the (natural) rosin-based fluxes, resin-based fluxes, and the organic acid (watersoluble) fluxes.* a. Natural Rosin-Based Fluxes. The natural rosin-based fluxes have been the primary chemistries used with the 37Pb–63Sn eutectic and 40Pb–60Sn off-eutectic compositions. These fluxes thermally degrade at temperatures in excess of approximately 275jC, thereby limiting their use to the eutectic and near-eutectic Pb–Sn alloys [43]. There are three categories of rosin-based fluxes distinguished by activity level [44]. The accepted classification framework for fluxes utilizes an alphanumeric code described in the ANSI/J-STD-004 specification [45]. (1) R (‘‘rosin’’) fluxes are the least active, containing pure water-white rosin with a solids content in the range of 10% to 40%, depending upon the desired strength level. (2) RMA (‘‘rosin-based, mildly activated’’) fluxes are moderately activated with 0.1% to 0.3% activator and contain a 10% to 40% solids content. (3) RA (‘‘rosin, activated’’) fluxes have 1.0% to 1.5% activator to maximize corrosion activity and are composed of 10% to 40% solids. Two types of rosin-based fluxes were developed for Pb–Sn solders to eliminate the use of solvents for removing residues. Those flux types are the no-clean fluxes and the low-solid (residue) fluxes. The no-clean fluxes have solids contents typical of the rosin-based flux groups (35% to 45%). However, they have been engineered so that when exposed to Pb–Sn soldering temperatures (typically 215j to 260jC), the organic molecules polymerize and harden, effectively sealing the activators from outside moisture to render the residues noncorrosive. A processing advantage of no-clean fluxes is that they perform like the traditional rosin-based fluxes. Low-solid fluxes have, as the term implies, chemistries with solid contents reduced to low levels, typically in the range of 2% to 8%. Their activities range between a typical R flux and an RMA flux, depending upon the level of activator(s) added to the base chemistry. The reduced solid content leaves little or no residue remaining on the assembly after completion of the soldering process. The consequence of a lower solid content is that the process window is considerably narrower when compared to rosin-based or no-clean fluxes. Low-solid fluxes have been specifically engineered for 37Pb–63Sn and 40Pb–60Pb soldering processes. The process window of low-solid fluxes can be widened to some degree through the use of an inert gas (e.g., N2) in the reflow furnace. b. Resin-Based, Synthetic Fluxes. The resin-based fluxes are synthetic materials having activator additions that give them activities between the RMA and RA categories of natural rosin-based fluxes. The designation for these fluxes is resin-based, synthetic, activated (RSA). The

* There is also a fourth category of soldering flux, the inorganic acid fluxes. These highly active fluxes are not used in electronic packaging or assembly applications because of potential corrosion damage posed by their residues. They are typically used in structural soldering applications such as plumbing and jewelry repair.

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particular attribute of RSA fluxes is that they are more compatible with higher processing temperatures; thermal breakdown occurs at temperatures of 350j to 375jC [43]. These fluxes can accommodate the high-Pb, 90Pb–10Sn, and 95Pb–5Sn solder processes. c. Organic Acid or Water-Soluble Fluxes. The organic acid fluxes, also called watersoluble fluxes or ‘‘intermediate fluxes’’ have strength levels commensurate with RMA and RA fluxes, and have been used extensively with eutectic and off-eutectic Pb–Sn solders. Thermal breakdown for these materials occurs at temperatures in the range of 250j to 275jC so that they are of marginal use with the high-Pb solder processes. The organic acid fluxes utilize aqueous or semiaqueous cleaning materials, thereby eliminating the need for environmentally damaging, solvent-based cleaning processes. 3. Controlled Atmospheres a. Passive Atmospheres Passive controlled atmospheres assist the flux by excluding oxygen (O2) during the heating cycle, thereby limiting reoxidation of the base metal and molten solder surfaces. A secondary benefit of passive atmospheres is that they reduce the rate of thermal degradation of the flux components (that is also a type of ‘‘oxidation’’ process but, which occurs to polymer chains). Typical passive atmospheres include nitrogen (N2) and argon (Ar). Nitrogen is preferred for controlled atmosphere soldering [46,47]. It is the least expensive gas to produce. Nevertheless, the use of a N2 atmosphere extols a significant cost penalty on the assembly process. Specially equipped furnaces are required to support a N2 reflow environment. Additional plant costs are incurred for providing the N2 supply as well as to transport N2 gas to the furnace location. Residual O2 levels less than 100 to 200 parts per million (ppm) and, preferably, in the range of 20 to 50 ppm, can significantly reduce Pb–Sn solder defects that occur during assembly. Of course, the controlled atmosphere that most effectively excludes O2 from the process is a vacuum. However, the use of vacuum furnaces for large-volume soldering assembly processes is usually cost prohibitive. b. Reactive Atmospheres. Reactive, controlled atmospheres actually reduce base metal or molten solder surface oxides. Strictly speaking, inert and vacuum atmospheres can be reactive by reducing metal oxides due to the low O2 partial pressure above the metal surfaces. Unfortunately, the thermodynamics of the reduction reaction are not favorable for relevant base metals such as Cu, Ni, Sn, or Pb at the relatively low temperatures used in soldering processes.* Similarly, the effectiveness for hydrogen gas (H2) or carbon dioxide (CO2) to reduce metal oxides is extremely low for common base metals exposed to Pb–Sn solder process temperatures. An alternative approach to reactive, controlled atmospheres for Pb–Sn solder processes has been to mix acid vapors with an inert carrier gas (N2) [48–50]. The acid vapors provide the fluxing action that reduces oxides on the base metal and molten solder surfaces. The inert gas also prevents reoxidation of those new, pristine surfaces. Gas mixtures of 5% formic acid or 3% acetic acid in Ar have been shown to provide adequate solderability of 37Pb–63Sn solder on Cu with little remaining residue. Although residues are considerably less than those that remain after the use of conventional Pb–Sn solder pastes, cleanliness tests should be performed to assure that residue levels meet product acceptance criteria.

F. Heating Methods Used for Assembly with Pb–Sn Solder Generalized time and temperature parameters have been identified for a number of Pb–Sn solder assembly processes. Those parameters are listed below for reflow furnace soldering, vapor phase soldering, wave soldering, and hand soldering.

* The two elements having a thermodynamic potential for oxide reduction under partial pressure O2 at soldering temperatures are Ag and Pd. However, the common surface tarnish (film) that forms on Ag is the sulfide, Ag2S. In the case of PdO, the rate kinetics for oxide reduction are too slow for a vacuum to improve solderability under typical processing conditions.


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FIG. 22 Schematic diagram of the generic time–temperature profile used for the in-line furnace processing of printed circuit assemblies.

1. Furnace Heating The preferred technique for large-volume soldering is the in-line reflow furnace. A schematic diagram showing the time–temperature heating profile is shown in Fig. 22. The process parameters are the heating and cooling rates (DT/Dt) as well as the preheat and reflow temperatures and times (Ti, ti). Some typical parameters for 37Pb–63Sn solder reflow using an RMA flux are listed in Table 6. The parameters will vary with the size of the work piece. However, they are similar for processes conducted in air or inert atmospheres. Also, the same process profiles are targeted for batch-type furnaces, albeit the time–temperature profile is governed by power changes to the heating (and/or) cooling coils rather that the multiple heat zone temperatures and conveyor speed of an in-line furnace. 2. Vapor Phase Heating The second process used in the assembly of printed wiring boards with Pb–Sn eutectic or neareutectic solder is referred to as vapor phase or condensation soldering. This method uses the heat of vaporization released when a vapor condenses into its liquid phase, to heat the assembly and reflow the Pb–Sn solder. Vapor phase reflow is a batch process [5]. The Pb–Sn solder is printed as a paste on the circuit board; the components are placed on the solder paste deposits; then, the assembly is placed into the equipment. A preheat step is performed on the assembly, typically in a separate chamber. Entry of the assembly into the vapor results in relatively fast heating rates for the solder joints, approximately 7j to 10jC/sec. Similarly, the joints experience fast cooling rates upon withdrawal from the fluid. The working fluid vapor is not a very efficient barrier against O2. 3. Wave Soldering The third process that was developed for Pb–Sn eutectic and near-eutectic solders is referred to as wave soldering. Wave soldering was developed as a large-volume process for through-hole printed wiring assemblies; it continues to be used for mixed assemblies that have both throughTABLE 6 Typical In-Line Furnace Process Parameters for 37Pb–63Sn Solder and an RMA Flux per the Diagram in Fig. 22 Process parameter DT/Dt T1 t1 T2 t2

Typical values 1–10jC/sec 150–170jC 60–180 sec 214–245jC 30–60 sec

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hole and surface mount components [5,51–53]. The printed wiring assembly is coated with flux and then passed through a preheat zone. Then, the assembly is exposed to the molten solder wave. The time–temperature cycle resulting from contact with the solder wave closely matches that of a hand soldering process. That is, there is a very rapid temperature rise followed by a short contact period (5 to 10 sec) between the assembly and molten solder. The subsequent cooling rate after a joint has left the wave is likewise, very fast. The wave soldering process has been precisely engineered for Pb–Sn eutectic solder. The type of flux, the take-off angle of the assembly as it exits the wave, the horizontal (belt) speed, and the shape of the wave, have all been specifically engineered to minimize ‘‘skips’’ and ‘‘bridges,’’ the common defects associated with wave soldering. 4. Hand Soldering The hand soldering of electronic assemblies, whether performed as the initial manufacturing step or as a subsequent repair or rework operation, is performed with an electric iron. The soldering step is typically completed in a matter of 2 to 7 sec. The relatively high temperature of the soldering iron (370j to 480jC), and the capability of the iron to concentrate heat into the immediate joint region, allows for the rapid heating of the solder and base metal so that solder flow, gap filling, and fillet formation are quickly completed without thermal damage to nearby materials.

IV.

INTERCONNECTION LEVELS

A.

Level 1 Interconnections

Level 1 interconnections refer to connections made between an integrated circuit (IC) die and the surrounding package. There are two variations to Level 1 soldered interconnections. In the first technique, high-Pb solder (97Pb–3Sn or 95Pb–5Sn) is typically deposited by evaporation or electroplating on the silicon chip bond pads in the wafer state. Upon singulation (dicing), the chips are placed and then reflow attached to a chip carrier. These solder joints are often referred to as controlled collapse chip connections, or ‘‘C4’’ solder joints. These high-Pb solder joints do not remelt during subsequent eutectic PbSn soldering processes. This direct, z-axis solder attachment of a chip to the next level of assembly is referred to as flip chip technology. This technique is used to attach the silicon chip to ceramic carriers because the ceramic carriers can withstand the high reflow temperatures required of the high-Pb solders (340j to 360jC). Alternatively, a low-melting-temperature solder (e.g., the eutectic Pb–Sn alloy) may be utilized to attach a silicon chip (having high-Pb, Pb–Sn solder bumps) to organic printed circuit boards. The circuit boards are not damaged by the low-temperature solder attachment process. This version of flip chip assembly is referred to as direct chip attach (DCA) or flip chip on board (FCOB) technology. A second flip chip technique is realized by the use of eutectic Pb–Sn or other low-meltingtemperature solder bumps in place of high-Pb solder bumps. These chips can be directly reflow attached to the chip carrier using low temperature. In this application, the chip carrier can be either ceramic or an organic laminate.

B.


The widely recognized use of Pb–Sn solders in electronics assembly is in the Level 2 interconnections. Level 2 interconnections are solder joints that attach device packages to the printed circuit board. The solderability and, ultimately, the reliability of Level 2 interconnections are determined by the materials used for the package input/output (I/O) and those used for the circuit board. 1. Package Input/Output Materials The materials used for the leads of component packages for through-hole assembly are typically Cu, or Fe–Ni alloy having a Cu or Ni solderable finish with a Sn or Pb–Sn protective layer.


195

Peripherally leaded, surface mount packages use lead materials of Cu or an Fe–Ni alloy with a Ni solderable finish and an electroplated Pb–Sn or Sn protective layer. Leadless chip devices have terminations comprised of a fired-on Ag thick film conductor that is overplated with a Ni or Cu solderable coating, followed with an electroplated Pb–Sn protective finish. Leadless ceramic chip components (LCCC) use castellated terminations with a thick-film Au finish. 2. Organic Laminate Substrate Materials Typical organic laminate substrate materials include epoxy–glass (FR-4), polyimide–quartz, and PTFE (Teflonk)–glass composites. The circuit board bond pad (base metal) is typically Cu (electrodeposited Cu over a Cu foil). The Ni (solderable finish)/Au (protective finish) system over Cu has been used in numerous applications to improve solderability, limit Cu dissolution under prolonged process cycles, or to prevent excessive Cu–Sn intermetallic compound layer growth during long-term service at elevated temperatures. The two Pb–Sn solders used exclusively for Level 2 interconnections are the eutectic 37Pb–63Sn and near-eutectic 40Pb–60Sn compositions. Their liquidus temperatures allow for process conditions that are compatible with organic circuit board laminates and the molding compounds used for packages. Flux and solder paste technologies have been so well designed around eutectic Pb–Sn solder processes that defect rates have been reduced to ppm levels, even for the low cost, high-volume assembly of consumer electronics. 3. Ceramic Substrate Materials A number of applications require the use of ceramic substrates. The ceramic materials include several grades of alumina–glass composites designated by the alumina content (94%, 96%, and 99%) as well as low-temperature, co-fired ceramic (LTCC) materials. Ceramic-based assemblies are referred to as hybrid microcircuits (HMCs). The conductor features on HMCs that form the base metal of Level 2 interconnections are thick film layers. The thick-film conductor structure between 37Pb–63Sn solder and a 76Au–21Pt–3Pd thick film layer is shown in Fig. 23. The thick

FIG. 23 (a) Hybrid microcircuit (HMC) assembly. (b) A scanning electron micrograph showing the cross section of a thick film solder joint from the HMC. The solder is 37Pb–63Sn and the thick film composition was 76Au–21Pt–3Pd. (Courtesy of Sandia National Laboratories.)

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film layer begins as a screen-printed paste composed of metal particles, glass particles, and some thixiotropic components for printability. The printed pattern is fired onto the ceramic substrate at temperatures of 800j to 900jC in air. The metal particles sinter together to form a 7- to 15-Amthick conductive layer that serves as the base metal of the solder joint. The glass component migrates to the metal/ceramic interface where it bonds the sintered metal layer to the ceramic chip carrier. Thick-film metal compositions are elemental metals Ag, Pt, Au as well as alloy compositions: Au–Pd, Au–Pt, Au–Pt–Pd, Ag–Pd, and Ag–Pt. (A limited number of HMC assemblies use Cuand Ni-based, thick-film systems.) Lead–tin solder joints made to these metal compositions present several challenges. First, the elemental films, Ag and Au, exhibit relatively fast dissolution rates in Pb–Sn solder. Excessive dissolution can risk complete consumption of the thick film and loss of the conductor pad. Second, rapid solid-state intermetallic compound layer growth occurs between the thick-film compositions and Pb–Sn solders, even at modest service temperatures [54– 57]. The consequences of solid-state aging are a thick, brittle intermetallic compound layer and loss of the thick film layer.

C.


The Pb–Sn solders, and primarily the eutectic 37Pb–63Sn and near-eutectic 40Pb–60Sn alloys, are used in the fabrication of Level 3 interconnections. Level 3 interconnections include cable– connector assemblies used to transmit signals between printed wiring assemblies as well as hard wired solder joints such as lead-to-turret or lead-to-eyelet interconnections. Examples of connectors are shown in Fig. 24. The Pb–Sn solder can be used in the actual construction of the connector, such as in the case of the coaxial connector shown in Fig. 25, or in the attachment

FIG. 24 Assortment of connector assemblies that represent the challenges of Level 3 interconnections. (Courtesy of Sandia National Laboratories.)


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FIG. 25 Cross-section view of a cable–coaxial connector assembly illustrating the use of 37Pb–63Sn solder joints in its construction. (Courtesy of Sandia National Laboratories.)

of a connector to the printed wiring assembly. Because the connector structures are Au plated for corrosion protection, assurances must be made to prevent the Au embrittlement of associated Pb–Sn solder joints [42].

V.

PHYSICAL AND MECHANICAL PROPERTIES OF LEAD–TIN SOLDER AND SOLDERED JOINTS

This section reviews the physical and mechanical properties of Pb–Sn solder and Pb–Sn solder joints.* As expected, the majority of Pb–Sn properties have been obtained for the 40Pb–60Sn and 37Pb–63Sn compositions due to their predominance in Level 2 interconnections. The low solidus temperatures of Pb–Sn solders in general, and the 40Pb–60Sn and 37Pb–63Sn compositions in particular, causes these alloys to readily undergo microstructural changes that, in turn, alter their mechanical properties. Tensile strength and hardness data from bulk 37Pb–63Sn solder, which illustrate a decrease in both properties as a consequence of room temperature aging, are shown in Fig. 26 [5,58]. Although the elastic modulus is not significantly impacted by microstructural changes, the time-dependent properties (e.g., creep and stress relaxation) are sensitive to these factors. An annealing treatment of 100jC, 16 to 24 hr will often ‘‘stabilize’’ the microstructure against further changes. Finally, it must be reiterated that the mechanical properties of Pb–Sn solders are extremely stain rate sensitive. Therefore, it is preferred that strength data be accompanied by testing rate parameters (strain rate or displacement rate and sample geometry).

A.

Physical Properties

The room temperature physical properties of (1) density, (2) electrical conductivity, (3) thermal conductivity, (4) coefficient of thermal expansion, and (5) elastic modulus are shown in Table 7 of selected Pb–Sn compositions [59]. The Poisson’s ratio of 40Pb–60Sn solder is 0.4 and is not highly temperature sensitive [29]. This value also provides a reasonable approximation for other Pb–Sn solders [60]. The surface tension of molten Pb–Sn solders at 300jC are provided in Table 8 [61].

* A more detailed compilation of these properties and other solder properties can be found in the third edition of the Soldering Handbook (American Welding Society, 2000).

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FIG. 26 The effect of room temperature aging time on (a) lap shear strength (0.2-mm gap, rate unknown) of 37Pb–63Sn/Cu joints subsequent to a soldering operation; and (b) Vickers hardness (100g load) for bulk 37Pb–63Sn solder after solidification. (From Ref. 5; excerpted from Ref. 58.)

The temperature dependence of the modulus of elasticity has been determined for 40Pb– 60Sn solder using acoustic techniques [60]. The behavior can be represented by the following equation: E ðGPaÞ ¼ C0 þ C1 T þ C2 T2 þ C3 T3

ð7Þ

where Co = 3.8305, C1 = 0.01105, C2 = 0.000112, C3 = 0.345 107, and T is temperature (jC).

TABLE 7

Physical Properties of Selected Pb–Sn Solders

Tin content (wt.%) 10 20 40 60 70 Source: Ref. 59.

Density (g/cm3)

Electrical conductivity (% of IACS)

Thermal conductivity (W/m-K)

Coefficient of thermal expansion (K1)

Elastic modulus (GPa)

10.5 10.0 9.28 8.52 8.17

8.2 8.7 10.1 11.5 12.5

35.8 37.4 43.6 49.8 —

27.9 26.5 24.7 21.6 20.7

19 20 23 30 35


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TABLE 8 Surface Tension Values at 300jC for Selected Pb–Sn Solders Solder (wt.%) 80Pb–20Sn 70Pb–30Sn 40Pb–60Sn

Surface tension (dyne/cm) 460 470 503

Source: Ref. 61.

B.

Monotonic Strength Properties

1. Time-Independent Deformation (Stress–Strain Response) The 0.2% offset, room-temperature yield strength values for selected Pb–Sn compositions are listed in Table 9 [5]. The strength of Pb–Sn alloys increase with Sn content. Both the yield strength and ultimate tensile strength increase with lower test temperature and decrease with higher test temperature [59,62]. The data provided in Ref. 62 demonstrate the strain rate sensitivity exhibited by the ultimate tensile strength of Pb–Sn solders, even at cryogenic temperatures. Other investigations have similarly illustrated the strain rate and temperature sensitivities of the time-independent deformation shown by the eutectic and near-eutectic Pb–Sn solders [63,64]. Fewer mechanical properties studies have been performed on the high-Pb, Pb–Sn solders (e.g., 90Pb–10Sn and 95Pb–5Sn). Constant strain rate (6.2 105 sec1) tensile properties were measured for 90Pb–10Sn and 97Pb–3Sn solders by Cole and Caulfield [65]. Selected 0.2% offset yield stress, the ultimate tensile stress, and the static Young’s modulus were excerpted from that study and listed in Table 10. The yield and ultimate tensile strengths were generally less than values corresponding to eutectic Pb–Sn solder when the tests were performed at 0j and 50jC. However, the high-Pb solder provided comparatively better performance at temperatures exceeding 100jC. The mechanical properties of bulk Pb–Sn solders are generally reproduced at the solder joint level. However, the actual strength values of Pb–Sn solder joints will differ from those of bulk solder. In particular, solder joint strength is affected by the gap size. An increase in joint strength with smaller gap sizes is most evident when joints are loaded in tension [66]; this trend is mush less pronounced when joints are loaded in shear [67]. The shear strengths recorded for Pb–Sn solder joints made to Cu base metal are listed in Table 11 [62]. The typical deformation mode that accompanies time-independent deformation in the eutectic and near-eutectic Pb–Sn solders is grain boundary sliding [68]. Fracture occurs in the solder, near to the intermetallic compound layer at the solder/base metal interface.

TABLE 9 Room Temperature 0.2% Offset Yield Strength of Selected Bulk Pb–Sn Solders Solder (wt.%) 90Pb–10Sn 40Pb–60Sn 37Pb–63Sn 20Pb–80Sn

Yield strength (MPa) 13.9 14.2 16.1 29.6

The testing rate was not specified. Source: Ref. 59.

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TABLE 10 Bulk Properties of 90Pb–10Sn and 97Pb–3Sn Solders at 0j, 50j, and 100jC Tested Under a Constant Tensile Strain Rate of 6.2 105 sec1 Solder composition (wt.%)

Test temperature (jC)

Elastic modulus (GPa)

0.2% Offset yield strength (MPa)

Ultimate tensile strength (MPa)

90Pb–10Sn

0 50 100 0 50 100

10 9.4 5.6 14 9.5 5.8

14.3 12.6 9.4 6.6 5.8 6.0

23.2 17.4 6.6 13.0 9.8 7.6

97Pb–3Sn

Source: Ref. 65.

2. Time-Dependent Deformation (Creep Response) The use of Pb–Sn solders at high temperatures (relative to their solidus temperature) causes timedependent deformation, or creep, to be a significant contributor to mechanical degradation in electronic interconnections. Creep studies have been performed on bulk eutectic and neareutectic Pb–Sn solder in tension, torsion, compression, and by indentation [69–73]. Minimum compression creep rate (de/dtmin) data can be represented by a ‘‘sinh law’’ expression over the temperature range of 55j to 125jC [72]: de=dtmin ¼ 2:48 104 ½sin hð0:0793rÞ3:04 expð56; 944=RTÞ

ð8Þ

1

where de/dtmin is the minimum creep rate (sec ), r is the applied stress (MPa), R is the universal gas constant (8.314 J/mol K), and T is the temperature (K). The microstructural features resulting from creep deformation in 37Pb–63Sn and 40Pb–60Sn solders are a combination of Sn– Sn grain boundary sliding, Pb–Sn interphase boundary sliding, and coarsening of the Pb-rich phase [74–76]. The creep behavior of Pb–Sn solder joints show similar trends as are observed in the bulk material [77,78]. A study examined the effect of cooling rate on the subsequent creep deformation of 40Pb–60Sn solder [79]. A faster cooling rate resulted in a finer starting microstructure. The resulting impact on creep behavior was most significant under low stress (low strain rates) where the steady-state strain rate increased with the finer microstructure.

C.

Fatigue and Microstructural Effects

The importance of understanding fatigue in soldered interconnects was brought to light by the premature failure of some surface mount solder joints made to organic laminate, printed wiring

TABLE 11 Effect of Pb–Sn Solder Composition on Double Lap Shear Strength at a Constant Strain Rate of Approximately 0.33/ sec Solder composition (Wt.%) 90Pb–10Sn 80Pb–20Sn 40Pb–60Sn Source: Ref. 62.

Shear strength (MPa) 20 23 27


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FIG. 27 Optical micrograph showing a fatigue failure of 37Pb–63Sn solder joining a surface mount, ceramic package to an organic laminate, printed circuit board. (Courtesy of Sandia National Laboratories.)

assemblies early in the technology [80–82]. The fatigue failure morphology of 37Pb–63Sn solder joints is illustrated in Fig. 27. The source of the fatigue failures was the strain generated in the solder during thermal cycling due to the coefficient of thermal expansion mismatch between ceramic packages and the organic laminate, printed circuit boards to which they were attached.* In fact, it became clear that time-dependent deformation had a critical role in the fatigue cycle; this was the so-called creep-fatigue phenomenon [83,84]. The term thermal mechanical fatigue (TMF) was introduced to describe this ‘‘temperature cycle plus thermal expansion mismatch’’ degradation mode in surface mount solder joints. This phase emphasizes that the source of the cyclic deformation as being temperature variations rather than applied cyclic loads (e.g., vibration), thereby distinguishing TMF from mechanical fatigue. The commonly observed, microstructural signature of TMF deformation is a localized coarsening of the Pb-rich phase [85]. This phenomenon is illustrated in Fig. 28 for a chip resistor component. The development of the coarsened Pb-rich phase particles is subsequently followed by microcracks that form along the Sn–Sn grain boundaries and Sn–Pb interphase boundaries. The microcracks join into a single discontinuity at which point, the load bearing capacity of the solder is lost. Studies have also examined the TMF behavior of high-Pb solder (97Pb–3Sn and 95Pb–5Sn) interconnections. These joints are typically made between a silicon chip and a ceramic chip carrier since the materials have similar coefficients of thermal expansion. As such, they do not develop a large degree of fatigue damage under temperature cycling conditions. In addition, the chips are of a sufficiently small size that shear strains developed in the joints are likewise of a lesser magnitude. The intrinsic fatigue resistance of these solders appears to be similar to that of eutectic solder [78].

* Thermal mechanical fatigue was not a prevalent in surface mount, hybrid microcircuit products because of the closer match between the coefficients of thermal expansions of ceramic packages and ceramic substrates. Also, TMF was not intensely investigated for through-hole, Pb–Sn solder joints. Although fatigue damage occurs in the latter, the larger bonding area reduces the impact of degradation below levels sufficient to cause electrical failure.

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FIG. 28 (a) Optical micrograph of a surface mount, 37Pb–63Sn solder joint having undergone TMF. (b) High-magnification view of the solder fillet showing the Pb-rich phase coarsening that accompanies TMF. (Courtesy of Sandia National Laboratories.)

D.

Fatigue Failure Studies and Models

1. Coffin–Manson Equation A number of low-cycle, isothermal fatigue studies have been performed on 37Pb–63Sn and 40Pb– 60Sn solders and their solder joints. The data were found to be well represented by a Coffin– Manson-type relationship: Dcp ðNf Þa ¼ h

ð9Þ

where Dcp, is the plastic strain range (shear or tension–compression), Nf is the number of cycles to failure, a is the fatigue exponent, and h is a constant. The Coffin–Manson equation was determined for 40Pb–60Sn solder tested in shear at temperatures in the range of 55j to 125jC, and cycle frequencies greater than or equal to 0.3 Hz [86]. The underlying failure criterion for Nf was a 50% load reduction. This resulting equation is: Dcp N f0:51 ¼ 1:14

ð10Þ

2. Frequency-Modified, Coffin–Manson Equation At frequencies below 0.3 Hz, the slower ramp rates allow a greater extent of damage to occur, resulting in the acceleration of degradation. Then, the frequency-modified, Coffin–Manson equation becomes applicable: Dcp ðNf =mÞ0:51 ¼ 1:14

ð11Þ

where m is the cycle frequency (cycles/sec). A more complex Coffin–Manson equation was developed by Engelmaier for predicting the fatigue life of 37Pb–63Sn solder under an applied shear deformation [87]. The cited work placed the frequency dependency of fatigue damage in the fatigue exponent, a. The resulting modified Coffin–Manson equation had been determined to be accurate for frequencies in the range of 0.00001 V m V 0.01 Hz. 3. Coffin–Manson Approach Based on Plane Strain Levels Isothermal fatigue predictions using the Coffin–Manson relationship have also been developed, but categorized according to the plastic strain levels to which the solder is exposed [88]. Guo et al. developed two equations to predict cycles to failure at 0.5 Hz, based upon a 50% load reduction criterion. The equations refer to two, tension–compression plastic strain ranges, Dep (%): ¼ 12:2; Dep N2:08 f

0:07 V Dep V 0:22%

ð12Þ

SOLDERS FROM THE LEAD–TIN SYSTEM Dep N1:35 ¼ 162; f

0:28 V Dep V 2:5%

203 ð13Þ

Plane strain values in the range of 0.22% to 0.28% would likely require the use of the equation for Dqp closest to the particular case at hand. 4. Strain Energy Approach Energy-based methodologies have also been used to predict fatigue damage in eutectic and near-eutectic Pb–Sn solder interconnections. This approach uses a generalized scalar, the strain energy, as a metric describing the extent of fatigue damage that has occurred in the solder. As a scalar quantity, fatigue degradation resulting from various conditions are simply be added together to define the total amount of degradation. The strain energy approach was utilized by Clech et al. in their development of the comprehensive surface mount reliability model (CSRM) [89]. The strain energy approach allows for the superposition of local thermal expansion mismatch strains, that is, those strains due to the thermal expansion mismatch between the package material, base material, and solder, on traditional global thermal expansion mismatch strains. These strains result from the thermal expansion mismatch between the microelectronic package and the printed circuit board laminate structure. The inclusion of the local expansion mismatch degradation improved the accuracy of TMF predictions compared to other methodologies. 5. Other Factors Each of the aforementioned fatigue prediction techniques are based upon the presumption that the solder in a joint undergoes uniform shear. This condition is most closely met in the gap region of solder joints for surface mount chip resistors or leadless chip carrier packages. However, the impact of the solder fillet and its particular geometry—concave or convex (‘‘bulbous’’)—is not addressed by these approaches because the stress and strain state is much more complex there. Unfortunately, the fillet accounts for the majority of the fatigue life of the solder joints. Predicting fatigue degradation in the fillets of solder joints, or other complex joint geometries, requires the use of computational models based upon constitutive equations and finite element techniques. The modeling approach is discussed in the following section. Generally, voids do not have a predictable impact on the fatigue life of Pb–Sn solder joints. First, the high ductility of Pb–Sn solders often mitigates any ‘‘stress concentration’’ effects caused by a void. Second, the solder deformation of a joint undergoing TMF occurs largely as a straincontrolled event. As such, the extent of plastic strain damage (or strain-energy damage) is the same per unit volume of material. However, voids do represent a loss of material through which a fatigue crack must propagate and thus, may detract from the fatigue life of a joint if suitably located in the crack path.

E.

Computational Modeling to Predict Solder Joint Reliability

1. Constitutive Models for Predicting Thermal Mechanical Fatigue Deformation It was recognized within the electronics community that given the variety of Level 2 solder joint geometries, package materials, and chip carrier materials used in surface mount interconnections, it was necessary to move beyond the Coffin–Manson-based fatigue prediction methodologies for Pb–Sn solder interconnects, and explore the use of computational models. The computational modeling approach requires that two tasks be addressed. The first task is to develop the appropriate constitutive equation that relates strain, e, or strain rate, de/dt, in the Pb–Sn material to the applied stress (r) and temperature (T) conditions: de=dt ¼ fðr; TÞ

ð14Þ

A number of material properties (e.g., elastic modulus, Poisson’s ratio, yield stress, activation energy of creep, etc.) are included as parameters within such an equation. The second task is to map the magnitude of fatigue deformation throughout the solder joint configuration. The latter requirement is accomplished by setting up finite element meshes of the various solder

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joint geometries. Finite element meshes for surface mount, through-hole, and connector (Level 3) interconnections are illustrated in Fig. 29. The critical task is the development of an appropriate constitutive equation for the Pb–Sn solder that is of interest. Current efforts in constitutive equation development have assumed the solder to be a continuum. Deformation is represented by a unified viscoplastic (or creep plasticity) constitutive equation [90,91]. The advantage of the unified viscoplastic approach is that both time-dependent deformation (creep) and time-independent deformation (plasticity from the stress–strain curve) are included in a single equation, thereby greatly facilitating the subsequent numerical computations. Recent advances in the development of TMF predictive models have included the incorporation of solder microstructure as a state variable in the constitutive equation [92,93]. A contour map (Fig. 30) shows the Pb-rich phase coarsening predicted in a leadless chip solder joint subjected to six thermal cycles having temperature limits of 50j and 80jC, ramp rates of 6jC/ sec, and dwell times of 10 min at both limits. The microstructure-based, viscoplastic constitutive equation for Pb–Sn solder improves the accuracy of solder fatigue strain predictions throughout the interconnection geometry. It provides a ‘‘real-time’’ adjustment of the local solder mechanical properties resulting from local changes in the solder microstructure (i.e., the Pb-rich phase). 2. Damage Models of Predicting TMF Crack Development The aforementioned constitutive equations predict the ‘‘continuum’’ deformation in the solder. However, the total functional lifetime of a solder interconnection also includes the initiation and propagation of fatigue cracks. Crack development is addressed by so-called damage models. The

FIG. 29 Examples of finite element meshes for Level 2 and Level 3 soldered interconnections: (a) gullwing leaded package, (b) chip device, (c) pin-in-hole, and (d) connector. (Courtesy of Sandia National Laboratories.)


205

FIG. 30 Contour map of Pb-rich phase coarsening in a 37Pb–63Sn chip capacitor solder joint predicted by a computational model. The Pb-rich phase size (coarsening) is the metric of fatigue deformation in the solder. The cycling conditions were: 6 cycles; temperature limits of 50jC and 80jC; ramp rates of 6jC/sec, and dwell times of 10 min. (Courtesy of Sandia National Laboratories.)

goal of reliability modeling is to couple continuum models and damage models so as to predict the total fatigue lifetime of an interconnection. The development of damage models is difficult. Cracks are discontinuities in the solder microstructure and, as such, present formidable challenges for identifying an appropriate constitutive equation as well as numerical code development. Nevertheless, significant progress has been made by investigators toward developing a computational representation of damage in Pb–Sn solders [94–96].

F.

Corrosion Resistance

Both Sn and Pb exhibit relatively good corrosion resistance. Tin is slightly more susceptible to general corrosion in seawater and acid media than is Pb [97,98]. As such, a higher Pb content usually improves the corrosion resistance of the solder. Fundamentally, galvanic-assisted corrosion should be of particular concern for soldered joints because the latter are comprised of dissimilar metals. However, by and large, Pb–Sn solder electronic interconnections do not exhibit a particularly strong susceptibility to galvanic corrosion. Flux residues do not cause severe corrosion, general or galvanic assisted, on Pb–Sn solder or their solder joints, respectively. The limited corrosion activity results primarily from the tenacious oxides of Pb and Sn, in particular, those of Pb. However, the presence of an applied voltage can potentially accelerate corrosion activity in Pb–Sn solder joints. The predominant phenomenon is electromigration, which is illustrated in

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Fig. 31 [99]. Electromigration occurs when a voltage difference between neighboring solder joints sets up a cathode/anode pair. A corrosion cell is formed when an electrolyte is present. For example, the combination of moisture and flux residues can comprise an electrolyte. Metal ions generated by an oxidation reaction at the anode travel to the cathode where they are reduced to their elemental metal. The elemental metal forms filaments that grow back toward the anode. Once the filament contacts the anode, a short circuit occurs between the two joints. The tenacious oxides of both Pb and Sn slow the electromigration phenomenon compared to other metals that do not form tenacious oxides, the most notorious being Ag.

VI.

SUMMARY

The Pb–Sn solders continue to play a critical role in the Level 1, Level 2, and Level 3 interconnections of commercial and military electronics. The widespread availability of Pb and Sn ores and the relative ease of winning the elements from these ores have contributed to their low primary metal costs. However, the real attributes of the Pb–Sn system lie in the physical and mechanical properties of these alloys:

Low liquidus temperatures allow for reduced processing temperatures that limit thermal damage to organic materials such as printed circuit board laminates and device molding compounds. The simple eutectic Pb–Sn system minimizes the sensitivity of the liquidus temperature to slight compositional variations and contamination effects. The Pb–Sn system provides

FIG. 31 Schematic diagram depicting the electromigration process. Metallic filaments grow from the cathode towards the anode. (From Ref. 5.)


207

both high-melting-temperature alloys (90Pb–10Sn and 95Pb–5Sn) for Level 1 interconnections, and low-temperature compositions (37Pb–63Sn and 40Pb–60Sn) for Levels 2 and 3 interconnects. Together, these alloys comprise a step-soldering process methodology or hierarchy used in a variety of electronic assemblies. The low surface tension of the Pb–Sn solders provides excellent solderability (wetting, spreading, and capillary flow behaviors) by the molten Pb–Sn solders in surface mount and through-hole interconnections. The Pb–Sn solders exhibit a limited susceptibility to all forms of corrosion thereby minimizing the need to implement corrosion mitigation measures such as coatings and desiccation. An in-depth understanding of the physical and mechanical metallurgy of Pb–Sn solders and associated interconnections has been established. In particular, the numerous studies on the TMF of Pb–Sn solders has allowed for the development of computational models that can predict the long-term reliability of their interconnects under various service environments. Lastly, an 80-year history of Pb–Sn solders in electronics has culminated into an unprecedented data base of information that supports both design and manufacture activities for today’s electronic products as well as the development of the next generation of advanced electronic packages and assemblies.

ACKNOWLEDGMENTS The author wishes to thank A. Kigo, R. Wright, and G. Bryant, whose optical metallography expertise provided the micrographs found in this chapter; G. Zender and B. Ritchy for the SEM work; and especially J. Rejent as the chief technologist in the soldering program at Sandia National Laboratories. The author also wishes to thank E. Lopez for his careful review of the manuscript. Sandia is a multiprogram laboratory operated by Sandia Corporation, a Lockheed Martin Company, for the United States Department of Energy under Contract DEAC04-94AL85000.

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7 Physical Basis for Mechanical Properties of Solders Morris E. Fine Northwestern University, Evanston, Illinois, U.S.A.

I. A.

INTRODUCTION High Homologous Temperature

Solders are high-temperature materials as their homologous temperatures at room temperature (TR/TM, where TR and TM are the room and melting temperatures, respectively, in Kelvin), are above 0.5. For example, the homologous temperature of Sn at room temperature is about 0.6. Tin–silver-, tin–zinc-, tin–indium-, and tin–bismuth-based solders have even lower homologous temperatures. Therefore, the mechanical properties at ambient and elevated temperatures are highly dependent on strain rate and dwell times. This impacts the design of solder interconnects in electronic devices for high reliability. The models for reliability prediction are based on the hightemperature behavior of solder. Thermally activated plastic flow (e.g., creep) dominates the stress–strain curve except at very low strains and very rapid strain rates. Elastic strains and plastic strains that are not thermally activated are very small compared to thermally activated plastic flow. Thermally activated plastic flow dominates both isothermal and thermomechanical fatigue (TMF).

B.

Coefficient of Thermal Expansion Mismatch

Components interconnected in electronic devices have different coefficients of thermal expansion and this leads to thermal stresses during thermal cycling of a device on turning it on and off. For example, the coefficients of expansion of Si and Al2O3 differ by about 6 ppm/jC. The difference between Al2O3 and the usual printed circuit board material is about 15 ppm/jC. In interconnection arrays such as flip chip and ball grid array joined components, the thermal stresses are almost pure shear. Tensile testing results of bulk solders can be converted to shear, as plastic deformation in crystalline materials (except for such processes as grain boundary sliding) occurs by motion of dislocations on shear planes. However, testing of actual joints in shear and in tension may give different results because fracture at interfaces between solder and an intermetallic compound or an IMC and the metallization it is in contact with will occur differently in shear and in tension.

211

212 C.

FINE Microstructural Stability

High homologous temperatures also result in microstructural instability. The grain and phase structures may coarsen with time even at room temperature. For example, the yield and flow stresses of cast Sn–Pb eutectic solder at constant strain rate decreases with time for many years [1]. These stresses rapidly decrease at first but the rate of decrease slows with time. The flow stress at 105 sec1 strain rate decreases from the 7-day value of 25–15 MPa after 7 years. Plastic deformation that occurs during fatigue from thermal cycling increases the rate of such aging. Further, as solute solubilities decrease with a decrease in temperature, a stable microstructure at ambient temperature contains a larger volume fraction of a second phase than at the extreme temperature of the thermal cycle in a device.

II.

THEORY OF THERMALLY ACTIVATED PLASTIC FLOW

This chapter discusses the theory of thermally activated plastic flow including creep theory, components of the stress strain curve vs. strain rate and temperature and, isothermal and thermomechanical fatigue. Finally, the physical basis for modeling reliability of interconnects under thermal cycling is discussed.

A.

Movement at the Atomic Level

Plastic flow by dislocation motion in crystalline materials occurs by thermally activated processes in the sense that an energy increase is needed to move a dislocation from one stable position to another as discussed below. There are a variety of energy barriers that need to be overcome. The energy of thermal vibrations of the atoms assists this process. The probability that a thermally activated process occurs is proportional to me raised to the (DG*/kT ) power, where m is the thermal vibration frequency (about 1013 times per second), DG* is the activation Gibbs free energy for the process, k is Boltzmann’s constant, and T is the absolute temperature. First, there is the Peierls stress; that is, the dislocation sits in a low-energy crystallographic location and energy is required to move it to the next location. This energy is vanishingly small in face-centered cubic crystals such as lead, but that is not true for body-centered tetragonal tin. Mechanical twinning transformation, which is a thermally activated process, occurs during plastic deformation of tin. This gives rise to a creaking sound, referred to as tin cry, which is heard when bulk tin is bent near one’s ear. Except in perfect crystals, there are many dislocations present and many of these go through the slip plane of an active dislocation, and must be cut by the active dislocation to achieve plastic flow. This is also a process of small activation energy and only impedes plastic flow at ordinary strain rates at very low temperatures. Thus a pure metal such as Cu has its highest flow stress at absolute zero temperature. The flow stress at constant strain rate rapidly decreases on cooling, reaching a plateau where the effect of changing temperature or changing strain rate (except at extremal values of strain rate) is small. This is the case for Cu at 0jC (homologous temperature of about 0.2). There are still obstacles to dislocation motion present that are longrange or with high activation energies. The long-range stress field from neighboring dislocations, for example, must be overcome by the applied stress before a dislocation can move. Any increase in dislocation density is the source of work hardening. The theory of the plateau yield stress is outside the scope of this chapter.

B.

Critical Normal and Shear Stresses

Fig. 1 from the seminal work of Kloske [2] shows the critical resolved shear stress (CRSS) of some Ag dilute Au alloys at temperatures down to 4.2K. The decrease in CRSS on heating is evident. The change in CRSS particularly for the 1 at.% Ag alloy at room temperature is quite small. The temperature dependence of CRSS at low temperatures increases with alloying. This is a general behavior in fcc alloys and is less in Au–Ag alloys than, e.g., Cu and Al alloys. There are interactions between solute atoms and dislocations giving rise to another energy barrier that must be overcome. The result is an increase in the CRSS as the temperature is lowered.

MECHANICAL PROPERTIES OF SOLDERS

213

FIG. 1 The critical resolved shear stress of silver-based gold alloys is plotted vs. temperature. Note that the slope at low temperatures increases with gold content. (From Ref. 2.)

Plastic flow in crystals occurs by dislocations moving on slip planes that slide over one another. Most experiments with single crystals to determine the CRSS are performed with tensile specimens. The shear stress on the active slip plane is computed by resolving the tensile stress onto the active slip plane. With polycrystalline tensile specimens, plastic flow still occurs on slip planes in the individual grains and these are grouped into wide shear bands that traverse the specimen at 45j from the tensile axis. The shear stress is usually calculated approximately for the whole specimen by dividing the tensile stress in half. Failure of a ductile tensile specimen usually occurs by initiation of a shear crack that is followed by a mode I crack that is normal to the tensile axis. Sudden final failure occurs when the true tensile stress exceeds the mode I fracture stress.

C.

Tin Embrittlement Due to Allotropic Modification

The flow stress of body-centered tetragonal Sn cannot be explored at cryogenic temperatures because it is not thermodynamically stable below 13jC where tin transforms to the brittle diamond cubic structure. This transformation is the cause of the well-known condition referred to as ‘‘tin pest.’’ The tin breaks into powder form because of the high internal stresses that result from the large volume change during transformation. The activation energy for nucleation of diamond cubic Sn is high, giving rise to significant undercooling. Tin pest usually occurs at much lower temperatures than 13jC and is suppressed by additions of small amounts of elements such as antimony. A detailed discussion of tin pest and other degradation phenomena involving tin are presented in Chapter 22.

214 D.

FINE Strain Rate Effects on Flow Stress

At high homologous temperatures, there is a decrease in yield stress from the plateau because the rate of diffusion of atoms from one place to another in the lattice becomes appreciable. This allows dislocations to climb over obstacles such as second-phase precipitates. Recovery then occurs, resulting in the reduction of dislocation density and work hardening. Recovery is a time-dependent process because it is determined by the rate of diffusion and the time allowed for diffusion, such as the period of a fatigue cycle. The flow stress at a given strain rate is a balance between recovery and work hardening. The flow stress, which increases with strain rate, is thus very much a function of strain rate. Stress–strain curves for eutectic Sn–Ag and Sn–Pb solders are shown in Fig. 2 (a) and (b), respectively [3] for strain rates varying from 0.001 to 0.0001. For these alloys, the true elastic limit or yield stress is very small. An apparent linear region after yielding is observed. The slope of this apparent linear region increases with strain rate and has caused confusion in values of modulus for these alloys reported in the literature. The low-stress value for eutectic Sn–Pb solder is 35 GPa and for eutectic Sn–Ag solder is 56 GPa. There are many mechanical properties related to thermally activated plastic flow at high homologous temperatures besides reduction of flow stress and the large dependence of flow stress on strain rate. Creep, i.e., plastic flow at a constant stress, and stress relaxation at a constant strain are those of most interest to solder applications. These are all discussed later in this chapter.

E.

Effect of Brittle Phases in Interconnections

Intermetallic compounds that form between solder and metallized substrate features such as terminal pads have much higher melting temperatures than the solder. Room temperature is not a high homologous temperature for them. Further, their crystal structures are much more complex than fcc, bcc, or bct. They have much higher Peierls stresses. The yield stress is so high that it is close or even higher than the fracture stress. Bismuth in Sn–Bi eutectic solder has a complex rhombohedral structure; the Peierls stress is so high that Bi is brittle at strain rates usually used in a stress–strain test or on impact loading. However, because Bi has a low melting temperature, at very slow stress rates such as during thermal cycling of an electronic component during application, thermally activated plastic flow is sufficiently large so that brittle fracture does not occur. Brittle materials at interfaces are susceptible to mode I fracture under tensile loading, but much less susceptible to brittle fracture under shear loading as a result of the coefficient of thermal expansion (CTE) mismatch between the components connected by an interconnection. Dispersed intermetallic compounds such as Ag3Sn in eutectic Sn–Ag solder and SbSn in tin– antimony solders increase the strength and creep resistance of the solder, but the alloys are quite ductile as subsequently discussed. The interface between these intermetallic compounds and the tin-based solid solution is not a source of brittle fracture.

III.

THEORY OF CREEP

A.

The Three Stages of Creep

The theory of creep will be briefly reviewed. The reader is referred to texts [4,5] on the subject for deeper coverage. When a static constant stress (single or combined) is imposed on a metal or alloy component or specimen and the temperature is sufficiently high (i.e., a homologous temperature greater than 0.5), the specimen or component undergoes deformation that increases with time. If the constant stress on the specimen is tensile, the specimen continues to elongate as time goes on. This phenomenon is called creep. A plot of specimen gage length with time shows three stages. An initial transient stage called primary creep, a steady state stage called secondary creep where the rate of elongation is constant with time, and finally tertiary creep where the rate continues to


215

FIG. 2 (a) Stress–strain curves for Sn–3.5Ag solder and (b) Sn–37Pb solder are plotted for three strain rates (0.001, 0.0001, and 0.00001) per second. The deviation from Hooke’s law occurs at lower and lower stress as the strain rate is reduced. The plastic flow stress also decreases with reduction in strain rate. (From Ref. 3.)

increase as the specimen elongates toward failure. The three stages of creep are schematically shown in Fig. 3. Most of the published research has been on steady state creep, which is the main interest in this chapter.

B.

Mechanisms

There are a number of creep mechanisms that may be active at the same time. If they are acting in parallel, the slowest essentially controls the creep rate; if they are acting in series, then the most

216

FINE

FIG. 3 The three stages of creep are schematically shown as they occur over time. The symbols en and tn are the corresponding rupture strains and time, respectively. Subscripts 1 and 2 represent the ends of stages I and II, respectively.

rapid is rate controlling. Creep mechanisms in crystals may be divided into two categories depending on whether dislocation motion is involved or not. 1. Diffusional Creep Mechanism At low stresses, creep may be due to diffusion of atoms alone. Under a tensile stress, the grains tend to elongate, setting up a stress gradient. In response, the atoms diffuse to relieve the stress gradient. In the Nabarro–Herring mechanism, bulk diffusion is rate controlling. In the Coble creep mechanism, grain boundary diffusion is rate controlling. In two-phase alloys, diffusion along interphase boundaries may contribute to creep. Diffusion along dislocations is much faster than bulk diffusion and such dislocation–core diffusion may also play a role. Of course, both grains and grain boundaries must deform or the material will begin to fracture. The accommodation may occur by grain boundary sliding, which is sometimes considered to be viscous flow. Because there is no amorphous phase at grain boundaries, the mechanism is closer to bulk plastic flow along slip bands involving defects such as grain boundary dislocations. In applications where the interconnections are under low stresses over a long accumulated time period, they must exhibit a very high degree of dimensional stability. Design against mass diffusional creep must be considered. Accelerated tests at higher stresses cannot be extrapolated into a low-stress regime. The diffusional creep strain rate equation for a pure metal usually used is de A1 rX ¼ Deff dt kTd2

ð1Þ

where Deff ¼ Dv þ

pdDb d

ð2Þ

In these equations, Dv and Db are the bulk and grain boundary diffusivities, respectively, in m2/ sec, X is the atomic volume in m3, r is the applied tensile stress in MPa, d is the average grain size in m2, A1 is a constant, and d is the grain boundary thickness in m. Obviously, a larger grain size results in a smaller diffusional creep rate. In solid solution alloys, Dv for the slowest diffusing


217

species controls the creep rate if its atomic fraction is sufficiently large, a few atomic pct. In twophase alloys, an interphase boundary diffusion term must be added. Of course diffusion is an activated process governed by an equation such as Eq. (3) D ¼ ameð kT Þ Q

ð3Þ

where a is a constant including the entropy term and a geometric factor with the same units as D, and Q is the appropriate activation energy in J/mol. The experimentally determined value of Q helps determine whether the diffusion controlling the creep is in the bulk, in an interface or grain boundary, or along the dislocation cores. The activation energy for bulk creep is the largest among these. 2. Dislocation-Related Creep At higher stresses, thermally activated dislocation motion such as dislocation climb out of its slip plane limits the rate of steady state creep. When a lattice vacancy exchanges position with an atom in a dislocation core, a jog is formed and in the process an element of the dislocation moves to a parallel slip plane one atom distance away. Dislocations can get around obstacles by repeated climb events. The direction of climb is controlled by the applied stress state, but, of course, the rate increases with stress as the material acts to reduce its free-energy state. The rate of climb under constant stress is controlled by the rate of diffusion. The theory of dislocation creep mechanisms begins with the well-known Johnston–Gilman equation e ¼ gqbv

ð4Þ

where g is a geometric factor relating tensile strain to shear strain, q is the mobile dislocation density in number per square meter, b is the Burgers vector in m, and v is the dislocation velocity in m sec1, which is a function of stress. The dislocation climb velocity is equal to the probability for a successful climb event times the frequency factor times the applied stress times the area swept out by each climb event (roughly b2), i.e., approximately Q 2 ð kT Þ v ¼ am r lb e

ð5Þ

where a is a constant and l is the shear modulus. The dislocation density is usually considered proportional to (r/l)2. The result is the creep equation 3 Q ´ ð kT Þ eY ¼ A r ð6Þ l e The constant A contains many terms including b, the frequency factor, the entropy term, etc. Experimentally, the exponent on stress is observed for pure metals and solid solution alloys to be in the range 3–8. There is not a general explanation for the variation in n to the author’s knowledge, and discussion of this subject is beyond the scope of the present chapter. The creep equation is generalized allowing the exponent on stress to be a variable giving the well-known Dorn equation n ´ ð kTQ Þ ð7Þ eY ¼ A r l e At stresses below the flow stress where the dislocation density is independent of stress, q is constant and n equal to one is predicted. This is referred to as the Harper–Dorn creep. a. Threshold Stress. In precipitation- and dispersion-strengthened materials, a much higher value of n is sometimes observed. A useful empirical equation is obtained by introducing a threshold stress into the equation subtracting from the applied stress. The threshold stress is a back stress or a stress that must be overcome before dislocation creep can occur. The resulting equation is n Q ´ r rt eð kT Þ ð8Þ eY ¼ A l

218 C.

FINE Tin Creep

Steady state creep data at constant temperature is usually presented as log–log plots of strain rate vs. stress. The slopes of these curves give the values of the exponent n. The activation energy Q is determined from the variation of steady state strain rate with temperature. Because all equations presented contain l, the shear modulus, its variation with temperature must be included in the analysis. The shear modulus and Young’s modulus decrease linearly with an increase in temperature in the vicinity of room temperature. The decrease of the shear modulus of tin is approximately 2000 ppm/jC. The variation with temperature decreases as the temperature is lowered toward 0K and must be zero there, a requirement of the third law of thermodynamics. Log–log steady state creep strain vs. stress curves for tin at three temperatures [6] are given in Fig. 4. There appears to be three stages in the creep behavior. At low stresses, the slope is smaller perhaps because of Harper–Dorn creep, then with an increase in stress there is a liner region where a slope of 8.6 fits the data. At still higher-stresses n decreases to 6. The line drawn for tin in the figure is an extrapolation of the 8.6 slope. Data points are seen to fall below this line. In the midregion, the activation energy is 73 kJ/mol but is 36 kJ/mol at higher temperatures. The activation energy for self-diffusion in tin is about 100 kJ/mol. This suggests that dislocation pipe diffusion may be controlling in the midregion. To the author’s knowledge, there is no good explanation for the high stress behavior. A suggestion is that it is because of the very rapid rate of recovery at high stresses.

D.

Sn–Ag and Sn–Zn Eutectic Alloys

Creep tests were conducted on Sn–Ag and Sn–Zn eutectic solders at 25jC and 80jC over the stress range from 10 to 22 MPa and the creep data of log steady state strain rate vs. log stress were fit to straight lines as shown in Fig. 5. The resulting activation energies were 82 and 68 kJ/mol, respectively [7]. These values are lower than the activation energy for creep of Sn. The exponents n were determined to be 11 and 6, respectively. The value 11 is high and may indicate thresholdstress-type behavior. The creep rates in the alloys are lower than that observed in pure Sn. This is mainly attributed to changes in the preexponential term, A, which is a function of microstructure. The theory for this is not well in hand.

FIG. 4 Steady state, stage II, creep of tin at three temperatures against stress. (From Ref. 6.)


219

FIG. 5 Steady state creep rate plotted against stress for Sn–Ag and Sn–Zn eutectic solders at 25jC and 80jC. (From Ref. 7.)

IV.

STRESS RELAXATION AT CONSTANT STRAIN

If a constant strain is imposed on a metal or alloy, the stress relaxes with time as the system reduces its free energy. Dislocations are annihilated and the remaining dislocations move to lower-energy configurations. This is the nature of the recovery process. At higher temperatures, diffusional processes equivalent to creep occur. This phenomenon is very important for solder joints in electronic devices because the device spends much time at the strain extremes when it is turned on and off or put into a sleep mode and returned to active duty. Stress at constant strain vs. time curves for Sn–3.5Ag solder at 25jC and 80jC, and at 0.3 % strain maximum are given in Fig. 6(a,b). The stress as a result of the coefficient of thermal expansion mismatches initially decreases very rapidly with time to a more or less steady state value. At 25jC, the steady state value is about 15 MPa and is rather independent of the initial stress value; however, at 80jC, the stress relaxes to zero. Thus when an electronic device is turned on, the thermal stress will relax to a low value possibly zero during use.

FIG. 6 Stress relaxation for a constant 0.3% strain plotted against time for Sn–Ag eutectic solder at (a) 25jC and (b) 80jC. (From Ref. 7.)

220 A.

FINE Phase Transformation Analog

Because stress relaxation at constant strain may be considered to be a phase transformation from a higher to a lower free energy state, the stress relaxation data may be treated by the standardphase transformation equation F¼

ðr rf Þ t m ¼ 1 eð s Þ ðri rf Þ

ð9Þ

Here F is the fraction transformed, r is the instantaneous stress, the subscript i refers to the initial stress, the subscript f refers to the final stress, t is time, s is the relaxation time, and m is a constant depending on the transformation mechanism. This is a good approximation when F is not too close to 1. Letting t equal s, F equals 0.63, thus s may be determined from curves of F vs. t. As determined by Mavoori, all relaxation times were quite small, less than 120 sec, s and m depend on the details of the stress relaxation process. Along with other factors, s is inversely related to the activation energy.

B.

Rate of Energy Reduction Model

Mavoori et al. [7] modeled the relaxation process on the basis of the rate of energy reduction, e(dr/dt)=(constant)(driving energy)(frequency factor)(no. of active sites)(energy reduction per event)(probability of a successful event). This leads to the following equation: e

dr ¼ CðerÞðmqÞðabrÞexpðQ=RTÞ dt

ð10Þ

In Eq. (10), m is the atomic vibration frequency, q is the dislocation density, a is the area swept out by a dislocation during an event, b is the Burgers vector, and Q is the activation energy. [A mistake in the reference has been corrected in Eq. (10).] Following the development of the Dorn creep equation, Eq. (10) can be generalized to dr ¼ Arn eðQ=RTÞ dt

ð11Þ

If a threshold stress is present, then the stress relaxation equation becomes dr ¼ Aðr rt Þn eðQ=RTÞ dt

ð12Þ

1. Parameters for Eutectic Sn–Ag Alloy Mavoori et al. [7] analyzed their stress relaxation at constant strain data for Sn–3.5Ag solder using Eq. (12) and determined the activation energy, Q, to be 34 kJ/mol and a stress exponent of 6. The threshold stress was found to be 10 MPa. How stress relaxation occurs in this solder is insufficiently known to attempt to interpret these numbers.

V.

STRESS–STRAIN RELATIONS

A.

Effect of Strain Rate

Because of the high homologous temperatures, the stress–strain relations for tin-based solders are much affected by strain rate except at very low stresses and strains. As already discussed, this presents difficulties in determining the modulus from a stress–strain curve. Fig. 3(a) from Ref. 3 gives a set of tensile stress–strain curves for Sn–3.5Ag solder at three different strain rates. The elastic line is also given. Deviation from this line is evident even at e = 0.0002. The deviation becomes less as the strain rate is increased.

MECHANICAL PROPERTIES OF SOLDERS B.

221

Three Strain Components

There are three components to the strain: eel, the elastic strain; epi, the time-independent plastic strain; and ept, the time-dependent plastic strain (creep strain). Plotting the inverse secant moduli vs. inverse strain rate separates these 1 etot eel epi ept 1 1 1 ¼ þ þ ¼ ¼ þ þ Esec Eel Epi Ept r r r r

ð13Þ

The time-dependent plastic strain decreases as the strain rate increases and it vanishes as the strain rate approaches infinity. The inverse secant modulus is plotted against square root of the inverse strain rate in Fig. 7(a,b) for three different stress values for Sn–Ag and Sn–Pb eutectic

FIG. 7 Inverse secant modulus vs. the square root of the inverse strain for (a) Sn–Ag eutectic solder and (b) Sn–Pb eutectic solder. (From Ref. 3.)

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FINE

FIG. 8 Stress strain curve for Sn–Ag eutectic separated into its three components for the three strain rates of Fig. 2. (From Ref. 3.)

solders [3]. At de/dt = 0, ept is zero and the extrapolated inverse secant modulus would equal the true elastic modulus if there were no epi. This is the case for bulk Sn–37Pb solder but there is a time-independent plastic strain component for bulk Sn–3.5Ag solder. The elastic strain is, of course, obtained from the elastic modulus. Then stress–strain curve is separated into its three components for Sn–3.5Ag solder in Fig. 8 [3].

C.

Effect of Strain Rate on Flow Stress

The Dorn equation can be inverted to give T de ðQ=RTÞ 1=n e r¼ A dt

ð14Þ

As shown in Fig. 9, after initial work hardening, the flow stress is independent of strain out to large values for bulk Sn–3.5Ag solder. The actual flow stress is a function of strain rate. Constant flow stress means that the work hardening rate equals the recovery rate. This occurs at increasingly higher stresses as the strain rate is increased. Plots of flow stress vs. strain rate from 105 to 102 at 25jC and 80jC were well represented by straight lines. From these data, n was determined to be 12 and Q, the activation energy, was determined to be 110 kJ/mol. Why there is variation in the n and Q values determined from creep, stress–relaxation, and stress–strain relations is not well known and needs much more study. However, empirically determined values are useful for design purposes.

VI.

ISOTHERMAL FATIGUE

Under cyclic loading or cyclic straining, metals fail after a few cycles (low-cycle fatigue, up to several thousand cycles), and after a large number of cycles (high-cycle fatigue, up to several million cycles). The cycles-to-failure increase as the stress amplitude or strain amplitude are reduced on continued lowering; a value is observed where failure does not occur no matter how large the cycle number—the endurance limit. All metals do not have an endurance limit but the first stress or strain amplitude on lowering amplitude that does not give failure after several


223

FIG. 9 Stress–strain curves at 25jC and 80jC for Sn–Ag eutectic solder are shown to approximately 4% strain. Strain rate varied from 102 to 105 sec1. (From Ref. 7.)

million cycles is taken as the endurance limit. Most isothermal fatigue experiments are carried out under constant amplitude cycles and the data is plotted as stress or strain amplitude vs. cycles-tofailure. The cycles can be complex with interruptions or dwell times at set places in the cycle such as at maximum stress or strain. Spectrum loading is important for many applications such as an airplane wing, and spectrum-loading fatigue has been extensively studied. Fatigue failure of a metal or alloy may be divided into three stages: fatigue crack nucleation, microcrack propagation and coalescence, and macrocrack propagation to failure. A macrocrack is a few millimeters or larger in length while microcracks can be submicron is size. Because of the small size of interconnections in electronic devices, only the first two stages are of interest. Of course, if a crack is already present, then the fatigue process consists of microcrack propagation and coalescence.

A.

Fatigue Crack Nucleation and Theory

1. Flaws Act as Stress Risers In the absence of a preexisting crack, the fatigue failure process begins with nucleation of one or more cracks. Stress risers or flaws are not cracks; the stress is larger at defects or stress risers. The latter may be part of a designed interconnection but still a fatigue crack needs to be nucleated. In a study of the effect of flaws and inclusions on the fatigue life of Al alloy castings [8], both pore and inclusion size and distance from the surface were found to be important in establishing a fail-safe criterion. A small pore or inclusion near the surface had more effect than a larger one further away. Fatigue failure of interconnections may sometimes be decreased by paying attention to stress risers in the design of the interconnection. For example, stress concentrators are inadvertently designed into ball grid array interconnects. The ball shape results in a stress concentrator near the top and bottom of the interconnect. A geometry that eliminates these stress risers should give longer lifetime against thermomechanical fatigue. 2. Fatigue Crack Nucleation Energy The energy barrier that prevents spontaneous formation of a crack on the first cycle, such as Griffith’s theory for the fracture of brittle solids, is the energy required to form new surfaces. Mura and coworkers [9,10] were the first to theoretically treat this process. In the Mura analysis, dislocations are assumed to accumulate on parallel slip bands during fatigue cycling. When the energy of the accumulation reaches a critical value, a fatigue crack is spontaneously assumed to nucleate. At high homologous temperatures, recovery reduces the rate of dislocation accumu-

224

FINE

lation, but thermal activation aids the nucleation process. This was treated using the concepts of phase transformation theory [11,12]. The energy release or reduction in internal energy when a fatigue crack forms results from the loss of lattice defects at the nucleation site W1 and from the release of stored elastic energy W2. Thus an equation for the free energy change on forming a crack of area A with c being the surface energy required to create a unit crack area can be stated DG ¼ W1 W2 þ cA

ð15Þ

a. Activation Free Energy. Because two surfaces are created when a crack forms, c equals 2cs, twice the surface energy per unit area. For a penny-shaped crack, A is pa2. Because of recovery balancing, steady state work hardening is achieved after a few fatigue cycles. At a sufficiently high homologous temperature, the dislocation density becomes constant and a defect energy density function, cd, assumed constant, can be defined. For a penny-shaped crack, W2 ¼

8ð1 m2 Þ 2 3 r a 3E

ð16Þ

In Eq. (16), m is Poisson’s ratio, E is Young’s modulus, and r is the maximum stress in a cycle. The change in free energy is then DG ¼ pa2 ðc cd Þ

8ð1 m2 Þ 2 3 r a 3E

ð17Þ

Differentiating DG with respect to a and setting the result equal to zero gives the critical crack size a* a* ¼

Epðc cd Þ 4ð1 t2 Þr2

ð18Þ

Substituting a* into Eq. (17) gives the activation free energy DG* DG* ¼

p2 E2 ðc cd Þ3 48r4 ð1 m2 Þ2

ð19Þ

If c equals cd, a crack spontaneously nucleates because DG* is zero. This is the nucleation condition in the Mura theory. From nucleation theory, the number of cracks, n, nucleated per unit volume per cycle, N, is given by dn DG* ¼ gtðkT=hÞexp ð20Þ dN kT where D is the number of potential nucleation sites per unit volume, t is a time in seconds equal to that portion of the cycle where the stress is near maximum, k is the Boltzmann constant, h is Planck’s constant, and kT/h is the frequency factor. Solving for DG* gives h dn ð21Þ DG* ¼ kT ln kTgt dN b. Other Considerations. The kinetics of fatigue crack nucleation can be studied without knowing the critical crack size as long as the cracks are not interacting and the crack propagation rate is fast compared to the nucleation rate. This concept was used to calculate the unknowns in these nucleation equations [11,12]. This is subsequently discussed. At lower homologous temperatures, cd is not constant but increases with fatigue cycling; there is an increase in dislocation density. Because cd is not measurable, it needs to be eliminated from the equations. This is also subsequently discussed.


225

Fatigue Crack Morphology

Fatigue cracks in bulk solder specimens and in interconnections are observed to form in shear bands [7,13–15]. Fig. 10 shows the surface of a fatigued Sn–Ag bulk solder specimen where crisscrossing shear bands are observed and consist of arrays of small cracks and slip steps. The small cracks lie along grain boundaries and interfaces between the intermetallic compound and the Sn solid solution. Lin et al. [10] observed that in near-eutectic Pb–Sn solder joints fatigue tested in shear, the plastic deformation was restricted to a shear band located in the solder close to the substrate. The microstructure in the shear band was much coarser than elsewhere in the solder joint, further illustrating the instability against plastic deformation of the initial microstructure. Coarsening of the microstructure in Sn–Ag and Sn–Zn eutectic solders during isothermal or thermal cycling is expected to be less because Ag3Sn or zinc phases are more effective in stabilizing grain size than a lead-based solid solution phase.

C.

Macrocrack Propagation

Macrocrack propagation itself is divided into three stages as described below. On fatigue loading at low values of crack intensity range, DK, such cracks remain constant in length; however, on increasing DK, a threshold is reached. Above this value of DK, the rate of crack propagation per cycle rapidly increases with increase in DK to a region where da/dN = A(DK )n, the Paris relation, holds. Here N is the cycle number, a is the crack length, n and A are constants depending on the material. Microcracks propagate intermittently, i.e., not during every cycle, at values of DK below the threshold. Thus it is not valid to assume a ‘‘large’’ crack if an interconnection is already present and predict fatigue lifetime of the interconnection using the Paris relation with A and n determined for long cracks in bulk specimens of an alloy. Even if a crack is already present, if it is below the macrocrack size, the Paris relation does not hold.

D.

Definition of Failure

Most relatively large scale fatigue tests define failure when the specimen becomes completely separated into two pieces. With electronic interconnections, failure is defined on the basis of unacceptable electrical resistance values or unreliable electrical connections. The sequence of events for fatigue failures in parts whose dimensions are in millimeters, as already discussed, are nucleation of microcracks, usually at flaws or stress risers, propagation, and link up of microcracks to form a macrocrack that grows following the Paris relation. Typically, interconnection sizes are so small that failure occurs in the microcrack regime. At some small size, but large compared to the size of the interconnection, the crack becomes unstable at the peak stress of a

FIG. 10 Isothermal fatigue damage at room temperature on the surface of bulk Sn–Ag eutectic solder showing criss-crossing shear bands. Strain range was 0.3% and the ramp time was 1 sec. (From Ref. 7.)

226

FINE

cycle and catastrophic failure occurs. This may cause only intermittent electrical failure because of the mechanical connection between the crack surfaces. A failure criterion is needed that is early in the fatigue process. 1. Reduction in Maximum Stress Details of the stress–strain hysteresis loops in mechanical fatigue tests are helpful. A schematic hysteresis loops for solders [7] is shown in Fig. 11 for constant tensile strain varying from zero to a maximum value. Important parameters describing the hysteresis loop are defined in the figure. Although the strain is controlled to vary from zero to a maximum tensile value, the stress goes into compression. The maximum stress rmax in a strain-controlled fatigue cycle varies during fatigue straining for several reasons. Work hardening occurs because of an increase in dislocation density from plastic strain cycling. However, recovery as a result of reduction in dislocation density also takes place. In addition, there is slow reduction in rmax as a result of microcracking and rapid reduction when a dominant crack begins rapid growth. Microstructural instability also affects the maximum stress. The sizes of the phases and the grain size in solders increase with time as the homologous temperature is high even at room temperature. A relatively stable microstructure is established in tin-based eutectic solders early in the fatigue process. A balance between work hardening and recovery to give a steady state at the maximum stress in a cycle is also established after a few cycles in these solders. The increase in rate of reduction of maximum stress when the final stage of cracking is reached is a useful criterion. Many other parameters describing hysteresis loops may also be used as failure criteria. Some of these will be shortly discussed. Initially, hysteresis loops are rather symmetrical but as fatigue failure is approached, a kink occurs in the compression part of the cycle as a result of crack closure effects. As subsequently discussed in lead-based tin solders such as 96.5Pb–3.5Sn, work-hardening continues to increase the maximum stress for many cycles although extensive cracking is observed. Eventually, the reduction in rmax due to cracking overcomes the contribution due to work hardening, causing the rmax vs. N curve to go through a maximum. This provides another failure criterion [16]. The variation in maximum stress of the fatigue cycle vs. cycle number for bulk Sn–Ag eutectic solder is shown in Fig. 12 [7]. After an initial transient, the negative slope of the rmax vs. N curve is rather constant for many cycles followed by a rapid decrease in rmax. The initial transient is attributed to establishing a nearly steady state dislocation structure at every stage in the cycle. Because microcracks were observed early in the fatigue test, the constant negative slope stage was attributed to microcracking. Establishment of a fatal crack, which then rapidly propagates, results in a rapid decrease in drmax/dN.

FIG. 11 Important fatigue parameters are defined on a schematic hysteresis loop. rmin is the maximum compressive stress, rmax is the maximum tensile stress, Det is the total strain range, and Deh is the hysteretic strain range.


227

FIG. 12 Failure criteria are defined as abrupt decreases in (a) rmax and (b) rmax/rmin. The abrupt changes are noted by an arrow in each case. Note that the former is a more conservative criterion at this total strain range. (From Ref. 7.)

2. Abrupt Change in rmax/rmin Ratio In strain-range controlled fatigue tests, the ratio of rmax/rmin is often used as a failure criterion and is also plotted vs. N in Fig. 12. This ratio is roughly constant but then abruptly decreases when a significant portion of an interconnection has cracked. The maximum compressive stress is larger because the crack closes in compression but opens in tension. In iron, this ratio has been observed to decrease when 15% of the final fracture surface had cracked [17]. For convenience, the 50% drop in maximum stress has been used as a failure criterion in Sn–Pb eutectic-type solders. Fig. 13 compares these two failure criteria to the 50% reduction in rmax. The data were linearized to fit an equation of the form Y = Ax + B. In the high-cycle regime, both are more conservative than the 50% reduction in rmax criterion. As predicted by the empirical equation, when the strain range achieves a 50% reduction in load at 50,000 cycles, the abrupt change in slope occurs at about 40,000 cycles, and the drop in the ratio occurs at about 44,000 cycles. 3. Higher Homologous Temperature Systems Bulk tin–lead solders with high lead content, such as 97Pb–3Sn cyclically harden for many cycles before the maximum stress in the cycle decreases, although cracking is observed early in the fatigue experiment [16] as shown in Fig. 14. The dislocation density increases with N; therefore cd in Eq. (17) is not constant. Thermal energy still assists crack nucleation except at extremely low homologous temperatures. The rate of recovery and the decrease in maximum stress from cracking are less than the rate of hardening. The maximum stress in a cycle when plotted against N goes through a maximum that indicates the beginning of rapid cracking leading to failure. This

228

FINE

FIG. 13 The rmax and rmax/rmin failure criteria are compared to the 50% reduction in rmax criterion. The equations Y = AX + B are plotted, where X is Nf for the 50 rmax criterion and Y is Nf for another criterion. For abrupt decrease in rmax, () A=0.72 and B=977. For the drop in rmax/rmin(o) criterion, A = 0.72 and B = 1915. The correlation coefficients are better than 0.98. The temperature was 25jC and the ramp time was 1 sec. (From Ref. 7.)

type of behavior is observed in many metals and alloys with low homologous temperatures. The maximum in the curve was taken as the failure criterion as is the practice with other metals and alloys [16]. It indicates the presence of a fatal crack or crack array. 4. Electrical Criteria It would be advantageous to establish an electrical criterion for final failure. Precise monitoring of the electrical resistance should provide this needed criterion because the resistance slowly increases as the cracking increases.

E.

Effects of Strain Range and Cycle Period

Fatigue life is the total time-to-failure including ramp times and dwell times in the cycle. It is equal to the number of cycles-to-failure times the period of the cycle. Thus the number of cycles-tofailure must be examined as a function of the cycle period including the ramp and any dwell times as well as the strain range. This was examined in great detail in bulk 96.5Pb–3.5Sn solder [16].

FIG. 14 Peak tensile stress is plotted against cycle number for 96.5Pb–3.5Sn solder. Failure was defined as the cycle number that exhibited the maximum peak stress as indicated by the arrow. The total strain range was 0.75% and temperature was 25jC. To stabilize the microstructure, the sample was heated to 150jC for 2 hrs and then aged at room temperature for 9 days. (From Ref. 16.)


229

FIG. 15 Cycles-to-failure vs. plastic strain range for temperatures varying from 5jC to 100jC. (From Ref. 16.)

1. High Lead–Low Tin Solder Increasing the strain range decreases the number of cycles to fatigue failure in materials. Such data for bulk 96.5Pb–3.5Sn solder is given in Fig. 15 for temperatures from 5jC to 100jC. The tensile plastic strain range is plotted vs. Nf on log–log scales. There was no hold in the cycle and the ramp time of the saw-toothed cycles varied from 0.1 to 2.5 sec. On the scale plotted, there is very little difference in the numbers of cycles-to-failure over this range of ramp time. The data were best fit by two straight lines with a break at about 0.3% plastic strain except for 5jC where there were few data points. The empirical Coffin–Manson relation, Nfhep = C (b and C are constants), does not hold over the full tensile strain range tested. Different constants are needed above and below the transition near 0.3% plastic strain. This was attributed to a change in fracture mode from intergranular at low strain ranges to mixed intergranular–transgranular at high strains. This points up the danger of using accelerated tests under extreme conditions that causes failure modes that do not occur under application conditions of the solder to predict reliability. There must be no difference in failure mode between an accelerated test and application conditions for the solder under consideration. As the temperature is increased from room temperature to 80jC, the cycles to failure at a given strain range decrease; however, there is little change on increasing the temperature to 100jC, an unexpected effect. Also unexpected, the cycles to failure decreased on cooling from room temperature to 5jC. 2. Lead-Free Solders The number of cycles to failure vs. total strain range are compared at room temperature for Sn– Pb, Sn–Ag, and Sn–Zn bulk eutectic solders in Fig. 16 over the strain range 0.2–2% [7]. Over most

FIG. 16 The isothermal cycles-to-failure for Sn–Ag, Sn–Zn, and Sn–Pb eutectic solders compared at 25jC vs. strain range. (From Ref. 7.)

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FINE

FIG. 17 Cycles-to-failure vs. total strain range in Sn–Ag eutectic solder at 25jC and 80jC. Ramp time was 1 sec. (From Ref. 7.)

of the strain range investigated, the lead-free eutectic solders have longer cycle-to-failure values than Sn–Pb eutectic solder. The bulk eutectic solders are much less sensitive to a change in temperature from 25jC to 80jC than bulk high-lead solder. Data for bulk Sn–Ag eutectic solder are shown in Fig. 17. 3. Effects of Frequency The effects of frequency or cycle period, 1/f, on cycles to failure and fatigue lifetime, tf, are examined in Fig. 18 for bulk 96.5Pb–3.5Sn solder [16]. The cycles-to-failure increases as the frequency increases but a plateau is reached at about 0.05 Hz. There is little increase in Nf as the frequency is further increased. The time-to-failure of course changes in the opposite direction and rapidly decreases at higher frequencies where a plateau in cycles-to-failure occurs. If Nf is constant as frequency is increased, then tf must decrease. The increase in Nf as frequency increases is expected because of the decrease in time for activated events to occur as fatigue failure is an activated process. When the period is long enough for the activated event to occur with high probability, then a further increase in time is less effective. In bulk eutectic solders, Nf changes little as f is increased from 0.002 to 0.5 Hz (Fig. 19) [7]. The plateau begins at lower frequencies for these solders than the high-lead solder. Of course, the time-to-failure decreases as the cycle time decreases. As shown in Fig. 20, the cycles-to-failure decreases in the bulk eutectic solders to a plateau value as the hold time at maximum strain increases [7]. There is a similar behavior in the bulk high-lead solder. The plateau, as in the previous paragraph, occurs because at some value of tc,

FIG. 18 Increasing the fatigue cycle frequency of 96.5Pb–3.5Sn solder at 25jC increases the life (cycles-to-failure) but decreases the time to failures. Strain range was 0.75%. (From Ref. 16.)


231

FIG. 19 Cycle number to failure is plotted vs. frequency for Sn–Ag and Sn–Zn eutectic solders. Data for 62Sn–36Pb–2Ag solder is also shown. The temperature was 25jC and total strain range was 1%. (From Ref. 7.)

the cycle time, the probability for occurrence of the activated event approaches 1. The cycles-tofailure moves in the opposite direction to the time-to-failure (Fig. 21). The time-to-failure is given as tf ¼ Nf ð2tr þ th Þ

ð22Þ

where (2tr + th) is the cycle time. Except for an initial transient, the relation between hold time and time-to-failure is linear [12] tf ¼ A þ Bth

ð23Þ

FIG. 20 Effect of hold time at maximum strain on cycles-to-failure is shown for eutectic Sn–Ag, Sn–Zn, and Sn–Pb solders. Temperature was 25jC and Det was 0.6%. Ramps time was 1 sec. (From Ref. 7.)

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FIG. 21 Ref. 7.)

F.

The same data in Fig. 20 is plotted as time-to-failure instead of cycles-to-failure. (From

Determining Cycles-to-Failure Utilizing Fatigue Crack Nucleation Theory

In eutectic Sn–Pb solder, fatigue cracks initiate at Sn–Pb interfaces and Sn–Sn grain boundaries and grow to the extremities of the individual facets with little interaction between neighboring cracks [15]. Based on micromechanics, Stolkarts [18] defined the microcrack density per unit volume, N = n/r3, where n is the number of cracked facets and r is the average radius of the microcracks. The change in microcrack density per unit volume per cycle, DNc, is related to the ratio Eeff/E, where E is Young’s modulus and Eeff is the effective Young’s modulus of a solder with an assembly of cracks. As the total strain amplitude in the cycles was constant in the tests conducted, Eeff/E is approximately equal to r/reff. Here r is the maximum stress during a cycle in the uncracked solder and reff is the maximum stress in the solder with an array of cracks for a given cycle number. There is also a relation between the strain amplitude and the increase in density per cycle. Cutiongco’s [15] data for Sn–Pb eutectic solder fit the relation DNc = 133De3 and dn/dN then equals r3(133De3). To estimate DG*, it is necessary to estimate D, the number of potential nucleation sites per unit volume, in Eq. (21). The number of shared interfaces or potential nucleation sites was estimated from the microstructure to be 1.4 1016 m3. In the model, the just-initiated cracks are assumed to rapidly grow to the extremities of the facets and the reduction in stress is due to facetsized cracks about 3 lm in radius. With a strain range of 0.006, the number of cracks initiated in one saw-toothed cycle from 0 to maximum strain is estimated to be 1016, which is about 10,000 times smaller than the number of sites. Site saturation is not expected for several thousand cycles. Taking t, the time during the cycle that the stress is near maximum, to be 0.1 sec or a tenth of the cycle time, DG* was computed to be 0.4 1020 J at room temperature. Substituting this value into Eq. (19) and with E = 35 GPa, r = 25 MPa, m = 0.37, ccd computes to 104 J/m2. Because the surface energy, c, is approximately 1 J/m2, cd, the energy of the dislocations, must essentially equal this value. If cd were zero, DG* computes to 2.5 109 J. This is so much larger than kT at room temperature that dn/dN would be zero [see Eq. (20)]. A substantial dislocation pile up at an interface is needed to nucleate a fatigue crack. As already mentioned in the Mura theory, which neglects thermal activation, the fatigue crack nucleates when the dislocations pile up at a singularity until c = cd. The present theory assumes that at high homologous temperatures, a


233

steady state dislocation distribution is established after some cycling. This dislocation distribution must be very high to give the dn/dN value estimated from the data. Using ccd = 104 J/m2, a*, the critical crack radius, computes to 6 nm from Eq. (19). Lin [10] observed fatigue cracks as small as 200 nm along slip bands in copper, and Kwon [19], using replicas, observed fatigue cracks that were only 100 nm deep along slip bands in copper. The justnucleated cracks in copper could have been smaller because of the limitations of the observation techniques. If ccd = 103 J/m2, a* is predicted to be 60 nm. The model thus predicts a ‘‘ball park’’ value for the critical crack radius from the observed decrease in rmax per cycle.

VII.

THERMOMECHANICAL FATIGUE

The thermomechanical fatigue (TMF) effect is the change in fatigue life by superimposing a temperature cycle onto a mechanical cycle. As already discussed, this arises in electronic components from turning them on or off or putting them into a ‘‘sleep cycle.’’ The components, being joined electrically and mechanically, have different coefficients of thermal expansion (CTE), which cause thermal strains and the associated thermal stresses. These, of course, are reduced if the mismatches in CTE of the components attached to each other are reduced. This should be a first consideration in designing a device where reduction in interconnection thickness is a goal. The temperature–strain cycles are complex because there are hold times at each temperature extreme in a cycle whether it is a full cycle between on and off, or a minicycle between the ‘‘sleep temperature’’ and the maximum temperature in a full cycle.

A.

Thermomechanical Behavior in Solders

1. High Lead–Low Tin Solder During a cycle, both temperature and strain are in phase because the temperature differential between the materials that are joined causes the strain. Their instantaneous temperatures determine the expansion or contraction of each material and thus the instantaneous strain. While the thermomechanical fatigue effect is a function of phase angle at least in high lead–tin solder [20], the concern here is mainly with zero phase angle. The thermomechanical fatigue of bulk high lead–tin solder has been studied much more thoroughly than any other solder [20]. The TMF effect is very complex. With in-phase cycling between the 25jC and 80jC and a cycle period of 200 sec, the TMF effect was substantially negative, but the effect diminished with cycle time. The TMF cycles-to-failure (saw-tooth waveform) increased with cycle time while the 80jC isothermal values slightly decreased. The two curves crossed at a 500-sec cycle time; for longer times, the TMF effect is positive, as shown in Fig. 22. Introducing a hold time at the maximum strain and temperature reduced the cycles-to-failure as with isothermal testing. Certainly more work is needed to fully explain all of this. Lawson [20] hypothesized that crack growth occurs by vacancies interacting with an oxide layer and he developed a model that was consistent with experimental observations of the effects of frequency, hold time, strain range, and phase angle changes. The complexity of the effects certainly means that design of devices and microelectronic components utilizing lead-free solders requires an in-depth investigation of the bulk and joint solder characteristics to understand the reliability of these materials. The many years of experience with Sn–Pb solders does not exist for lead-free solders. 2. Lead-Free Solders The thermomechanical cycles-to-failure for bulk specimens of Sn–Ag and Sn–Zn eutectic solders are compared in Fig. 23 over the total strain range from approximately 0.2% to 2% [7]. The thermal cycle range was between 20jC and 80jC without a hold time in the cycle and a ramp time of 50 sec. Two failure criteria were used, 50% reduction in load and abrupt drop in load. At least under these conditions, the Sn–Zn eutectic solder was more resistant to TMF failure over all of the strain ranges tested. The overall strain range is almost equal to the total strain range because the elastic strain range is very small. Further, as the plastic strain is inhomogeneous, the plastic strain range in the shear bands is very much higher.

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FIG. 22 Cycles-to-failure for thermomechanical (TMF) 25–80jC and isothermal fatigue at 80jC in 96.5Pb–3.5Sn solder vs. cycle period. The TMF effect is negative for periods less than 60 sec but positive at longer cycle periods. (From Ref. 20.)

The TMF effect for the 50-sec ramp time can be evaluated by comparing the lead-free solder curves in Fig. 23 with isothermal cycles-to-fatigue failure data at 20jC and 80jC (Fig. 17) [7]. For both Sn–Ag and Sn–Zn eutectic solders at 1% total mechanical strain range, the cycles-to-failure in bulk samples are about 4000 cycles at 25jC, and 3000 cycles at 80jC. These results are little affected by a change in ramp time from 0.5 to 500 sec. The TMF cycles-to-failure at 1% strain and 50-sec ramp time is about 200 cycles for bulk specimens of Sn–Ag eutectic solder, and about 300 cycles for Sn–Zn eutectic solder. The TMF effect is in the range of 100–200%. A similar effect occurs with Sn–Pb eutectic solders but appears to be somewhat smaller [15]. Predicting the cyclesto-failure by strain-range partitioning, i.e., integrating over a cycle using the isothermal data and time spent at each temperature, gives a value of cycles-to-failure for thermomechanical fatigue much larger than the true value at the 50-sec ramp time investigated.

FIG. 23 Cycles-to-thermomechanical-fatigue failure (20–80jC) plotted against total strain range for Sn–Ag and Sn–Zn eutectic solders. The ramp time was 50 sec. The mechanical and thermal cycles were in phase. The plot also allows a comparison to be made between the 50% reduction from smax and abrupt load drop failure criterion. (From Ref. 7.)


235

Superimposed Hysteresis Loops

The isothermal hysteresis loops for Sn–Ag eutectic solder at 25jC and 80jC and the thermomechanical fatigue cycle between 25jC and 80jC are superimposed on each other in Fig. 24, which represents the first complete cycle for 1% mechanical strain. The TMF heating-load cycle initially follows the 25jC curve but then drops smoothly to the 80jC curve; then, during the cooling–unloading cycle, it initially follows the 80jC curve but eventually returns to the 25jC cycle curve. Both Sn–Zn and Sn–Pb eutectic solders exhibit similar behavior [7,15].

C.

Thermomechanical Fatigue in Joints and Components

There have been several published studies of TMF of led-free solders utilized in joints and components. Choi et al. [21] characterized the cracking in Sn–Ag eutectic solder and also Sn– 4Ag–0.5Cu solder joints between alumina and copper that were cycled between 40jC and 125jC. The joints were first examined after 250 cycles. Grain boundary cracks adjacent to the Ag3Sn intermetallic layer and adjacent to the alumina substrate were the ones that led to failure. They propagated along grain boundaries in the solder and not at the interfaces. There was no evidence of grain coarsening as has been observed with Sn–Pb solders. Palmer et al. [22] recorded hysteresis loops during TMF of bulk specimens of Sn–Ag, Sn–Bi, and Sn–Pb. The strain during the temperature cycle was applied from the attachments to the specimens having different coefficients of thermal expansion. It was suggested that the accumulated energy absorption could give an insight into the failure process. All but a tiny fraction of the mechanical work is stored as damage in a specimen, the rest is heat loss. However, a moving crack driven by an applied strain and the associated stress does plastic work at the crack tip. The associated thermal energy is lost to the surroundings. The accumulated summation of the plastic work of crack propagation is a measure of the extent of cracking. The accumulated hysteretic energy obtained from the hysteresis loop areas is a measure of the heat loss, and it should be a measure of the extent of cracking. This is a promising approach.

D.

Solubility Effects

Thermomechanical fatigue of solders may be further complicated because the compositions and volume fractions of the phases present vary with temperature. In contrast to Sn–Pb solders, this is likely to only be a minor problem with Sn–Ag eutectic solder because the solubility of silver in solid tin is very small at all temperatures. This is also true for Sn–Cu-based solders. With Sn–Znbased solders, the effect of temperature on the microstructure is minimal because the maximum solubility of Zn in Sn is only 0.6 at.% at the eutectic temperature. Although the phase

FIG. 24 Thermomechanical fatigue at 25–80jC thermal cycle and isothermal fatigue hysteresis loops at 25jC and 80jC are compared. Mechanical strain range was 1%. There was no hold in the 50-sec ramp time cycle. (From Ref. 7.)

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composition and volume fractions are constant, continued temperature cycling of high homologous temperature solders with added strain cycling significantly coarsens their microstructures. As with isothermal fatigue, the plastic deformation in TMF is inhomogeneous. In bulk specimens of Sn–Ag and Sn–Zn eutectic solder, criss-crossing shear bands are seen and fatigue cracks form in these shear bands [7]. The particle sizes of the minor phases coarsen, and there is some evidence for cavitation particularly in Sn–Zn eutectic solder during 25–80jC and return. As with joints made with Sn–Pb solders [13,14], Sn–Ag and other Sn-based, Pb-free solder joints are expected to develop thin shear bands parallel to the shear direction. However, because the solubility of zinc or silver in tin varies little with temperature and the microstructure is expected to be more resistant to coarsening than Sn–Pb eutectic solder, grain coarsening and thus shear band formation is expected to be less of a problem.

VIII.

FINAL COMMENTS

While much is known about the mechanical properties of the Sn–Ag and Sn–Zn eutectic solders, the knowledge base on fatigue both isothermal and thermomechanical has many gaps. An interdisciplinary research program including theory, science-based modeling, microstructure, and properties is needed for whatever lead-free solders are selected for application. The many years of experience with Sn–Pb solders allows confidence in utilizing them in design but this is lacking for lead-free solders. The research must address actual solder joints themselves including size effect as well as studies on bulk specimens to obtain fundamental properties. Thermomechanical fatigue studies of solder joints where the strain ranges, temperature range, dwell times, ramp times, specimen size, and phase angle are varied and controlled are time-consuming, difficult, and costly, but necessary. However, such information correlated with microstructural studies is needed for a science-based methodology for reliability prediction. Otherwise, each selected solder and processing procedure will require a separate and unique evaluation.

REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13. 14.

Vaynman, S.; Ghosh, G.; Fine, M.E. Effects of palladium and solder aging on mechanical and fatigue properties of tin–lead eutectic solder. J. Electron. Mater. 1998, 27, 1223–1228. Kloske, R.A. Solid solution strengthening and yield drop effects in Au–Ag alloy single crystals containing 1 to 5 and 95 to 99 at. pct. Ag. Trans. Metall. Soc AIME 1969, 245, 217–225. McCabe, R.J.; Fine, M.E. Athermal and thermally activated plastic flow in low melting temperature solders at small stresses. Scr. Mater. 1998, 39, 189–195. Garofalo, F. Fundamentals of Creep and Creep—Rupture in Metals; Macmillan Series in Materials Science: Macmillan, New York, 1965. Weertman, J. Microstructural mechanisms in creep. In Mechanics and Materials; Meyers, M.A., Armstrong, R.W., Kirchner, H., Eds.; Wiley: New York, 1999 (chapter 13). McCabe, R.J. Creep of InSb and SbSn Tin-based Alloys: Effects of Precipitate Size, Shapes and Spacing, Ph.D. Thesis, Northwestern University, Dec. 2000. Mavoori, H.; Chin, H.; Vaynman, J.; Noran, B.; Keer, L.M.; Fine, M.E. J. Electron. Mater. 1997, 26,783–790. Seniw, M.E.; Conley, J.G.; Fine, M.E. The effect of microscopic inclusion locations and silicon segregation on fatigue lifetimes of aluminum alloy A356 castings. Mater. Sci. Eng. 2000, A285, 43–48. Mura, T.; Nakasone, Y. A theory of fatigue crack initiation in solids. J. Appl. Mech. 1990, 57, 1–6. Lin, M.R.; Fine, M.E.; Mura, T. Fatigue Crack Initiation on Slip Bands: Theory and Experiment, Ph.D. Thesis, Northwestern University, 1984. Acta Metall. 1986, 34, 619–628. Fine, M.E. Phase transformation theory applied to elevated temperature fatigue. Scr. Mater. 2000, 42, 1007–1012. Fine, M.E.; Stolkarts, V.; Keer, L.M. Fatigue crack nucleation assisted by thermal activation. Mater. Sci. Eng. 1999, A272, 5–9. Morris, J.W., Jr.; Mei, Z. Solder Mechanics, A State of the Art Assessment, EMPMD Monograph Series; Frear, D.R., Jones, W.B., Kinsman, K.R., Eds.; TMS: Warrendale, PA, 1990; Vol. 1, 239–270 (Chapter 6: Toward New Solders with Improved Fatigue Resistance). Frear, D.; Grivas, D.; Morris, J.W. A microstructural study of the thermal fatigue failures of 60Sn– 40Pb solder joints. J. Electron. Mater. 1988, 17, 171–180.

MECHANICAL PROPERTIES OF SOLDERS 15.

16. 17. 18. 19. 20. 21. 22.

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Cutiongco, E.C. Fatigue of Near-Eutectic Lead–Tin Alloy for Surface Mount Technology, Ph.D. Thesis, Northwestern University, 1991; Cutiongco, E.C.; Jeannotte, D.A. Proc. Morris E. Fine Symposium; Liaw, P.K., Weertman, J.R., Marcus, H.L., Santner, J.S., Eds.; TMS: Detroit, Oct. 1990; 435–444. Vaynman, S. Isothermal Fatigue of 96.5Pb–3.5 Sn Solder, Ph.D. Thesis, Northwestern University, 1988; Vaynman, S.; Fine, M.E.; Jeannotte, D.A. Isothermal fatigue of low tin lead based solder. Metall. Trans., A 1988, 19A, 1051–1059. Cooper, C.V.; Fine, M.E. Fatigue microcrack initiation in polycrystalline alpha–iron with polished and oxidized surfaces. Metall. Trans. A 1985, 16, 641–649. Stolkarts, V.; Keer, L.M.; Fine, M.E. Damage Evolution Governed by Microcrack Nucleation with Application to the Fatigue of 63Sn–37Pb, PhD Thesis, Northwestern University, J. Mech. Phys. Solids 1999, 47, 2451–2468. Kwon, I.B; Fine, M.E.; Weertman, J. Microstructural Studies on the Initiation and Growth of Small Fatigue Cracks at 298, 77 and 4.2 K in Polycrystalline Copper, Phd Thesis, Northwestern University, Acta Metall. 1989, 37, 2927–2936. Lawson, L.R.; Fine, M.E.; Jeanotte, D. Thermomechanical Fatigue of a Lead Alloy, PhD Thesis, Northwestern University, Jour. of Materials Research 1993, 8, 745–757; Metall. Trans. A 1991, 22A, 1059–1070. Choi, S.; Subramanian, K.N.; Lucas, J.P.; Bieler, T.R. Thermomechanical fatigue of Sn–Ag solder joints. J. Electron. Mater. 2000, 29, 1249–1257. Palmer, M.A.; Redmond, P.E.; Messler, R.W. Thermomechanical fatigue testing and analysis of solder alloys. J. Electron. Packag. 2000, 122, 48–54.

8 Sn–Ag and Sn–Ag–X Solders and Properties Karl J. Puttlitz IBM Corporation, East Fishkill, New York, U.S.A.

I. A.

INTRODUCTION Use and Compatibility

Lead–tin (Pb–Sn) solders have been utilized in electronic packages and assemblies for more than 40 years to attach components to printed circuit boards (PCBs) owing to their unique combination of properties of low melting point, wetting, and mechanical characteristics, and are relatively inexpensive. In addition, these materials are compatible with component lead and finish materials, and allow processing in a temperature range in which card materials and components maintain their integrity even when exposed to multiple thermal excursions as happens during rework operations. Lead–tin solders have demonstrated a high level of compatibility with electronics assembly processing and materials over the years with only a few exceptions, most notable is the effect of even relatively minor gold concentrations to cause embrittlement within solder joints [1].

B.

Migration to Lead (Pb)-Free Solders

There are two main drivers motivating the migration to lead-free solders: legislative-based issues and health concerns related to lead (Pb)-containing materials including solder. Also, there is pressure to support the growing worldwide movement to remove lead from processes and products to demonstrate environmental consciousness and maintain a competitive posture in the market place. However, assuring the reliability of solder joints made with lead-free materials is difficult due to the lack of field experience and lack of fundamental property data upon which to construct reliable service life models [2]. 1. Tin–Silver-Based Systems of Key Interest The lead-free solder alloys attracting the greatest attention are the tin-rich alloys based on the tin– silver (Sn–Ag) and tin–copper (Sn–Cu) systems and their ternary alloys, Sn–Ag–Cu [3–5]. Numerous studies have been reported on Sn–Ag–Cu near-ternary alloys investigating processing and mechanical property behavior, and also their solidification characteristics [6–9].

C.

Challenges for Lead-Free Solder Alloys

Each of the roles and functions of solder joints in electronic assemblies becomes increasingly more difficult to achieve owing to the lighter, faster, smaller, cheaper trend, particularly portable,

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and hand-held electronics products. This is perhaps best exemplified by Moore’s Law that predicts an exponential growth in the number of transistors on a single chip, doubling the number every 18 months. This, in fact, has been the case for over a decade. The impacts, thus challenges, of accommodating the large increase in transistors by traditional methods of increasing the die size and decreasing the interconnection (i.e., solder joint) pitch [10] are discussed in the following sections. 1. Increased Die Size Although a reduction in feature sizes referred to as shrink is one of the ways to accommodate an increased transistor count, it is often not sufficient to provide the required I/O counts, necessitating an increase in die size as well. However, as the strain that a solder joint experiences is directly proportional to the distance from the neutral point (DNP, i.e., the temperature invariant point), this can have a significant impact on the reliability of solder joints located in the outer rows of area-array, flip chip devices. An increased die size may result in a maximum DNP that poses a potential reliability concern for some applications, requiring that additional precautionary steps be taken, among them enhanced cooling capability, chip underfill, and increased inherent fatigue resistance of the solder. 2. Decreased Solder Joint Pitch Another measure to accommodate an increased transistor count is to decrease the interconnection (i.e., solder joint) pitch [10]. A decreased pitch is typically accompanied by a reduction in solder joint diameter. The presence and influence of flaws, such as voids, and particulate debris whose size remains relatively constant becomes more important with a reduced solder joint cross section. Stresses that a solder joint experiences due to a mismatch in the coefficient of thermal expansion (?CTE) between the materials a solder joint is attached to also increase with a pitch reduction. This arises from the fact that the stress is inversely proportional to the height of the solder joint that typically decreases with a pitch reduction as well. 3. Increased Joint Current Density and Temperature Even if the solder joint size did not decrease, chip power requirements would increase simply because of the increased number of transistors on a chip, causing an increased thermal dissipation. Even with cooling enhancements, chip temperatures are anticipated to raise over time. Increasing temperature serves to accelerate the mechanisms that give rise to failure in solder joints. a. Creep-Induced Thermomechanical Fatigue. Thermal fatigue is considered the primary threat to the reliability of solder joints because they typically experience many thermal excursions during the course of their applications due mainly to equipment on/off cycles. Stress is generated during the course of each excursion or thermal cycle due to both CTE mismatches as noted earlier, and inhomogeneities in the pattern of heating that cause thermal gradients. The dominant stresses tend to be in the shear mode. Because solders are utilized at high homologous temperatures (ratio of the ambient or operating temperature to the melting point in absolute temperature, i.e., Kelvin), they tend to undergo significant creep deformation, especially at the hightemperature part of the cycle, resulting in failure by creep-induced fatigue [11,12]. b. Thermo and Electromigration. Attendant with the exponential increase in the number of transistors over time as predicted by Moore’s Law is the near-exponential increase in power as well [10]. Therefore future packages will operate at much higher currents and temperatures, so lead-free solder joints will be required to exhibit good thermo and electromigration resistance properties, and have thermally stable microstructures [13]. Reductions in solder joint cross sections will further aggravate the situation, and particularly so of small area-array components such as flip chips and fine-pitch chip scale packages (CSPs). Severe current crowding can be anticipated in the via region of flip chips with high current flow (i.e., high current flux densities), making them vulnerable to both thermo and electromigration whose mechanism and details are discussed in Chapter 20.

SN–AG AND SN–AG–X SOLDERS AND PROPERTIES D.

241

Response to Stresses

The empirically generated field life prediction relationships that have been utilized for Pb–Sn solders are not automatically applicable to Pb-free solders. Solder joints as noted experience stresses due to thermal excursions, excessive electrical current, and also due to various mechanically applied loads. The tolerance or resistance and responses to stress as a result of these sources are typically different for Pb-free solders compared to the benchmark eutectic Sn– Pb solder. As noted, stresses in a solder joint can result due to a coefficient of thermal expansion (CTE) mismatch between package materials, but can also be due to temperature fluctuations caused by power transitions or environmental changes given the complexity of some modern microelectronic package structures and the environments they are placed in respectively. The ability of solder joints to endure mechanically induced stresses will also be an important factor for Pb-free solder joints due to the increased number and popularity of hand-held electronics continuously introduced into the market place. The solders must be resistant to mechanical influences such as shock impact, bending, and twisting [14].

E.

This Chapter

Taking the foregoing concerns into account this chapter discusses in detail the key metallurgical aspects including as-solidified microstructure, thermal stability characteristics, effects of elemental additions, and thermal aging, etc. as they relate to the Sn–Ag, Sn–Cu and their ternary alloys (Sn–Ag–Cu) into account. These alloy systems are the subject of much attention as solder replacement candidates for solders of the Pb–Sn system, particularly eutectic Sn–Pb. In addition, the chapter discusses the major mechanical properties of interest for electronic packaging applications as a consequence of their metallurgical characteristics.

II.

TIN–SILVER (SN–AG) SYSTEM

Several binary eutectic or near-eutectic alloys have been suggested as replacements for eutectic Sn–Pb, among them are Sn–Bi, Sn–In, and Sn–Ag, which have all been utilized in electronics applications [15–17]. However, the Sn–Ag system has a limited solubility of Ag in Sn, making it more resistant to microstructural coarsening, and the Ag3Sn precipitate is uniformly distributed, resulting in an overall structure that has the attributes of being more reliable [18,19]. Eutectic tin–silver (3.5% Ag) is also attractive because it is free of toxic Pb, offers better mechanical properties than eutectic Sn–Pb [20], is better in high-temperature applications due to its higher melting temperature, and exhibits good wetting behavior on copper and copper alloys [21]. Several higher temperature solutions have emerged in the Sn–Ag system which typically include the elemental additions of Cu, Bi, or Sb [22].

A.

Microstructural Aspects

1. As-Solidified Microstructure The as-solidified microstructure of eutectic tin–silver is relatively coarse and nonuniform. It consists of two phases, large globules of essentially pure tin (less than 0.04% Ag in solution) dendrites ranging in size from about 20 to 200 mm. The second phase is an intermetallic compound (IMC), Ag3Sn, referred to as epsilon (q) phase comprising less than a 5% volume fraction. The eutectic structure consists of the rod-like (q) phase in a nearly pure tin matrix. The lamellar morphology apparently is not routinely observed for eutectic Sn–Ag compared to eutectic Sn–Pb [23]. It has, however, been observed that the region between the Sn dendrites consists of a lamellar two-phase eutectic structure of a Sn matrix and Ag3Sn IMC phase with an elongated morphology [16]. The presence of the Sn dendrites is indicative of a nonequilibrium cooling rate. However, the structure is noted to be stable as the Sn dendrites were observed to be present after months of room temperature aging [16].

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In nonequilibrium cooling, characteristic of solder joints, the rod-like (q) phase which typically forms, tends to minimize the surface free energy. This occurs in phase transformations when the relative volume fraction of one phase is less than 1/p (i.e., approximately 0.3%) [24]. The size and morphology of q-phase depends on the cooling rate during solder solidification.

B.

Effect of Copper Dissolution

In solder joints, molten Sn–Ag solder is typically in contact with Cu termination pads and component leads. The copper dissolves into the solder and forms Cu–Sn IMCs both in the bulk ´ solder and at the Cu interfaces. Two compounds can form Cu6Sn5 (D) or an ordered version D at room temperature, and Cu3Sn (q). However, the addition of a small amount of Cu to a Sn–Ag-based solder alloy can help prevent the dissolution of Cu base metal [25]. At the same time it is possible to seed a fine, uniform dispersion of small precipitates of ternary or quaternary alloying elements to produce fine and homogenous microstructures [26]. 1. Mobility of Copper It has been reported that the diffusion coefficient of Cu at 25 jC along the a and c axes of Sn is about 0.5108 and 2106cm/sec, respectively [27]. These high diffusion rates are an indication that the nucleation and growth of Cu–Sn IMCs can easily occur at elevated temperatures.

C.

Microstructural Stability

The formation, growth, and morphology of the Cu–Sn IMCs formed greatly depend on the soldering process or any process with varying thermal conditions [28]. But it has been demonstrated that the various microstructures that can be generated through several soldering processes (i.e., heating and cooling rates) will not survive an exposure to elevated temperatures, although it has been suggested microstructural integrity can be maintained throughout the service life [29]. As noted in the previous section, the Sn–Ag microstructure remains relatively stable when exposed to relatively low operating temperatures, or for short periods at higher temperatures. However, the as-soldered microstructure becomes significantly altered if exposed to high temperatures, or for long times at moderate temperatures [23].

D.

Intermetallic Compound Layers and Morphology

The formation of Cu6Sn5 at copper interfaces such as lead frames and chip carrier pads occurs more rapidly in the case of eutectic Sn–Ag compared to eutectic Sn–Pb. However, the addition of 0.5% to 1.7% Cu to eutectic Sn–Ag reduces the IMC formation at interfaces and also lowers the melting point of the alloy [30].

E.

Effect of Heating Rate

Heating rate also plays an important role. A study was conducted comparing standard infrared furnace heating to a laser heating process. A peak reflow temperature of 235jC with a 200-sec hold period was conducted with an infrared furnace. The soak period was generally representative of manufacturing practice. It was observed that this heating resulted in (q) phase (Ag3Sn) rods oriented perpendicular to the joint interface, indicative of the solidification direction of the solder. A much faster heating rate was achieved utilizing a Nd:YAG continuous wave laser, resulting in very fine Ag3Sn IMC formed, only visible at a very high magnification (about 4000). The cooling rate was also very rapid, 104jC/sec compared to about 10jC/sec under IR reflow conditions [31]. Rapid cooling promotes nucleation but severely limits growth. The laser heating method superheated the solder to high local temperatures, typically 400–600jC compared to the IR reflow (240–260jC). Superheating increases the copper dissolution rate from termination pads, resulting in an increased copper concentration in the solder. The dissolution of

SN–AG AND SN–AG–X SOLDERS AND PROPERTIES

243

copper is more dependent on peak temperature than time at temperature, even for short time periods. The significant copper dissolution in the solder results in the precipitation of greater quantities of Sn–Cu IMCs. Those Sn–Cu IMCs formed at the solder/Cu interfaces were observed to be much thinner in the case of solder joints that were created by laser heating compared to standard IR-heated joints. This resulted from the fact that the reaction time (i.e., time above the melting temperature) was so short, coupled with the rapid cooling rate that limited growth.

F.

Melt Temperature

Additions of the elements Cu, Bi, Sb, or In can be used to reduce the melt temperature of eutectic Sn–Ag. Eutectic Sn–Ag has a high melt temperature (221jC), so assemblies must be exposed to temperatures as high as 240jC to achieve adequate wetting. However, typically adding a constituent to lower the melting temperature also results in creating a melting range. Consider the alloy Sn–3.4Ag–4.8Bi that melts between 208jC and 215jC. Although it melts at a somewhat lower temperature than eutectic Sn–Ag, it does so over a 7jC temperature range. This and other lead-free alloys have been found suitable for reflow attaching conventional surface mount components. But fillet lifting in the region of the annular ring of pin-in-hole (PIH) components during wave soldering is a significant concern for alloys that freeze over a temperature range of 5– 10jC or more. Fillet lifting susceptibility also tends to increase with substrate thickness and feature size, and also as CTE mismatch increases [32].

G.

Localized Microstructure Coarsening

Thermal fatigue and mechanical creep in solders occur more rapidly when elevated temperature cycles are involved as discussed in Chapter 19. Accordingly, one of the methods to improve solder joint reliability is to modify the system to achieve and maintain a fine and uniform microstructure. Systems whose major constitutes exhibit a significant mutual solid solubility for each other, especially at elevated temperatures (as is the case with the Pb–Sn system) do not maintain a refined microstructure. That is, these alloys often are not stable over time, resulting in microstructural coarsening. Instability is often confined to localized regions within a solder joint, brought on by strains concentrated near the terminal pads of solder joints. These coarsened bands serve as crack initiation sites from which fatigue cracks propagate and eventually result in failure. An advantage of the Sn–Ag system, which has a limited solubility of Ag in Sn, is that it is intrinsically more resistant to coarsening than Pb–Sn solders [16].

H.

Aging

Aging studies are conducted as an indicator of long-term microstructural stability under field conditions, i.e., how the microstructure evolves with time and temperature. The IMC Ag3Sn that is dispersed throughout the bulk solder coarsens, whereas the dendritic Cu3Sn IMC in the bulk solder was observed to decrease. A layer of Cu3Sn (q-phase) forms nearest to the copper [33]. 1. Temperature Effect Because of kinetic considerations, joints aged at a lower temperatures exhibit slower microstructural development and evolution. That is, both the IMC phases (i.e., Ag3Sn, Cu6Sn5) show little signs of coarsening in the bulk solder even after a year, accompanied by a very moderate increase in IMC thickness at Cu interfaces. The effect of temperature is apparent in a study conducted by Yang, et al. [23] who evaluated Sn–Ag solder joints of 132-pin, quad flat pack (QFP) components that were surface mount attached to FR-4 boards. The components were of the gull-wing type with copper alloy leads spaced on a 25-mil pitch. The solder joints were aged at several temperatures (40jC, 90jC, 140jC, and 190jC) for periods ranging between 1 and 300 days. The results of those aging studies are summarized in Fig. 1, which shows little or no structural changes in solder joints aged at 40jC, but very significant changes in those aged at higher temperatures, e.g., 140jC and 190jC.

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FIG. 1 Effect of aging time and temperature on the microstructure of Sn–Ag solder joints. Test specimens were aged at: a) 40jC, b) 90jC, c) 140jC, and d) 190jC. (From Ref. 23.)

2. High Temperature Stability Based on eutectic Sn–Ag’s higher melting point, and better mechanical properties (discussed in Section 2.J) at room temperature, the alloy is expected to exhibit better high-temperature performance than eutectic Sn–Pb. Of particular interest are applications such as under-the-hood electronics in automobiles, where the service temperature can reach 180jC. Although better than eutectic Sn–Pb, it is apparent from Fig. 1 that the Sn–Ag microstructure is also unstable at higher temperatures. 3. Interface Intermetallic Compounds It is important to note the differences observed in the nature of the IMC layer formed at the copper interface. At room temperature, and when aged at both 40jC and 90jC, the IMC layer only consisted of Cu6Sn but was found to consist of both Cu6Sn5 and Cu3Sn when aged at higher temperatures (140jC or 190jC) on the printed circuit board, but not component leads. Alloying elements such as Ni, Si, and Mg may hinder the formation of q phase, while enhancing the growth of the D-phase [34].

I.

Growth Kinetics of Intermetallic Compound

The growth kinetics of the Cu6Sn5 (D) and Cu3Sn (q) IMC phases has been found to follow a parabolic-growth-kinetics model [34,35], W ¼ wo þ kt1=2

ð1Þ

where W is the thickness of the IMC layer at any time, wo is the initial thickness of an IMC layer, k is the growth rate of the IMC layer, and t is time. 1. An Example The growth rate for D is taken to be, k(D)=8.99109 m/sec 1/2, and q is k(q)=4.43109 m/sec1=2 [35,36].

SN–AG AND SN–AG–X SOLDERS AND PROPERTIES J.

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Mechanical Properties

The microstructure consists of mostly pure Sn phase, which is much softer than the eutectic structure. The nonuniform distribution of this phase is believed to contribute to the eutectic Sn– Ag alloy’s creep and fatigue behavior. In those cases where a layered IMC structure forms at a copper interface, the presence of the Cu3Sn (q) phase may have an adverse effect on the reliability of solder joints including Sn–Ag, because of the hard, brittle characteristics of that IMC [37]. This can occur as a consequence of strains generated due to a CTE mismatch among the key elements comprising a solder joint: the copper or copper alloy termination pad, component lead, Cu–Sn IMC at the interface, and the solder. If the thermally induced strains are sufficient, cracks are initiated at the hard and brittle Dphase which can grow and propagate.

K.

Wetting Characteristics

1. Effect of Gold The enhanced solder spreading characteristics observed for eutectic Sn–3.5Ag solder on gold surfaces compared to copper are thought to be due to the greater affinity that tin-based alloys have in reacting with gold, as evidenced by how readily Au–Sn IMCs are formed [21]. 2. Effects of Atmosphere The reflow atmosphere has also been found to have a significant effect on the wetting time of Sn– 3.5Ag. It was determined that the wetting time could be reduced by a factor of approximately 1.5 depending on the nature of the reflow atmosphere. However, in the same study it was observed that the benchmark eutectic Sn–Pb solder displayed a greater sensitivity to reflow atmosphere, and that the Sn–3.5Ag alloy generally exhibited better wetting characteristics. Accordingly, it is anticipated that the wetting behavior of eutectic Sn–3.5Ag would be affected less than eutectic Sn–Pb to drastic changes in oxygen content of the reflow atmosphere [21].

III.

TIN–SILVER–BISMUTH (SN–AG–BI) SYSTEM

A.

Microstructure

A wide range of compositions in the Sn–Ag–Bi system has been reported [38–40]. When the Bi content is low (approximately 1%) it is primarily located within the lattice sites of the Sn matrix as a substitutional solid solution. However, when the Bi content exceeds about 1%, it precipitates out as a second phase at room temperature [41].

B.

Effect on Melting Point

Elemental additions of Bi convert the Sn–Ag binary alloy, which melts at a single temperature (221jC), to a ternary alloy that melts over a range of temperatures. For example, Sn–3.3Ag–4.7Bi melts between 210jC and 215jC. The use of bismuth additions is not an efficient way to reduce the melt temperature of Sn–Ag, as it is estimated that a 1% Bi addition results in about a 1.7jC temperature depression for Sn–Ag alloys in the 160jC to 230jC melt temperature range [41].

C.

Presence of Low-Melt Constituents

1. The Concern As the migration toward lead-free microelectronic assemblies moves forward, there will be a hybrid transition period. Some surface-mount-component lead frame terminations will be plated with Sn–Pb solder coatings. The nominal lead (Pb) content in a typical solder joint depending on the plating thickness for Sn–Pb solder compositions is given in Fig. 2. The effect of intentionally adding 3% and 6% Pb on the melting characteristics of several Sn–Ag–Bi-based alloys, on Sn–

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FIG. 2 Effect of component termination pad plating thickness on the nominal lead (Pb) content in surface mounted solder joints for several alloys. (From Ref. 42.)

Ag–Cu, and Sn–Bi is shown in Fig. 3. The concern is the creation of low-melting phases either due to the addition of Bi, Pb, or a combination of the two that would cause a significant reduction in mechanical properties. Of particular concern is a reduction in fatigue resistance resulting from localized deformation due to an increased creep rate relative to the higher melting point material. 2. Effect of Bismuth No low-melting point constituents were detected by differential scanning colorimetry (DSC) studies in any of the alloys shown in Fig. 3 that did not contain lead if the Bi content was less than 10%. This was also reported to be the case by Huang and Lee for Sn–Ag–Bi alloys with less than 7.5% Bi [43]. Bismuth lowers the melting point of Sn–Ag alloys, but a concentration of 5% Bi or more is required to make a reasonable difference in the reflow temperature [42]. It is encouraging that no evidence of the low-melting (138jC) Sn–Bi binary eutectic was noted to have formed due to inhomogenous solidification at the low Bi-content levels (i.e., process C, indicating the direct

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effect of void fraction on those properties. The composite solder made by process C using a composite solder paste did not show any strengthening even with the addition of up to 30 vol.% Cu6Sn5 particles. This suggests that the large volume fraction of porosity in the composite solder prevented strengthening by reinforcements. It was also observed that the average strain rate of the composite solder made by process A can be reduced by a factor of about 8 with the addition of 30 vol.% Cu6Sn5 particles. This composite solder exhibits comparable ductility with Sn–Pb solder with no reinforcements. Marshall and Calderon [31] compared two processing methods and concluded that the processing method significantly affects the microstructure of a composite solder. The first method involved blending powdered reinforcement particles with eutectic Sn–Pb solder paste. In the second method, the particles were slowly added to molten eutectic Sn–Pb solder, while the melt was stirred under an inert atmosphere. The particles utilized in both methods included Ni, Ag, Cu, Cu3Sn, and Cu6Sn5. The blended particle method provided the composite solders with a larger fraction of porosity than the particles addition to molten solder method for all types of particles tested. Increasing the volume fraction of reinforcement particles produced larger volume fraction of porosity in the composite solder. Remelting of the composite solders, produced by the first method, under an argon-purged high-vacuum environment greatly reduced the porosity in the composite solders. 2. Porosity Formation Mechanisms Clough et al. [33] proposed that a possible mechanism for the formation of pores is through flux or gas bubbles trapped due to the viscosity of the melt. Since each particles in the solder paste is coated with flux, flux can be internally trapped when the solder paste melts. On solidification, these bubbles are left as pores within the solder when there is insufficient time for the bubbles to escape. Porosity is generally detrimental to strength and ductility. Pores can serve as sites of stress concentration resulting in the formation of cracks, leading to failure. It is apparent that the method of mixing the particles with the molten solder containing the least amount of flux results in minimal porosity and higher strength compared to the methods utilizing solder paste. Based on buoyancy, the upward velocity (v) of gas bubbles in the melt can be described by the Stokes Law v¼

gðDqÞd2 ; 18g

ð1Þ

where g = 9.8 m/sec2 is the gravitational constant, Dq = qsqb is the difference between the density of the solder and the bubble, d is the bubble diameter, and g is the viscosity of the melt. The viscosity of molten solder at 280jC is about 0.13 Pa sec. The density of the solder is 8880 kg/ m3. When d = 0.5 Am, the velocity of the gas bubble is 9.3 103 Am/sec. On the other hand, if the bubbles are 50 Am in diameter, the velocity will be 93 Am/sec. Smaller bubbles have a lower velocity. Thus smaller bubbles tend to remain and form pores on solidification. The viscosity of the melt has been observed to increase with an increasing amount of added particles [33]. The velocity of bubble motion then decreases due to increased viscosity of the melt as evidenced by the above formula, resulting in larger fraction of residual porosity. There is no quantitative analysis of this phenomenon in solder available in published literature.

C.

Effects of Solubility and Diffusivity of In Situ Formed Reinforcements

The possibility of producing composite solders containing closely spaced, fine reinforcements by rapid-solidification processing of induction melting/gas atomization (IMGA) was investigated by Sastry et al. [34]. First, rapidly solidified powders of Sn–Pb–Cu and Sn–Pb–Ni alloys were produced by induction melting and gas atomization. Powders less than 180 Am in size were coldpressed and extruded into rods at 120jC to consolidate into bulk solder specimen. The extruded rods were reflowed with a maximum temperature of 200jC to produce microstructures simulating those found in solder joints. The dispersoids in Sn–Pb–Cu and Sn–Pb–Ni composite solders had a composition close to that of Cu-rich Cu6Sn5 and Ni-rich Ni3Sn4, respectively.

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In the rapidly solidified and extruded condition, the Sn–Pb–Cu composite solders exhibited 0.1 to 1 Am spherical Cu–Sn dispersoids in Sn–Pb solder matrix with 5–10 Am grain size. Reflow of the composite solders increased grain size of Sn and Pb phases and drastically changed morphology and size of Cu–Sn dispersoids. The Cu–Sn dispersoids underwent significant coarsening and changed their morphology to platelike shape. In the rapidly solidified and extruded condition, the Sn–Pb–Ni composite solders contained Sn–Pb solder matrix with finer grains of 1–5 and 0.05–0.5 Am spherical Ni–Sn dispersoids. Upon reflow, limited coarsening of grains and dispersoids resulted. As compared to Sn–Pb–Cu composite solders, the Ni–Sn dispersoids remained spherical and were less than 1 Am in size. As proposed by Sastry et al. [34], the difference in stability of dispersoids in both these composite solders was attributed to higher solubility and diffusivity of Cu in Sn–Pb matrices than Ni. Sn–Pb–Ni composite solder thus was expected to provide relatively greater strengthening than Sn–Pb–Cu composite solder. When tested in tension at 25jC and with a strain rate of 2 103/sec, these composite solders showed 35–180% higher yield stress and 12–90% higher Young’s modulus than those of eutectic Sn–Pb solder processed by the same method. As expected, Sn–Pb–Ni composite solder exhibited 20% higher yield stress and Young’s modulus than those of Sn–Pb–Cu composite solder.

D.

Effects of Dispersoids on Microstructural Stability

One of the important considerations to increase the reliability of the solder joints used in microelectronic industry is to engineer and stabilize microstructures of solders after reflow and homogenize the deformation during service. The effect of dispersoids on microstructural stability was investigated using dispersion-strengthened composite solders made by mechanical alloying [18]. Particles of Ni3Sn4 and agglomerates less than 20 Am in size were blended with 60Sn–40Pb solder and mechanically alloyed to introduce 2.2 wt.% dispersoids in the composite solder. Solder joints were prepared by reflowing the composite solder between two Cu blocks on a hot plate at 220jC and were subjected to aging at 100jC for 29 hr. The grain size of the composite solder was finer in both the reflowed and aged conditions compared to the Sn–Pb solder without dispersoids. The finer microstructure in the as-reflowed condition was attributed to the Ni3Sn4 dispersoids that acted as heterogeneous nucleation sites in the Sn–Pb solder matrix. The Ni3Sn4 dispersoids are also considered responsible for retarding grain growth during aging by pinning the interphase boundaries between Sn and Pb phases. As a result, the composite solders with finer grain size exhibited higher strength, ductility, and a greater resistance to both mechanical and thermomechanical fatigue damage compared with conventional solder without dispersoids.

E.

Effects of Ultrafine Oxide Dispersions

In optoelectronic packaging, strong creep resistance of the solder joints is required because of the need to maintain positional accuracy of the components over a long period of time. Gradual misalignment of the components due to time-dependent creep deformation results in a reduced intensity or even a complete loss of signal transmission. Higher melting point solders such as eutectic Sn–Au (MP 278jC) are often utilized in applications requiring creepresistant solder joints to ensure dimensional stability. However, the higher processing temperatures can cause deterioration of optical fibers and optoelectronic components. Thus it is highly desirable to use low melting point solders with high creep and mechanical strength for these applications. 1. A Novel Processing Method to Achieve Ultrafine Oxide Dispersions Dispersions of nanosize Al2O3 (10-nm average) and TiO2 (5-nm average) oxide particles in Sn–Pb solders were attempted by Mavoori and Jin [26,34] to develop composite solders with superior mechanical properties. Nanosized Al2O3 and TiO2 oxides particles were chosen because they are nonreacting, noncoarsening, and relatively undeformable. Thus they can enhance mechanical properties by impeding dislocation motion and grain boundary sliding. A uniform distribution of dispersoids in a solder matrix was achieved through a novel approach of particle coating and repeated plastic deformation to rearrange the oxide particles [26,34]. First, powders of eutectic

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Sn–Pb solder (nominally 35-Am size particles) and nanosized TiO2 and Al2O3 particles were dispersed and mixed in ethanol with constant stirring to form a slurry comprised of Sn–Pb particles covered with 3 vol.% oxide particles. After drying, they were compacted into pellets at room temperature. The pellets were flattened due to repeated compressive deformation at 120jC in an inert atmosphere. Such pellets were cut into six pieces, stacked, and sealed in a heattreatment envelope filled with argon, followed by heating to 120jC and pressing to obtain a sixfold reduction of the stack. This process was repeated six times to obtain a calculated interparticle spacing of 0.8 nm. Finally, the specimens were annealed to remove the residual stress. The dispersoid-free Sn–Pb solder control specimens were prepared by the same method, resulting in a 2-Am grain size. It was determined that superior creep resistance of other solder systems such as Sn–Au, Sn– Ag, or Sn–Bi solders can also be achieved by employing this technique. Mavoori and Jin suggested that either a preform or solder paste might be used in applications of these composite solders [34]. Fast soldering operations such as laser soldering were recommended to minimize possible segregation of the dispersoids during melting or solidification. It was also suggested that the choice of proper dispersoid particles density which is matched to the solder matrix is beneficial. 2. Effects of Fine Oxide Particles on Microstructure and Mechanical Properties It is ideal to have interparticle spacing less than the size of dispersoids for maximum strengthening (2 103/sec). It was also observed that the composite solders were more ductile than the Sn–Pb solder control at higher strain rates (>103/sec). It was tentatively attributed to the pseudosuperplastic behavior present in the ultrafine composite solders (200-nm grain size). This suggests that the composite solders can be engineered to exhibit high ductility over a wide strain rate range by altering the volume fraction of dispersoids or the grain size to make solders useful for high strain rate applications.

F.

Effects of Reinforcements with Low Coefficient of Thermal Expansion

Higher coefficient of thermal expansion (CTE) of solders as compared to the other entities present in a solder joint is a root cause for producing thermal strain in the solder joints and failure of electronic parts. Thus there is a need for a solder with a low CTE. Solders can be tailored to possess a desired thermal expansion coefficient with proper choice of type and volume fraction of reinforcements.


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1. Carbon Fiber Reinforcements Incorporation of carbon fibers was attempted to develop a composite solder with a low CTE [35]. Carbon fibers were selected because they have nearly a zero CTE with good electrical and thermal conductivities. A continuous form of carbon fibers was also expected to prevent nonuniform distribution of carbon fibers after remelting and solidification. A carbon-fiber reinforced solder was fabricated by squeeze casting 60Sn–40Pb solder with continuous Cu-coated, 10-Am diameter carbon fibers. About 1-Am thick Cu layer was electroplated on carbon fibers to ensure good wetting to the solder and the fiber. Composite solders containing uncoated fibers were also prepared for comparison. The CTEs of the composite solders were measured at temperatures ranging from 25jC to 105jC along directions parallel and normal to the fibers. The CTE values parallel to the fibers were lower than those perpendicular to the fibers. The CTEs in the parallel direction drastically decreased with increasing fiber content, while the reduction of CTEs in the normal direction was not significant. The composite solders showed CTE values close to 0 in the parallel direction at fiber contents above 42 vol.%. Increasing temperature caused an increase in CTE values for a given fiber content and fiber orientation. 2. Mechanical Properties In tensile tests where the load was applied along the fiber direction, the tensile modulus and strength increased with increasing volume fraction of carbon fibers, while the ductility decreased with increasing fiber content. The composite solders containing coated fibers exhibited higher strength and ductility than those with uncoated fibers. This result was attributed to the improved wetting between the solder matrix and fiber provided by the Cu coating. On the other hand, the modulus was not affected by copper coating. Thermal fatigue testing was performed using solder joints made with alumina (Al2O3) and Kovar (Fe–29Ni–17Co) by cycling between room temperature and 100jC. Alumina and Kovar were coated with Au (20 to 40 A˚ thick) to promote wetting prior to soldering. The start of debonding identified by slight cracking between the composite solder and alumina or between the composite solder and Kovar was used for estimation of the fatigue life. Among the fiber volume fractions tested, the composite solder joints containing about a 29.42 fiber vol.% exhibited the longest fatigue life. This result was attributed to the fact that the composite solder containing 29.42 fiber vol.% exhibits a CTE (4.3 106/jC at 25jC) closest to alumina (6.7 106/jC at 25jC) and Kovar (5.3 106/jC at 25jC) parallel to fiber orientation.

G.

Growth of Interfacial Intermetallic Compound Layers

1. Typical Growth Without Reinforcements Two layers consisting intermetallic compound (IC) of Cu3Sn and Cu6Sn5 form at the interface of Sn-based solders and Cu circuit pads. The Cu-rich Cu3Sn layer forms next to the Cu pads and Snrich Cu6Sn5 layer forms adjacent to the solder. Immediately after soldering, only the Cu6Sn5 layer is observed and aging is required to form the Cu3Sn layer. During aging, the interfacial IC layers grow by consuming Sn from the solder and Cu from the bond pad [32]. 2. Effect of Reinforcement Type on Solder Interfacial IC Growth Wu et al. [32] and Pinizzotto et al. [36] compared the growth of interfacial IC layers in eutectic Sn– Pb composite solder joints with similar joints without reinforcements at several aging temperatures ranging to 160jC for periods up to 64 days (1536 hr). The composite solders were prepared by mixing the eutectic Sn–Pb solder paste with particles of Ni, Ag, Au, Cu, Cu3Sn, and Cu6Sn5. The solder joints were produced via reflow on a hot plate preheated to 250jC for 1 to 1.5 min. Although none of the particle additions strongly affected the initial thickness of the Cu6Sn5 layer, they had significant influence on the growth kinetics of the interfacial IC layer. Compared to the standard eutectic Sn–Pb solder (i.e., no reinforcements), additions of Ag and Au particles retarded, and Ni particles completely suppressed the growth of the Cu3Sn layer, while enhancing the growth of the Cu6Sn5 layer. In contrast, Cu-based particles (Cu, Cu3Sn, and Cu6Sn5) impeded Cu6Sn5 layer growth but promoted Cu3Sn layer growth.

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3. Interfacial IC Growth Kinetics Since the Cu substrate/Sn-based solder interface is planar in nature, the growth of Cu–Sn interfacial IC layers in the solid state occurs in a planar manner. Tin and copper diffuse in opposite directions, and then react with each other. The growth kinetics of the interfacial IC layers is modeled in many studies by assuming a diffusion-controlled growth [32]: pffiffiffiffiffiffi x ¼ Dt; ð2Þ where x is the interfacial IC layer thickness at time t and D is the IC layer growth rate constant expressed by D ¼ Do expðQ=RTÞ;

ð3Þ

where Do is the preexponential constant corresponding to absolutely maximum possible layer growth rate (in the absence of an activated process, where Q = 0), Q is the activation energy for the growth of IC layers, R is the universal gas constant, and T is the temperature in degrees Kelvin. It has been determined that the diffusion-controlled growth model is valid for interfacial IC layer growth over the temperature range studied (RT to 160jC) [32,36]. Table 4 lists the activation energy values for intermetallic layer growth in Sn–Pb eutectic and composite solders. The incorporation of Cu-based particles (Cu, Cu3Sn, and Cu6Sn5) increased the activation energy for Cu6Sn5 layer growth while decreasing the activation energy for Cu3Sn layer growth in eutectic Sn–Pb solder. This is consistent with findings of a retarded growth of Cu6Sn5 layer and enhanced Cu3Sn layer growth as a consequence of adding Cu-based reinforcements. On the other hand, Ag increased while Au decreased the activation energy for Cu6Sn5 layer growth compared to eutectic Sn–Pb solder. Tin–lead eutectic composite solders containing Ag and Au particles produced lower activation energies for Cu3Sn layer growth than without those additions, thus exhibit a more rapid Cu3Sn growth rate. Nickel particles drastically increased the Cu6Sn5 thickness with higher activation energy for the growth of Cu6Sn5 layer than that in the eutectic Sn–Pb solder. However, the measured activation energy values for intermetallic layer growth in composite solder containing Ni and Ag reinforcements appear to be in contradiction with the observed enhanced growth of Cu6Sn5 and Cu3Sn layer. 4. Proposed Growth Mechanisms Based on the findings of activation energies for growth, and growth behavior of the interfacial IC layers, possible mechanisms by which reinforcement particles affect the growth of intermetallic layers have been proposed [32]. The proposed mechanisms are based on an overall reduction of tin diffusion through the solder matrix required for growth of intermetallic layers at the solder/Cu interface. First, Cu and Cu3Sn particles act as Sn sinks. During aging, both Cu and Cu3Sn TABLE 4 Activation Energies for the Growth of Cu6Sn5 and Cu3Sn Intermetallic Compounds in Eutectic Sn–Pb and Composite Solders Activation energy (kJ/mol)

Eutectic Sn–Pb and additions

Cu6Sn5 IC layer

Cu3Sn IC layer

Eutectic Sn–Pb 20 wt.% Cu6Sn5 20 wt.% Cu6Sn5 7.6 wt.% Cu 4 wt.% Au 4 wt.% Ag 4 wt.% Ni

77 118 117 96 54 131 209

163 92 114 108 107 114 –

Source: Refs. 32 and 36.


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particles continue to react with Sn until the preferred terminal phase Cu6Sn5 is formed. They remove Sn from the solder and reduce the amount of Sn available to form intermetallic layers. Removing Sn enriches Cu, so the kinetic balance is directed toward favoring the formation of Cu3Sn over Cu6Sn5. However, the mechanism of a Sn sink is not applicable to Cu6Sn5 particles because Cu6Sn5 does not directly react with Sn. It was postulated that Cu6Sn5 particles impede Sn diffusion by serving as a partial barrier, limiting the Sn supply to the solder/Cu substrate interface. This promotes the growth of the Cu-rich Cu3Sn phase and deters the growth of Sn-rich Cu6Sn5. However, the role of reinforcement particles present in a solder matrix on the growth of Cu6Sn5 and Cu3Sn intermetallic layers at the solder/Cu interface is uncertain at present.

H.

Further Investigations and Implementation

Other studies which address the effects of reinforcement particle additions on the microstructure, solderability, and mechanical properties of composite solders have also been reported [16,17,25,27,31,37]. In general, composite solders tend to render improved material properties. All the reported investigations were basically exploratory in nature, and further detailed studies are necessary. However, the driving factors for continued development and ultimately widespread implementation of composite solders are heavily dependent on environmental and economic considerations. Most investigations on composite solders have been focused on leadbearing systems, and studies on lead-free composite solders commenced much later.

VI.

STUDIES ON LEAD-FREE COMPOSITE SOLDERS

Several lead-free solders such as binary Sn–Ag, Sn–Bi, ternary Sb–Ag–Cu, etc. are being considered as potential alternatives for lead-bearing solders. Composite approaches used in lead-bearing solders can also be applied to lead-free solders. Although studies on composite solders utilizing lead-free solders are somewhat limited compared to those for lead-bearing solders, developments of lead-free composite solders are essential based on the reliability considerations.

A.

Bulk Solder Without Reinforcements

Eutectic Sn–Ag solder possesses a microstructure consisting of almost pure Sn phase and the eutectic, a mixture of Sn and Ag3Sn intermetallic phases. The eutectic Sn–Ag solder was found to exhibit microstructural instability during isothermal aging as observed with lead-bearing solders. Fig. 1 shows the microstructures of eutectic Sn–Ag solder joint made with Cu substrate before and after aging at 150jC for 100 hr [5]. During aging, Ag3Sn particles in the solder matrix coarsen, resulting in an increased particle size and a reduction in particle count. After aging, Ag3Sn particles are primarily found along Sn–Sn grain boundaries. It is considered that coarsening of Ag3Sn particles took place along grain boundaries providing a fast diffusion path. Cu6Sn5 intermetallic phase in the solder joint is also shown in Fig. 1, resulting from the dissolution of Cu from substrate and reaction with Sn in solder. The aging temperature has been found to dramatically affect the coarsening of Ag3Sn particles. The instability of the Ag3Sn phase provides the impetus for identifying suitable particle-strengthening strategies.

B.

Solders with Reinforcements Introduced by In Situ Method

Figs. 2 and 3 provide microstructures of eutectic Sn–Ag solder joints containing Cu6Sn5, Ni3Sn4, and Fe–Sn reinforcement particles in the as-joined condition and after long-term aging [38,39]. Intermetallic particles of Ni3Sn4 and Fe–Sn incorporated into the solder matrix by the in situ method were found to be relatively stable even after 1400 hr of aging at 180jC as compared to Cu6Sn5 intermetallic particles, an attribute desirable for a composite approach. The Cu6Sn5 intermetallic reinforcements were observed to coarsen under long-term conditions, 5000 hr at

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FIG. 1 Light photo micrographs depicting microstructural coarsening of the eutectic Sn–Ag solders, where the dark phase is Sn and white particles are Ag3Sn. (a) Before aging, (b) after aging at 150jC for 100 hr. (From Ref. 4.)

FIG. 2 Scanning electron microscope secondary electron emission photographs of the initial microstructure of Sn–Ag eutectic solder joints containing 20 vol.% (a) Cu6Sn5, (b)FeSn2, and (c)Ni3Sn4 reinforcement particles. (From Refs. 38 and 39.)


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FIG. 3 Scanning electron microscope secondary electron emission photographs of the microstructure of in situ composite solder joints after aging (a) Cu6Sn5 composite solder at 150jC for 5000 hr, (b) FeSn2 composite solder at 180jC for 1400 hr, and (c) Ni3Sn4 composite solder at 180jC for 1400 hr. (From Refs. 38 and 39.)

FIG. 4 Scanning electron microscope secondary electron images depicting microstructural stability of Cu6Sn5 particles during short-term aging of a composite Sn–Ag solder (a) before aging, (b) after aging at 150jC for 100 hr. (From Ref. 4.)

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temperatures above 120jC. They did not coarsen during short-term aging consisting of a few hundred hours as shown in Fig. 4. This may indicate different mechanism of coarsening at low and high temperatures. 1. Interfacial Intermetallic Compound Layers The effect of the Cu6Sn5, Ni3Sn4, and FeSn2 particles on the solder/substrate interfacial IC layer growth in the eutectic Sn–Ag composite solder joints was investigated by carrying out aging studies up to a few thousand hours at several temperatures [38,39]. Fig. 5(a) shows an example of the total growth of interfacial IC layers (Cu6Sn5 and Cu3Sn) during aging at 150jC up to 5000 hr. The solder joints exhibited only Cu6Sn5 intermetallic layer at initial condition, and the Cu6Sn5 and Cu3Sn intermetallic layers are clearly visible after aging as shown in Fig. 3. The interfacial IC layers normally exhibited rapid growth at the initial stage of aging, but the growth rate decreased as the layer thickness increased. The reinforcement particles were observed to affect the formation of interfacial intermetallic layers during reflow, resulting in a thinner initial interfacial layer compared with the same solder without reinforcements. Fig. 5(b) shows IC layer growth rate constant of the Cu6Sn5 intermetallic layer in the eutectic Sn–Ag and Cu6Sn particulate-reinforced Sn–Ag solder joints during aging at 150jC up to 5000 hr at five temperatures. The Cu6Sn5 particulate-reinforced Sn–Ag solder joints showed higher IC layer growth rate constants at temperatures above 120jC, while the IC layer growth rate constants were lower below 120jC. A similar behavior was observed for Ni3Sn4 and FeSn2 particulate-reinforced Sn–Ag solder joints. This was a temperature-dependent effect in that the reinforcement particles were effective in retarding the interfacial intermetallic layer growth at temperature below 120jC, but not above this temperature, as shown in Fig. 5(b).

C.

Solders with Mechanically Added Reinforcements

The stability of mechanically incorporated metallic reinforcement particles in the Sn–Ag composite solders was examined after multiple reflows and during aging. The composite solders were prepared by mechanically blending nominally 15 vol.% of micron-sized Cu, Ag, or Ni

FIG. 5 (a) Graph which depicts the effect of aging time at 150jC on the total thickness of interfacial IC layers (Cu6Sn5 + Cu3Sn) for eutectic Sn–Ag and composite Sn–Ag solder joints which contain Cu6Sn5 intermetallic particles introduced by the in situ method. (b) Graph which depicts the IC layer growth rate constants at five different temperatures for the Cu6Sn5 intermetallic layer in the eutectic Sn–Ag solder joints and Cu6Sn5 particulate-reinforced Sn–Ag solder joints during aging at 150jC up to 5000 hr. (From Ref. 38.)


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particles with eutectic Sn–Ag solder to form a mixture with a uniform distribution of reinforcement particles [40,41]. Figs. 6–8 illustrate the microstructure of Cu, Ag, and Ni particles in asfabricated solder joints, after aging at 150jC for 500 hr, and after three reflows at 280jC, respectively. 1. As-Fabricated Microstructure The intermetallic layers formed around the particles due to the reaction of Sn in the solder matrix with the metallic particle during the first reflow can be seen in Fig. 6. The Cu–Sn intermetallic layers consist of Cu3Sn and Cu6Sn5 layers formed on the surface of Cu particles, and Ag3Sn intermetallic layer formed on the surface of Ag particles. In contrast, a ternary Cu–Ni–Sn intermetallic layer formed around the Ni particles. This indicates that Cu diffused to Ni particles from the Cu substrate during reflow and participated in the formation of the ternary Cu–Ni–Sn intermetallic layer. It was also observed that binary Cu–Sn and ternary Cu–Ni–Sn intermetallic

FIG. 6 Scanning electron microscope secondary electron images depicting the as-fabricated microstructure of Sn–Ag eutectic solder joints with (a) Cu reinforcement, (b) Ag reinforcement, and (c) Ni reinforcement, which were mechanically added. In (a), a padal-like layer of CuSn IC with a light appearance surrounding the dark-appearing Cu particles was added. In (b), the added Ag particles have a gray appearance and surrounded by a thin layer if light-appearing Ag3Sn. The Ni particles added in (c) appear dark and surrounded by light-appearing Cu–Ni–Sn intermetallics. (From Refs. 40 and 41.)

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FIG. 7 Scanning electron microscope secondary electron images depicting the thermally aged microstructure of Sn–Ag eutectic solder joints with (a) Cu reinforcement, (b) Ag reinforcement, and (c) Ni reinforcement, which were mechanically added. The solder joints were aged at 150jC for 500 hr. (From Refs. 40 and 41.)

layers were much thicker on Cu or Ni particles than the Ag3Sn intermetallic layer on Ag particles. Similar differences in the thickness of the intermetallic layers around particles were also noted after isothermal aging at 150jC or reflow at 280jC as shown in Figs. 7 and 8. 2. Aged Microstructure As noted in Fig. 7, the growth of intermetallic layers around particles during aging was in the following order: Ni particles>Cu particles>Ag particles. In the study conducted by Guo et al. [40], Cu particles were largely consumed by Sn in the solder matrix after aging at 150jC for 1000 hr, while there were no significant changes in the case of Ag particles under similar conditions. In contrast, Ni particles were completely converted to ternary Cu–Ni–Sn intermetallics after aging for 500 hr [41]. These studies attribute the difference in growth kinetics of intermetallic layers that surrounded Ni, Cu, and Ag particles to their difference in diffusion coefficients. Table 5 gives a comparison of diffusion distances of metallic elements in solid Sn under the same condition. As indicated in the table, the diffusion of Ni, Cu, and Ag in Sn is in the following order: Ni>Cu>Ag under the same conditions. This large difference in diffusion capabilities may account for the observed difference in intermetallic layer growth around reinforcement particles. Since Ag particles consume less amount of Sn, it is expected that change in composition of solder matrix would be least in Ag-reinforced composite solders compared to composite solders containing Cu


319

FIG. 8 Scanning electron microscope secondary electron images depicting the multiple reflowed microstructure of Sn–Ag eutectic solder joints with (a) Cu reinforcement, (b) Ag reinforcement, and (c) Ni reinforcement, which were mechanically added. The solder joints were reflowed three times at 280jC with a total time of about 30 sec above the melting point of eutectic Sn–Ag (MP = 221jC). (From Refs. 40 and 41.)

TABLE 5

Comparison of Diffusion Distance for Ag, Cu, and Ni in Sn

Diffusion species and direction Ag in Sn ?c Ag in Sn //c Cu in Sn ?c Cu in Sn //c Ni in Sn ?c Ni in Sn //c a

D0 (cm2/sec)

Q (kJ/mol)

0.18 77 0.0071 55.1 0.0024 33.1 D =2 106 cm2/seca 0.0187 54.2 0.0192 18.1

(Dt)1/2, 100 hr (Am) 25jC

75jC

100jC

150jC

0.454 7.5 369 8,485 — 21,549

4.2 37 964 — — 36,421

— — 1,414 — — 44,915

45 200 2,660 — 369 —

D0 and Q values are not available in the references individually. ?c indicates diffusion in the direction perpendicular to the c-axis of a tin unit cell. //c indicates diffusion in the direction parallel to the c-axis of a tin unit cell. Source: Refs. 4 and 40.

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and Ni reinforcements. Ag particles also exhibit higher resistance to coarsening than Cu and Ni reinforcement particles due to its lower diffusion coefficient in a Sn matrix. 3. Effect of Multiple Reflows on Microstructure On the other hand, Fig. 8 illustrates that the growth of intermetallic layers around particles during multiple reflow was: Cu particles>Ni particles>Ag particles. The intermetallic layers present around metal reinforcement particles appear to control the growth of the intermetallic layers under such conditions.

D.

Effects of Reinforcements on Mechanical Properties/Other Characteristics

1. Examples of Enhanced Creep Resistance a. Fe in Sn and Sn–Bi. Composite solder containing Fe particles introduced in the presence of a magnetic field to achieve uniform distribution of reinforcements in solder was found to possess greater creep resistance than unreinforced solders, as reported by McCormack et al. [29]. The Fe powders were chemically coated with a submicron layer of Sn before processing to improve wetting and prevent oxidation during handling and processing. Powders of Sn and the Sn–Bi alloy were mixed with 2.5 wt.% of approximately 2-Am size Fe powder in hexane to form a slurry. After drying the slurry, the dry powder mixture was melted and solidified in a quartz tube under a magnetic field. The choice of Fe powders as reinforcements is based on the facts that Fe has a low solubility and diffusivity in Sn-based solders, avoiding their coarsening. The Sn composite solders containing 2.5 wt.% Fe powder exhibited a 60% to 100% higher ultimate tensile strength than the dispersion-free control solder. More importantly, the creep resistance of the composite solder at 100jC was 20 times greater than the dispersoid-free solder. In the case of the eutectic Sn–Bi composite solder with Fe particles, the tensile strength was a modest 10% higher, and the creep resistance was 5 times better than the dispersoid-free specimens under similar conditions. b. Cu6Sn5 in Sn–Ag. Results of creep testing performed on Sn–Ag composite solder joints containing 20 vol.% Cu6Sn5 particles introduced by the in situ method in as-fabricated and preaged conditions for 100 hr at various temperatures have been reported by Subramanian et al. [42] and Choi et al. [43]. Creep testing was carried out in an interrupted manner where the specimens were periodically unloaded to obtain micrographs of a particular scratch placed on the side of a solder joint to quantify the shear displacement in the specimen. Fig. 9 illustrates creep behavior of the composite solder joints in as-fabricated and pre-aged condition along with data of eutectic Sn–Ag solder joints from Darveaux et al. and McDougall et al. The creep resistance of the composite solder joints was about 2 to 3 orders of magnitude better compared to noncomposite solder in the as-fabricated condition at room temperature. Furthermore, the composite solders were capable of sustaining a creep strain similar to lead–tin solders, but lower than the dispersoidfree control solders can be sustained [42]. It was also determined that pre-aging of the composite solder reduced its creep strength. Depending on the aging temperature and time, its behavior can approach that of the noncomposite solder. In dispersoid-hardened composite alloys, the particle size and interparticle spacing are important parameters affecting the mechanical properties of composite solders. In the pre-aged condition, the intermetallic particles maintained their initial morphology, while Ag3Sn particles coarsened, resulting in a larger particle size and interparticle spacing. A coarsened microstructure apparently allows dislocations to bypass particles more easily, thus rendering them as ineffective dislocation barriers, resulting in reduced creep resistance. c. Cu6Sn5, Ni3Sn4, and FeSn2 in Sn–Ag. Choi et al. [44] have performed constant-load creep tests on composite solder joints containing 20 vol.% in situ Cu6Sn5, Ni3Sn4, and FeSn2 reinforcement particles. Fig. 10 compares the normalized creep behavior of eutectic Sn–Ag and each composite solder joint, earlier creep data obtained from cyclic creep testing, and the data of Darveaux and Banerji [44]. For the eutectic Sn–Ag solder joints, the activation energy for creep


321

FIG. 9 Log plot of normalized shear strain rate vs. normalized shear stress data showing creep behavior of Sn–Ag composite solder joints in the as-fabricated, pre-aged condition, along with data from Darveaux et al., and McDougall et al. (From Refs. 42 and 43.)

was determined to be 60 kJ/mol, which is about half the activation energy for tin self-diffusion. This suggests that grain boundary or dislocation pipe diffusion may control deformation of Sn– Ag eutectic solder joints. A value of 120 kJ/mol was found in composite solder joints containing Cu6Sn5 or Ni3Sn4 reinforcements, but only 38.5 kJ/mol for composite solder joints with FeSn2 reinforcements. Fig. 10 shows that Sn–Ag composite solder joints typically exhibit the same properties as Sn–Ag eutectic solder joints at elevated temperatures, but not at lower temperatures. Cu6Sn5 and Ni3Sn4 composite solder joints appear to be weaker at room temperature, while FeSn2 composite solder joints are effectively stronger than the eutectic solder joints at equivalent normalized stresses. 2. Creep Resistance Factors Data from Choi et al.’s [43] prior work on the same in situ Cu6Sn5 composite solder joints showed substantially better creep resistance even under pre-aged conditions as shown in Fig. 10(a). These data points were obtained by unloading the specimen to obtain micrographs of a particular scratch placed on a solder joint to quantify the shear displacement. Recovery that can occur during these periods of interrupted unloading apparently reduces the mobile dislocation density available for deformation on reloading the sample. Quantitative evaluation of global and localized creep deformation behavior of solder joints has been developed by mapping the distortion of a straight-line ablation pattern imprinted by an excimer laser on the surface of shear lap solder joints [45]. Fig. 11 is an illustration of distortion due to creep deformation of a laser ablation pattern placed on a solder joint [46]. a. Effect of Temperature. Damage based on the onset of tertiary creep was examined in shear creep deformation of eutectic Sn–Ag single shear lap solder joints with and without Cu6Sn5 reinforcements. These studies indicated that at elevated temperature, local tertiary creep commenced at similar strains for both composite and noncomposite specimens, and their fracture surfaces were also similar. However, composite solder joints generally strain more homogeneously than noncomposite counterparts at elevated temperatures. At room temperature, where the creep resistance of the composite solders was found to be much greater than noncomposite parts, the composite solders generally exhibited a more uniform strain distribution. However,

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FIG. 10 Graphical depiction of normalized steady-state creep behavior of eutectic Sn–Ag and composite solder joints. Data are presented for (a) the eutectic and Cu6Sn5 composite solders, (b) the eutectic and Ni3Sn4 composite solders, and (c) the eutectic and FeSn2 composite solder joint. Note that (a) contains data from Choi for solder joints in the pre-aged condition. (From Ref. 44.)


323

FIG. 11 Shape of laser ablation patterns placed on a solder joint. (a) Before creep and (b) after creep. (From Ref. 46.)

local tertiary creep was observed to commence at a lower creep strain compared to noncomposite solder joints. It is considered that since the composite material is more creep-resistant at lower temperatures, the strain hysteresis may be smaller during thermal cycling, preventing development of large strains. However, it is not clear whether this benefit outweighs the fact that less strain before the onset of tertiary creep is tolerated locally in a composite solder at higher strain rates and lower temperatures. b. Strain for Tertiary Creep. Eutectic Sn–Ag solder joints exhibited a significantly higher strain for the onset of tertiary creep than all the composite solder joints tested. Among the composite solder joints, the tertiary creep of in situ Cu6Sn5 composite solder joints occurred at strain value comparable to eutectic Sn–Ag solder joints. The relatively higher strain value required for the onset of tertiary creep observed in in situ, Cu6Sn5-reinforced composite solder joints is attributed to the weak interfacial bonding between Cu6Sn5 particles and the solder matrix identified from nanoindentation tests [51]. It was concluded that the weak interface of uniformly distributed Cu6Sn5 particles in the solder matrix provided multiple initiation sites for deformation throughout an entire solder joint. This condition promoted homogeneous deformation, leading to the observed higher ductility of in situ compared to composite solder joints prepared by mechanical mixing. c. Particle Reinforcement Type. A summary of the creep properties of composite solders reinforced with Cu, Ag, and Ni metallic elements introduced by mechanical mixing methods is given in Table 6, as are eutectic Sn–Ag and in situ Cu6Sn5 composite solders for comparison [46– 48]. Solder joints made with Cu-particle reinforced composite solders are noted to exhibit a lower steady-state creep rate at all testing temperatures compared with Ag particle-reinforced composite solder joints. The creep behavior of the Ag particle-reinforced composite solder joints is comparable to eutectic Sn–3.5Ag solder. Ni particle-reinforced composite solder joints are about 5 times more creep-resistant than composite solder joints reinforced with Cu and about 30 times better than eutectic Sn–Ag and composite solder reinforced with Ag. Even at 105jC, the Nireinforced composite solder joints were stronger than Cu- and Ag-reinforced composite solder joints.

E.

Stress Relaxation

Stress relaxation describes the time-dependent deformation behavior of a material under constant displacement, which is evaluated by measuring the stress reduction with time. The evaluation of the stress relaxation process is necessary in order to assess the effect of hold periods in the thermomechanical fatigue behavior of solder joints. Understanding of the stress relaxation mechanisms is important because the deformation processes involved in stress relaxation are fairly representative of the stress history experienced by solder joints at the temperature extremes of thermal cycling experienced in service. Thus the stress relaxation behavior may be a better

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TABLE 6 Summary of Creep Properties of Eutectic Sn–Ag and Composite Solders Prepared by In Situ and Mechanical Mixing Methods Steady-state creep rate (S1) (at 17 MPa nominal stress) Solders Sn–3.5Ag solder paste Cu particle-reinforced composite solder (15 vol.% Cu) Ag particle-reinforced composite solder (15 vol.% Ag) Ni particle-reinforced composite solder (15 vol.% Ag) In situ Cu6Sn5reinforced composite solder (20 vol.% Cu6Sn5)

105jC

Average strain for the onset of tertiary creep (at 17 MPa)

2.62 10

4

1.50 10

1.90 103

0.54

—

4.07 106

2.03 105

3.95 104

0.08

M

1.91 105

8.37 105

1.80 103

0.19

M

8.02 107

1.08 105

1.54 104

0.11

M

7.6 106 (25 MPa)

85jC

125jC

0.45

I

9.8 105 (13 MPa)

5.8 104 (12 MPa)

25jC

65jC 5

Methods M-Mech. I—in situ

Source: Refs. 46–48.

metric rather than steady-state creep behavior in developing constitutive relationships required for modeling thermomechanical fatigue behavior of solder joints. 1. Effect of Plastic Deformation History Stress-relaxation tests were performed using single shear lap solder joints made with eutectic Sn– Ag and Sn–Ag composite solder containing 20 vol.% Cu6Sn5 particles introduced by in situ methods at 25jC and 150jC, as reported by Jadhav et al. [49]. The solder joint thickness was about 100 Am and soldered area was approximately 1 mm2. The Cu half dog bones, about 0.5 Am thick, are with reduced gauge sections of 1 mm in width and 11.5 mm in length. Stress-relaxation tests were conducted after imposing three different displacement levels, which correspondingly created three stress states: before peak stress, at peak stress, and beyond the peak stress. It was determined that the percent stress reduction (%SR) increased with increasing plastic strain imposed prior to relaxation as illustrated in Fig. 12. The %SR also increased with temperature for all the pretest stress levels examined. These trends were observed for both eutectic Sn–Ag and composite Sn–Ag solder joints. It was also observed that the %SR depends on the accumulated plastic strain, indicating a dependence on the prior deformation history. Substrates to which solder joints are attached can significantly affect the stress-relaxation behavior of those solder joints. For example, it was determined that solder joints of the configuration and dimensions described earlier in this paragraph did not exhibit complete stress relief even after several hours, while the stress in bulk solder relaxes to zero with time depending on the temperature [52]. This indicates that constraints imposed by a substrate and specimen geometry play important roles during the stress-relaxation process of solder joints. 2. Creep Mechanism The creep parameters required to assess deformation mechanisms were evaluated utilizing stressrelaxation data [49]. The stress exponent values obtained during relaxation were large, i.e., above 7, which are commonly observed for dispersion-strengthened alloys. The experimental results from the stress-relaxation tests were interpreted on the basis of dislocation recovery processes. During relaxation at high temperatures (>0.5 Tm), dislocations are annihilated by thermally activated rearrangement processes, resulting in a reduced mobile dislocation density during the


325

FIG. 12 Stress relaxation curves of three different composite solder joints ranging in thickness from 110 to 115 Am and tested at 25jC. Arrows indicate the beginning of stress relaxation. Cross-head speed is 0.5 mm/min. Initial plastic shear strains of 0.08 (A), 0.48 (B), and 1.5 (C) correspondingly resulted in 10%, 53%, and 77% stress reduction after 1 hr. (From Ref. 49.)

stress relaxation process. Therefore the dislocation density prior to relaxation determines the plastic strain level achievable prior to dislocation annihilation. Accordingly, when a high initial mobile dislocation density is created (e.g., by inducing plastic strain), it correspondingly leads to more plastic strain during relaxation, as these dislocations are annihilated, resulting in a greater %SR.

F.

Solderability

The wettability of in situ composite solders was observed not to degrade up to 20% volume fraction of Cu6Sn5, and the limit was investigated [42]. The contact angle is a measure of the wettability of solder to the substrate. Several micrographs of the spherical caps of the solders produced by reflow of disk-shaped solder preforms on Cu substrate were analyzed with PhotoShopR to arrive at average contact angles [42]. The measured contact angles for composite and noncomposite solders on a Cu substrate are given in Table 7. The eutectic and composite eutectic solders exhibited similar contact angles, indicating no substantial difference in wettability.

TABLE 7 Contact Angle Measurements of Eutectic Sn–Ag Solder and 20 vol.% In Situ Cu6Sn5 Reinforced Sn–Ag Composite Solders on a Copper Substrate Contact angle measurement

Specimen

Actual contact angle

Composite Sn/Ag solder 5-disk stack; disk radius = 1500 Am; disk thickness = 165 Am Eutectic Sn/Ag solder 5-disk stack; disk radius = 1500 Am; disk thickness = 190 Am

#1

16.62, 17.15

#2 Average #1

21.17, 19.55 18.63 18.52, 19.35

#2 Average

19.67, 19.62 19.29

Source: Ref. 42.

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The comparison of average wetting angles of mechanically incorporated Cu, Ni, and Ag particle-reinforced composite solders measured after four reflows is given in Table 8 along with the data for the eutectic Sn–Ag solder and in situ Cu6Sn5-reinforced composite solders [41,47]. The wettability of composite solder with 6 vol.% mechanically incorporated Cu particles is comparable to 15 vol.% Ag particle-reinforced composite solder produced by a similar method and 20 vol.% of in situ Cu6Sn5-reinforced composite solder. This was attributed to the fact that the volume fraction of Cu–Sn intermetallics was nearly doubled after four reflows, resulting in an increased solder viscosity and wetting angle during reflow. On the other hand, the wetting angle of the 15 vol.% Ni-reinforced composite solder (12.5j) was comparable to eutectic Sn–Ag solder (10.5j), but significantly better than 15 vol.% Ag (21.1j) and 15 vol.% Cu (47.3j) reinforced solders. The total effective volume fraction of Ni reinforcements was observed to remain nearly unchanged during multiple reflow processes.

G.

Fracture Characteristics

The fracture behavior of in situ composite and noncomposite solder joints was studied in isothermal mechanical fatigue (i.e., a bending mode) using a mechanical vibration exciter [50]. The fracture surface of the composite solder joints exhibited cleavage of the Cu6Sn5 particulate reinforcements and ductile mode I fracture of the eutectic matrix where fracture initiated at several locations, suggesting homogeneous ductile fracture. Mode I failure refers to fracture in a plane perpendicular to the loading direction. Meanwhile, the fracture surface of noncomposite counterparts exhibited ductile and mixed fracture behavior of modes I and II and step-type fatigue striations that originated at a local region. Mode II failure is due to shear directed parallel to the fracture surface.

H.

Weakly Bonded Reinforcements

Nanoindentation experiments have shown that the interfacial strength between particles and solder matrix can be estimated from indenting particles at an off-center location so that they

TABLE 8 Average Wetting Angles of Metallic Particle-Reinforced Composite Solders, Along with Eutectic Sn–Ag Solder, and Sn–Ag In Situ Composite Solders after Four Reflows on a Copper Substrate

Solders Eutectic Sn–Ag solder paste Cu particle-reinforced composite solder Ag particle-reinforced composite solder (15 vol.% Ag) Ni particle-reinforced composite solder (15 vol.% Ag) In situ Cu6Sn5 reinforced composite solder (20 vol.% Cu6Sn5) Source: Refs. 41 and 47.

Methods M-Mech. I—in situ N/A M

Average wetting angle after four reflows on Cu substrate 10.5j

M

47.3j (15 vol.%) 18.0j (6 vol.%) 21.1j

M

12.5j

I

17.6j


327

FIG. 13 Scanning electron microscope secondary electron images depicting the rotation of Cu6Sn5 particle in a eutectic Sn–Ag matrix resulting from an off-axis load applied by a nanoindenter. (From Ref. 51.)

rotate about their initial position [51], as illustrated in Fig. 13 for a Cu6Sn5 reinforcement particle. Weak interfaces between Cu6Sn5 reinforcement particles and eutectic Sn–3.5Ag matrix determined by the nanoindentation method are responsible for increasing the ductility of the eutectic Sn–Ag solder, but without increasing the alloy strength significantly. A similar result was also reported for 60Sn–40Pb composite solders containing Cu9NiSn3 intermetallic particles [27], which exhibited a 40% increase in shear strain to cause failure, while the ultimate shear strength essentially remained unchanged. It is suggested that Cu6Sn5 particles in the Sn–Ag composite solder promote homogeneous deformation by creating strain heterogeneities in the interfacial regions due to weak interfacial strength, resulting in an increased ductility for the in situ Sn–Ag composite solder [42].

VII.

SUMMARY

Use of dispersoids is a viable means to improve the properties and service-temperature capabilities of solders to be utilized in microelectronic packages and assemblies. Composite approaches can provide improvements without significantly affecting current solder-joint fabrication practices. Dispersoids must be compatible with the solder matrix material and remain relatively stable when solder joints are in service. Reinforcement materials can be introduced to the solder by either in situ methods or by converting mechanically mixed elemental metallic particles into stable intermetallic compounds. The conversion into intermetallics occurs due to a chemical reaction when the solder is molten during solder joint formation or during reflow. The presence of dispersoids aids in stabilizing solder-joint microstructures by retarding the aging process. All dispersoids tend to improve solder creep strength by several orders of magnitude. An ideal dispersoid should enhance solder ductility without significantly strengthening it. Since the microstructure of solder joints typically is highly inhomogeneous, the deformation within a joint tends to be highly localized. Dispersoids create weakly bonded heterogeneities which promote the initiation of deformation at many locations within a solder joint, resulting in deformation which is much more homogeneous than similar nonreinforced solders. In other words, reinforcements aid in improving the overall ductility of a solder joint. These features render solder joints more compliant by increasing their capacity to accommodate stresses by relaxation while delaying the onset of tertiary creep. These factors can result in a significant enhancement of thermomechanical fatigue resistance of solder joints. Reinforcements introduced by in situ methods appear to offer the best opportunity for achieving this goal.

VIII.

FUTURE

When the current rush to identify suitable lead-free solder candidates to replace Pb-based solders succeeds, attention will be focused on improving their service-temperature capabilities. It is

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anticipated that activities related to lead-free composite solders will increase dramatically since they show promise in providing properties difficult to achieve but necessary for high-temperature applications. Although there is sufficient evidence at present to indicate the potential of these composite solders, they have not received sufficient attention. The technology to produce leadfree solders with dispersoids either by in situ or mechanical means already exists. However, implementation of these solders to produce joints in large-scale manufacturing settings still needs to be developed. These solders conveniently lend themselves to be utilized in either a preform or paste format. Atomization techniques must be developed to convert in situ composite solders into a powder form to make paste. However, in the mechanical mixing method, the particles can be added to the metal powder during paste-making process. These hurdles are manageable, and lead-free composite solders are likely to gain initial importance for solder joints that experience harsh environments, and then perhaps move globally to provide enhanced fatigue resistance in high-reliability applications.

ACKNOWLEDGMENTS Several of the results presented in the section on lead-free composite solders are based on the research activity at Michigan State University by the authors and their colleagues. The authors wish to thank Composite Materials and Structures Center at Michigan State University for financially supporting this effort.

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Con’f on Microstructural and Mechanical Properties of Aging Materials—Proceedings of Minerals, Metals and Materials Society (TMS) Symposium; Chicago, IL, 1993; 417–423. 18. Betrabet, H.S.; McGee, S.M.; McKinlay, J.K. Processing dispersion-strengthened Sn–Pb solders to achieve microstructural refinement and stability. Scr. Metall. Mater. 1991, 25, 2323–2328. 19. Marshall, J.L.; Calderon, J. Hard-particle reinforced composite solders, part 3: Mechanical properties. Solder. Surf. Mount Technol. 1997, 9 (3), 11–14. 20. Satoh, R.; Arakawa, K.; Hadara, M.; Matsui, K. Thermal fatigue life of Pb–Sn alloy interconnections. IEEE Trans. Compon. Hybrids Manuf. Technol. (CHMT) 1991, 14 (1), 224–232. 21. Liu, C.Y.; Chen, C.; Liao, C.N.; Tu, K.N. Microstructure-electromigration correlation in a thin strip of eutectic SnPb solder stressed between Cu electrodes. Appl. Phys. Lett. 1999, 75(1), 58–60. 22. Lead-Free Solder Alternatives, Report by NCMS (National Center for Manufacturing Sciences), Ann Arbor, MI, 1998. 23. Richards, B.P.; Levoguer, C.L.; Hunt, C.P.; Nimmo, K.; Peters, S.; Cusack, P. Lead-Free Soldering; Department of Trade and Industry (DTI): UK, 1999; 39 pp. 24. Richards, B.P.; Levoguer, C.L.; Hunt, C.P.; Nimmo, K.; Peters, S.; Cusack, P. Lead-Free Soldering; Department of Trade and Industry (DTI): UK, 1999; 44 pp. 25. Wasynczuk, J.A.; Lucey, G.K. Shear creep of Cu6Sn5 eutectic composites. Proceedings of the Technical Program, National Electronic Packaging and Production Conference NEPCON WEST: Des Plains, IL, 1992; Vol. 3, 1245–1255. 26. Mavoori, H.; Jin, S. New creep-resistant, low melting point solders with ultrafine oxide dispersions. J. Electron. Mater. 1998, 27 (11), 1216–1222. 27. Betrabet, H.S.; McGee, S. Towards increased fatigue resistance in Sn–Pb solders by dispersion strengthening. Proceedings of the Technical Program, National Electronic Packaging and Production Conference; NEPCON WEST: Des Plains, IL, 1992; Vol. 3, 1276–1277. 28. Marshall, J.L.; Sees, J.; Calderon, J. Microcharacterization of composite solders. Proceedings of the Technical Program, National Electronic Packaging and Production Conference; NEPCON WEST: Des Plains, IL, 1992; Vol.3, 1278–1283. 29. McCormack, M.; Jin, S.; Kammlott, G.W. Enhanced solder alloy performance by magnetic dispersions. IEEE Trans. Compon. Packag. Manuf. Technol. (CPMT), Part A 1994, 17, (3), 452–457. 30. Mavoori, H.; Jin, S. Dispersion strengthening for dimensional stability in low-melting-point solders. JOM 2000, 52 (6), 30–32. 31. Marshall, J.L.; Calderon, J. Hard-particle reinforced composite solders, part 1: Microcharacterisation. Solder. Surf. Mount Technol. 1997, 9 (2), 22–28. 32. Wu, Y.; Sees, J.A.; Pouraghabagher, C.; Foster, L.A.; Marshall, J.L.; Jacobs, E.G.; Pinizzotto, R.F. The formation and growth of intermetallics in composite solder. J. Electron. Mater. 1993, 22 (7), 769– 777. 33. Clough, R.B.; Patel, R.; Hwang, J.S.; Lucey, G. Preparation and properties of reflowed paste and bulk composite solder. Proceedings of the Technical Program, National Electronic Packaging and Production Conference; NEPCON WEST: Des Plains, IL, 1990; 1256–1264. 34. Sastry, S.M.L.; Peng, T.C.; Lederich, R.J.; Jerina, K.L.; Kuo, C.G. Microstructures and mechanical properties of in-situ composite solders. Proceedings of the Technical Program, National Electronic Packaging and Production Conference; NEPCON WEST: Des Plains, IL, 1990; Vol. 3, 1266–1275. 35. Ho, C.T.; Chung, D.D.L. Carbon fiber reinforced tin–lead alloy as a low thermal expansion solder preform. J. Mater. Res. 1990, 5, (6), 1266–1270. 36. Pinizzotto, R.F.; Wu, Y.; Jacobs, E.G.; Foster, L.A. Microstructural development in composite solders caused by long term, high temperature annealing. Proceedings of the Technical Program, National Electronic Packaging and Production Conference; NEPCON WEST: Des Plains, IL, 1999; Vol. 3, 1266–1298. 37. Marshall, J.L.; Calderon, J. Hard-particle reinforced composite solders, part 2: Solderability. Solder. Surf. Mount Technol. 1997, 9 (3), 6–10. 38. Choi, S.; Bieler, T.R.; Lucas, J.P.; Subramanian, K.N. Characterization of the growth of intermetallic interfacial layers of Sn–Ag and Sn–Pb eutectic solders and their composite solders on Cu substrate during isothermal long-term aging. J. Electron. Mater. 1999, 28 (11), 1209–1215. 39. Choi, S.; Lucas, J.P.; Subramanian, K.N.; Bieler, T.R. Formation and growth of intermetallic layers in eutectic Sn–Ag solder and its composite solder joints. J Mater Sci. Mater. Electron. 2000, 11 (6), 497–502. 40. Guo, F.; Choi, S.; Lucas, J.P.; Subramanian, K.N. Microstructural characterization of reflowed and isothermally-aged Cu and Ag particulate reinforced Sn–3.5Ag composite solders. Solder. Surf. Mount Technol. 2001, 13 (1), 7–18. 41. Guo, F.; Lee, J.; Choi, S.; Lucas, J.P.; Bieler, T.R.; Subramanian, K.N. Processing and aging characteristics of eutectic Sn–3.5Ag solder reinforced with mechanically incorporated Ni particles. J. Electron. Mater. 2001, 30 (9), 1073–1082.

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Subramanian, K.N.; Bieler, T.R.; Lucas, J.P. Microstructural engineering of solders. J. Electron. Mater. 1999, 28 (11), 1176–1183. 43. Choi, S.L.; Gibson, A.W.; McDougall, J.L.; Bieler, T.R.; Subramanian, K.N. Mechanical properties of Sn–Ag composite solder joints containing copper-based intermetallics. Design Reliability of Solders and Solder Interconnections—Proceedings of Minerals, Metals Materials Society (TMS) Symposium; Orlando, FL, 1997; 241-245. 44. Choi, S.; Lee, J.G.; Guo, F.; Bieler, T.R.; Subramanian, K.N.; Lucas, J.P. Creep properties of Sn–Ag solder joints containing intermetallic particles. J. Miner. Metals Mater. Soc. 2001, 53 (6), 22–26. 45. McDougall, J.L.; Choi, S.; Bieler, T.R.; Subramanian, K.N.; Lucas, J.P. Quantification of creep strain distribution in small crept lead-free in situ composite and non composite solder joints. Mater. Sci. Eng. 2000, A285, 25–34. 46. Guo, F.; Lucas, J.P.; Subramanian, K.N. Creep behavior in Cu and Ag particle-reinforced composite and eutectic Sn–3.5Ag and Sn–4.0Ag–0.5Cu non-composite solder joints. J. Mater. Sci. Mater. Electron. 2001, 12 (1), 27–35. 47. Guo, F.; Choi, S.; Lucas, J.P.; Subramanian, K.N. Effects of solder reflow on wettability, microstructure and mechanical properties. J. Electron. Mater. 2000, 29 (10), 1241–1248. 48. Guo, F.; Lee, J.; Choi, S.; Lucas, J.P.; Subramanian, K.N.; Bieler, T.R. Creep properties of eutectic Sn–3.5Ag solder joints reinforced with mechanically incorporated Ni particles. J. Electron. Mater. 2001, 30 (9), 1222–1227. 49. Jadhav, S.; Subramanian, K.N.; Bieler, T.R.; Lucas, J.P. Stress relaxation of tin–silver solder joints. J. Electron. Mater. 2001, 30 (9), 1197–1205. 50. Gibson, A.W.; Subramanian, K.N.; Bieler, TR. Comparison of mechanical fatigue fracture behavior of eutectic Sn–Ag solder with and without Cu6Sn5 intermetallic particulate reinforcement. J. Adv. Mater. 1998, 30 (2), 19–24. 51. Lucas, J.P.; Gibson, A.W.; Subramanian, K.N.; Bieler, T.R. Nanoindentation characterization of microphases in Sn–3.5Ag eutectic solder joints. Proc. Mater. Res. Soc. Conf. 1998, 522, 339–345; (Pittsburgh, PA). 52. Mavoori, H.; Chin, J.; Vaynman, S.; Moran, B.; Keer, L.; Fine, M. Creep, stress relaxation, and plastic deformation in Sn–Ag and Sn–Zn eutectic solders. J. Electron. Mater. 1997, 26 (7), 783–790.

11 Solder Wetting and Spreading Timothy J. Singler State University of New York at Binghamton, Binghamton, New York, U.S.A.

Stephan J. Meschter BAE Systems, Johnson City, New York, U.S.A.

James Spalik IBM Corporation, Endicott, New York, U.S.A.

I.

INTRODUCTION

During the Second World War, a situation occurred—in retrospect rife with irony given the current subject of this book—when wartime demands for tin (Sn) led to an effort to develop Snfree solder alloys. The results of this effort were not entirely successful and produced lead (Pb)rich alloys that exhibited poor wetting properties in comparison to their Sn–Pb predecessors that were Sn-rich. A deep appreciation developed for the excellent wetting characteristics with which Sn endows many solder alloys. Now, some 60 years later, a markedly different imperative has served to initiate an effort to identify a set of candidate Pb-free solder alloys capable of replacing the existing Pb–Sn alloy set. This effort is spurred by health concerns over Pb in the waste stream and by economic concerns that those with Pb-free microelectronics products would attain an edge in the market place based on the environmentally benign character of their products. The Pb–Sn system is one of the most extensively characterized low-melting-point metal systems because of its longstanding success in soldering of microelectronic components. This alloy system has a long history based on its adaptability to increasingly sophisticated soldering processes and demanding soldering applications. Current industry processes use the Pb–Sn eutectic alloy (63Sn–37Pb) and higher Pb content alloys (e.g., 90Pb–10Sn, 97Pb–3Sn, etc.) extensively. The popularity of the Pb–Sn system grew out of its many desirable material properties [e.g., low melting point, limited dissolution of copper (Cu) substrates, etc.]. Recently, a variety of Pb-free alternative alloys have been investigated (e.g., Sn–Ag, Sn–Ag–Cu, Sn–Cu, Sn–In, Sn–Sb, Sn–Bi, and Sn–Zn) [1]. There is a general consensus that a single-alloy system may not have sufficient versatility to serve as a basis for drop-in replacements for the present Sn–Pb alloys, particularly eutectic Sn–Pb. However, one characteristic of the new Pb-free solder formulations that has emerged with some measure of certainty is that of high-Sn compositions.

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II.

CONCEPTS OF WETTING

A.

Definition of Wetting

The term wetting is generally applied to the fluid dynamic process of a liquid displacing another fluid (gas or liquid) along a solid surface (Fig. 1). The locus of points where solid, liquid, and fluid intersect is referred to as the contact line. Wetting is generally described by the position of the contact line, its rate of advance, and the contact angle, the local angle formed at the contact line between the tangents to the liquid–fluid and liquid–solid interfaces and measured within the liquid. General categorizations of wetting performance are best illustrated by considering a small drop of liquid in a state of equilibrium on a flat, horizontal surface (Fig. 1). A liquid is said to be nonwetting when hE > 90j (Fig. 1a) and to wet the solid if the equilibrium contact angle hE80

Bridge

D

D

D

D

Sn–2.5Ag– 5Bi–0.5Cu

Sn–0.7Cu

Sn–3.5Ag– 0.7Cu Sn–37Pb O

O

O

O

Plating compatibility

O

D 10%

O

D 20%

Void

0 = preferred. D = Marginal, O = acceptable, and O Courtesy of Panasonic.

Solder

Wett (% land coverage)

D

D

D equip dep

O

Icicle

TABLE 8 Summary of the Wave Soldering Evaluation Results

O

O

O

O

Overall bond strength

JW—O QFP—O JW—D QFP—O

JW—O QFP—O

JW—O QFP—D

Fracture mode

O

D

O

0 O

Crack, % elongation

JW—O QFP—O JW—D QFP—O

JW—O QFP—O

JW—O QFP—D

Strength

Bond strength and thermal aging

D 2/9 top side 0 O

X 9/9 both sides D 1/9 top side only

Fillet lifting

0 O

D Cu concentration dross D

D 0.2% Pb

Impurity res.

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Pb-FREE COMPONENT CONVERSION

621

TABLE 9 Comparison of the Joint Strength of Several Lead-Free Alloys and the Benchmark Eutectic SnfPb Solder Wave soldering alloy Snf2.5Agf5Bif0.5Cu Snf0.7Cu Snf3.5Agf0.7Cu Snf37Pb

Melting point (jC)

QFP pull strength (kgf)

Jumper wire pull strength (kgf)

Elongation (%)

199f218 227 216f219 183

1.5f3.7 1.8f3.4 2.0f3.8 2.3f3.3

6.1f7.7 4.4f6.9 5.1f7.4 3.8f5.1

23 57 49 59

Courtesy of Panasonic.

b. Monitoring Bath Impurity Levels. As noted earlier, fillet lifting is influenced by impurity levels. Also, impurity levels can have a big influence on the strength of solder joints formed by some lead-free wave solder baths, thus an important consideration. In general, realtime monitoring and impurity control are not necessary with Sn–Pb baths as they can accommodate larger quantities of impurities. However, simple bath maintenance regime can easily be adopted to manage the control of impurities. Panasonic determined the standard practice of controlling a wave solder bath by simply replenishing solder was impractical and costly in the case of lead-free solders. Therefore a real-time monitoring system was developed to measure impurity concentrations in the bath on the production floor. The monitor, developed in cooperation with Malcom Industries, operates on the principle of differential temperature analysis or DTA [8]. Fig. 52 illustrates the system in operation over the range of 0% to 0.3% Pb impurity. c. Real-Time Control of Impurities. Panasonic has also phased in its manufacturing operations a three-zone refining process as a method to control impurity levels in real time. The method uses a technique, very similar to the methods used in IC crystal growing, whereby a controlled cooling and heating rate concentrates the impurities into a low-melting point alloy [9]. The process is based upon controlling the rate at which the solid interface traverses across a billet of solder (Fig. 53). d. Principle of Zone-Refining. As the metal alloy solidifies, the solute or impurities extrude from the crystal and the excess solute disperses into the solvent or liquid tin. Fig. 54 illustrates the diffusion of the liquid phase. Here diffusion is considered only vertical to the solidification interface and can be ignored into the solid area. The amount of solute flowing ahead of the solid interface is given by, D d2 C=dx2 ðcm2 =secÞ

TABLE 10

ð1Þ

Summary of Wave Solder Alloy Evaluations

Lead-free solder attribute Wetting Melting point Fillet lifting Single alloy choice Courtesy of Panasonic.

Summary Reduced for all lead-free materials sampled Significantly higher than that of eutectic Sn–Pb (of those evaluated) Can be an issue with lead-free solders, especially with bismuth containing solders (in general). No single alloy proved best in all categories

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FIG. 37 Thermal cycle tests for a 0.8-mm pitch QFP using several lead-free wave solder alloys. Thermal cycling conditions were 40jC to 85jC for 30-min dwell times. Electrode finishes tested were a) Sn–Pb finish, b) Sn–Bi finish, and c) Pd finish. The Sn–Ag–Cu lead-free alloy is generally observed to be the most resilient to electrode finish materials, outperforming even the eutectic Sn–Pb solder in most cases. (From Ref. 24)

where D is the coefficient of diffusion in liquid phase. Using R (cm/sec) as the interface migration speed, the steady state is given as, D d2 C=dx2 þ R dC=dx ¼ 0

ð2Þ

Solving Eqs. (1) and (2) yields, CL ¼ Ca eðRXV=DÞ þ C0

ð3Þ

where Ca is the excessive concentration of liquid phase at the interface, C0 is the initial composition, XV is the distance from the interface, and CL is the concentration at XV. The distribution coefficient (k0) is the ratio of the solid phase concentration (C0) to the liquid phase concentration (CL), k0 ¼ C0 =CL

ð4Þ

As the state of the solid is C0 at some distance from the interface, CL can be stated as CL ¼ C0 þ Ca

ð5Þ


623

FIG. 38 Thermal cycle tests for a single-sided jumper wire using several lead-free and eutectic Sn–Pb wave solder alloys. Thermal cycling conditions were 40jC to 85jC for 30-min dwell times. Wire plating tested were a) Sn–Pb, b) Sn–Bi, and c) Sn. With jumper wires, the Sn–Ag–Bi wave solder outperforms the remaining materials over a wide range of finishes. (Courtesy of Panasonic.)

Using Eq. (5), Eq. (4) can be rewritten to be, Ca ¼ C0 =k0 C0

ð6Þ

From Eq. (6), Eq. (3) can be rewritten to provide the concentration profile of the liquid phase in front of the solid interface in steady state (Fig. 55) [9], h i ð7Þ CL ¼ C0 1 þ fð1 k0 Þ=k0 geðRXV=DÞ The concentration profile for the solid phase after the interface in steady state is represented as [10], n h i o k RXV=D Cs ¼ C0 ð1 k0 Þ 1 e0 þ k0 ð8Þ A graphical representation of Eqs. (7) and (8) can be seen in Figs. 56 and 57. As the interface migration speed is reduced, the wider the clean or purified area becomes. The above discussion provides the theoretical basis for zone refining, a method used to remove impurities through controlled solidification and segregation.

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FIG. 39 Comparison of the wettability of several lead-free solder baths and the eutectic Sn–Pb benchmark solder for various chip components and finishes. The lead-free alloys each show varying degrees of wettability with the various component electrode finishes, whereas the eutectic Sn–Pb solder completely wet for each test condition. (Courtesy of Panasonic.)

FIG. 40 Comparison of jumper wire wettability for various lead-free alloys and eutectic Sn–Pb as the benchmark. Eutectic Sn–Pb wet 100% of the land (terminal pad) while all the lead-free alloys exhibited a reduced level of coverage. (From Ref. 25.)


625

FIG. 41 Wetting balance results for eutectic Sn–Pb and two lead-free alloys. The test consists of monitoring the time required to change the force exhibited on a coupon from compressive (solder is pushing on the coupon) to a tensile force (solder is pulling on the coupon). This change in force direction (or vector) is called the ‘‘zero force time.’’ (From Ref. 26.)

FIG. 42 Wave soldering results for Sn–Ag–Bi–Cu alloy that show the effect of increasing the wave temperature or reducing the conveyor speed to improve wetting. As the wave temperature is increased from 245 to 270jC, the percent coverage increases. In addition, as the wave speed is reduced from 1.0 to 0.75 m/min, the percent coverage is increased while the bath temperature is kept constant at 250jC. (Courtesy of Panasonic.)

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FIG. 43 Wave soldering results for Sn–0.7Cu alloy that show the effects of increasing the solder wave temperature or reducing the conveyor speed to improve wetting. As the temperature of the bath is increased or the speed of the PCB is slowed, the percent coverage of the terminal pad was observed to increase. (Courtesy of Panasonic.)

e. Verification of Zone Refining of Lead-Free Solders. Panasonic conducted tests to determine the optimal rate, billet length, and other parameters to utilize the zone refining technique to remove impurities from lead-free solder. Fig. 58 shows the system used to evaluate the Sn–3.0Ag–0.5Cu solder alloy that was initially contaminated with 0.18% Pb. As the solder was processed in the zone refining system (Fig. 58), the Pb impurity was reduced in the solidified solder to a maximum of 0.07% when the migration rate was set to the slowest rate. At faster rates, the impurity levels were higher (from 0.16% to 0.18% at a rate of 1jC/5 min). The results of migration speed and impurity removal efficiency are shown in Fig. 59. Also, as predicted by Eqs. (7) and (8), the concentration levels in the liquid phase increased to a higher level ranging from 0.24% to 0.71% Pb depending on the solidification speed. In addition to varying the migration speed, the desired solder purification level can also be achieved utilizing a multipass method. In Fig. 60, a Sn–3.0Ag–0.5Cu alloy is purified at a high rate of migration (1jC/5 min) but with multiple passes to achieve the same effect as fewer or a

FIG. 44 Frames captured from video observations of Sn–3.5Ag solder of a 14-pin DIP IC during the soldering process. (From Ref. 27.)


627

FIG. 45 Illustration of the lift-off phenomena sequence. As the solder solidifies, a low-melting-point alloy forms near the surface of the land (terminal pad) just under the solidified solder. This lowermelting-point material forms a crack near the surface and allows the solidifying solder on the outside to pull up and away from the land. (From Ref. 27.)

FIG. 46 Change in solid phase percentage in solidification of Sn–0.8Cu–0.1Pb–(0–0.6)Bi. The higher the concentration of impurities, the longer the liquid phase or the lower the melting point (calculated with Sheil’s equation). (From Ref. 28.)

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FIG. 47 As the impurity level due to a) lead (Pb) and b) copper (Cu) increases, the lift-off rate was observed to increase for the lead-free Sn–0.7Cu solder. (From Ref. 29.)

single pass at slower rates. As the material is repeatedly purified in the zone refining system, the impurity levels continue to drop (Fig. 60). Fig. 61 shows the factory system implemented in production testing. The results of the initial trials show that a 150-kg Sn–3.0Ag–0.5Cu solder bath can be purified from a 0.25% Pb to 0.18% Pb in about 10 days. The cycle rate is approximately 30 kg every 2 days with a bath being turned over every 10 days (five turns). The discharge concentration was 0.8% Pb. With this system, the bath level can be maintained continuously at a level of less than 0.2% Pb [11]. f. Practical Examples of Wave Soldering. With a Sn–0.7Cu wave, a high dross rate and potential tin whisker formation and growth are a concern. Therefore trace amounts of Ni are added to the Sn–0.7Cu eutectic bath to reduce the drossing rate and eliminate the Sn whiskering problem in this bath (Fig. 62). This Sn–Cu-based bath from Nihon Superior, called SN100C, is very stable with makeup solder-type SN100Ce. It is widely used by Panasonic using a bath temperature set at 255jC for products such as video tape recorders (VTR) and DVDs. A Sn–3.0Ag–0.5Cu solder from Senju (M705) is utilized at Panasonic’s TV factory for both wave and selective soldering baths (Figs. 63 and 64). It required approximately 4 months to integrate lead-free soldering and completely phase-out the use of Sn–Pb solders. The wave solder bath is maintained using a simple temperature probe to monitor the solder level. When the bath level drops below the probe height, an automatic wire feeder is energized to feed in the solder wire

FIG. 48 The incidence of fillet lifting can be significantly reduced by reducing the size of the land on the top of the PCB. (Courtesy of Panasonic.)


629

FIG. 49 Several changes were made to the wave solder nozzles to improve lead-free soldering. From a), the distance between the turbulent and laminar flow sections was reduced from 100 to 50 mm, and b) the multihole pattern was changed from a two-row design to a three-row design. (Courtesy of Panasonic.)

of a Sn–3Ag+0.008P (Senju M708) solder (Figs. 65 and 66). The typical bath life of the Sn–Ag– Cu bath is about 3 weeks depending on production volume.

G.

Summary and Conclusions

1. Successful Entry to Lead-Free Manufacturing With the introduction of lead-free products such as the MiniDisk players, Panasonic has gained broad experience with lead-free soldering. From this and other products, Panasonic learned it is possible to use lead-free solders in consumer electronics manufacturing without sacrificing quality or production volumes. The quality levels were noted to have improved with the MiniDisk

FIG. 50 Photographs of the wave nozzles. (a) Showing the overall view of the two nozzles and (b) a close-up view of the mountain-like shape of the multihole patterned nozzle. Further enhancements to the nozzle design include a four-hole pattern (not shown). (Courtesy of Panasonic.)

630

STALTER ET AL.

FIG. 51 Series of photographs that show the superior wetting of the new nozzle design compared to the standard nozzle in both air and in an inert atmosphere. Here, a) are photographs of plated through hole without any component lead, and b) a plated through hole with a component lead present. (Courtesy of Panasonic.)

player even as the lead-free system (Sn–Ag–Bi–In) was in its initial stages of manufacturing compared to production with Sn–Pb (Fig. 67). Similar results were obtained during VTR production utilizing Sn–0.7Cu wave solder. Here quality levels saw an increase in defect levels after the introduction of lead-free solder. However, after improvements such as nozzle design enhancements, the defect level was reduced to one-half the eutectic Sn–Pb solder level (Fig. 68). Panasonic has experienced similar results when lead-free solder assembly was phased in for TV production. The defect levels dropped from a historical rate of 3% with eutectic Sn–Pb to a 0.07% level with Sn–3.5Ag–0.5Bi–8In solder paste and Sn–3.0Ag–0.5Cu wave solder. a. Many Challenges. Success was only possible after a significant number of challenges were addressed, some of which persisted into the manufacturing stage. For example, the Panasonic factory at Utsunomiya worked for more than 1 year to utilize Sn–3Bi–8Zn solder paste on various products. This solder alloy has proven difficult to work with because of its

FIG. 52 A real-time monitor based on DTA techniques was developed in conjunction with Malcom Industries to automatically monitor the impurity levels in a lead-free wave soldering bath. a) The monitor indicates the b) initial melting point of the solder bath alloy and c) lower melting point due to lead (Pb) contamination. (Courtesy of Panasonic and Malcom Industries.)


631

FIG. 53 As the molten area migrates down the sample, the level of impurities increases in the molten or liquid phase, while the impurity level decreases in the solid phase that had previously been heated. The technique is based on a well-known concept referred to as zone refining. (Courtesy of Panasonic.)

shorter stencil life and intermittent nonwetting defects (voids) to ball grid array (BGA) balls consisting of Sn–Ag–Cu [12]. Therefore the Utsunomiya factory switched over to the bismuth and indium containing solder alloy in early 2003. 2. Monitoring Field Performance Figs. 69, 70, and 71 show that the use of lead-free solder (Sn–Ag–Bi–In) has not resulted in reliability problems. However, some joints have indicated a concern with the potential precipitation of bismuth from the alloy over time. Therefore the bismuth content of the solder used to assemble the MD player will be reduced from 2.5% bismuth to 0.5% bismuth with 6% indium.

FIG. 54 An illustration of the diffusion principle of the solute (impurity) into the solvent (molten tin). The concentrations of the impurities increase at the interface. (Courtesy of Panasonic.)

632

STALTER ET AL.

FIG. 55 The concentration profile as described by Eqs. (3), (4), (5), and (6). Immediately after solidification starts, when the interface is moved slightly, the impurities in the solid phase are decreasing and the impurities are increasing in the liquid phase. (From Ref. 30.)

3. Future Lead-Free Products Further reductions in the melting point are being realized with the Sn3.5Ag8.0In0.5Bi alloy (Fig. 72). This alloy is geared toward desktop and notebook computer assembly applications. It is anticipated that rolling this out to the manufacturing floor will pose some significant challenges. Among these may be modifications to existing wave soldering equipment and new reflow oven designs. Some ‘‘tweaking’’ of AOI equipment may be necessary to compensate for the duller or matte appearance characteristic of these solders. 4. Cost Material cost was another issue of concern. With the introduction of lead-free, economies of scale worked against the MD player product. No other manufacturer was using a lead-free solder in significant volume. When first introduced in 1998, the cost of the lead-free solder paste was about $150/kg. Then in 2000, it dropped to $100/kg. By 2003, the cost of the lead-free solder paste was only marginally higher than eutectic Sn–Pb with the downward trend continuing.

FIG. 56 Result of Eq. (8), the concentration profile of the solid phase of a binary, Sn–0.5% Pb alloy at various solidification migration speeds, R of 1.2, 3.0, and 6.0 mm/min. At the slower rates, the level of impurities (Pb) drops from the solid tin closer to the interface. The distribution coefficient, k0, is 0.074, the coefficient of diffusion, D, is 5105 (cm2/sec). (Courtesy of Panasonic.)


633

FIG. 57 Result of Eq. (7), the concentration profile of the liquid phase of a binary, Sn–0.5% Pb alloy at various solidification migration speeds, R of 1.2, 3.0, and 6.0 mm/min. At the slower migration speeds, the impurity levels are concentrating at the solid-to-liquid interface more heavily than at faster migration speeds. (Courtesy of Panasonic.)

IV.

OTHER EXAMPLES OF INDUSTRY PRODUCT CONVERSIONS

A.

Flextronics International Ltd. Lead-Free Assembly

Flextronics International Ltd., with worldwide locations providing electronic manufacturing services, perceived early on the importance of environmentally friendly manufacturing. Flextronics began the journey to lead-free manufacturing in 1997 and as of 2000 had high-volume Pbfree production of 150,000 units per month across several products [33]. As of November 2002, they had shipped several million products with lead-free solder including cell phones, printers, and computer security devices, and have several factories in Asia, the Americas, and Europe using lead-free solder in volume manufacturing [34].

FIG. 58 A laboratory scale system of the zone refining equipment used in the Environmental Production Engineering Lab of Kadoma, Japan, to verify and evaluate the feasibility of the technique. (Courtesy of Panasonic.)

634

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FIG. 59 Testing of the zone refining process by purification of Sn–3.0Ag–0.5Cu solder with an initial state of 0.18% Pb contamination. As the solidification rate is reduced, the concentration of the Pb remaining in the purified solder is reduced more than with higher migration rates. (From Ref. 31.)

1. Considerations Flextronics separated the issues surrounding the conversion to lead free into three categories: impact to the process, impact to equipment, and cost impact. Using a test vehicle which included QFPs with pitches ranging from 0.3 to 0.5 mm, and CSPs with pitches ranging from 0.4 to 0.6 mm, a study was conducted to understand the manufacturing impact of three no-clean, lead-free solders from the Sn–Ag–Cu family compared to a standard no-clean eutectic Sn–Pb solder paste. The evaluation focused on printability, solder paste pot-life, wettability, reflow process window, and inspection [33].

FIG. 60 As the zone refining process is repeated, the impurity levels continue to concentrate into the liquid phase further reducing the contaminate in the solidified material. Here the impurity levels are reduced in a Sn–3.0Ag–0.5Cu alloy from 0.17% Pb to 0.07% Pb in five passes at a rate of 1jC/5 min. With a single pass and at the same rate of 1jC/5 min, the Pb level is only reduced to 0.11% Pb. (Courtesy of Panasonic.)


635

FIG. 61 Photograph of the zone refining system utilized in mass production scale-up trials. The Pb concentration level was maintained below the desired impurity levels of 0.2%. (From Ref. 31.)

a. Printability. Variables such as printing pressure, printing speed, and squeegee separation speed were first determined. It was then concluded from previous evaluations [35] that the ratio of stencil thickness to aperture opening, or so-called area ratio, was the critical parameter affecting print quality. Solder paste height measurements were plotted as a function of the area ratio, and comparisons of the lead-free solder alloys to eutectic Sn–Pb revealed that the area ratio of 0.5 must be achieved. It is above this value that the solder paste height was observed to remain consistent, but below a ratio of 0.5 height decreased linearly as a function of decreasing area ratio, independent of alloy (Fig. 73) [33].

FIG. 62 Additions of trace amounts of Ni to the Sn–0.7Cu bath have shown to dramatically improve the wave solder bath by reducing the dross rate, stabilizing the wave shape and height, and eliminating whisker defects. (From Ref. 24.)

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FIG. 63 Photograph of the selective soldering machine used at Panasonic’s TV factory in Utsunomiya, Japan. (Courtesy of Panasonic.)

b. PCB Surface Wettability. An experiment was conducted by Yi et al. to evaluate the wetting of several different solder pastes on different surface finishes as compared to a Pbcontaining paste [33]. HASL (Sn–Pb), OSP (Entek), and Ni/Au surface finishes were utilized. A 6mil-thick stencil with 22-mil round openings was utilized. The solder was screened and reflowed and the solder wetting diameter measured and cross sectioned (Fig. 74). It was concluded by the authors that the Ni/Au finish had the best spread and smallest wetting angle, while the Entek OSP

FIG. 64 Photograph of a TV tuner assembly where large subassemblies and odd-shaped connectors are soldered automatically in the selective soldering machine. (Courtesy of Panasonic.)


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FIG. 65 Photograph of Panasonic engineers demonstrating the lead-free wave soldering system at the TV factory in Utsunomiya, Japan. (Courtesy of Panasonic.)

FIG. 66 Photograph of a system that maintains the wave solder bath level by integrating a temperature probe to the wire feeder. As the bath lowers, the temperature of the probe is reduced. When the temperature is lowered to a predetermined level, the wire feeding mechanism is energized to feed makeup wire Sn–Ag+0.008P into the bath to maintain level. (Courtesy of Panasonic.)

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FIG. 67 Upon introducing lead-free soldering at the minidisk player factory, an improvement in manufacturing quality was observed immediately as noted by the comparison of a typical month of eutectic Sn–Pb production and the first three months utilizing lead-free soldering. (From Ref. 32.)

FIG. 68 Upon introducing lead-free soldering at the VTR factory, defect levels were significantly higher than eutectic Sn–Pb. However, after instituting several improvements to the wave soldering equipment, the defect rate of lead-free wave soldering was reduced to one-half that of eutectic Sn–Pb. The enhancements are: step 1, raising the preheat temperature from 100 to 125jC; step 2, use of the multihole patterned nozzle design; and step 3, reducing the pitch between nozzles. (Courtesy of Panasonic.)


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FIG. 69 Photograph of a minidisk player assembly that was returned from the field after having experienced approximately 3000 hr of operation. The joints were analyzed to determine the long-term effects on the solder alloy. (Courtesy of Panasonic.)

performed the worst of the three, having the smallest spread area and the largest wetting angle. Among the several Pb-free solder pastes evaluated, none of them individually performed the best across all of the surface finishes (Fig. 75). c. Reflow Process Window. To evaluate the process window during reflow of the Pb-free solder pastes, peak temperatures of 220jC, 235jC, and 245jC were tested. The highest peak temperature resulted in the formation of a dark residue on the joint suggesting charring of the flux in the paste. At 220jC, insufficient wetting was observed. The authors concluded that although a F10jC process window is desirable and suitable for eutectic Sn–Pb solder, a tighter process window is required when forming solder joints with Pb-free solder pastes.

FIG. 70 SEM cross sections of the QFP depicted in Fig. 69. a) Overview of a lead that experienced 3000 hr of operation in the field. b) An enlargement of the area designated as A, showing the lead/ solder interface. c) An enlargement of the area designated as B, showing the solder/land (terminal pad) interface. No problems were noted in relation to the Sn–Ag–Bi–In alloy. (Courtesy of Panasonic.)

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FIG. 71 SEM cross sections of the SOP depicted in Fig. 69. a) An overview of a lead that experienced 3000 hr of operation in the field. b) An enlargement of the area designated as A, showing the lead/ solder interface. No problems were observed. c) An enlargement of the area designated as B, showing the solder/land (terminal pad) interface. Although this particular cross section shows no observed issues, it has been noticed that a slight amount of elemental Bi was observed to have precipitated from the Sn–Ag–Bi–In solder alloy in some cases. (Courtesy of Panasonic.)

d. Conclusion. The result of these and several other studies conducted at Flextronics [33,34] resulted in the successful implementation of a lead-free soldering process capable of volume production. Although there is not an extensive database of experience on Pb-free soldering, companies are performing the necessary evaluations, and implementation of Pb-free assembly processes are slowly becoming pervasive in the electronics industry.

B.

Visteon and the Challenge of Lead-Free in Automotive Systems

Implementing lead-free assembly processes in the automotive industry gives rise to unique requirements that must be met for successful product introduction. The most challenging

FIG. 72 Further reductions in melting point coupled with reductions in Bi content are the future trends for lead-free alloys with Panasonic products that use the Sn–Ag–Bi–In system. (Courtesy of Harima Chemicals.)


641

FIG. 73 Solder paste height vs. area ratio (AR) for a) Pb-free and Sn–Pb solder alloys. (From Ref. 33.)

application environments necessitate very stringent electronic product demands. For example, many products must be capable of high thermal resistance because of the location of the parts positioning in the engine compartment resulting in thermal-induced strains and material degradation. Depending on season and geography, the electronics in a vehicle may experience very severe temperature fluctuations undergoing approximately a 200jC change in a matter of minutes and exposure to high humidity conditions as well. Also, differences in road conditions give rise to engine vibration conditions in addition to the environmental challenges as noted. Visteon Automotive Systems has performed extensive engineering analysis on Pb-free alloys since the early 1990s. Their objective was to launch the first lead-free soldering process into volume production on a commercial automotive electronics product [36]. 1. Alloy Selections Considerations Given the wide selection of lead-free solder alloys with higher processing temperatures than eutectic Sn–Pb solder, a choice required a solder that had equivalent or better mechanical properties, adequate availability, and one with suitable soldering characteristics. The focus was

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FIG. 74 Cross sections depicting wetting angles of Pb–Sn (type A) and lead-free solders (types B–D) as a function of board finish. (From Ref. 33.)

on developing a no-clean process. The tin–silver binary eutectic (96.5Sn/3.5Ag) was chosen based upon metallurgical stability, mechanical properties, and widespread availability in paste, bar, and wire form. 2. Product Selection Several criteria were evaluated in selecting a product for conversion, as illustrated in Table 11. Each criterion was weighted and the best candidate with the simplest design and manufacturing process was chosen (Fig. 76). Table 12 provides the details of the selected product. 3. Process Development Immersion silver plating was chosen as the board surface finish because of the consistency in plating thickness and surface topography. The major unknown that had to be evaluated was the suitability of the surface finish. Because there is a significant schedule delta among component


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FIG. 75 Bar-graph illustrating the wetting diameter of Pb–Sn alloy (type A) and lead-free alloys (types B–D) as a function of board finish. (From Ref. 33.)

manufacturers in their conversion to lead-free surface finishing, mixed component finishes were utilized. The product had to meet the existing reliability standard [36]. In choosing a part with the simplest reflow process, Visteon chose a single-sided assembly requiring only one pass through the reflow furnace. A tin–silver eutectic paste was evaluated with several no-clean flux formulations. A temperature profile similar to the eutectic Sn–Pb soldering profile was utilized consisting of a preheating stage, a temperature soak, and a temperature spike. Solder wetting characteristics, component skewing, extraneous solder balls, component tombstoning, PCB warpage, and PCB delamination were evaluated. First attempts at lead-free soldering revealed an increased occurrence of solder ball formation between the component

TABLE 11 Visteon Automotive Product Selection Criteria for Pb-Free Conversion Candidates Selection criteria . . . . . . . .

Perceived customer interest Low consumer risk product Product engineering office acceptance Long production life Less product design complexity Low manufacturing process complexity Medium-to-high production volume Low line mix

Source: From Ref. 36.

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FIG. 76 Photo of Visteon’s lead-free conversion product. a) Photo depicting panel with 56 array of cards. b) Single card fully populated with components. (From Ref. 36.)

terminations, as well as tombstoning of the 0805 capacitors. Solder ball formation was attributed to the print characteristics utilizing a common stencil resulting in more solder volume being deposited with the lead-free solder paste compared to the eutectic Sn–Pb solder paste. It was concluded that the conventional dual-ramp reflow profile was not satisfactory for lead-free processing this product as it increased the incidence of tombstoning. Paste printing parameters, in conjunction with profile modifications, were made to improve solder joint integrity, reduce the occurrence of solder ball formation, and decrease component tombstoning [36]. 4. Testing It was noted that the penetration by the test probe in the bed-of-nails tester resulted in less penetration into the immersion Ag as compared to penetration into the Sn–Pb HASL finish, as shown in Fig. 77(a). A higher spring force and a three-sided chisel probe were chosen to improve

TABLE 12

Visteon Selected Product Attributes

Variable . . . . . . . . . . . .

Product application Printed circuit board material Panel dimension Panel thickness Number of cards per panel Card size Number of copper layers Track width/spacing Surface finish Component count Component types Solder joint count


Product attribute Automotive electronic module FR4 194mm 214mm 0.062 in. 30 cards 30mm 30mm two layers 0.008 in./0.008 in. Sn/Pb HASL 33 SMD per card, single-sided application SOIC, 0805, and 1206 chips, SOD, SOT, crystal 92 per card


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FIG. 77 In-circuit testing on immersion silver pads. a) SEM of a low-force spear head test-probe indentation. b) SEM of a high-force three-sided test-probe indentation. (From Ref. 36.)

on yields and eliminate false rejections. Typical indentation marks with the three-sided probe are shown in Fig. 77(b). 5. Qualification Rigorous testing was performed which included accelerated environmental stress testing at each level of assembly. The testing was designed to simulate 10 years/150,000 miles endurance. [36] 6. Final Steps to Production In addition to passing the electrical, mechanical, and reliability testing, there were other considerations by the Visteon team that had to be addressed prior to launching the product. Close control of production would be required as products in the line switched between the standard lead-containing process and lead-free processing. The hardware and materials would have to be distinguishable to the production operators to insure a mix-up did not occur. To mitigate these risks, Visteon worked with their paste supplier to visually change the packaging of the cartridge with a colored band to represent the lead-free paste. In addition, two identification holes, pictured in Fig. 78, were added to the PCBs, that are automatically detected at each subassembly step by optic sensors. Three check gates were set up within the surface mount assembly line to perform the checking and verification. Operator training was also key to adding precaution to the automatic checks with respect to visual identification, lead-free solder joint criteria, and handling of the materials to prevent contamination [36]. 7. Summary Volume production of the lead-free assembly has been underway since June 1, 2001. Visteon produces about 7000 units a day with lead-free processing. The next level of complexity for future products will incorporate double-sided reflow, selective soldering, and wave soldering process implementation. Visteon received approval from Ford Motor Company to manufacture and supply its Passive Anti-Theft System transceiver module using lead-free solder [37].

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FIG. 78 Photo of a printed circuit board (PCB) with lead-free identification holes. (From Ref. 36.)

C.

IBM and the Conversion of High-Complexity SMT Components to Lead-Free Technology

One of the biggest advantages to complex components that are lost with the conversion to leadfree processing is the ease of incorporating a solder hierarchy that is critical to many interconnection solutions. IBM uses a dual melt solder termination system for its ceramic column grid array (CCGA) components. Solder columns are used as the interconnection between the ceramic chip carrier and PCB. They provide a method of accommodating the strain induced due to the mismatch in coefficient of thermal expansion between the carrier and the board. This is an important feature, especially as the ceramic carrier increases in size, thus increasing the distance to neutral point [38, 39]. In addition to providing a method of attaching the component to the board, the columns also provide the electrical pathway between the chip carrier and the PCB. The solder columns fabricated by IBM are one of three configurations: a high lead (90Pb/10Sn) column attached to both the chip carrier and PCB with a low-melt (63Sn/37Pb) solder fillet; a high lead (90Pb/10Sn) column ‘‘cast’’ to the chip carrier by melting the column directly onto the NiAu pad of the ceramic chip carrier; and a column last attach solder process (CLASP) [40]. In the latter case the high melt column is attached last, or post chip join and underfill, to the chip carrier with a 63Sn/37Pb solder doped with palladium such that after reflow the solder mixes and forms a higher melting alloy (Fig. 79). In order to replicate this structure in a lead-free configuration, it would be necessary to find a suitable lead-free alloy in the 300jC range to provide the solder hierarchy necessary to prevent the collapse of the connection during PCB assembly. 1. CCGA Replacement Criteria In addition to replicating the mechanical properties of the column such that the thermal fatigue resistance is equivalent to the existing CCGA versions described above, it is critical that the interconnections have equivalent electrical performance. Manufacturability aspects must also be considered to assure that the columns can withstand the higher temperatures required of lead-free processing. The advantages of the column last configuration were desirable in converting to a lead-free interconnect in that the part could experience first-level assembly of the active and passive components, the bare Ni/Au I/O electrically tested, and then experience a low-temperature reflow to attach the column resulting in a high-melting solder joint. 2. Column Selection A column made out of pure copper was selected to replace the 90Pb–10Sn column primarily because of its high melting temperature, coupled with electrical performance characteristics. The column had to be optimized for both electrical and mechanical properties, in addition to selecting


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FIG. 79 Ceramic column grid array (CCGA) configurations. a) Single melt eutectic Pb–Sn fillet, b) cast column with eutectic Pb–Sn on card pad, c) column last attach solder process (CLASP) with Pddoped solder on chip-carrier pad and eutectic Pb–Sn fillet on card pad.

FIG. 80 Actual loop inductance data vs. predicted results for different Pb-free column geometries. (From Ref. 42.)

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TABLE 13 Elastic Limit of 0.36-mm Cu Wires Annealed in N2 for 1 hr at Temperatures Increasing from A to E Anneal temperature Nonannealed A C D E

Elastic limit, MPa 81.3 56.5 54.4 41.4 39.3


a lead-free alloy to attach the column to the chip carrier. Attachment to the PCB would be with the preferred alloy of most industry consortia, SAC, in an effort to support the commonality with second-level assembly suppliers [41]. a. Electrical Optimization. The electrical performance of a column improves with height reduction (lowers loop inductance) and increased column diameter (improves d.c. resistance). However, the taller and thinner the column, the more compliant the structure, and thus the better the thermomechanical fatigue resistance. Several combinations of height and diameter were modeled [42] and compared to the standard 0.5-mm-diameter, 2.2-mm-high 90Pb–10Sn column. Parts were assembled to validate the best candidates from the modeling results and are depicted in Fig. 80. A column length of 1.52 mm was selected to offer the best trade-off between electrical and mechanical performance for the specific application. b. Mechanical Optimization. Taking into account the electrical performance constraints described above, the properties of the Cu column were optimized with respect to thermomechanical behavior. Standard 90Pb–10Sn columns fatigue in the column itself because of the ductility of the material. A copper column, if fabricated by a drawing operation, was significantly more rigid than the Pb–Sn column. A more rigid column may transmit strain to the fillet causing fatigue cracking, and inevitably failure. To optimize the column, annealing was performed to reduce the stiffness. Solder columns exhibit approximately 1% flexural strain, therefore stress at and immediately following yielding was of great interest. Although annealing did not significantly change the ultimate strength of the wire, it did alter the elastic limit (Table 13) [41]. The lifetime of the joints as a function of anneal temperature was predicted and verified by experimental thermal cycle data as shown in Table 14. Cross sections of the copper column at a corner location after annealing and 0jC to 100jC thermal cycle testing are shown in Fig. 81. An annealing temperature was chosen to balance the improved reliability with the potential for handling damage because of the softer columns. Depending on the particular electrical requirements of the application, column dimensions (height, diameter) would be chosen to meet the electrical requirements while maximizing the lifetime.

TABLE 14 Observed Lifetime of Annealed Copper Column Interconnects. All Columns Were 2.2mm in Length and 0.25mm in Diameter Anneal temperature B D F Source: From Ref. 41.

Cycle life 2954 2575 4370


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FIG. 81 Annealed copper column grid arrays (CuCGAs) in a corner location after 4370 cycles at 0 to 100jC. (From Ref. 41.)

3. Fillet Alloy The copper columns were tin-plated to maintain the wettability of the copper by the fillet alloy during both attachment to the chip carrier and board assembly. Having a solder fillet at the chip carrier location that has a slightly higher melting temperature than the card fillet is desirable for card assembly robustness and reworkability. As mentioned earlier, based on the considerations for commonality at card assembly, the SAC alloy was chosen for the card fillet. This dictated the options for a higher melting lead-free alloy required for the copper column-to-component attachment for which a Sn–Ag alloy was chosen [43]. The resulting structure is shown in the cross section in Fig. 82. The figure indicates that the Sn–Ag uniformly wets the tin-coated copper column. Tin-copper intermetallic layers are observed at the interface of the column, and tin–silver precipitates are noted in the solder fillet. 4.

Qualification

a. Moisture Sensitivity Testing. Capless and direct lid attach CuCGA modules were subjected to J-STD-020B [44] moisture sensitivity testing at moisture sensitivity level (MSL) 1 and

FIG. 82 The copper column grid array (CuCGA) structure shown a) optically and b) in cross section as attached to the I/O pad of the chip carrier. (From Ref. 41.)

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TABLE 15 Chip Carrier and Printed Circuit Board Attributes Utilized in Thermal Fatigue Testing of the Copper Column Grid Arrays Ceramic TV Ceramic size Ceramic thickness Encapsulation Column dimensions Column material Column pitch Column array Driclad test card Card dimensions Card cross section Card layout Card surface finish Card pad diameter

White alumina (TV 852) 42.5 42.5mm 2.55 mm, 17 layers Mechanical chip with underfill, direct lid attach 0.25mm diameter, 1.5mm in length Tin-plated copper 1.0 mm 41 41, 1657 total I/O 8640789 229 279 2.36 mm 12S10P five sites OSP 0.70mm design, 0.65mm measurement


qualified for three reflows at 245jC. Underfill integrity was verified by C-mode scanning acoustic microscopy, and electrical testing verified the integrity of the flip chip solder joints that experienced the higher lead-free reflow temperature at the card-level assembly operation. The packages were mounted on cards, subjected to, and passed 2000 deep thermal cycles (DTC) from 55 to 125jC, 96 hr of highly accelerated stress testing (HAST) at 130jC with 85% relative humidity, and 1000 hr of high-temperature storage (HTS) at 150jC. b. Thermal Fatigue. Thermal fatigue testing was performed on special CuCGA daisy chain test vehicles assembled to Driclad PCBs (Table 15). The parts were tested with and without column rework at the chip carrier level, and with and without board rework at the card level. A subset of the parts was subjected to shock and vibration testing. The parts were then thermal cycled from 0jC to 100jC at a frequency of two cycles per hour (Fig. 83). The failure data were fit

FIG. 83 Ceramic chip-carrier and PCB fillet fractures in high-DNP columns after 3480 thermal cycles of 0 to 100jC. a) Low-magnification photo, b) magnified fillet on chip-carrier side, c) magnified fillet on PCB side. (From Ref. 41.)


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TABLE 16 Summary of Copper Column Grid Array Thermal Fatigue Data from Thermal Cycling at 0 to 100jC, 2 cph Cell

Readout

Parts

First fail (cycle)

N50 (cycle)

Sigma

Primea Primea Column rework Card rework

In situb Benchb Bench Bench

30 30 10 10

1410 1660 2700 2360

2200 2410 2760 2770

0.27 0.26 0.15 0.10

a

Half of prime parts preconditioned with impact shock and vibration. Prime parts monitored both in situ and bench. Source: From Ref. 41.

b

to a lognormal distribution to calculate the N50 and sigma. These laboratory fatigue test results are summarized in Table 16 [41]. 5. Summary IBM developed and qualified a lead-free column grid array solution that delivers equivalent electrical performance and a superior lifetime compared to the standard tin–lead technology. The qualification of the lead-free technology supports the manufacturing conditions of 1 column rework with up to two hot oil submersions for column removal. The CuCGA solution can be customized to support various specific application requirements.

D.

Sony Lead-Free Soldering Experience

Green Management 2005 is Sony’s mid-term environmental action program, aimed at specific targets for environmental performance. Under the plan, Sony set restrictions and reduction targets for hazardous materials including lead (Pb) and intends to introduce lead-free solder for use in its main products, and phase-out Pb in most products by March 31, 2005 [45]. Table 17 [46] lists some examples of the progress Sony has made toward that end. Fig. 84 shows a sampling of products with Pb-free printed wiring board assembly introduced by Sony [45]. Sony has been using lead-free solder primarily in consumer products such as televisions, audio products, mobile phones produced in Japan and other locations in Asia since March 31, 2002. By replacing the plating material used for external component contacts, Sony has eliminated 70% of lead from its products [47]. 1. Progress at Sony’s Kohda Division The Kohda Division of Sony that manufactures the Sony Vaio laptop computer as well as camcorders, digital cameras, and the Playstation 2 is one of the most active companies in the movement toward lead-free soldering. These products are assembled on a total of 62 lines in three shifts that run 7 days a week [48]. a. Range of Complexity. From double-sided motherboards to small cards for handheld cameras containing 1500 components, Sony was intent on converting even its most complex designs to lead-free. b. Commencement. Activity started by developing a lead-free solder paste and lead-free reflow profile. The solder selected by Sony is Sn–2.5Ag–0.5Cu–1Bi. Many Japanese OEMs prefer a Bi addition to improve the wetting characteristics of the solder and to slightly lower the melting point. This particular lead-free solder is molten at approximately 222jC. The challenge for Sony was to reach the peak temperature quickly while at the same time maintaining a small temperature delta across the board. Fig. 85 illustrates a typical reflow profile used at Sony for Sn–Pb solder assembly compared to the lead-free solder assembly profile preferred by most manufacturers [48].

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TABLE 17 Categorical List of Products Currently Manufactured at Sony Corporation with LeadFree Solder Assembly Category Car audio Handheld computer except cameras Handheld computer including cameras MD Walkman MD Walkman Digital still camera Digital camcorder Digital camcorder Notebook PC Notebook PC Color television Color television Color television Color television Computer CRT display Color LCD for computer LCD video projector MD system Personal audio system DVD player Video cassette recorder Video cassette recorder AIBO Mobile phone Mobile phone

Lead-free solder

Halogen-free flame retardant

Packaging

XR-CA430X PEG-NR70

PWB-CC PWB-CC

PWB-E PWB-E

NSF NSF

PEG-NR70V

PWB-CC

PWB-E

NSF

MZ-E909 MZ-N1 DSC-P9 DCR-TRV50 DCR-TRV40 PCG-SR9M/K PCG-R505FR/D KV-14DA1 KV-29DS55 KV21DA55 KV-32DX550 CPD-G220 SDM-N51AV VPL-CX4 MDS-JE780 CMT-P555DVD DVP-NS515 SLV-N77 SLV-SE220 ERS-311/312 C413S C1002S

PWB-SC PWB-SC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC PWB-CC

PWB-E PWB-E PWB-E PWB-E PWB-E PWB-E PWB-E PWB PWB PWB PWB PWB-E PWB-E PWB-E PWB-E PWB-E PWB-E PWB-E PWB-E PWB-E PWB-E PWB-E

NSF NSF NSF NSF NSF NSF NSF NSF RSF NSF RSF NSF NSF NSF NSF NSF NSF NSF NSF NSF NSF NSF

Model

Lead-free solder: PWB-SC—All printed wiring boards and some components; PWB-CC—All printed wiring boards and circuit components; PWB—Main printed wiring board. Halogen-free flame retardants: PWB—Eliminated from printed wiring board; PWB-E—None in printed wiring board or enclosure. Packaging: NSF—No styrene foam; RSF—Recycled styrene foam using limonene. Source: From Ref. 45.

c. Reflow Equipment Issues. The 8-year-old reflow ovens at Sony were not capable of meeting the rigorous demands of the lead-free profile and the environmental requirement of fluxfree emissions. The Kohda facility began a worldwide search to find a supplier to develop such a reflow oven that could meet these demands [48]. The Kohda team evaluated the reflow ovens of numerous manufacturers. The evaluations consisted of furnace loading capability, temperature deltas across assemblies, repeatability, and profile drift. A baseline was established for comparison purposes to determine the potential of each supplier. d. Solution. The criteria employed by the Sony Kohda team are detailed in Table 18. The most important requirement was the ability to maintain a tight process window. In addition, removing flux emissions was a requirement, and the use of filters had been ineffective because the disposal eventually contaminated landfills. Maintenance of these filters also resulted in unaffordable tool down time. Some of the innovations developed by Sony in conjunction with the chosen


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FIG. 84 Representative product examples at Sony Corporation manufactured with lead-free assembly. a) MD walkman, b) color television, c) Playstation 2, d) Network Handicam digital camcorder, e) DVD player, f) mobile phone, g) VAIO-SR notebook personal computer, h) AIBO entertainment robots. (From Ref. 45.)

reflow oven manufacturer were tight gas flow circulation management and a filterless flux removal system with a self-cleaning mode. A high-temperature blower assembly was also developed, as well as a shorter center board support and a reduced length to minimize floor space. The importance of tight gas flow circulation control is that there could exist significant temperature deltas between adjacent furnace heating zones without significant air movement to alter neighboring zone temperatures as occurs in a forced convection oven system. With this, the board could move from a 180jC soak zone to a 290jC ramp zone and then to a 235jC reflow zone. The number of heating zones was increased from the standard three to four zones, to eight or nine shorter zones to improve precise control. Also helping to maintain the tight temperature control was the shorter center board support (CBS) that spanned only the high temperature

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FIG. 85 Reflow soldering profile comparison. a) Typical eutectic Sn–Pb solder profile, b) lead-free solder profile preferred by Japanese manufacturers. (From Ref. 48.)

zones. This is utilized to prevent board warpage but can also block air flow. In reducing the length of the CBS to those zones where PCB warping would occur, there is less obstruction to air flow and less need for special profiling. Fig. 86 illustrates the CBS. Lastly, a condensation/precipitation method was utilized to remove flux from the oven tunnel prior to its cooling, eliminating the need for filtering. The flux is transported to a series of trays where the concentrated flux byproduct runs through a self-clean cycle that operates

TABLE 18 Criteria for Reflow Oven Selection for Lead-Free Solder Assembly Employed at Sony Corporation Criteria . . . . . .

Tight temperature control Very low temperature delta Rapid ramp to meet lead-free profile Flux-free and filter-free operation Minimal maintenance Short machine footprint to conserve floor space



655

FIG. 86 Reflow oven with shortened center board support spanning only the high temperature zones. (From Ref. 48.)

automatically during production where the flux is reliquified and captured in collector trays (Fig. 87). The ‘‘self-clean’’ cycle takes 20 min and the collector trays have a capacity to hold the flux that accumulates over approximately 1 year. At that time the trays can be removed and cleaned, resulting in only about 10 min of down time over a 1-year period (Fig. 88) [48]. e. Goal Achieved. As a result of this cooperative effort between the supplier and manufacturer, Sony was able to accomplish its goal of implementing lead-free soldering with flux-free emissions.

E.

Seiko-Epson Lead-Free Soldering for PCBS

Seiko-Epson established a Lead-Free Promotion Committee in June of 1999 to hasten the transition to lead-free equipment. The goal of the committee was to ‘‘gradually reduce the use

FIG. 87 Flux collection system for air reflow including dual internal and external exhausts. (From Ref. 48.)

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FIG. 88 Flux collection trays showing a) flux accumulation after 30 days of processing (56 kg solder paste), b) after 20-min automatic self-cleaning cycle. (From Ref. 48.)

of solder lead in circuit boards and achieve the lead-free goal by the end of the fiscal year 2001’’ [49]. 1. Criteria Seiko-Epson’s criteria for implementation were that the interconnect exhibit characteristics equal or superior to those of the current 63Sn–37Pb alloy utilized. In addition, to utilize to the maximum extent the existing production equipment and technologies already established. Lastly, to develop the lead-free processes in room air conditions. Thermal fatigue performance, soldering performance, and visual inspectability equal to that of eutectic Pb–Sn solder were the front-up approach. It was also a primary consideration that the alloy be nonhazardous and readily available [50]. Seiko-Epson’s approach focused on defining lead-free solutions for the three major soldering processes utilized to assemble electronic units: surface mount reflow soldering, wave soldering, and manual soldering for rework operations. Preliminary investigations narrowed the alloy selection to the following candidates to evaluate: Sn–Ag solder (with Cu, no Bi and with Cu and a small amount of Bi, and with Cu and a larger amount of Bi) as well as Sn–Cu, Sn–Zn, and Sn–Bi solder. Final selections for each soldering method were made mainly on the basis of wetting characteristics and bonding reliability. Based on this experimentation, Seiko-Epson researchers chose the alloys in Fig. 89 for reflow, flow (wave), and rework processing. The optimum production conditions were developed, using Sn–Bi plating for electrical components. By the end of 2000, this process was completed in Japan, and by the end of the following year for the remainder of the world [51]. 2.

Surface Mount Reflow Soldering

a. Wetting. The solder system selected for reflow was Sn–Ag–Cu with a melt temperature of about 220jC using a temperature profile with a peak temperature of 235jC and a dwell time (i.e., time above 220jC) of at least 30 sec in air. Seiko-Epson utilized a test vehicle consisting of a 1.6-mm-thick FR-4 board to which was mounted 208-pin QFP components with a 0.5-mm pitch. Two types of component lead frames were evaluated: Alloy 42 with 90Sn–10Pb plating, and Cu with 98Sn–2Bi plating. The wetting was determined adequate for both the Ob and Pb-free plated leads. b. Solder Joint Characterization. Microstructural analysis confirmed that sound solder joints were achieved. Some Ag3Sn IMC was formed scattered throughout the Sn–Ag–Cu bulk solder. The microstructure did not coarsen when exposed to elevated temperatures. The joint mechanical properties were judged to be very good.


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FIG. 89 Seiko Epson lead-free solder selection for reflow, wave, and manual (repair) soldering. (From Ref. 50.)

Pull strength was evaluated after subjecting parts to thermal cycle conditions consisting of 40jC to 125jC for up to 2000 cycles. It was determined that parts soldered with Sn–Ag–Cu (90Sn–10Pb plated leads) exhibited better pull strength than similar parts reflowed with either 63Sn–37Pb or Sn–Ag–Cu–Bi solders used as a benchmark. Also, it was observed that Sn–Ag–Cu has a much lower tendency to develop cracks during accelerated temperature cycle (ATC) testing than eutectic Sn–Pb as is evident in the cross-section photographs of the solder joints depicted in Fig. 90.

FIG. 90 Cross-sectional photo illustrating leads soldered with eutectic Sn–Pb and Sn–Ag–Cu solder before and after thermal cycle testing (2000 cycles). (From Ref. 50.)

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c. Which Sn–Ag–Cu Alloy. Seiko-Epson evaluated several Sn–Ag–Cu alloys in comparison to eutectic Sn–Pb as noted in Table 19. All of the compositions evaluated were judged to provide satisfactory wetting using the test vehicle described earlier in this section. In addition, the pull strength of leads bonded with any of the Sn–Ag–Cu compositions evaluated was equivalent or superior to the eutectic Sn–Pb benchmark parts after exposure to ATC testing (i.e., 40jC to 125jC, 2000 cycles). d. Other Factors. The Sn–Ag–Cu alloys are known to exhibit slightly inferior wettability compared to eutectic Sn–Pb. Although a study was conducted to investigate various preheat and peak temperature combinations to determine one that resulted in an optimum front fillet wetting height, all combinations provided reliable solder joints although there was some variation in fillet height. Seiko-Epson also investigated utilizing Sn–Ag–Cu to attach area array packages such as BGAs and CSPs. ATC testing (40jC to 125jC, 30 min/cycle) confirmed that parts attached with Sn–Ag–Cu exhibited more than three times the fatigue life compared to conventional eutectic Sn– Pb solder. e. Conclusion. Based upon all the foregoing, it was concluded that Sn–Ag–Cu alloys were a viable choice over a range of both Cu and Ag composition to assemble Seiko-Epson products. The Sn–Ag–Cu family of alloys was capable of producing adequate solder reflow joints whose properties were equivalent to or better than the benchmark eutectic Sn–Pb solder under a broad variety of processing and lead finish conditions [50]. 3. Wave Soldering The Sn–Cu system was selected for wave soldering, with the eutectic m.p.=227jC, which is about 40jC higher than eutectic Sn–Pb (m.p.=183jC). Although a significant difference in melt temperature exists, Seiko-Epson determined that a wave soldering capability equivalent to eutectic Sn–Pb could be achieved only by raising the bath solder temperature for Sn–Cu by about 10jC, which was typically 240jC to 250jC for eutectic Sn–Pb. To offset the inferior wettability of CuSn relative to eutectic Sn–Pb, Seiko-Epson optimized the post flux by modifying the active agents and increasing the solids content to improve the wettability. Using a solder bath temperature of 250 (F5) jC, a molten solder exposure time of 2 to 5 sec, and the modified post flux, the solder wettability in through holes and component leads was found to be similar to eutectic Sn–Pb [51]. a. Evaluation/Characterization. In evaluating the pull-out strength of pins wave soldered in through holes of a PCB, Sn–Cu soldered pins exhibited a higher pull-out strength initially and after ATC testing (40jC to 125jC) than eutectic Sn–Pb. Large cracks were observed in the eutectic Sn–Pb fillets after ATC testing, but only very slight cracks were observed in the case of the Sn–Cu fillets as shown in Figs. 91 and 92. Pins of large and heavy components or connectors soldered in through holes are a solder creep concern. Creep is a gradual plastic deformation process that takes place over time due to an applied load. As indicated in Fig. 93, both Sn–Ag–Cu and Sn–Cu exhibited vastly superior results in a standard creep test at both 125jC and 150jC compared to eutectic Sn–Pb. Thus both these lead-free solders are adequate for bonding relatively large and heavy components.

TABLE 19

Lead-Free Solders Under Evaluation at Seiko-Epson for Reflow Soldering Alloy composition

Sn–Ag–Cu (type A) Sn–Ag–Cu (type B) 63Sn–37Pb solder Source: From Ref. 50.

Sn

Ag

Cu

Pb

Remaining portion Remaining portion

3 to 3.8 wt.%

0.5 to 1.2 wt.%

0.1 wt.% or less 37 wt.%


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FIG. 91 Cross-sectional photo illustrating posts soldered with eutectic Sn–Pb and Sn–Ag–Cu solder before and after thermal cycle testing (2000 cycles). (From Ref. 50.)

b. Fillet Lifting Phenomena. One of the key concerns with the use of lead-free alloys for wave soldering applications is associated with a phenomenon referred to as ‘‘fillet lifting’’ discussed earlier in studies conducted by Panasonic (Section 4.6.2). It is the partial separation of the soldered area from the land of a PCB through hole immediately after wave soldering. It does occur with lead-free solders, including Sn–Cu as shown in Fig. 94. There is no known solution to prevent the phenomenon [51]. c. Bond Strength. Seiko-Epson conducted testing on double-sided PCBs with components attached using both Cu–Sn and eutectic Sn–Pb wave soldering. Although fillet lifting is certainly an undesirable feature, it is interesting that the pull strength of through-hole joints was at approximately the same strength level for both solders, although fillet lifting was observed in the case of the Sn–Cu joints, but not the eutectic Sn–Pb joints. This was true for joints evaluated both before and after ATC testing (40jC to 125jC, 2000 cycles) [51].

FIG. 92 Cross-sectional photos illustrating enlargements of cracked soldered area after thermal cycle testing (2000 cycles). (From Ref. 50.)

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FIG. 93 Creep characteristics of eutectic Sn–Pb solder compared to lead-free solder alloys at 150jC. (From Ref. 50.)

d. Electrical Characteristics. Although the bond strength appears adequate even in cases with fillet lifting, it was not immediately obvious that the electrical characteristics of these joints were sufficient and remained so. However, Seiko-Epson, under the same conditions discussed in the previous section, confirmed that the solder joint resistance values remained at about the same value in Sn–Cu joints as the conventional eutectic Sn–Pb joints. e. Conclusions. Based on the testing Seiko-Epson conducted, they adapted Sn–Cu as their lead-free wave solder material for volume manufacturing [51]. 4. Manual (Repair) Soldering Sn–Ag–Cu was selected by Seiko-Epson as the lead-free solder of choice for rework applications, the same as used for surface mount reflow operations. Unlike eutectic Sn–Pb, solder spreading on Cu land areas was found not to vary much, independent of the soldering iron tip temperatures. However, at tip temperatures of 350jC or below, the alloy is difficult to melt. The optimum tip temperature was found to be approximately 365jC to 375jC. At 380jC and above, flux charring and land separation from the PCB is observed [51].

FIG. 94 Fillet lifting on pins post a wave solder operation utilizing Sn–Cu. (From Ref. 50.)


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a. Need for Quick Repairs. Even with a tip temperature at about 370jC, which is desirable from a workability standpoint, repair operations must be carried out quickly to avoid damage to the PCB lands. For example, component and bond strength with joints created using a 370jC tip temperature is optimum with a soldering time of about 3 sec, but decreases rapidly thereafter. The tendency to create thermal-induced damage or peeling of PCB land features follows the same trend. When manual soldering is performed using Sn–Ag–Cu solder at a suitable temperature and soldering duration, the soldering characteristics are very similar to eutectic Sn– Pb as noted in Fig. 95 [51]. b. Effect of Mixed Technologies. A key concern to Seiko-Epson was the effect of a hybrid system, i.e., mixed technologies. That is, what is the impact, if any, on the pull strength where a lead-free solder in a joint becomes mixed with solder containing lead (Pb). An evaluation was conducted addressing the issues from several perspectives. It was determined, as depicted in Fig. 96, that the pull strength of QFP leads remained relatively unaffected when various solders were used to repair a soldered component originally mounted with lead-free solder, Sn–Ag–Cu (either type A or B). The reverse was also true, that the pull strength was relatively unaffected when a lead-free solder was used to repair QFPs originally mounted with eutectic Sn–Pb. It was therefore concluded at Seiko-Epson that no problems would result by utilizing lead-free Sn–Ag–Cu for their solder repair operations independent of whether a SMT component was attached with eutectic Sn–Pb or Sn–Ag–Cu [51]. 5. Lead-Free Status at Seiko-Epson Seiko-Epson has taken action to make in-house processes (including overseas manufacturing affiliates and subcontractors) lead-free, make purchased components lead-free, and implement procedures to obtain approval from customers to switch custom products to lead-free manufacturing. As of January 2002, the solder used in in-house production at Seiko-Epson was 70.8% lead-free [51]. (Not included in that list of products are products that are exempt; products that

FIG. 95 Manual soldering of QFP utilizing Sn–Ag–Cu and eutectic Sn–Pb solder on Sn–10Pb-plated leads. (From Ref. 50.)

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FIG. 96 Bond strength comparison for QFP initially bonded with Sn–Ag–Cu (type A), Sn–Ag–Cu (type B), and eutectic Sn–Pb solder, in the no-repair and repaired condition. (From Ref. 50.)

ceased production before FY2001, and their repair parts; products for which Seiko-Epson has been able to get permission of the customer to switch to lead-free materials.) In April of 2002, Seiko-Epson stated that 98.6% of the solders used in in-house manufacturing processes were lead-free and that 40% of the total number of purchased parts, under this program, were lead-free [50]. a. Source of Supply. As with most major manufacturers, Epson procures the majority of its electronic components from other suppliers, including PCBs. To facilitate the logistics of insuring that its supplier’s parts were lead-free, Seiko-Epson set up a database compiling the leadfree status of over 60,000 parts from over 300 manufacturers, and made it available in both English and Japanese. This database is used worldwide by the Epson Group of companies. [50] b. Custom Products. Seiko-Epson’s business is also greatly composed of made-to-order semiconductors and liquid crystal displays. Permission is therefore required to convert these products to lead-free. Seiko-Epson has approached approximately 6000 of its customers with data and samples, but this process has taken longer than expected. Conversion of these products will take place on a separate schedule in line with the requirements of each customer. [50]

V.

SUMMARY

This chapter offers just a few of the many examples of manufacturers that are well underway in the conversion to lead-free electronics and has highlighted some of the challenges they encountered along the way. There has been substantial progress in both technology and acceptance of the reality of the lead-free movement. It has been over a decade since its inception,


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and manufacturers and their suppliers must be prepared to support customers globally, in response to legislation or ‘‘green marketing’’ influences. The few examples demonstrate, however, that the database surrounding lead-free solders is growing, and that the momentum has shifted from a technology issue to understanding the manufacturing implications that are driven by the conversion to lead-free components and processing. NEMI has demonstrated that manufacturing in most cases can be implemented without major changes to production tools and equipment, and produces reliable solder joints. However, as these examples illustrate, the conversion still requires a significant effort on the part of manufacturers to produce reliable, cost-effective products in high volume. The examples also illustrate that the task is achievable.

ACKNOWLEDGMENTS Authors Tom Baggio and Kenichiro Suetsugu would like to thank the following Panasonic members for their contributions to this work: S. Hibino, A. Furusawa, and T. Ikari of Environmental Production Engineering Laboratory, Kadoma City, Osaka Japan; E. Abe, H. Saitoh, M. Ojima, Y. Hayasa, and Y. Takaku of the AVC TV Products Division, Utsunomiya, Japan; H. Kitahara, S. Kumamoto, and M. Aihara of Harima Chemicals, Inc., Kakogowa, Japan.

REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13. 14. 15. 16. 17. 18. 19. 20.

Seelig, Karl; Suraski, David. A study of lead-contamination in lead-free electronics and its impact on reliability. Proc. SMTA International Chicago, IL; Sept 22–26, 2002. Judd, Mike; Brindley, Keith. Soldering in Electronics Assembly—Second Edition; ButterworthHeinermann: Oxford, 1999. Suetsugu, K. Development of flow and reflow soldering technology for lead-free solder. Proc. of Jisso/Protec Forum; Japan; November 2002; 163 pp. MEI, Environmental Sustainability Report (Covering FY2001), June 2002, 28 pp. Baggio, T. Case studies of lead-free technology implemented in Panasonic electronic products. Proc. Nepcon West, TS-13; March 2000, 5 pp. Suetsugu, K. Advanced applications of lead-free soldering by identifying lift-off mechanism. Proc. 10th Micro Electronics Symposium, Japan; 214 pp. Suetsugu, K. Advanced applications of lead-free soldering by identifying lift-off mechanism. Proc. 10th Micro Electronics Symposium; Japan, 2000; 215 pp. Suetsugu, K. Development of flow and reflow soldering technology for lead-free solder. Proc. of Jisso/Protec Forum; Japan, November 2002; 159 pp. Suetsugu, K. Development of flow and reflow soldering technology for lead-free solder. Proc. of Jisso/Protec Forum, Japan, November 2002; 155 pp. Suetsugu, K. Development of flow and reflow soldering technology for lead-free solder. Proc. of Jisso/Protec Forum, Japan, November 2002; 156 pp. Suetsugu, K. Development of flow and reflow soldering technology for lead-free solder. Proc. of Jisso/Protec Forum, Japan, November 2002; 158 pp. Suetsugu, K. Development of flow and reflow soldering technology for lead-free solder. Proc. of Jisso/Protec Forum, Japan, November 2002; 162 pp. Baggio, T. The Panasonic mini disk player, turning a new leaf in a lead-free market. Proc. IPC Works ’99, October 1999; pS-05-4, 4 pp. MEI, Environmental Sustainability Report (Covering FY 2001), June 2002; 28 pp. Baggio, T. The Panasonic mini disk player, turning a new leaf in a lead-free market. Proc. IPC Works ’99, October 1999; pp. pS-05-4, 4 pp. Baggio, T.; Suetsugu, K. Guidelines for lead-free processing. SMT Magazine, September 1999; 64 pp. Baggio, T.; Suetsugu, K. Guidelines for lead-free processing. SMT Magazine, September 1999; 62 pp. Baggio, T. Case studies of lead-free technology implemented in Panasonic electronic products. Proc. Nepcon West, March 2000; TS-13 8 pp. Baggio, T.; Suetsugu, K.; Okumura, T. Challenges and solutions for lead-free soldering of large PCB assemblies. Proc. IPC Apex, P-MT3-1, March 2000; 2 pp. Baggio, T. Case studies of lead-free technology implemented in Panasonic electronic products. Proc. Nepcon West, TS-13, March 2000; 10 pp.

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Baggio, T. Case studies of lead-free technology implemented in Panasonic electronic products. Proc. Nepcon West, TS-13, March 2000; 11 pp. 22. Baggio, T. Case studies of lead-free technology implemented in Panasonic electronic products. Proc. Nepcon West, TS-13, March 2000; 12 pp. 23. Suetsugu, K. Development of flow and reflow soldering technology for lead-free soldering. Proc. Jisso/Protec Forum, Japan, November 2002; 160 pp. 24. Baggio, T.; Suetsugu, K.; Okumura, T. Challenges and solutions for lead-free soldering of large PCB assemblies. Proc. IPC Apex, P-MT3-1, March 2000; 4–5. 25. Baggio, T. Case studies of lead-free technology implemented in Panasonic electronic products. Proc. Nepcon West, TS-13, March 2000; 14 pp. 26. Baggio, T. Case studies of lead-free technology implemented in Panasonic electronic products. Proc. Nepcon West, TS-13, March 2000; 13 pp. 27. Suetsugu, K. Advanced applications of lead-free solder by identifying lift-off mechanism. Proc. 10th Micro Electronics Symposium, Japan, 2000; 214 pp. 28. Suetsugu, K. Advanced applications of lead-free solder by identifying lift-off mechanism. Proc. 10th Micro Electronics Symposium, Japan, 2000; 215 pp. 29. Suetsugu, K. Advanced applications of lead-free solder by identifying lift-off mechanism. Proc. 10th Micro Electronics Symposium, Japan, 2000; 216 pp. 30. Suetsugu, K. Development of flow and reflow soldering technology for lead-free soldering. Proc. Jisso/Protec Forum, Japan, November 2002; 156 pp. 31. Suetsugu, K. Development of flow and reflow soldering technology for lead-free soldering. Proc. Jisso/Protec Forum, Japan, November 2002; 158 pp. 32. Baggio, T. The Panasonic mini disk player, turning a new leaf in a lead-free market. Proc. IPC Works ’99, October 1999; pS-05-4, 7 pp. 33. Yi, Sammy; Geiger, David; Wang, Mei; Singh, Haprpuneet; Lee, Alan; Wong, Harris A case study of lead-free assembly implementation in EMS environment. Proc. SMTA Lead Free Symposium, Boston, MA, June 2000. 34. Flextronics Press Release, Singapore, Nov. 15, 2002. 35. Wang, M.; Nakajima, K.; Lewis, A.; Kurwa, M.; Bhat, R.; Yi, S. Solder paste printing optimization for CSP. Proc. SMTA International; San Jose, CA, September 1999. 36. Wong, Stephen; Nester, Myron. High Volume Lead Free Production, CMAP Lead Free Workshop, Toronto, Canada, May 22, 2002. 37. Visteon press release, Dearborn, MI: Dec. 18, 2000. 38. Master, Raj, et al. Ceramic column grid array for flip chip applications. Proc. 45th Electrical Components and Technology Conference; IEEE: Las Vegas, NV, 1995; 925–929 pp. 39. Ray, Sudipta, et al. Ceramic column grid array for a high performance workstation application. Proc. 47th Electrical Components and Technology Conference; IEEE: San Jose, CA, 1997; 319–324. 40. Ray, Sudipta, et al. CLASP ceramic column grid array technology for ceramic chip carriers. Proc. Semicon West, San Jose, CA, July 1999. 41. Interrante, Mario, et al. Lead-free package interconnection for ceramic grid arrays. Proc. Semicon West, San Jose, CA, July 2003. 42. O’Connor, Daniel, et al. Electrical modeling and characterization of packaging solutions utilizing lead-free second-level interconnects. Proc. of the 53rd Electrical Components and Technology Conference, New Orleans, LA, May 2003. 43. Interrante, Mario; et al. High Density Column Grid Array Connections and Method Thereof, U.S. Patent 6,429,388, 2002. 44. J-STD-020B, Moisture/Reflow Sensitivity Classification for Non-hermetic Solid State Surface Mount Devices. IPC/JEDEC, July 2002. 45. Sony Corporation, Environmentally Conscious Products and Services, Social and Environmental Report, 2002. 46. Sony Corporation, http://www.sony.net/SonyInfo/Environment/environment/communication/ report/2002/index.html. 47. Sony Corporation. Sony semiconductor’s environmental efforts—towards a circulatory society. Sony CX-News, Feb. 2003; 29. 48. Kikuchi, Atsushi. The Greening of the Reflow Process, Through Partnership, http://www.sehouk. com/Sony_article.doc, 1/2/2002. 49. Seiko Epson Corporation. Epson declares solders 98.6% lead free*. Epson Environmental Newsline, Dec. 2002. 50. Seiko Epson, The Road to Lead-Free Soldering (for PCB mounting), http://www.eea.epson.com/ aboutEEA/PCBMounting.pdf. 51. Press Release, Seiko Epson Announces Lead-Free Status of Products, March 11, 2002.

17 Major International Lead (Pb)-Free Solder Studies Carol A. Handwerker and Frank W. Gayle National Institute of Standards and Technology, Gaithersburg, Maryland, U.S.A.

Erik E. de Kluizenaar Philips CFT, Eindhoven, The Netherlands

Katsuaki Suganuma Osaka University, Osaka, Japan

I. INTRODUCTION Beginning in 1991, the microelectronics community worldwide became increasingly aware of the possibility of being required, by law, by tax, or by market pressure, to replace tin–lead eutectic solders in electronic assemblies. Over the next 10 years that followed, separate groups formed in the United States, Europe, and Japan to examine solder-based alternatives to tin–lead eutectic solder and to understand the implications of such a change before it becomes necessary. Lead-free solder research projects from industrial, academic, and national laboratory groups are described in the Lead Free Soldering report sponsored by the UK Department of Trade and Industry [1]. The time line for comprehensive national or regional studies, shown in Figure 1, illustrates the long-term commitment to develop lead-free solder assemblies. The following studies are featured in this chapter: the National Center for Manufacturing Sciences (NCMS) Lead-Free Solder Project and the NCMS High-Temperature Fatigue-Resistant Solder Project from the United States, the Japan Institute of Electronics Packaging (JIEP) and Japan Electronic Industry Association (JEIDA) projects from Japan, and the IDEALS Lead-Free Solder Project from the European Union (EU). Based on these studies, the microelectronics community gained sufficient experience with the performance of lead-free solders to begin addressing lead-free assembly to assess lead-free issues, including manufacturing yield, process windows for complex boards, and component survivability. Results from these five studies are compared with subsequent data obtained by Motorola and IBM. Progress on more recent national and/or industrial sector groups investigating lead-free solders, such as the National Electronics Manufacturing Initiative (NEMI) Lead-Free Task Force and the High Density Packaging Users Group (HDPUG) is also outlined. Updates and information obtained by these groups can be found at http://www.nemi.org, http://www.leadfree.org, and http://www.lead-free.org. In addition to these studies, which focused on implementation issues, primarily with respect to manufacturing and reliability, other studies yielded materials property data for lead-free solders that give insight into phase transformations, solderability, and thermomechanical fatigue (TMF) resistance. For example, over 20 universities in the United States, Canada, and the

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FIG. 1 Timeline of several lead-free solder studies.

European Union have conducted lead-free solder research programs supported by their national science foundations, companies, and national laboratories. Since 1982, the International Tin Research Institute (ITRI), now known as Soldertec, has performed research on lead-free solders, beginning with plumbing solders. Information on the Soldertec data can be found on at the website http://www.lead-free.org.

II. NATIONAL CENTER FOR MANUFACTURING SCIENCES (NCMS) PROJECT A. Project Organization and Participants From 1992 to 1996, the National Center for Manufacturing Sciences (NCMS), a not-for-profit cooperative research consortium of more than 200 North American manufacturers, national laboratories, and universities, carried out a multiyear effort to develop Pb-free alternatives to Pbbased solders used for electronics assembly [2]. The proposed Reid bill, banning or taxing the use of lead in electronics, sensitized the microelectronics industry to the possible need to remove Pb from solders, and led to the formation of this project. The total cost of the research project was $10.6 million, with approximately $1 million coming from Department of Defense funds and the balance from in-kind contributions. Participating organizations and primary technical contacts are presented in Table 1.

B. Project Objectives The goal of the NCMS project was to determine whether safe, reliable, nontoxic, and costeffective substitutes exist for lead-bearing solders in electronics manufacturing. The scope of the project was based on the following premise: the production of durable, reliable, safe, and affordable electronic products with lead-free solders will require the manufacturer to understand material properties, manufacturing processes and equipment, toxicological effects, alloy cost, long-term availability, and reliability.

C. Down-Selection Process Comprehensive manufacturing assessments and thermal cycling stress tests could only be performed on a limited number of candidate lead-free alloys: seven alloys plus Sn–Pb eutectic control alloy. The Project’s initial list of over 70 candidate alloys was quickly reduced by: (1) eliminating candidates containing toxic elements, (2) eliminating all but one of a family of

MAJOR INTERNATIONAL LEAD (PB)-FREE SOLDER STUDIES

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TABLE 1 Participating Organizations and Primary Technical Contacts in the National Center for Manufacturing Sciences Project on Lead-Free Solders Participating organizations AT&T/Lucent Technologies (Lucent) Electronics Manufacturing Productivity Facility (EMPF) Ford Motor Company (Ford) General Motors–Delco Electronics (GM–Delco) General Motors–Hughes Aircraft (GM–Hughes) Hamilton Standard, Division of United Technologies Corporation (Hamilton) National Center for Manufacturing Sciences (NCMS) National Institute of Standards and Technology (NIST) Rensselaer Polytechnic Institute (RPI) Rockwell International Corporation (Rockwell) Sandia National Laboratories (Sandia) Texas Instruments Incorporated (TI)

Primary technical contacts Iris Artaki and Donna Noctor John Greaves Tsung-Yu Pan and Howard Blair Gordon Whitten and Yun Zhu Jerald Rosser Charles Desantis and Willie Desaulnier Duane Napp Carol Handwerker and Leonard Smith Lawrence Felton and Mark Palmer John Mather and Scott Schroeder Paul Vianco and Darrel Frear Joe Felty

candidates having very minor composition differences, (3) eliminating candidates falling outside certain manufacturing, reliability, and economics/availability criteria established by project participants, and (4) applying judgment criteria set forth in a weighted decision matrix for those alloys that passed the first three screenings. The process from the initial listing of alloys through assessing reliability is shown schematically in Figure 2. 1. Toxicology Cadmium is clearly toxic in many forms and is severely restricted in Europe and was, therefore, eliminated as a potential alloying element. The data on all other elements was incomplete, and

FIG. 2 Diagram of the NCMS lead-free solder project alloy down-selection process.

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sometimes contradictory. Given the uncertainty in the toxicology of other possible alloying elements, no other elements were eliminated. The elements remaining for consideration were Sn, Ag, Bi, In, Cu, Sb (as a minor constituent), and Zn. Small additions (6 wt.%) is the discontinuous precipitation structure of Bi at the grain boundaries [5]. The deleterious effects of high Bi content on the plasticity cmax were observed earlier at f4.66 wt.% Bi (Fig. 20b). The isothermal fatigue life of bulk Sn–3.5Ag–(2–10)Bi solder alloys is remarkably reduced with increasing Bi concentration due to a decrease in ductility caused by the coarse, platelike, and irregularly shaped Ag3Sn phase throughout the microstructure [31,32]. However, the reliability of Sn–Ag–Bi bulk solders can be equivalent to eutectic Sn–Pb if 2 wt.% Bi is taken as an upper limit. However, the reliability of Sn–Ag–Bi bulk solders can be maintained as equivalent to eutectic Sn– Pb if 3 wt.% Bi is taken as an upper limit. Nikander [57] reported stress–strain curves of Sn– 3.5Ag–1.5Bi and Sn–3.5Ag–6Bi showing the tensile strength of Sn–3.5Ag–6Bi to be much higher than Sn–3.5Ag–1.5Bi, which is similar to the results reported by Kariya and Otsuka [31,32]. 3. Sn–Ag–In Alloys The compositions of several Sn–Ag–In alloys were investigated, as noted in Table 8, along with their melting ranges [15,31,50,56,58,59]. Because indium is very expensive and scarcely available, it may only be considered for alloys with small additions to Sn–Ag solder alloys. The melting point of Sn–Ag alloys decreases with indium additions. Tensile strength increases but the ductility decreases with increasing indium content, as shown in Fig. 21. Microstructural observations suggest that indium atoms dissolve in the h-Sn solid solution of Ref. 32; thus, the increase in tensile strength accompanying indium additions is most likely the result of solution hardening. The loss of ductility of Sn–3.5Ag–In alloys is ascribed to the irregularly shaped second phase (Ag3Sn) (Fig. 22). The fatigue life of Sn–3.5Ag alloy decreases with increasing indium content (Fig. 1). Eutectic Sn–3.5Ag alloy exhibits a relatively high fatigue resistance, and indium additions of less than 2 wt.% were found to have no practical effect on fatigue life. In one study, Sn–Ag–In alloys were found to be superior to Sn–Pb eutectic even with 5 wt.% In [32]. Although Sn–Ag–In solders exhibit good mechanical properties at ambient temperatures, their application in consumer electronics is limited due to their low eutectic temperature, and the strong segregation of indium as reported by Korhonen and Kivilahti [60]. The system has a eutectic reaction L!Ag2In+g+h at about 113jC. A solder alloy whose composition is within the Ag2In–g–h three-phase triangle at 113jC starts to melt or solidify at this temperature, greatly limiting the upper service temperature of products that utilize these solder alloys. The potential problem of this solder system is that, even if a solder composition is well outside the Ag2In–g–h three-phase triangle, the alloy is not completely solidified above 113jC as confirmed by DSC measurements [60]. This is due to the fact that the diffusion of indium in the solid g-phase during

FIG. 20 Effects of bismuth (Bi) content on the (a) ultimate shear strength (smax) and (b) plasticity (cmax) of Sn–Ag–Bi solder joints with Ag=3.3(F0.1) wt.%. (Courtesy of ASME, Ref. 56.)

RELIABILITY ASPECTS OF LEAD-FREE SOLDERS TABLE 8

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Melting Temperatures of Several Sn–Ag–In Alloys

Alloy composition

Melting temperature (jC)

Sn–(1–6)Ag–(4–35)Sn Sn–3.5Ag–5In Sn–3.5Ag–(1,3,5)In Sn–3.19–Ag–881In Sn–3.5Ag–(1.5)In

(167.9–179)–(212–214) f210 210–216 205–215 210–217

References 50 58 59 56 31

solidification is very slow, causing indium atoms to be segregated within that phase, leaving a greater amount of the liquid phase along the grain boundaries (i.e., it remains in the liquid state until 113jC and then undergoes solidification). Conversely, when the service temperature reaches 113jC, the liquid phase starts to form at the grain boundaries and renders an interconnection susceptible to hot cracking.

C. Sn–Zn–X Alloys Tin–zinc alloys are generally considered corrosion-prone owing to the oxidation potential of zinc. They have poor wettability and are not well suited for wave soldering due to high dross formation. 1. Sn–Zn–In Alloys The eutectic Sn–9Zn alloy melts at 199jC, but it can, together with near-eutectic alloys, be processed to maintain a stable fine-grained microstructure [36,61]. In contrast, alloys based on the eutectic Sn–Bi (139jC) and Sn–In (120jC) systems are not as stable microstructurally under the same conditions. A partial list of studies conducted on Sn–Zn–In alloys, whose findings are reported in the literature, is listed in Table 9 [15,50,58,62–66]. An addition of 5 wt.% In to Sn–9Zn decreases the melting point down to about 188jC, only 5jC higher than the eutectic Sn–Pb, while significantly improving the wetting characteristics of the eutectic Sn–Zn alloys [15,64]. When using an organically activated (OA) flux, the wetting behavior of this Sn–Zn–In alloy approaches eutectic Sn–Pb. However, OA fluxes require cleaning and are not always practical to use.

FIG. 21 Effect of In on the tensile strength and ductility of the Sn–3.5Ag–In solder at 25jC (e = 5 103). (Courtesy of TMS, Ref. 31.)

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FIG. 22 Effects of In content and plastic strain range on the fatigue life of the Sn–3.5Ag–In solder. (Courtesy of TMS, Ref. 31.)

Reflow soldering utilizing Sn–Zn–In alloy pastes on printed circuit boards has been studied using a no-clean, RMA-type flux in both air and N2 atmosphere [40]. After reflowing in air on Auplated Cu pads at 235jC, the solder balls were observed to be shiny and covered the pads well. After reflowing in N2 on pretinned Cu pads at 235jC, a device with pretinned leads was soldered to 1-in. square perimeter arrays of Sn–9Zn–5In solder pastes deposited on a few of these pads. The joints were observed to be well shaped. TABLE 9

Studies of Some Sn–Zn–In Solders

Compositions (wt.%) Sn–9Zn–(5–10)In

Sn–8Zn–5In–(0.1–0.5)Ag

Sn–8Zn–5In Sn–8Zn–5In–0.5Ag Sn8.8Zn–(0–10)In

Sn–9Zn–(2–5)In–(1–9)Bi Sn–6Zn–(2–5)In–(1–9)Bi

Investigations

References

Compressive creep and wetting properties compared with 40Pb–60Sn; oxidation and corrosion problems discussed V0.1 wt.% Ag dramatically improved mechanical properties of Sn–8Zn–5In; microstructures, tensile stress–strain properties, fracture surface in Ag-free and 0.1 wt%. Ag containing alloys were compared Effects of four fluxes on solderability, melting, and wetting behavior were investigated Tm=187jC; Cu was found to be beneficial for microstructural improvement Phase equilibrium calculation shows that a small amount (2 or 5 wt.%) of In is desirable in this alloy; Tm is about 19jC Low Bi content is preferred (1–4 wt.%); A phase transition occurs at around 183jC and may have practical merit; Tm decreases first and then increases with increasing Bi content; microstructures, and mechanical and wetting properties were studied; The 6 wt.% Zn alloy series had the highest reliability

62 63 64 65 58

15 50 66

66

RELIABILITY ASPECTS OF LEAD-FREE SOLDERS

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a. Microstructure. It should be noted that indium additions result in the formation of a rather undesirable microstructure (i.e., a much coarser and less uniform microstructure containing a platelike phase distributed in a dendritic matrix structure). However, this structure can be modified by small (0.1 wt.%) additions of Ag, as depicted in Fig. 23. The ductility of a Ag-doped Sn–9Zn–5In alloy is significantly improved, as shown in Fig. 24. b. Issues. Some issues with In-containing solders that need to be addressed are listed in Table 10. 2. Sn–Zn–Bi Alloys Tin–zinc alloys have been doped with other elements such as bismuth and indium to improve their mechanical and thermal fatigue characteristics. Tensile deformation properties of bismuth-doped and indium-doped alloys, and estimates of thermal fatigue resistance have been reported. Strain rate sensitivity and microstructures of as-cast samples aged at 120jC for 12 days were studied using cast and machined alloys. Bismuth-doped alloys exhibit their maximum strength at relatively large strains compared to either indium-doped or eutectic Sn–Zn alloys. Bismuth additions have a greater impact on strain rate sensitivity, in comparison to indium additions. A 5% addition of bismuth to a Sn–Zn alloy was found to remarkably increase the tensile strength and to lower the elongation from 40% to 20%, whereas the addition of indium had little or no effect on elongation. Thermal aging has an imperceptible effect on both elongation and tensile strength. Tin–zinc alloy systems were estimated to have better thermal fatigue resistance than Sn– Pb alloys, but were not comparable to Sn–Ag systems. Alloys doped with bismuth or indium display considerable coarsening of the acicular zinc phase [67]. Sn–Zn solders and zinc-bearing alloys have stable surface oxides that are difficult to remove.Alloys containing indium have the disadvantage of high cost, which precludes them from most mass soldering applications. However, Sn–Zn–Bi alloys are good candidates as lead-free solders in mass soldering applications for temperature-sensitive components. These alloys are commercially available and typically have melting points in the range of 190–200jC. Sn–Zn alloys have been used as solders for many years, primarily for soldering aluminum assemblies. The alloy has also found use in diverse applications such as high-performance corrosion-resistant coating for steel and as sprayed coatings for capacitor plates. Their use as interconnection alloys in electronics, however, is a more recent development [68]. a. Solidification Aspects. Most Sn–Zn–Bi solders investigated for electronics are close to the Sn–Zn binary eutectic composition with minor additions of bismuth [44]. The alloy is a

FIG. 23 Fracture surfaces of (a) Sn–8Zn–5In and (b) Sn–8Zn–5In–0.1Ag solder alloys. The Ag-free alloy displays undesirable platelike fracture features, and the Ag-containing alloy displays the more desirable ductile dimples on the fracture surface. (Courtesy of TMS, Ref. 64.)

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FIG. 24 Tensile stress–strain curves for Sn–8Zn–5In and Sn–8Zn–5In–0.1Ag Pb-free solders. The curve for the binary Sn–Zn solder alloy is also shown for comparison. (n IEEE 1995, Ref. 58.)

homogeneous liquid at soldering temperatures, with tin-rich and zinc-rich solid solution phases formed upon solidification, often as alternating lamella. Bismuth has appreciable solubility in tin, particularly at elevated temperatures, and may solidify as part of the tin matrix at low concentrations. However, bismuth solubility decreases rapidly as the solidified joints cool down, resulting in the precipitation of discrete Bi-rich particles [69]. The addition of bismuth to tin–zinc seems to improve wetting and corrosion performance, but, unfortunately, it also causes the potential for a problem often referred to as ‘‘hot shortness’’ [70]. Under certain conditions, pockets of Bi-rich liquids that have lower freezing temperatures than the matrix can be formed during solidification (Bi–Sn eutectic=139jC and Sn–Zn eutectic=199jC). These molten pockets remain in the liquid state long after the matrix around them has solidified. The combined effects of stresses due to contraction on cooling and the residual pockets of molten solder can give rise to cracking, or ‘‘hot tears.’’ The defect has been observed on wave-soldered joints with this alloy, but not on surface mount assemblies [70], possibly because the joint volumes are smaller. Reducing the bismuth content reduces the incidence of hot tears. b. Reactivity of Zinc. In most Sn-based solders, it is the tin that is the active element and dominates the interfacial reactions between the solder and adjacent metallized terminal pads. As a result, tin-based solders share many common properties. For example, oxides, which form at the surface of molten Sn-based solders, are almost exclusively those of Sn because they are

TABLE 10

Issues with Indium-Containing Solders

Issues . . . . .

Their wetting behavior in air and inert atmosphere with different fluxes. Mechanism by which Ag beneficially affects their microstructures. Sn–8Zn–5In–0.1Ag alloy requires investigation of its mechanical properties, microstructural evolution, wettability, and melting temperature and is compared with the Sn–Pb eutectic solder. Sn–6Zn–(2–5)In–(1–9)Bi alloys require further investigation to verify the results of the thermodynamic calculations in Ref. 66. The applicability of the Sn–Zn–In alloys to industrial mass production.


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thermodynamically more stable. However, because zinc is very active, the addition of zinc completely changes the picture. It has been observed that the outermost surface of molten 86Sn– 8.5Zn–5.5Bi solders is not dominated by tin oxide but by zinc oxide. Therefore, flux compatibility and drossing rate should be taken into account when using the Sn–Zn–Bi solder [69]. c. Reactions with Copper. It is generally known that Zn reduces the wettability of Cu by Sn-based solders. In addition to oxidation, zinc prevents tin from reacting with copper, which reduces the mechanical integrity of Zn-containing solder alloys with copper conductors [71,72]. During wetting balance tests of etched copper wires either in air or under a high-purity (H2O+O21000 hr) at 235jC, the metastable (Cu,Ni)6Sn5 phase decomposes into the ternary compound Sn45Cu29Ni26 and the stable IMC (Cu,Ni)6Sn5 [84]. However, in most practical applications, the Ni content of the metastable (Cu,Ni)6Sn5 is far above its equilibrium value, and there are indications that this compound becomes more brittle with an increasing nickel content [86]. Furthermore, a detailed microstructural characterization utilizing TEM was conducted on the relatively complex interfacial reaction layers generated between an ENIG finish and a eutectic Sn–Ag–Cu solder [86]. As shown in Fig. 30, the columnar two-phase (Ni3P+Sn) layer structure is formed between the amorphous Ni(P) and Ni63Sn25P12—the latter being in local metastable equilibrium with the (Cu,Ni)6Sn5 intermetallic layer. There are numerous voids in the vicinity of the ternary Ni63Sn25P12 phase. This ‘‘porous’’ Ni63Sn25P12 phase region is thought to make solder joints susceptible to brittle fracture known as ‘‘black pad.’’

FIG. 29 Metastable equilibria of the Sn–Cu–Ni system at 235jC. (From Ref. 84.)


805

FIG. 30 TEM micrograph showing the reaction layers formed during reflow soldering of an ENIG finished board with a eutectic Sn–3.4Ag–0.8Cu solder at 240jC.

D. Component Termination Finish 1. Lead Finish/Mold Compound Adhesion Component lead finish is another important aspect that must be considered to achieve total leadfree implementation. The matter of component and PWB finish is discussed in detail in Chap. 12. Both Ni/Pd and Ni/Pd/Au are among the major surface finish candidates for lead-free component leads. Important aspects to consider are the adhesion of a lead finish to the molding compound and process compatibility. The typical thicknesses for these finishes are listed in Table 13. Inadequate adhesion can lead to delamination, followed by moisture ingress and possible popcorn effect. 2. Ni/Pd-Based Finishes The finishes listed in Table 13 constitute the finish for lead frame materials such as copper CDA194. It is intended that the palladium and gold layers in the thin film stack are to be completely dissolved into the solder and a metallurgical bond established by forming Ni–Sn intermetallic compounds. It is important to recognize that x-ray fluorescence measurements may not be adequate to accurately measure gold thickness in the 0.003-Am range [87].

TABLE 13

Typical Layer Thickness for Ni–Pd and Ni–Pd–Au Finishes

Finish

Ni (Am)

Pd (Am)

Au (Am)

Ni/Pd Ni/Pd/Au

1 0.5–2.0

0.075–0.25 0.02–0.075

— 0.003–0.010

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3. Other Finishes Other possible lead finishes such as pure Sn, Ag, Ag–Bi, Ag–Pd, etc. are available. However, it is important to address such issues as tin whisker formation and silver migration before deciding on a given lead finish composition. Potential reliability-related aspects of finishes must be understood. The mechanisms and kinetics of potential tin-related reliability issues are discussed in detail in Chaps. 21 and 22.

E. Process Atmosphere Reactions The wetting of substrate metal by a solder alloy is strongly affected by the flux composition, partial pressures of oxygen and water in the atmosphere, chemical reactions between the solder and the substrate, dwell time above the melting temperature of the solder, and cooling rate. The primary reactions that take place during soldering are reduction–oxidation (red-ox) reactions between gaseous components, flux components, elements of a solder alloy, and substrate metallizations. There are sufficient active flux components in pastes to effectively reduce the oxidized surfaces of solder particles, components, and board metal finishes (except Au) to achieve adequate wettability so that typically only minor oxide film fragments enter into solder joints. However, this state of affairs can change markedly, as the solder joints become much smaller as occurs in reflow soldering of 0201-type passive components. Moreover, most lead-free solders are somewhat more strongly oxidizing due to the higher chemical activity of tin in these alloys and lead-free finishes such as pure tin. Therefore, the fluxes may need to be more reactive in lead-free solder pastes. However, owing to decreasing volumes of printed solder pastes, the flux solvents evaporate more rapidly during the initial heating phase of reflow soldering operations. The fluxes are therefore less effective in removing oxide skins of metal particles in solder pastes, and this impairs proper fusion of the particles during reflow in air (Sec. VI.A). Therefore, the reflow soldering of small components on high-density PWBs will start gradually to require the employment of inert gases such as nitrogen (discussed in detail in Chap. 15) with low oxygen and moisture contents. Expenses will increase with the level of inerting purity needed to produce consistent solder joints. In such a situation, the small addition of a reducing component such as hydrogen into an inert gas can be a viable option. For example, in the case of Sn–Pb–Ag and most lead-free systems, the most stable oxide is SnO2(s), which can be reduced by nitrogen/hydrogen mixture according to the following reaction: SnO2 ðsÞ þ 2H2 ¼ ½Sn þ 2H2 O

ð12Þ

provided that the pressure ratio pH2O/pH2 at the activity level of the metal [Sn] in the alloy is sufficiently low at the soldering temperature. 1. Effect of Gas Contaminants and Mixtures In principle, a pure reducing gas has an immense reducing power. In practice, however, all gases contain some moisture and other gas components that impact their reducing ability relative to oxide films on the surfaces of solder particles, liquid solders, and solid metal substrates. The limiting conditions for oxide reduction by a gas can be considered using metal oxide equilibria and the corresponding equilibrium composition of the gas mixture. It can be shown that hydrogen or its mixture with some inert gas such as nitrogen (or argon) can reduce an oxide film if the moisture content of the gas is: pffiffiffiffiffiffi Kp ð1 xN2 Þ xH2 O V pffiffiffiffiffiffiffi pffiffiffiffiffiffi aSn þ Kp

ð13Þ

where XN2 is the mole fraction of nitrogen, aSn is the activity of tin in solders, and Kp is the equilibrium constant [71]. The water content of a 20% H2/80% N2 gas mixture must be very low in order to satisfy the thermodynamic conditions necessary for the reduction of oxides to be fulfilled at 250jC without any flux.


807

2. Effect of Cooling Rates Dwell time above the melting temperature will increase the dissolution of minor elements into the solder and can increase the intermetallic compound growth rate in the solder joints. Further growth of intermetallic compounds will occur with slow cooling rates. An increase in the growth rate of Ag3Sn plates in Sn–Ag–Cu solders has been reported as significant with cooling rates slower than 3jC sec1 [88]. Cooling rapidly past the solidification temperature minimizes the formation of Ag3Sn plates. At cooling rates slower than 1.0jC sec1, the primary Ag3Sn plates can subtend the entire cross section of a BGA solder joint. Cooling issues are compounded by the thermal capacity of each solder pad. At sufficiently high cooling rates, the time interval during which the Ag3Sn plates can grow in the liquid phase before final solidification is sufficiently reduced to limit plates to a small size. The critical cooling rate for the desired plate size suppression may vary with the size of the solder joint and the nucleation kinetics of the solder mass. Solder joints that are connected to a ground plane can have a much significantly slower cooling rate due to the thermal lag of cooling the ground plane. This results in a disparity in the amount and size of intermetallic compounds formed in adjacent solder joints. Cooling rates should be sufficiently fast to minimize the intermetallic growth rates in spite of the differences in the thermal properties of solder joints across the assembly. More effective cooling also tends to result in a finer solder joint microstructure.

VII. BOARD LEVEL RELIABILITY A. Ceramic BGA (CBGA) Components with Pb-Free Interconnections Farooq et al. conducted a comparative thermal cycle test involving the reliability of ceramic BGA assemblies thermal cycled in the range of 0–100jC on 1.5- and 2.9-mm-thick FR-4 boards evaluating eutectic Sn–Pb, Sn–3.8Ag–0.7Cu, and Sn–3.5Ag–3.0Bi solders. The experimental matrix consisted of CBGAs with standard 10Sn–90Pb, Sn–3.8Ag–0.7Cu, as well as Sn–5Sb solder balls. The card attach solder paste was Sn–3.8Ag–0.7Cu with 63Sn–37Pb solder paste used as a control. Thus, the study consisted of both single-melt and dual-melt BGA joints. The interconnection reliability data were normalized to a standard joint height. The fatigue life of these lead-free solders is reported to be 1.5–1. 6 for Sn–Ag–Cu and 1.9–2.0 times better for Sn–Ag–Bi interconnections compared to eutectic tin–lead for the board thickness evaluated in the test. The thin boards or cards exhibited a 1.7 greater life compared to thick boards, apparently due to their lower stiffness. A thin layer (2 Am) of Ni–Cu–Sn intermetallic compound was observed to adhere to the Ni surface, with Ag3Sn and Cu6Sn5 dispersed throughout the solder matrix. Crack propagation was found to occur within the bulk solder of the joint instead of along the intermetallic compound zone [89]. In another study, Lee et al. [90] investigated the microstructure, joint strength, and failure mechanisms of Sn–3.5Ag and Sn–3.5Ag–0.5Cu vs. 62Sn–36Pb–2Ag ball grid array solder joints on electroless nickel/gold surfaces 6 and 1 Am thick, respectively. In addition, the samples were aged at 150jC for 1000 hr. The key findings were: In the Sn–Pb–Ag system, AuSn4 was found throughout the ball and some AuSn3 was also present. In contrast, Ag3Sn was observed to distribute along the Sn-rich phase boundaries in lead-free systems as fine particles or elongated structures. A thin layer of Ni–Sn intermetallics formed at the interface in the case Sn–Ag and Sn–Pb–Ag solders. On the other hand, Ni–Sn–Cu intermetallics are formed in the case of Sn–Ag–Cu solders. Most of the AuSn4 formed during the initial joint formation in the Sn–Pb–Ag system disappeared from the bulk after aging and tended to resettle at the substrate interface as (Au,Ni)Sn4 pseudoternary compound layer as reported by Zribi et al. [91], Ho et al. [92], and Minor and Morris [93]. The ternary intermetallic constitutes a contiguous layer on top of the Ni3Sn4 layer after 48–96 hr of aging at 150jC. In the Sn–Ag system, the layers adjacent to the pad were found to be Ni3Sn4 and AuSn4; together, they transform to a platelet compound (Au,Ni)Sn4. Thermal aging results in the

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formation of a quaternary intermetallic compound with a nominal atomic composition of 48Cu– 39Sn–10Ni–2Au. The intermetallic compounds AuSn4 and Cu–Sn–Au are both found in the bulk solder, whereas Cu–Au–Sn tends to settle toward the Sn-rich phase boundary and AuSn4 is dispersed throughout the bulk of the joint. Sn–Ag and Sn–Ag–Au joints tend to fail due to plastic deformation within the bulk solder even after 1000 hr of thermal aging at 150jC. In Sn–Pb–Ag systems, the initial failures are ductile fractures, but after aging, they are due to brittle fractures at the interface. Thus, in both lead-free systems, the failure mechanism is attributed to be a purely ductile fracture.

B. Plastic BGA (PBGA), Thin Small Outline Package (TSOP), and Other Components Mounted with Sn–Ag–Cu Solder Woosley et al. investigated the board level reliability of area array packages with pitches in the range of 0.8–1.27 mm, with various ball compositions, a range of substrate thickness (0.26–0.56 mm), leaded packages such as TSOP, and leadless passive components. Liquid-to-liquid thermal cycling (55jC to 125jC) tests were conducted with 388 I/O PBGA assemblies and air-to-air thermal cycling in temperature ranges 0jC to 100jC, 40jC to 125jC, and 50jC to 150jC. In a 40jC to 125jC accelerated thermal cycle (ATC) test of 324 I/O PBGAs, Sn–Pb–Ag solder balls joined with eutectic tin–lead paste and Sn–Ag–Cu paste performed almost identically based on the two-parameter Weibull plot with g values of 4555 and 4575 cycles, and b values of 15.19 and 14.76, respectively. On the other hand, assemblies with Sn–Ag solder balls attached with Sn– Ag–Cu paste performed the best with an g value of 8690 cycles and a b value of 8.1, as shown in Fig. 31 [94]. In a separate comparative study, accelerated thermal cycle tests (40jC to 125jC, 14-min dwell time at the extremities, with a 3-min transition) indicated that Sn–3.8Ag–0.7Cu exhibited equivalent or slightly better reliability compared to Sn–Pb–Ag solder interconnections, as shown in Fig. 32. The components (289 I/O, 0.8-mm pitch, PBGAs) were assembled on FR-4 boards

FIG. 31 Two-parameter Weibull plot of board level thermal cycle test results for a 324 I/O PBGA component. (Courtesy of Motorola Lead-Free Team Ref. 94.)


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FIG. 32 Temperature cycle test results for a 289 I/O PBGA component mounted on a FR-4 RC Cu board with via-in-pad configuration utilizing several solder/pad combinations.

using OSP and Ni/Au surface finishes. Typical failures in Sn–Pb and Sn–Ag–Cu solder joints were at the component/solder interface [95]. Prasad et al. conducted board level reliability studies of lead-free packages with several ball metallurgies assembled utilizing Sn–3.9Ag–0.6Cu and eutectic tin–lead for comparison purposes. The lead-free ball metallurgies included Sn–3.5Ag, Sn–0.7Cu, Sn–4.0Ag–0.5Cu, and Sn–2.5Ag– 1.0Bi–0.5Cu. The results indicated that the assemblies with the Sn–Ag–Cu alloy ball metallurgies outperformed the traditional Sn–Pb solder assemblies by a factor of two. Typical Weibull plots for 208 I/O and 72 I/O (M2CSP), 0.8-mm-pitch, chip-scale packages assembled on 1.5-mm-thick polyclad boards are shown in Figs. 33 and 34 for the two temperature ranges [96]. In the case of 48 I/O TSOP components with Sn–Pb-plated Alloy 42 lead, interconnections with eutectic tin–lead solder seem to outperform those with Sn–Ag–Cu joints in all the temperature ranges investigated, namely, 55jC to 125jC, 40jC to 125jC, and 0 to 100jC, respectively, as shown in Fig. 35 [94]. Comparison of thermal cycle tests performance utilizing various temperature ranges to evaluate 2512 resistors indicates that eutectic Sn–Pb and Sn–Ag–Cu pastes exhibit a similar performance at higher temperature ranges. In the 0–100jC temperature range, lead-free solder interconnections exhibit a slope that is less steep (2.84 vs. 7.1), and a slightly larger g value (3063 vs. 2256), as shown in Fig. 36 [94].

C. Effect of Aging on Several Types of Lead-Free PBGA Joints The effect of alloying metals on the aging behavior is not well understood. In addition, thicker intermetallic compound layers due to higher processing temperatures and attendant fragility of the solder joints are concerns. It has been reported in a thermal aging study conducted at 125jC and 150jC with 388 I/O plastic ball grid array package assemblies that the intermetallic compound layer growth is comparable to that in eutectic Sn–Pb solder joints. The activation energy for several Pb-free solders is given in Table 14. The activation energy is lower when the

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FIG. 33 Weibull plot for a 208 I/O CSP accelerated thermal cycle tested in the temperature range from 40jC to 125jC. (Courtesy of Chip Pac, Alpha Metals, Sanmina, Hewlett Packard, Ref. 96.)

silver content is higher. Sn–Ag and Sn–2.5Ag–1.0Bi–0.5Cu alloys exhibited fracture loads in the range of 600–700 N and those containing Cu, namely, Sn–Ag–Cu and Sn–Cu, exhibited fracture loads in the range of 900–1050 N. The fracture caused pad pullout from the epoxy at the glass/ epoxy interface. The initial IMC thickness in the case of both eutectic Sn–Pb and lead-free interconnections was in the range of 1.6–2.3 Am. Based on limited data, the IMC layer growth rate does not appear to be greater than for eutectic Sn–Pb solder. Due to higher reflow tem-

FIG. 34 Weibull plot for a 72 I/O M2CSP accelerated thermal cycle tested in the temperature range from 0jC to 100jC. (Courtesy of Chip Pac, Alpha Metals, Sanmina, Hewlett Packard, Ref. 96.)


811

FIG. 35 Weibull plot for a 48 I/O TSOP with Alloy 42 leadframe material and various solder paste combinations, and thermal cycle temperature ranges. (Courtesy of Motorola Lead-Free Team, Ref. 94.)

FIG. 36 Weibull plot for a 2512-sized 0-V resistor for various solder paste and accelerated thermal cycle test conditions. (Courtesy of Motorola Lead-Free Team, Ref. 94.)

812 TABLE 14

VISWANADHAM ET AL. Activation Energies for Several Solder and Ball Compositions

Solder ball solder

Attachment solder

63Sn–37Pb Sn–2.5Ag–1.0Bi–0.5Cu Sn–4.0Ag–0.5Cu Sn–3.5Ag

63Sn–37Pb Sn–4.0Ag–0.5Cu Sn–4.0Ag–0.5Cu Sn–Ag–Cu

Activation energy (kJ/mol) 45 68 33 31

peratures, the IMC layer at the lead-free solder interface is slightly thinner compared to eutectic Sn–Pb solder. The joint strength is equivalent to Sn–Pb joints as determined by four-point bend tests [97].

D. Effect of Bi and Cu on Quad Flat Pack (QFP) Attached with Eutectic Sn–Ag Karia et al. studied the effect of additional elements such as Bi and Cu on the thermal fatigue life of QFP with Sn–3.5Ag solder joints. The 100-pin quad flat packs consisted of Alloy 42 leads with a Sn–20Pb solder finish. It was determined that the addition of bismuth (in the range of 2–5%) to the Sn–3.5 Ag solder degrades the thermal cycle reliability of the solder joints in the 100jC and 160jC DT range, resulting in a 20% reduction in joint pull strength. Cracks were reported after 200 thermal cycles. Solder joints with Sn–3.5Ag–(0.5–1%)Cu showed an initial higher joint pull strength by 11–22% compared to the eutectic Sn–Pb solder. The Pb phases at the lead/solder interface and bismuth in the Sn–Ag solder may promote crack propagation at the interface and reduce thermal fatigue life. From an isothermal fatigue point of view, an upper limit of 2% bismuth was suggested for Sn–Ag solders to achieve the same fatigue resistance as eutectic tin– lead solders [98].

E. Passives Attached with Pb-Free Solder Vianco and May evaluated large 1825 and 1205, and smaller 0805 chip capacitors in a study using Sn–3.5Ag, 58Bi–42Sn, and Sn–3.3Ag–4.8Bi solders. The component termination configuration was 100% Sn-plated finish over Ni or Cu barrier layers. A 127-Am-thick stencil was used for solder paste printing and test boards were reflowed in a four-zone oven using a peak temperature between 20jC and 30jC above the melting point of the solders. Wetting of the package terminations was reported to be excellent and, overall, the manufacturing feasibility analysis determined that there were no catastrophic assembly problems. A higher propensity for component misregistration was observed with Bi–Sn solders, was attributed to poor self-alignment during reflow, and was consistent with the lower surface tension of the Bi–Sn solders. Increased voiding was observed in solder joints on the larger packages. Mechanical strength of the solder joints was evaluated by applying shear loads on components as fabricated and after thermal cycling (55jC to 100jC, 66-min cycles) and liquid-to-liquid thermal shock (155–100jC, 15-min dwell). The strength values of the solder remained largely unchanged for the larger chips. However, test results for the 0805 chips indicated that the position of the components on the assembly and their proximity to other components can impact solder joint reliability [99].

F. Mixed Composition Systems Dunford et al. investigated the metallurgical and reliability aspects of Pb-free mixed technology electronic assemblies utilizing area array packages, leadless ceramic chip passives, and small outline integrated circuits with conventional Sn–Pb terminations on a thin PWB with Au/Ni surface finish utilizing Sn–Ag–Cu–Sb solder as the interconnection alloy. The Sn–Au and Sn–Ag


813

intermetallic compounds were generally well dispersed in the solder joints. Solder joints made with gold-terminated components were observed to have regions of Sn–Cu intermetallic compounds separating the two gold-rich regions, as shown in Fig. 37. A greater concentration was observed to exist on the component side compared to regions on the board side. Detectable amounts of W, Co, Ni, Pt, Pd, and Fe in the bulk solder of these solder joints suggested dissolution and dispersion of the component terminations into the joint due to the higher reflow temperature compared to the Sn–36Pb–2Ag solder. The joints generally exhibit a mottled surface finish and surface voids, which are quite different from those found in conventional eutectic tin– lead solder joints. Thermal cycling and mechanical drop testing indicated that the Sn–36Pb–2Ag control samples experienced failures earlier than the lead-free samples [17].

G. Effects of Process Parameters and Metallurgical Interactions Second-level reliability was evaluated by Bartelo et al. [36] using Sn–Ag–Cu and Sn–Ag–Bi solders. Pad metallurgies employed are generally Ni/Au or Cu with OSP. Ag3Sn and Cu6Sn5 intermetallic compounds seem to interfere with the formation of a smooth shiny surface during the solidification process, resulting in a grainy and matte appearance. Some of the pad metallurgies at each of the joint interfaces (Ni or Cu) dissolved into molten solder and tended to migrate from one interface to the other, forming Cu6Sn5 on board pads metallized with nickel, probably substituted for Cu. In addition, copper from the solder alloy and that due to the dissolution process found their way to the nickel interface. Thus, intermetallic compounds containing Cu, Ni, and Sn can be found on the nickel surface. The Sn–Ag–Bi alloy was determined to be devoid of copper. Copper from the board pads generally finds its way into the bulk solder and forms Cu6Sn5. Reliability studies indicate that the number of cycles to failure decreases with an increase in temperature range and peak temperature, and a decrease in cycle time. Fatigue failure times (N50) in the 0–100jC range for the lead-free solder investigated was determined to be approximately twice that of eutectic tin–lead solder. Significant differences exist in the reliability data obtained in the 0–100 and 40jC to 125jC range. At larger dwell times (240 min/cycle), the lead-free solder joints performed worse than the control eutectic Sn–Pb samples,

FIG. 37 Vertical cross section of an LCCC solder joint showing two distinct zones within the solder joint separated by a Cu/Sn layer (500). (Courtesy of Nokia Corp.)

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only exhibiting 75% of fatigue life of Sn–Pb samples. Some variations in the results of the various studies suggest differences in the module and board characteristics (causing cyclical strains) and stress states to be different. The higher temperature range of 40jC to 125jC, or the increased temperature range of 165jC, can have a dramatic effect on the relative fatigue behavior: a larger strain range and increased fatigue damage, or both. Cross-polarizer optical microscopy revealed thermomechanical stress-induced microstructural changes in the vicinity of fatigue cracks generated during ATC testing in lead-free alloys. These changes are believed to be the result of plastic deformation and were verified by finite element analysis.

VIII. FLIP CHIP-RELATED ISSUES A. Characteristics of Some Lead-Free Solder Bumps There is no readily available drop-in replacement for wafer scale bumping of either the high-Pb or eutectic Sn–Pb solders. As explained before, the process can be further complicated by the metallurgical interactions of the solder alloy with the silicon device UBM, metal finishes on PWBs, and thermal compatibility issues. The possibility of the UBM forming low-melting eutectics or other phase mixtures with the constituents of lead-free alloys cannot be ruled out and must be carefully evaluated. Moreover, many of the metal finishes on PWBs dissolve rapidly into liquid solders and can therefore take part in the reaction between the ‘‘new’’ solder alloy and underbump metallizations, thus complicating the prediction of metallurgical reactions significantly. Several interconnection alloys have been explored recently to find a replacement for the eutectic tin–lead. These include binary alloys Sn–3.5Ag and Sn–5Sb, ternary alloys Sn–20In– 2.8Ag (Ind alloy) and Sn–10In–3.1Ag, and quaternary alloys such as Sn–10In–3.1Ag–1.0Cu, Sn– 20In–2.8Ag–1.0Cu, etc. A replacement for the high-temperature Pb–5Sn alloy is still an open issue. There have been some attempts to find a replacement but with limited success.

B. Effects of Assembly Processes Variables in the assembly process can have a significant effect on the long-term reliability of a Pbfree assembly. Most are associated with high-temperature soldering operations that can detrimentally affect the components. Secondary reflow and moisture sensitivity are initial concerns, along with parameters affecting intermetallic growth rates. Assembly issues are discussed in detail in Chap. 14. 1. Secondary Reflow Most surface assembly practices are likely to be centered around the Sn–Ag–Cu and Sn–Ag–Bi alloys with a melting temperature in the vicinity of 215jC. As a consequence, the melting points of solders internal to the package must be approximately 260jC to prevent the bumps from undergoing secondary reflow during the surface mount assembly process. Secondary reflow of the bump metallurgy can have undesirable consequences such as deformation and extrusion of solders within the underfill (i.e., solder migration), resulting in reduced assembly reliability and even electrical shorts. In these cases, thermal migration of solders from the component’s cold side to the hot side due to solid-state diffusion is not uncommon, and could result in voids or electrical shorts. 2. Other Factors and Issues Flip-chip assemblies using 10 10 and 5 5 mm chips attached to a halogen-free microvia board with Sn–3.5Ag–1.0Cu solder bumps (115 Am in diameter) were reported by Baynham et al. [100]. Nitrogen reflow was demonstrated to provide superior wetting. It was indicated that the solder shapes varied considerably in air reflow attendant with partial or incomplete wetting. Solder shapes, as shown in Fig. 38, can have a significant impact on reliability especially on joints at a large distance from the neutral point (DNP). It was also indicated that flip chips attached with lead-free solder can have a narrower process window compared to those attached with eutectic


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FIG. 38 Vertical cross section of a flip-chip attachment showing partial wetting of the pads. (Courtesy of Auburn University.)

Sn–Pb solders. This was attributed to the higher reflow temperature and the inadequacy of flux activity, joint starvation, pad oxidation, mask height tolerance issues, reflow atmosphere, etc. that can have significant influence on the structure of lead-free solder joints compared to chips joined with eutectic Sn–Pb solder. For example, at higher temperatures, flux may decompose or burn off prematurely. Alternatively, at higher temperatures, wetting quality may be affected [100]. Weise et al. [101] studied microstructure differences between flip-chip solder joints of Sn–Ag–Cu and samples six orders of magnitude bigger than flip-chip joints. The bulk samples exhibited nonuniform phase distribution and a larger grain structure. The flip-chip joints were more uniform with a much smaller grain size than the bulk solder because rapid solidification suppresses grain growth.

C. Effects of UBM and Lead-Free Solder Interactions 1. Sn–Ag and Sn–Ag–X Solders Interactions that a lead-free interconnection alloy may have with underbump metallurgy can have an impact on the integrity of flip-chip solder joints under both thermal and mechanical loading. Puttlitz and Totta [22] investigated the effect of TiW/Cu, Ni, and electroless nickel/gold underbump metallurgies with Sn–3.8Ag, Sn–3.8Ag–0.7Cu, and Sn–2Ag–2Bi solders on the system’s metallurgical characteristics, and on thermal and isothermal mechanical reliability. No large intermetallic compounds were observed to form with Sn–Ag–Bi solders and Ni–P UBM. Copper-containing solders produced an interfacial intermetallic compound containing Ni4Cu7Sn6, whereas those without Cu produced only Ni3Sn4. The presence of Cu generally affects interfacial adhesion. In joints without Cu, the needlelike Ni3Sn4 compound that formed caused a loss of adhesion and spalled into the solder. The adhesion is considered good in the case of Cu-plated compounds due to the formation of Cu3Sn and Cu6Sn5 at the interface. The bulk solder has Ag3Sn and Cu6Sn5 in the Sn–Ag–Cu solder system, but only Ag3Sn in the Sn–3.5Ag solder system. The interfacial intermetallic compounds in the Cu-containing solders are more uniform and stable than those that form with no Cu. In the case of Ni–P/Sn–3.5Ag bumps, the Ni3Sn4 IMC is observed to grow and it is also the region where fractures occurred. The case of Sn– 3.5Ag reflowing twice at 260jC and annealing at 150jC for 1000 hr resulted in a more extensive

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intermetallic compound formation compared to reflowing only once. The IMCs changed from a blocky to a planar morphology, suggesting differences in growth mechanisms between the liquid and the solid states. The solid-state IMC growth is considered to be predominantly an interfacial phenomenon [102]. 2. Sn–Bi and Sn–Pb Solders Jang and Park investigated the effect of a reflow process and UBM systems on the growth of intermetallic compounds for Sn–Bi and Pb–Sn solder bump/UBM interfaces. The UBM systems sputtered were Al/Ti/Cu, Al/Ti/electroless plated Cu, Al/NiV/sputtered Cu, and Al/electroless plated Ni/Au, and the solder bumps were eutectic Sn–Pb and Sn–Bi fabricated by electroplating. The microstructure and composition of IMC phases and their morphology were examined. The IMC compositions that formed were Cu6Sn5, Cu3Sn, Ni3Sn4, or Ni3Sn, depending on the UBM and solder compositions. The effect of IMC growth on bump adhesion strength was also investigated using a ball shear test. The ball shear strength was determined not to depend on processing conditions but on the solder/UBM interfacial adhesion strength in this case [103]. 3. Solid-State Interactions The growth mechanism in the solid state differs from the liquid state; the planar type is perhaps due to interfacial reaction. Lim et al. also investigated the tensile properties of Sn–3.5Ag, Sn– 0.7Cu, Sn–3.8Ag–0.7Cu, and eutectic Sn–Pb solders in the temperature range of 40jC to 180jC. It was reported that Sn–3.8Ag–0.7Cu showed the greatest strength at lower temperatures, but decreased significantly at higher temperatures. The tensile strength of Sn–3.5Ag decreased to the Sn–3.8Ag–0.7Cu level at higher temperatures. In addition, the tensile strength of Sn–0.7Cu is lower than the other two alloys at lower temperatures, but is slightly higher at higher temperatures. Results of tensile tests carried out on flip-chip bumps are reported to be similar to the tape samples. The differences between Sn–3.5Ag and Sn–3.8Ag–0.7Cu are insignificant. 4. Interfacial Failures The alloy Sn–0.7Cu exhibits the best fatigue life with a copper or nickel UBM. A Ni–P UBM with a Sn–3.8Ag–0.7Cu solder exhibited the lowest fatigue life due to interfacial failures also exhibiting similar failures with a Cu UBM. Failures observed for tin–lead solders occurred close to the solder/IMC interface, as is the case for the Sn–3.8Ag–0.7Cu and Sn–3.5Ag solders. In some cases, failures are through the IMCs themselves, or through the solder bulk for Sn–0.7Cu [102].

D. Effect of Creep on Pb-Free Flip-Chip Solder Joints The effect of composition differences on the microstructure of Sn–Ag–Cu solders is due to precipitates dispersed in the h-Sn matrix, which impart a higher creep resistance via precipitation strengthening or precipitation hardening. Steady-state strain rate data indicate that the Sn–Ag– Cu solder has a better high-temperature creep resistance than either Sn–Pb–Ag or Sn–3.5Ag. A large number of small particles offer greater resistance to the movement of dislocations, rather than a small number of large particles. Flip-chip Sn–Ag–Cu solder joints have been shown to exhibit much lower creep rates than bulk specimens. A comparison of creep-related parameters for flip-chip solder joints and bulk specimens is given in Table 15. The coefficient of thermal expansion (CTE) of Sn–Ag–Cu solder is 17.6 ppm/jC in the 20–150jC range, 16.7 in the 20– 100jC range, and 18.8 in the 100–150jC range [101].

TABLE 15 Creep-Related Parameters for Bulk Sn–Ag–Cu Solder and Flip-Chip Solder Joints Parameters Creep exponent Activation energy (kJ/mol)

Bulk solder 10 57.1

Flip chip 18 83.1


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E. Thermal Fatigue Characteristics of Some Lead-Free Solder Bumps ATC tests of flip-chip assemblies using cycles 50jC to 150jC and 40jC to 125jC have been reported in the literature. Flip-chip assemblies with Sn–3.5Ag solder bumps and no underfill materials were determined to be only 30–60% as reliable as corresponding assemblies with eutectic Sn–Pb eutectic solders. Quaternary alloy assemblies with Sn–2.5Ag–0.8Cu–1.0Sb solder bumps were about 80% as reliable as their eutectic Sn–Pb counterparts. Ternary Sn–20In–2.8Ag has been shown to form a low-melting Sn–In eutectic (113jC) that deteriorated the underbump metallurgy. The quaternary alloy Sn–10In–3.1Ag–1.0Cu, on the other hand, exhibited almost 1.5 times better reliability than the eutectic Sn–Pb system with no underfill enhancement. The absence of the low-temperature Sn–In eutectic was attributed to the low indium content of the alloy and the presence of copper. In addition, the soft and ductile nature of indium was suggested to contribute to the reliability [104]. However, DSC studies suggest that the improved reliability is predominantly due to the indium content, rather than the presence of copper. The reduction of indium from 20% to 10% and the addition of 1% copper may have shifted the reaction equilibrium in a manner that reduces the propensity for the formation of low-melting eutectic Sn–In. 1. Lead-Free Flip Chips on Ceramic and Organic Carriers Jung et al. investigated Sn–9.5Bi–0.5Cu (mp 198jC), Sn–3.5/Ag (mp 221jC), Sn–0.7Cu (mp 227jC), and Au–20Sn (mp 280jC) as potential wafer bumping alloys for flip-chip applications. The Au–Sn alloy was used for flip-chip attachment to a ceramic, whereas the other alloys were used with an organic laminate carrier. The underbump metallurgy in each case was Ni/Au, and the substrate surface finish for the organic laminate carrier was Ni/Au, but was a Pd–Ag–Au thick film metallization for the low temperature cofired ceramic (LTCC) carrier. The bump heights were in the range of 85–130 Am in all cases. All the assemblies were fabricated using a no-clean flux, underfilled and cured. Thermal shock was performed by utilizing a 55jC to 125jC range with a 10-min dwell at the extremities. The Sn–Ag assemblies showed failures earlier than 1000 shock cycles and could not be evaluated due to assembly defects of underfill delamination caused by flux contamination. The Au–Sn assemblies on LTCC lasted in excess of 4000 cycles. Preliminary results indicate promise for the use of these alloys for chip-scale packages and flipchip applications [105].

F. Evolution of Microstructures in Highest-Density Assemblies Constant demands for higher functionality and enhanced portability are driving manufacturers to increase the performance and to reduce size of their products. All these must be done without sacrificing reliably or increasing cost while meeting increasing technical challenges (e.g., in paste printing, component alignment, substrate technologies)—in addition to the environmental regulations. More fundamentally, increasing miniaturization will decrease solder joint volumes and reduce the thickness of metallization layers. There is a growing risk that the microstructure of small Sn–Pb solder bumps will result in a significant reduction in reliability in the field. Fig. 39a gives an example of such a decomposition of the originally uniform microstructure during the thermal cycling of flip-chip assembly. Fig. 39b shows crack propagation through the Pb-rich region of an LCC solder joint after 500 cycles from 40jC to 85jC. It is imperative that the effects of dissolution processes, intermetallic formation, and phase transformations be thoroughly understood and controlled, especially when the solder volume decreases to 104–105 mm3 or less. This is due to the fact that the solder joint strain is inversely proportional to the solder joint height. As the solder volume decreases, so does the joint height, resulting in a corresponding increase in solder joint strain. Another important reliability-controlling factor to be taken into account in designing and manufacturing the highest-density electronics in the future is the physical and chemical compatibility of the materials [106,107]. This is becoming increasingly more significant because the number of combinations of lead-free solder alloys, component termination metallizations, and board finishes is increasing—a fact that has been emphasized in this chapter.

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FIG. 39 (a) Flip-chip joint after 2500 thermal cycles from 25jC to 125jC, at 1 hr/cycle. (n IEEE 1998, Ref. 1a) (b) Microstructure of an LCC joint after 500 cycles from 40jC to 85jC, at 32 min/ cycle, showing crack propagation through the Pb-rich region. (Courtesy of Nokia Corp.)

1. An Example: Tin–Bismuth–Copper System When very small solder joint volumes are involved, the major component of the solder alloy (tin) reacts preferentially with the conductor metals (copper or nickel), gradually depleting tin from the alloy and thereby altering the original alloy composition. This occurs in most solder systems because the most commonly used alloying elements (lead, silver, or bismuth) do not react with either copper or nickel, and therefore are not consumed in solder reactions. In considering the Sn– 58Bi/Cu system, two cases can be distinguished: (a) the intermetallic compound reactions during reflow soldering, when the solder is in the liquid state, and (b) the further growth of the intermetallic compounds Cu6Sn5 and Cu3Sn during solid-state aging at ambient or elevated temperatures. a. Liquid State Interactions. During soldering, when the eutectic liquid Sn–Bi solder comes in contact with Cu (along the contact line, CL in Fig. 40), the local equilibrium is quickly

FIG. 40 Isothermal section of the Cu–Sn–Bi system at 235jC. Note the connection line (CL). (Courtesy of Helsinki University of Technology.)


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established between the saturated liquid A and the Cu6Sn5 (D) compound. The dotted line, from A to Cu6Sn5, indicates the local equilibrium. After Sn has diffused through the Cu6Sn5 layer, pure, solid Cu becomes saturated with Sn, resulting in the nucleation and growth of the solid Cu3Sn. These IMC reactions, being well established experimentally, are responsible for chemical bonding of solid solder to copper. What is not so well known is the sequence of events occurring with very small solder volumes, or, alternatively, during a very long time at temperatures where the diffusion of tin is limited [14,91,92]. If the amount of liquid is relatively small, it becomes enriched in bismuth because only tin is consumed in the IMC reaction. Therefore, the local composition of the liquid moves along the Cu6Sn5/liq phase boundary (i.e., liquidus projection) toward the Bi corner of the phase diagram. Accordingly, the diffusion path passes the apex of the three-phase triangle (liq+Cu6Sn5+Cu3Sn) (i.e., point C), and therefore the Cu6Sn5 reaction will stop and the Cu6Sn5 layer starts to disappear from the (liquid) joint structure, mainly by transforming into the Cu3Sn (q) compound [14,91]. With increasing reflow time, the liquid solder will be in equilibrium with pure bismuth so that the diffusion path is through the Bi corner (Fig. 40). According to this local equilibrium condition, a pure bismuth layer is formed between the liquid solder and the Cu3Sn phase. b. Solid-State Interactions. On the other hand, during solid-state aging of a larger joint, tin is consumed near the solder IMC interface because of slow diffusion of tin in the solid solder alloy. A layer rich in bismuth is formed between the Cu6Sn5 layer and the bulk solder. The bismuth layer acts as a diffusion barrier for tin and so retards further growth of the IMCs. The corresponding microstructure is generated when the solid eutectic Sn–Pb solder is in contact with copper at elevated temperatures. In fact, this is what happens during thermal cycling, as shown in the flip-chip solder in Fig. 39a. In this joint, practically pure lead layers are formed during thermal cycling (for 2500 cycles from 25jC to 125jC) between the bulk eutectic Sn–Pb solder bump and Cu UBM layer on the component side, and the Cu pad on the board side. It is interesting to note that the Cu3Sn layer (which only can be in local equilibrium with the Pb-rich phase) is thinner on the bump side where the crack is located. These examples point out the importance of understanding the local equilibrium in the microjoints and associated metallurgies in assessing reliability. This is especially true where solder bump volumes, thin metallizations, or kinetic constraints responsible for the redistribution of the second phase (e.g., Bi or Pb) in the solder/conductor system are involved.

G. Transfusion-Bonded Pb-Free Flip Chips A method of avoiding the decomposition of very small solder joints into brittle microstructures was reported recently by Kivilahti and Kuloja¨rvi [108]. It is essential that this Pb-free lowtemperature interconnection technique not be based on the formation of IMCs, but rather on the formation of high-tin microjoints when chemically coated tin and bismuth layers on Cu conductors are fused together. The silicon devices are bumped with pure tin followed by a thin chemical coating of bismuth, as are the conductor pads on flex or rigid chip carriers. The devices’ bumps are placed on the mating chip carrier pads and are heated to temperatures above 139jC. The fusion is initiated at both Sn–Bi interfaces in contact with each other as the eutectic temperature of the Sn–Bi system is attained. During solidification, tin-rich joints with finely dispersed Bi particles are formed. The transient nature of the fusion means the second macroscopical melting will occur at temperatures that are close to the melting point of pure tin. The mechanical properties such as strength, ductility, etc. of the joints are controlled primarily by the temperature of bonding and the cooling rate. The flip-chip joints with appropriate underfilling have been shown to survive 1000 cycles in thermal shock testing in the range of 40jC to 125jC without failures [109,110].

IX. SUMMARY A number of Pb-free binary, ternary, and quaternary alloys with combinations of potential candidate elements solder alloys have been surveyed and the differences in their mechanical

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properties were discussed. These included those containing copper, tin, silver, bismuth, indium, and germanium. Among the lead-free solders, the tin–copper–silver-based alloys seem to stand out as potential replacements for tin–lead solders, although none of them is a drop-in substitute. The composition of the interconnections formed can be significantly different from the initial alloy due to interaction with the component lead termination metallurgies and the printed wiring board surface finish metallizations. Although the technical literature on traditional tin–lead solders is quite exhaustive because this solder has been utilized in electronic assembly for about five decades, the same is not the case for lead-free solders. Lead-free solder interconnections are much more complex in their structure, and considerable research effort is needed to understand these multicomponent systems. However, the thrust for environmentally safe electronic interconnect materials is gaining ground, and the number of technical contributions has been increasing. Thermal cycling reliability studies with Sn–Ag–Cu-based alloys indicate acceptable reliability for many applications, and in fact suggest a 1.5–2 improvement over eutectic Sn–Pb solders in many cases. However, there are instances where the reliability of Pb-free solder interconnections is equivalent, or slightly lower than, conventional eutectic Sn–Pb solder joints. This is particularly the case with thin small outline packages or packages with stiff leads. Thus, it is important to recognize these variabilities and verify assembly reliability for specific applications. In addition, much of the reliability information pertains to performance under thermal loading conditions, and considerable effort is needed to understand the reliability under mechanical loading conditions such as drop, bend, and torque. The choice of the printed wiring board surface finish has been shown to have significant effects on the joint microstructure in the interfacial regions. The effect of minor elements contamination is an important aspect in the overall second-level assembly reliability. In general, Pb contamination reduces the reliability of solder joints from Pbfree alloys. Lead content in the range of 2–6% has more pronounced effects than at higher concentrations (125 Am) to cause an electrical short between two legs of a lead frame, as shown in Fig. 6(a). An enlarged scanning electronic microscope (SEM) image of a long whisker on a Sn–Cu finish is shown in Fig. 6(b). The whisker in Fig. 6(b) is about 3 Am in diameter and more than 200 Am in length. It has a straight, fluted surface similar to the grooved columns in some classical buildings. The root of the whisker is shown in the SEM image of Fig. 6(c). The top surface (i.e., the tip of the whisker) of this particular whisker is rough and similar in appearance to the surface of the finish. Many whiskers are kinked at sharp angles, as shown in Fig. 6(d). Short (f10 Am) whiskers on pure Sn finishes are shown in Fig. 7(a) and (b). The surfaces of the whiskers in Fig. 7(b) and (b) are faceted but not fluted, and their diameter is similar to

STRUCTURE AND KINETICS OF TIN-WHISKER

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FIG. 4 SEM micrograph of a bright-tin surface nodule or bud. (Courtesy of G. Galyon and L. Palmer, IBM Corporation.)

the grain size (3–5 Am) of the Sn matrix. The top surface of these short whiskers is flat. Clearly, the morphology of the whiskers grown on Sn–Cu and pure Sn Figs. 6 and 7, respectively) is different.

C.

Overview of Tin Whisker Studies

Any risk-assessment decision on electroplated Sn usage is hampered by the lack of scientific consensus on the growth mechanism involved in whisker formation. Experimental results are notoriously unpredictable and nonrepeatable. Whiskers may appear within days after electrodeposition or they may not appear for years. This delay in appearance, or incubation period, makes the study of whisker growth a frustrating laboratory endeavor. It has also been

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FIG. 5 Schematic diagram of the cross section of a lead frame on an electronic component mounted on a board. (From Ref. 47.)

frustratingly difficult to examine the physical structure of as-deposited electroplated films because of their small thickness and relative softness. As a result, there have been few published attempts to examine the metallurgical aspects (microstructure) of electroplated-Sn films. Experimental investigations are further hampered by the lack of a standard acceleration test that can be utilized to certify a plating process as whisker-free. Perhaps, the greatest variable is the electroplating process itself. Seemingly similar processes can produce sequential runs of electroplated films with

FIG. 6 SEM images of Sn whiskers grown on eutectic Sn–Cu finish on a lead frame. (a) A whisker creates an electrical short between two lead frame legs. (b) A fluted whisker on the Sn–Cu finish surface. (c) An enlarged image of the root of a whisker. (d) A few of the whiskers are bent at sharp angles. (From Ref. 42.)


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FIG. 7 SEM images of short whiskers on a pure Sn finish. The surface of the whisker is not fluted and the whisker diameter is similar to the Sn matrix grain size. (From Ref. 42)

measurably different whisker formation and growth characteristics. Nonetheless, various researchers have periodically brought new analysis tools to bear on the problem of whisker formation and growth and added new data, and occasionally new insights. Based on what is currently known, the principal facts that most Sn whisker theorists would probably agree on are listed in Table 1.

II.

TIN PLATING TYPES

There are several types of electroplated Sn, and each type exhibits different whisker behavior. The technical literature often does not define the type of Sn plating being discussed, but it is an important fact to know. Independent of the type of electrodeposited Sn, the substrate material on which the Sn is deposited has a significant effect on the formation and growth of whiskers. Much of the early work between 1950 and 1975 was on matte Sn deposited on steel, with or without an underlay material. Tin on steel tends to grow numerous whiskers of relatively long (>100 Am) lengths. Over time, Sn-whisker research extended to substrate materials of Cu and Cu alloys. In general, Sn plated onto Cu substrates has been observed to grow more whiskers than a comparable Sn film deposited on a steel substrate. Tin deposited on some Cu alloys, such as brass (CuZn), have been observed to grow even more numerous whiskers than Sn deposited on relatively pure Cu. However, substrate material affects whisker formation only in degree, not in any absolute manner. Much the same thing can be said about electrodeposition-process variables. Process variations affect whisker formation and growth in degree but cannot preclude whisker formation and growth. A brief description of the various types of Sn plate is presented in the following sections.

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TABLE 1

Summary of Whisker Attributes with a High Degree of Consensus

Assertion Statement Whiskers grow by atoms being added to the base of the whisker, not the tip. Whiskers are single crystals with growth directions parallel to dislocation glide vectors

Whisker growth is caused by the relief of internal compressive stresses.

Whisker growth rates are too fast to be explained by theories relying on bulk diffusion of Sn atoms in a Sn matrix. The two most prevalent theoretical approaches used by whisker modelers are dislocation theory and recyrstallization. A third approach involving a defective oxide concept has been postulated relatively recently.

A.

Commentary and Reference Asserted by Koonce and Arnold in 1953 [24] and never refuted Statistically true, but exceptions were noted by Ellis et al. [37] based on studies of kinks in whiskers at Bell Labs during the 1950s. The exceptions noted are critical to the critique of dislocation-based whisker growth theories A generally accepted statement. Early dislocation theories presumed that negative surface tension due to oxidation was the driving force. The clamp-pressure experiment of Fisher et al. [27] established the concept of compressive stress gradients as the driving force behind whisker growth. A generally accepted statement. All modern theories of whisker growth presume that the atoms migrate to the root of the whisker via grain boundaries or dislocation networks. Dislocation theories were first proposed in 1953 by Eshelby [26] and Frank [25]. Recrystallization was first proposed by Ellis et. al. in 1958 [37]. The cracked oxide concept was proposed by K.N. Tu in 1994 [54].

Matte Tin

A Sn plate whose appearance is relatively dull is referred to as matte Sn. A dull appearance is the result of a rough surface texture due to the relatively large (typically 5 Am or greater) grain structure of matte Sn in comparison to fine-grained (typically 1 Am or less) bright Sn. Matte Sn was the only type of Sn plating up through the mid-1960s, and all the published literature up to about 1968 was in reference to matte Sn even if it was not specifically stated to be the case. Modern matte Sn tends to have large, polygonal grains that extend through most, if not all, of the plating thickness. These large polygonal grain structures are achieved by additions of specific (and often proprietary) organic substances to the plating bath. Residual stress levels in asdeposited matte Sn depend on a number of factors, but are generally considered to be less than the residual-stress levels of comparable bright Sn films. Matte Sn does grow whiskers, but the whisker densities and lengths are generally somewhat less than those observed for comparable bright Sn finishes.

B.

Bright Tin

A Sn plating that appears relatively bright because of the high degree of reflection (low scattering) of light from the surface of the plating is referred to as bright Sn. The surface texture of bright Sn is smooth as a result of a relatively small (typically less than 1 Am) grain size. Bright Sn was developed in the mid-1960s for a number of reasons unrelated to whisker growth. The key to making bright Sn is the addition of certain organic additives to the plating bath. Residual stress levels in bright Sn are generally higher compared to matte Sn, and whisker growth is more prolific in bright Sn than matte Sn. Satin-bright Sn is a term describing an intermediate degree of


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brightness between a matte and bright Sn and the internal stress levels in satin bright tins are intended to be as close to zero as possible.

C.

Fused Tin

A fused-Sn plating is one that has been reflowed (melted and resolidified) subsequent to electroplating. Typically, an electroplated part is dipped into a hot-oil bath and held at a temperature above the melting point of Sn. The objective of the fusing operation is to allow the Sn to slowly solidify and, thereby, remove any internal stresses that may have resulted from the electrodeposition process. Fused Sn is reputed to not grow whiskers unless the surface of the tin is scratched, compressively stressed, or plastically deformed [2].

D.

Immersion Tin

Immersion Sn is an electroless-plating process. It is a slow process compared to electroplated Sn and is not used to plate lead frames. It is believed that immersion Sn has never been the focus of a published Sn whisker study, probably because there have never been any reported incidents of Sn-whisker growth on immersion Sn finishes [3–5]. Immersion Sn is used when it is necessary to obtain good edge coverage for purposes of corrosion protection as electroplated films do not provide good edge coverage. Copper heat sinks are occasionally immersion-Sn plated, as are printed-circuit boards (PCBs). Immersion-Sn coatings are considered to be a low-stress plating in comparison to electroplated Sn. Unlike electroplating, where the current and potential is supplied by an external source and the plating follows Faraday’s Law, electroless (or immersion) plating amounts to an exchange of one ion in the plating bath for one atom on the surface of the material being plated. This is a self-limiting process in that, when all the surface atoms have been exchanged, an equilibrium is established with the plating bath and the reaction ceases. It must also be noted that immersion Sn films can be made extremely thin (>1 Am), and this may have been a key factor in the reputed inability of immersion Sn films to form whisker structures. These very thin immersion-tin films may be completely consumed by the formation of intermetallics at the substrate/film interface. As a result, there may be little or no unconsumed Sn available for whisker formation. However, the absence of whisker formation on immersion-tin films is contradicted by the reported observations by one of the authors who found 10-Am whiskers forming within 30 days of depositing a 0.6-Am-thick immersion-tin layer [6]. It is thought to be unlikely that such a thin film has the capacity to grow whiskers to lengths over 50 Am. Such short whiskers are probably not a problem in most electrical component applications, but the assertion that immersion tin is fundamentally not capable of growing whiskers is in error. Expert advice is recommended for all immersion Sn film applications.

E.

Tin with an Underlay

Underlays are electroplates of another metal deposited before (underneath) a Sn film. The most common underplate is nickel (Ni), but some studies have been carried out with other underplate materials. An underplate has many purposes, one of which is to retard, and hopefully eliminate, whisker formation and growth [7]. Nickel appears to change the state of stress within a Sn film on a Cu substrate from a compressive to a tensile state, and it forms a different intermetallic compound (Ni3Sn) with the Sn plate [8]. Nickel also acts as a diffusion barrier between a Cu substrate and a Sn plating. A typical Ni-underlay thickness ranges between 1 and 2.5 Am, although there has been some work carried out with thinner films. Nickel diffuses into a Sn plating at a slower rate than does the Sn into Ni. This is in contrast to the very rapid diffusion of Cu into Sn compared to the diffusion rate of Sn into Cu. Thus, it is logical to presume that the addition of Cu to a Sn plated film will cause a compressive stress state, whereas the diffusion of Sn into Ni will result in a Sn film that is tensile.

860 F.

CHOI ET AL. Tin Alloys

An alloy is a material consisting of two more elements, the majority constituent being a metal. Some electroplated-Sn alloys have been shown to retard whisker formation, such as Sn–Pb [9], and some have been shown to accelerate whisker formation, most notably Cu, Fe, and Zn [10].

III.

TIN WHISKER RELIABILITY PROBLEMS

Whisker reliability problems were first noted on cadmium (Cd)-plated filters during World War II [11]. In 1948, the Bell Telephone System experienced whisker problems on cadmium-plated filters and initiated a program investigating metallic-whisker growth that was first reported in the technical literature in 1951 [12]. By 1956, Bell Laboratories had developed and published whisker mitigation strategies [1,2,9]. Nevertheless, Sn-whisker problems continued to proliferate and be reported in the technical literature for the next 45+ years. It is perplexing that a problem defined by the middle 1950s continues to be a serious problem into the 21st century, although the problem of whiskers has not been ignored in the intervening years. Review papers have periodically revisited the topic of whisker formation and new data has been presented in an attempt to define the fundamental mechanisms of whisker growth. Nonetheless, new hardware designs continue to ignore the whisker knowledge base and incorporate electroplatings prone to whisker formation and growth. As discussed in the following sections, the end results of these unfortunate design practices can be catastrophic.

A.

Field Problems

In 1975, the European Space Agency noted its experiences with Sn-whisker problems and strongly recommended that surfaces that may support stress-induced whisker growth, such as tin, cadmium, and zinc, be excluded from spacecraft design [13]. The agency also noted that ‘‘an alternative finish not been seen to support whisker growth is 60/40 tin–lead.’’ These recommendations were probably the first ever publicly made stating that whisker-forming metals (Sn, Cd, Zn) should be excluded from mission-critical applications such as spacecraft hardware. It is not known whether the European Space Agency followed its own recommendations, but it is a fact that Sn-whisker problems continued to plague some military and commercial spacecraft hardware over the following quarter century. In 1993, the U.S. Air Force published some of their Sn-whisker experiences, along with a detailed description of a disposition process for Sn-plated parts [14]. The NASA Goddard Space Flight Center has also listed a number of Sn-whisker related incidents (Table 2).

B.

Some Industry Practices Used to Cope with Tin Whiskers 1. Vacuum Clean—Northern Electric Corporation described their experiences with Sncoated wiring on filter assemblies and Sn-plated steel frames in a series of papers published between 1968 and 1975 [15–18]. As an interim solution, Northern Electric developed a vacuum tool to remove whiskers from Sn-plated steel frames while still in field service and found that whiskers did not regrow on the vacuumed parts. General Electric has reported Sn-whisker problems with wire-wrapped relay contacts [19] (Fig. 8a and b). GE used a vacuuming technique to deal with the affected hardware in a manner similar to that pioneered by Northern Electric. 2. Heat treatment/fusing—To provide a solution to the problem of whisker formation at the assembly stage, Northern Electric developed a thermal treatment based on heattreatment concepts, initially proposed by Glazunova and Kudryavtsev [10], as well as a fusing operation based on a recommendation initially proposed by S.M. Arnold [2]. However, in the mid 1960s Northern Electric was developing bright-Sn processes as a replacement for matte Sn, and they found that bright-Sn deposits cracked and spalled when heat-treated [18].

STRUCTURE AND KINETICS OF TIN-WHISKER TABLE 2

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Some Whisker-related Reliability Incidents

Problem Description Heart Pacemaker-Class 1 Product Recall—due to tin whisker problems F-15 Radar: electrical short inside hybrid package due to tin whiskers. U.S. Missile Program—tin whiskers on relays Phoenix Air-to-Air Missile: electrical short circuits due to tin whiskers. Galaxy IV/Galaxy VII: loss of satellite ops. Nuclear Utilities: relay electrical shorts due to tin whiskers. Patriot Missile II: whiskers from tin-plated terminals.

Reference Ref: http//www.fda.gov/gra/inspect_ref/itg/itg42.html Ref: B. Nordwall, Aviation Week and Space Technology, June 20, 1986, pp. 65–70. Ref: J. Richardson and B. Lasley, Proc. 1992 Govt. Microcircit Apps. Conf., Vol. XVIII, pp. 119–122, Nov. 10–12, 1992. Ref: L. Corbid, Proc. 3rd Inter. SAMPE Electronics Conf., pp. 773–779, June 20–22, 1989. Ref.: http://www.sat-index.com/failures/ Ref: http://www.nrc.gov/reading-rm/doc-collections/events: status/part21/1999/1999352.html Ref: http://www.anoplate.com/news/pastings/fall2000/tin.htm

3. Lead (Pb) additions—Because of cracking and spalling problems when annealing bright tin, Northern Electric adopted the practice of adding lead (Pb) to mitigate whisker formation. The addition of Pb to a Sn plating in the amount of 1–10% by weight became the most common whisker-mitigation practice largely as a result of the work by Bell Laboratories and Northern Electric. There are no known instances of Sn-whisker incidents in actual field service for tin–lead (Sn–Pb) platings, which have been the predominant lead-frame finish since the 1950s. Laboratory experiments with Sn–Pb films have only been able to grow whiskers when the films are put under clamping pressure or when they are subjected to stress testing (e.g., temperature cycling). 4. Nickel underlay—A major electronics manufacturer has reported over 20 years of success with bright Sn on a Ni underlay over Cu substrates [20]. Occasional manufacturing lapses that resulted in inadvertently omitting the Ni underlay resulted in whisker growth in the field requiring service action. Nickel (Ni) underlays are the predominant whisker mitigation practice for separable connectors where wear resistance is an important design consideration, and should be considered as a prime method for whisker mitigation of tin films for applications other than connectors. It seems that a Ni underlay results in a Sn film with a built in tensile stress that remains stable over time [21]. The great majority of theoretical models for whisker formation presume that a

FIG. 8 (a and b) Tin whiskers on relay pins. The spacing between the pins is about 50 mils. (From Ref. 19.)

862 TABLE 3 Ref. 7)

CHOI ET AL. Tin Whisker Mitigation Recommendations from the International Tin Research Institute (From

Sn-Whisker Mitigation Recommendation Sn on brass should have a Ni underlay, but Sn on steel may be better without an underlay Bright tin directly on brass must not be used. All bright tin coating should be accompanied by every possible safeguard where whisker growth could be damaging Coating thicknesses of tin electrodeposits, not flowmelted, should be at least 8 Am thick. Heat treat tin coatings after plating at 180–200jC for 1 hr. Care is needed in applying heat treatment to bright-tin coatings. If heat treating interferes with subsequent soldering, a protective nitrogen atmosphere should be considered. A copper undercoat may help to obtain better results from heat treatment. Storage conditions and, when possible, service environments should be controlled to avoid corrosion of the base metal since this may introduce harmful stresses to the coating. A hot dipped (or flow melted) tin coating is at far less risk than an unheat-treated, electrodeposited coating Tin–lead deposits, at least 8 Am thick, matte or bright, are probably safe and suitable for most purposes where whisker growth is a potential hazard. The use of a nickel or copper undercoat on brass is a useful additional precaution. Heat treatment is not necessary and can lead to undesired fusion of the tin–lead coating if not well controlled. A lead (Pb) content of 1% has been claimed as effective, but it seems better to select a tin–lead process giving a larger lead (Pb) content. All tin and tin–lead coatings may develop whiskers rapidly where they are subjected to local pressures. When the diminished solderability and ductility of tin–nickel alloy (with 65% tin) are not a bar to use, this coating will provide immunity from whisker growth. Although organic coatings of the thickness commonly used for protection cannot be relied upon to prevent emergence of whiskers, the use of thick layers of resin or the introduction of a solid insulating barrier between points in danger is effective. If, in spite of all precautions, whisker growth occurs, it may be possible to rehabilitate equipment by the physical removal of whiskers. . .a useful means of removal is a small head attached to a vacuum system.

Commentary The Zn in brass is blocked by the Ni and cannot diffuse into the Sn film. Bright tin films have relatively high internal stresses and very small grain sizes. They are very prone to whisker growth.

Bright tins are very difficult to heat treat. Northern Electric tried to heat treat bright tin films and abandoned the practice.

Testing at elevated humidity levels has consistently been shown to accelerate the formation of whiskers. Fused tin is essentially flow melted tin.

Tin–lead films have been the standard (most common) tin-based plating for over 40 years and have never been shown to have any whisker growth under ambient usage conditions. High clamp-stresses on parts like bolted connectors can result in whisker growth in tin–lead films.

This recommendation refers to bolted connectors.

Tin–nickel alloys are not in common usage.

Such practices are often used to mitigate the risk of using inventory, which has been compromised because of whisker formation by a bad choice of plating.

This idea was initially proposed by M. Rozen of Northern Electric in the 1960s, and it has been periodically used since then to recover a contaminated inventory of parts.


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compressive-stress state in the film is a necessary prerequisite for whisker formation and growth. While it is true that laboratory stress testing has produced whisker formation on Sn films over a Ni underlay, there are no known instances of whisker formation on Sn over Ni underlay in typical field-service environments. 5. Pure tin finishes—User communities are well advised to exercise caution when designing electrical hardware utilizing any electroplated material capable of growing whiskers. The European Union’s Reduction of Hazardous Substances (RoHS) legislation requires that Pb be eliminated in most electronic-hardware applications by July 1, 2006. Many component providers will attempt to meet the European Union regulations by simply modifying their Sn–Pb platings to pure Sn. Tin can, and has, been successfully used. For example, all surface-mount, multilayer ceramic capacitors (MLCCs) have used a pure Sn end-cap metallurgy since 1992 with no known field-service whisker problems (as of 2002). However, the MLCC process utilizes a barrel-plated Sn with a Ni underlay on a Ag-frit substrate, and almost all of the Sn on MLCC end-caps is reflowed during the subsequent board assembly operation. While no field-service whiskering has been reported on MLCCs, there has been some laboratory work performed where whiskers were grown on MLCC end-cap metallurgy by subjecting the device to temperature cycling between 40 and + 60jC [22]. Whiskers were observed to grow from the Sn surfaces at about 100 cycles. It appears that the success of the commercial application of lead-free MLCCs was attributable, in part, to the fact that normal commercial applications do not subject the part to temperature cycling as described above, and also because most of the end-cap metallurgy is reflowed during the solder attach process. However, aircraft applications are another matter, as temperatures are constantly cycled between 40 and +60jC. For aircraft applications, it is advisable to use MLCC end-cap metallurgies other than Sn. The user should clearly understand the data behind the claims of ‘‘whisker-free’’ platings before implementing Sn-electroplated parts. The user should understand that the decision to use any Sn-electroplated film is not risk-free, despite claims to the contrary.

C.

Recommendations

The Tin Research Institute in England, now known as the International Tin Research Institute (ITRI) [7] detailed 20 years of Sn-whisker observations based on work carried out in cooperation with Bell Laboratories. Recommendations, listed in Table 3, were made which summarized the best thinking at that time, and these recommendations can still be regarded as relevant to the careful user.

IV.

WHISKER FORMATION AND GROWTH THEORIES BASED ON DISLOCATIONS

Dislocations are line defects in three-dimensional crystalline solids consisting of extra (or missing) atoms arranged in line in an otherwise ideal crystalline-lattice structure. These line defects are the means by which all metals are plastically deformed, and they also play a role in many solid-state phenomena, such as diffusion and electrical resistivity. Dislocations were the basis of the first theories proposed for the growth mechanism of Sn whiskers. The first whisker formation and growth dislocation theory was published in 1952 [23]. Peach proposed that whiskers grew by Sn atoms migrating through a screw dislocation located at the center of a whisker and depositing themselves at the tip of the whisker. This particular theory was quickly invalidated by subsequent experimental data.

A.

Whisker Growth from Whisker Base

The first experimental data of relevance was the seminal observation by Koonce and Arnold (K–A) in 1953 that whiskers grew from Sn atoms continually added to the base, and not the tip,

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of a whisker [24]. This observation effectively negated the idea of a whisker with a centrally located screw dislocation acting as the conduit for diffusing Sn atoms to the tip of a whisker, as proposed by Peach.

B.

Early Dislocation Theories

Dislocation theorists quickly began to construct theories that were consistent with the Koonce– Arnold observations. These theories were based on the generation of dislocations and their subsequent movements. Frank [25] and Eshelby [26] independently proposed diffusion-limited dislocation mechanisms where the driving force for dislocation generation and motion were stress fields created by surface oxidation. For a dislocation to move it is necessary to have a stress field, either externally applied or internally created. To create a whisker, it is necessary for a dislocation to move to a surface and deposit its extra plane of atoms on that surface. Furthermore, any theory proposing that whiskers grow by the movement of dislocations that fails to reconcile the Koonce–Arnold observation would be invalid. The K–A observation has never been contradicted or questioned since it was first introduced in 1953. Eshelby’s model postulated an internal source of simple prismatic dislocations that expanded by climb to the boundaries of a whisker-forming grain and then moved by glide to the surface of the grain. Frank’s model involved a rotating edge dislocation pinned to a centrally located edge dislocation. Both Frank and Eshelby felt that bulk diffusion of Sn would support the observed growth rates for spontaneously grown whiskers, and they both claimed their models were consistent with the Koonce–Arnold observation.

C.

Effect of Externally Applied Forces on Whisker Growth Rates

An important contribution to the Sn-whisker knowledge base was a study conducted by Fisher et al., where they observed that a Sn-plated steel specimen held within a metallurgical clamp grew a prolific number of whiskers over a period of only a few days. In a more controlled experiment, Fisher et al. made the first whisker growth-rate measurements under conditions of an externally applied stress [27] (Figs. 9 and 10). At an applied compressive stress of 8000 psi, some whiskers were observed to grow at rates of 10,000 A˚/sec. Spontaneously grown whiskers (grown without any externally applied stress) grew at rates closer to a fraction of 1 A˚/sec. This data demonstrated

FIG. 9 A growth plot for three different tin whiskers (A, B, and C) that were grown on tin-plated steel with a clamp pressure of 8000 psi. (From Ref. 27.)


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FIG. 10 A growth-rate plot for tin whiskers grown on tin-plated steel with a clamp pressure of 8000 psi. The three plots (1, 18, and 100) represent the number of individual whiskers that correspond to the three plot lines. (From Ref. 27.)

that whisker growth was directly tied to the magnitude of compressive-stress gradients in an electroplated-Sn film. Fisher et al. explained the sudden stoppage of whisker growth (i.e., cessation) as the result of impurity atoms plugging the dislocations and inhibiting fast diffusion of tin atoms through the dislocation networks. Although Fisher’s results have not been specifically reproduced, the general relationship between compressive-stress gradients and whiskergrowth rates has been verified by other experimenters. Clamp-pressure experiments were conducted on tin-plated steel at pressures of 150 kg/cm2 that resulted in whisker-growth rates of 2.3 A˚/sec [28] (compared to the 10,000 A˚/sec reported by Fisher et al.). Additional clamppressure experiments were conducted on Sn-plated steel reported where the highest whiskergrowth rate achieved at clamp pressures of 8000 psi was 593 A˚/sec [29]. In this later study, it was observed that whisker-growth rates decreased over time, which was in contrast to the linear growth rates reported by Fisher et al. This later study showed that whisker growth ceased with time, as was also observed by Fisher et al. This study also commented about clamp-pressure experiments on hot-dipped Sn and 50Sn–50Pb platings on both steel and Cu substrates. While whisker densities were observed to decrease with the addition of Pb, considerable whisker growth occurred at high clamp pressures (several thousand psi) for the 50Sn–50Pb plating. It was also observed that a hot-dip Sn coating on Cu produced considerably fewer (about half) whiskers compared to the same hot-dip coating on steel substrates when subjected to high-clamping pressures.

D.

Dislocation Theories: Post Early Clamp-Pressure Results

The first whisker dislocation theory subsequent to the clamp-pressure results of Fisher et al. proposed that internal stress in the whisker-growing medium is relieved by whisker growth when that stress could not be relieved by other processes [30]. Whisker growth stopped when the internal stresses were relieved. This model proposed that a concentration of vacancies at the root of the whisker was maintained at a constant level by absorption of vacancies at edge dislocations situated at the end of a stress gradient. Using these assumptions, it was demonstrated that a linear relation should exist between growth rate and applied pressure, as reported by Fisher et al.

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However, predicted growth rates were still at least 1 order-of-magnitude smaller than those reported by Fisher et al. In addition, the pinning of dislocations by impurity atoms over time suggested that growth rates should tail off gradually and not abruptly as observed by Fisher et al. Some clamp-pressure experiments [29] were, in fact, reported that did show a gradual tailing off of whisker growth rates over time [27]. Another dislocation model proposed that helical dislocations in the whisker grain move to the surface by climb [31]. No experimental evidence for the existence of helical dislocations in Sn films was offered. It was stated that the helical dislocation model was consistent with the Koonce– Arnold observation, but there was no discussion of the helical-dislocation model in reference to the clamp-pressure experimentation of Fisher et al. [27].

E.

Dislocation Glide Theory

J. Franks [32] recognized that rationalizing the Fisher et al. observations appeared to be ‘‘difficult,’’ so he proposed a dislocation–glide model where the frequency of dislocation generation was dependent on diffusion processes. He proposed that whisker-generating dislocations, called Eshelby sources, are initially pinned due to lattice faults, and thereafter act as dislocation sources under the influence of either an applied or internal-stress field [33]. Once pinned, these dislocations expand so as to lie in glide planes, and then glide to the surface, piling up as they reach the surface. Eventually these piled-up dislocations break through the surface and grow a whisker. According to this theory, the material for whisker growth is supplied by a ratecontrolling diffusion mechanism. This model accounted for the observation that whisker growth suddenly stops at some fixed end point, regardless of the growth rate or applied pressure, by surmising that the dislocation sources eventually become fixed or plugged. That is, the dislocations become incapable of generating new dislocations because of faults introduced into the ‘‘region by the constant addition of material to the region at the base of the whisker, eventually locking the whisker-generating dislocations.’’ Franks’ theory attributed the stress fields to either internal or external forces, and not oxidation effects as was the case with the earlier Eshelby expanding-loop theory [26].

F.

Two-Stage Dislocation Models—Climb and Glide

U. Lindborg claimed that prior dislocation theories based on lattice diffusion achieved agreement with experimentally observed Sn-whisker growth rates because the estimates for Sn diffusion coefficients were in error on the high side by about 106 [34]. Lindborg proposed a two-stage dislocation model for the growth of whiskers in Zn, Cd, and Sn where the first stage was a dislocation loop-expansion stage, based on dislocation climb and vacancy diffusion, similar to the prior theories of Eshelby [26] and Franks [32,33]. The concepts of grain-boundary and dislocation-pipe diffusion were used to account for the whisker growth rates reported in the prior literature. A second stage was also postulated wherein dislocations, after they are created by the source and expand by climb, glide toward the surface where they deposit a layer of Sn atoms. Resistance to gliding dislocations was attributed to a network of forest dislocations within the grains. Either stage could be the rate-determining factor in the whisker-growth rate. It was speculated that the very high, stress-accelerated growth of whiskers was an example of the first stage (diffusion-limited growth) determining the growth rate. The clamp-pressure results of Fisher et al. [27] were thought to be a manifestation of the first stage in this model, where diffusion limits the growth rate. 1. Change in the State-of-Stress Lee and Lee proposed a model where prismatic dislocation loops in the whisker grain expand by climb up to the grain boundaries, and then glide toward the surface driven by a stress gradient. Whisker kinking (a change in growth direction) was attributed to prismatic dislocation loops expanding on new slip planes, and the internal stress gradients were attributed to Cu6Sn5 precipitation in grain boundaries. Lee and Lee reported the first direct measurements of residual stresses in Sn electroplate through the use of a cantilever-beam method. One side of a cantilever


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beam was plated, thereby causing the beam to deflect as a result of the stresses in the film. The deflection can be correlated to the average residual stress in the film. It was observed that immediately after plating the stress in the films was tensile. However, within days, the tensile stress changed to become compressive. This change in the state-of-stress was attributed to the formation of Cu6Sn5 at the Sn/Cu boundary and at the grain boundaries. 2. Crystallographic Orientation Lee and Lee [35] also determined the preferred-orientation indices of the deposited films by using X-ray diffractometry. Their Sn was plated on phosphor-bronze substrates from an acid-stannous bath maintained at room temperature. It was not stated whether the Sn was matte or bright, but the polygonized nature of the grain structure suggests that brighteners were not added to the plating solution. The average grain size of the electrodeposits was 1 Am, which is a relatively small grain size for matte-Sn plating. Unlike some previously published data, the grain size was observed to decrease with increasing electrodeposition-current densities [36]. The crystallographic orientation of individual whisker grains was determined by comparing the angle of whisker growth to the preferred orientation of the film, and found that whisker grains were different in orientation compared to the predominant texture of the as-plated films. It was conjectured that the differently oriented whisker-forming grain may have been the result of a recrystallization event, and that the different orientation of the whisker grain resulted in a weaker TABLE 4

A Summary of Whisker Growth According to Lee and Lee [35] with Commentary

Assertions Compressive stresses in the film are built up over time due to the diffusion of copper atoms from the substrate into the tin film, and the subsequent formation of the intermetallic Cu6Sn5. The as-deposited stress state for the tin film was net tensile (+11 MPa) and changed to a compressive stress (8 MPa) after a few days. Subsequent to whisker formation and about 50 days later, the stress level decreased (5 MPa). Tin whiskers grow from grains whose orientation is different from the major orientation of the tin film. . .the tin surface oxide film can be sheared along the boundaries of the grain. To release the compressive stress in the film, tin whiskers grow from the grain whose surface oxide is sheared. The whisker growth is controlled by the expansion of the prismatic dislocation loop on the slip plane by climb. . ., i.e., the operation of a Bardeen–Herring dislocation source. The dislocation loop expansion is restricted by the grain boundary. The loop then glides along its Burgers vector direction. As a result, the tin whisker grows by one atomic step. The continuous operation of the Bardeen–Herring dislocation source gives rise to whisker growth until the stress is relieved.

Commentary These data were the first direct measurements of tin film internal stresses, and the cantilever beam method was utilized.

The different orientation of the tin whisker grain had been seen by others prior to Lee and Lee’s work. It is a generally accepted fact that the whisker grain is different in orientation than the surrounding regions. It is generally accepted that tin whiskers are one type of stress relief mechanism. This proposal is an expansion of the original concept proposed by J.D. Eshelby in 1953 and expanded upon by others. It is not generally accepted as a proven concept mainly because there is no direct experimental evidence to support it.

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(i.e., thinner and/or cracked) oxide layer. A thinner oxide could facilitate gliding dislocations to break through the surface layer and cause the formation and growth of a whisker. Lee and Lee’s proposals integrated a number of prior concepts into a general concept of whisker formation and growth. It is, however, important to note they did not discuss their dislocation model with respect to the Koonce–Arnold observation. The major conclusions based on the work of Lee and Lee are summarized in Table 4.

G.

Critique of Tin Whisker Formation and Growth Based on Dislocation Theory

Ellis et al. stated in 1958 that some whisker-growth directions in kinked crystals were not dislocation glide directions, and that an alternative (nondislocation) mechanism would be required to rationalize these observations [37]. All dislocation–glide theories require that whisker-growth directions be a dislocation glide plane. Ellis, et al., pointed out that while most whisker-growth directions are parallel to dislocation glide planes, not all are. Ellis’ point was acknowledged at the time by Frank, one of the earliest advocates of a dislocation-based, whisker-growth theory. No one has ever refuted Ellis’ observations, even as they have been largely ignored. The importance of dislocation networks and grain boundaries as fast-diffusion paths for Sn atoms to reach the base of the growing whisker was acknowledged by Ellis. However, the major thesis of Ellis’ work was the proposal that recrystallization was a key step in whisker formation and growth. Ellis had no direct proof of recrystallization, but inferred from various data the reasonableness of such a proposal. Tin atoms would move out of the grain boundaries and into the recrystallized grains, thereby growing the volume of the recrystallized grain.

V.

ANALYTICAL TOOLS THAT PROVIDE FUNDAMENTAL UNDERSTANDING

Bell Laboratories was the first to use the electron microscope to investigate whisker growth problems in the early 1950s [38]. The scanning electron microscope (SEM) was introduced with the first micrographs in 1972 [17]. Auger analysis was initially utilized to investigate whiskers in the 1980s [39]. X-ray diffractometry has been used to identify chemical compositions and film stresses since at least 1996 [40]. Focused ion beam (FIB) microscopy was initially used to study Sn whiskers in 2001 [41], and microfocus X-ray diffractometry in 2002 [42].

A.

Need to Understand Structural Details

By the end of the twentieth century, it was clear that theories of whisker formation and growth had not been substantiated by sufficient experimental evidence and, therefore, were little more than informed speculation. A very comprehensive survey concluded that none of the then existing whisker-growth theories were supported by adequate experimental verification, and that it was critical to understand the crystalline microstructure of Sn-electroplated films and whiskers [43]. However, Sn-film microstructural analysis was extremely difficult to do because of the physical nature of Sn. Tin is a very soft material that is extremely difficult to polish and etch in the classical metallurgical manner. Furthermore, tin tends to readily recrystallize as a result of mechanical deformation from laboratory polishing operations, which means that the microstructure eventually observed after polishing and etching is not the actual microstructure. It was clear that new analytical tools were required to advance the fundamental understanding of whisker formation and growth. In particular, it was necessary to develop a means for examining the detailed microstructure of Sn whiskers and the surrounding region; and to gain an understanding of stress in the whisker region in detail, rather than just on average. Detailed stress measurements were made possible by the development of microfocus X-ray diffractometry that focused X-ray beams to a 1-Am2 spot size (previously only possible down to about 500 Am square). Microfocus X-ray provided a means to attempt the measurement of relative stress levels for 1-Am square regions around a whisker. Furthermore, the development of focused ion beam


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(FIB) microscopy provided a means of conducting in situ micromilling on Sn films without the necessity of mechanically abrading the Sn. FIB microscopy utilizes a rastered stream of accelerated gallium (Ga) ions to mill the surface of a target film and, at the same time, generate a stream of secondary electrons from the milled surface that can be collected to generate a visual picture of the structure. Some FIB microscopes incorporate a scanning electron microscope capability integrated into the basic FIB tool, the so-called dual-beam FIB. Microbeam X-ray and FIB diagnostic tools have been the primary means by which new insights into the fundamental mechanisms of whisker formation and growth have been determined.

B.

Microbeam X-Ray Diffractometry (Synchrotron Radiation)

1. Advantages and Method The microdiffraction apparatus at the Advanced Light Source in Lawrence Berkeley National Laboratory was used by Choi et al. [42], to study the stress around Sn whiskers grown on a eutectic Sn–Cu plated finish deposited on a Cu lead frame. An X-ray microdiffraction end-station was utilized which was capable of delivering a white X-ray beam (6–15 keV) focused to 1 Am via a pair of elliptically bent Kirkpatrick–Baez mirrors. The size of the beam (f1 Am) was smaller than the grain size (3–5 Am) of the deposited film, so that a single-grain white beam Laue pattern could be obtained for individual grains. The Laue patterns were collected with a large area (9 9 cm2) charge-coupled device (CCD) detector with an exposure time of 1 sec at each step scan, from which the orientation matrix and a deviatoric strain tensor for each illuminated grain were derived. Because of the low absorption of X-rays, several grains were illuminated at the same time through the thickness of the finish, but the grain of interest (the one closest to the surface) was discriminated from the rest by the intensity of its reflections. An area of about 100 100 Am was scanned by using 1.5-Am beam squares. The strain (and equivalent stress) in the Sn–Cu matrix grains was determined by using the lattice parameters of the whisker as a stress-free internal calibration reference. Fig. 11 is a low-magnification SEM micrograph depicting an area where a whisker and its surrounding region was scanned. A microbeam diffraction pattern of the whisker is shown in Fig. 12. The axis along the length of the Sn whisker was determined to be (001). Pole figures of (100), (110), (321), and (211) orientations are shown in Fig. 13(a)–(d), respectively. A high concentration of (321) at the center of the pole can be seen in Fig. 13(c), thus the surface of the finish had a (321) texture. 2. Deviatoric Stress State The above microdiffraction study showed that for 1.5-Am square areas immediately around a whisker the stress was not biaxially uniform; rather the stresses were highly inhomogeneous with grain-to-grain variations. The film stress was biaxial only in average. Fig. 14 shows a plot of rVzz, which is the deviatoric component of the stress along the surface normal, that was measured from deviations in crystal Laue pattern and which did not include any volume expansion corrections. Because r Vxx+r Vyy+r Vzz=0 (by definition), r Vzz is a measure of the in-plane stress. A positive value of r Vzz (=r Vxx+r Vyy) indicates an overall tensile stress, whereas a negative value indicates an overall compressive stress. However, the measured strains of less than 0.01% were only slightly larger than the strain detecting sensitivity of the white-beam Laue technique (the strain sensitivity of the technique is 0.005%). The key conclusion from the above study is that the compressive stresses in the finish were very low (less than 10 MPa) and no long-range macrostress gradients could be determined around the root of the whisker to indicate that whisker growth had relieved most of the local compressive stress in the surrounding grains. However, when Choi et al. assumed that the whisker was stress-free the microdiffraction data could be interpreted to mean that the surface layers of the Sn–Cu finish around the base of the whisker were compressively stressed (the calculated values of the stress mapping of the region in and around the whisker are shown in Table 5). The compressive stress values were quite low, on the order of several MPa (or several hundred psi), and there was a slight (i.e., short range) stress gradient from the whisker root to the immediately surrounding regions. This implied that the stress level just below the whisker was slightly less compressive than the surrounding area, which

FIG. 11 Low-magnification SEM image of the surface of a lead frame coated with a Sn–Cu finish that exhibited several Sn whiskers. The one selected for a microbeam diffraction analysis is circled. (From Ref. 42.)

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FIG. 12 Microbeam diffraction pattern of the whisker circled in Fig. 11. (From Ref. 42.)

is to be expected because the stress very near to the whisker is presumed to be relieved by whisker formation and growth. In Table 5, the arrows indicate the stress gradients. Some examples of adjacent blocks in Table 5 having the same stress level value are indicated by circled sets, which means that they belong to the same grain. The microdiffraction technique also provided the grain orientation distribution surrounding the root of a whisker, as shown in Fig. 15. Of particular interest is the grain underneath a whisker. 3. Grain Orientation Under a Whisker Choi et al. [42] intuitively felt that the microstructure of the whisker grain should be different from the surrounding grains. From the pole figures in Fig. 13, they determined that the surface layer of the Sn–Cu solder finish had a (321) texture, and from Fig. 15 they determined that the grain just below the whisker exhibited a (210) orientation. This (210) grain was, therefore, considered to be a discontinuity in the predominantly (321) microstructure. Choi et al. conjectured that this (210) structural discontinuity could be a location where the surface oxide could easily be broken by a whisker growth. 4. Summary From the studies conducted by Choi et al. [42] on Sn–Cu films, it was determined that the internal stresses, on a scale of 1.5-Am squares, were highly inhomogeneous with significant grain-to-grain

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FIG. 13 Several pole figures of grains in the Cu–Sn finish. (a) Pole of (100), (b) (110), (c) (321), (d) (211) orientations. A high concentration of (321) grains is noted at the center of the pole in (c). (From Ref. 42.)

variability. On average, the stresses around whiskers were compressive and generally smaller in magnitude at regions immediately adjacent to the whisker. a. Internal Stress Variation in Sn-Electroplated Films. Choi et al. established that internal stress variation in Sn–Cu electroplates were highly variable from grain to grain. However, vertical stress gradients were not addressed in this work. The Sn/substrate interface is a region of great metallurgical interest. Tin forms an intermetallic compound (IMC) with most substrate materials almost immediately after the electroplating process has been completed, and for Cu substrates the IMC is usually Cu6Sn5. It is well known, and generally accepted, that Cu6Sn5 creates a compressive stress state at the Sn/Cu interface. Therefore, the presumption is that there is a vertical compressive stress gradient in the Sn-plate films with the highest compressive stress located at the Sn/Cu boundary where the Cu6Sn5 IMC is located. This vertical compressive stress gradient would be further enhanced by the diffusion of Cu (and other impurity atoms) from the substrate into the Sn film. The stress state at the surface of the film will be less


873

FIG. 14 Deviatoric stress in the Z-direction for the area around the whisker root, with the contribution of the whisker deleted. (From Ref. 42.)

TABLE 5 Stresses and In-Plane Stress Gradient Map of the Region Surrounding a Tin Whisker Root (From Ref. 42)

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FIG. 15 Distribution of grain orientation around the root of a whisker determined by X-ray microdiffraction. (From Ref. 42.)

compressive than the film at the substrate interface. It is even possible, perhaps even likely, that the stress state of the film at the surface will be zero or even tensile. W. Stevens reported synchrotron X-ray diffractometry results on matte-Sn platings indicating that compressive stresses at the Sn/Cu substrate interface approach 40 MPa [44]. If confirmed, this would be the highest magnitude internal stress ever reported for a Sn film and it would add some weight to the speculation that there are positive (i.e., compressive to tensile) vertical stress gradients in Sn-electroplated films of relatively high magnitude. This X-ray result has to be considered very preliminary, but it is consistent with observations on cantilever beams that have been plated on one side and then etched [45]. Shortly after the cantilever beam was plated with a tin film a compressive stress was generated that deflected the beam. The tin film was then etched to remove the tin film but not the IMC that had formed between the tin film and the Cu substrate. It was observed that the deflection was relatively unchanged by the etching operation. The conclusion from this experiment was that compressive stresses in a Sn film were largely a result of IMC formation at the tin/substrate interface. It has also been theorized that Sn films over a Cu substrate have compressive stress gradients resulting from fast diffusing Cu atoms migrating into the grain-boundary network of the Sn film and forming precipitates of Cu6Sn5 IMCs, and that these compressive stress gradients were the principal driving force for whisker formation and growth [46].

C.

Focused Ion Beam (FIB) Microscopy

FIB studies were conducted that identified some important microstructural characteristics of Sn whiskers relative to their immediate surroundings [47]. FIB specimens were characterized by


875

using a selective area electron diffraction pattern, and the corresponding transmission electrons microscopic images of the whisker cross-section normal to its growth direction are shown in Fig. 16(a) and (b), respectively. The diffraction patterns were indexed to be a single crystal in the h001i orientation. 1. Focused Ion Beam Images of Sn Whiskers and Eutectic Sn–Cu Finishes Fig. 17 depicts a focused ion beam (FIB) image of a surface area on the surface of a Sn–Cu finish with two rectangular etched holes and a thin bridge of Sn–Cu left standing between the two holes. This bridge was subsequently removed intact and prepared for transmission electron microscopy (TEM). Utilizing a set of FIB trenches in this manner is a typical method of preparing a sample from a specific location for cross-sectional TEM study. Several Sn whiskers can be observed surrounding the two holes. Actually, there was a whisker located between the two holes before the FIB etching. This thin sample prepared by FIB enabled an image of cross-sections parallel to the whisker growth direction to be studied. Similarly, cross-sectional images of the grains immediately surrounding and below the root of a whisker could also be examined. Fig. 18(a) and (b) shows corresponding FIB and TEM images of a thin slice similar to the one in Fig. 17 In the cross-sectional FIB image of Fig. 18(a), a crack is apparent in the upper left part of the image. The same crack can also be found in the upper left part of the corresponding cross-sectional TEM image in Fig. 18(b). The top layer above the crack is a deposited protective layer needed to prevent etching the thin slice by the focused ion beam during sample preparation. The image of a whisker, which is below the top protective layer, is indicated by an arrow in Fig. 18(b). Although the whisker was a single crystal, it nevertheless contained observable defects. Most likely, these defects were introduced when etching the sample with the high-energy ion beam. a. Cu6Sn5 Precipitates in Grain Boundaries of Eutectic Sn–Cu. Sheng et al. [47], noted that there were many small precipitates of Cu6Sn5 in the Sn grain boundaries (Fig. 18a). They appeared as bright particles and their composition was confirmed by energy dispersive analysis to be Cu6Sn5. The presence of these Cu6Sn5 precipitates provided an explanation for the origin of the compressive stresses thought to be the driving force for whisker growth, and an explanation for why eutectic Sn–Cu finishes form more whiskers than pure Sn finishes. It was also reported [47] that most of the grain boundary Cu6Sn5 particles precipitated during the electroplating of the eutectic Sn–Cu alloy. Sheng et al. stated that if the eutectic Sn– Cu film was reflowed (i.e., melted), most of the Cu in solution during the liquid state would precipitate out as grain boundary Cu6Sn5 IMC particles during solidification. During solidification, the Sn becomes supersaturated with Cu atoms, which then precipitate out of solution. As the grain size of the resolidified Sn–Cu film was only a few microns, it took less than a second for Cu in the interior of a grain to diffuse to the surrounding grain boundaries to form an intermetallic (IMC) Cu6Sn5 precipitate. The interstitial diffusivity of Cu in solid Sn is extremely fast; more rapid than 108 cm2/sec at room temperature. The above solidification scenario is also true for electroplating. Supersaturated Cu atoms remaining in solution within the Sn film after plating quickly diffuse to the grain boundaries during storage at room temperature. Whether the diffusion of these remaining supersaturated Cu atoms to a grain boundary to form Cu6Sn5 would induce a compressive stress, or not, is controversial. The stress caused by a Cu6Sn5 precipitate in Sn films depends on whether the Cu comes from the Cu substrate or from Cu atoms in the grains that originated from the plating deposition process. Cu atoms from the substrate add volume to the film and result in a compressive stress, whereas Cu atoms coming out of solution and precipitating as Cu6Sn5 in the grain boundaries result in a tensile stress because of the difference between the partial molar volumes of 6 Cu and 5 Sn atoms in solution and in the Cu6Sn5 precipitate. Sheng et al. [47] and Choi et al. [42] believe that the IMC Cu6Sn5 phase is the reason why the addition of Cu to a Sn-electroplated film builds up a compressive stress that enhances whisker growth. However, even for the case of pure Sn deposited over a Cu substrate there is significant diffusion of Cu from the Cu substrate through the Sn film which will grow grain boundary precipitates resulting in the build up of a compressive stress in the Sn film.

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FIG. 16 Cross-sectional transmission electron microscopic images of a tin whisker. (a) A selected area diffraction pattern of a whisker in the 001 growth direction. (b) Cross-sectional images near a whisker bend thought to be images of screw dislocations. (From Ref. 47.)


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FIG. 17 Preparation of a thin slice sample utilizing a focused ion beam (FIB) for a transmission electron microscope (TEM) study. Two rectangular holes were etched on the surface of a eutectic Sn– Cu finish with a thin slice left standing between the two holes. The thin slice was used to perform the TEM analysis. (From Ref. 47.)

The morphology of the Cu6Sn5 particles shown in Fig. 18(a) and (b) suggested that the stress generated by their growth was not uniform. The Cu6Sn5 was distributed through the microstructure as discrete particles, and not as a thin layer, or wedge, in the grain boundaries. This observation was the basis for reasoning that there was a stress concentration gradient around each individual Cu6Sn5 particle. Furthermore, during room temperature aging, the Cu6Sn5 particles in the grain boundaries were observed to ripen (enlarge). In a closed system, with no additional (outside) availability of Cu atoms there can be no volume ripening. However, in the case of a Sn–Cu or pure Sn plated lead-frame the system is open, and Cu atoms from the leadframe substrate readily diffuse into the Sn–Cu finish to form increasing amounts of Cu6Sn5 over time. This continual ripening or growth in the volume of Cu6Sn5 within a Sn–Cu or pure Sn film produces a compressive stress. The existence of many Cu6Sn5 particles in the Sn–Cu finish was thought to be the main reason why eutectic Sn–Cu films had more whisker growth than was observed on pure tin finishes. b. Cu6Sn5 Precipitates in Grain Boundaries of Pure Sn. An FIB image of the cross section of a pure-Sn finish is shown in Fig. 19. While the Sn grains are clearly shown, relatively few grain-

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FIG. 18 (a) A thin slice of eutectic Sn–Cu finish prepared by FIB thinning. The whisker is indicated by an arrow. Many grain boundary precipitates can be observed that have been identified to be Cu6Sn5 by elemental analysis. (b) A corresponding TEM image. The whisker is indicated by an arrow. Although the whisker is a single crystal of Sn, it nevertheless contains numerous defects, most of which are dislocations. (From Ref. 47.)


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FIG. 19 An FIB image of a cross-sectional area of pure Sn finish prepared by FIB thinning. (From Ref. 47.)

boundary Cu6Sn5 precipitates are observed in comparison to those seen in Sn–Cu films (Fig. 18a), which is reflective of the major difference between the two finishes.

D.

Additional Features and Characteristics of Sn Whiskers Provided by Analytical Tool

1. IMC Particle Near Whisker Root One of the first published [41] FIB cross sections of a pure-Sn whisker (Fig. 20) clearly shows a floating particle, presumably Cu6Sn5, at the root of a whisker. There have been other published FIB x-sections of tin whiskers for matte, satin-bright, and bright Sn, in which the presence of an IMC particle was clearly evident directly at the root of the matte Sn and satin-bright whisker grains [48] (Fig. 21). 2. No IMC Particle Near Whisker Root However, in the above study [48], there was no indication of an IMC particle at the root of brightSn whiskers. This was thought to be attributable to the fact that bright-Sn whiskers originate

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FIG. 20 The first published FIB cross-sectional micrograph of a tin whisker grown on a copper substrate showing an intermetallic SnCu particle at the root of the whisker. (From Ref. 41.)

from a surface nodule rather than a surface grain. For bright-Sn films, a nodule appears to be a precursor state for the formation of whiskers. During bright-Sn aging experiments, nodules are observed before whiskers appear. 3. Effect of an Underlay Internal stress measurements using X-ray diffraction (XRD) and cantilever-beam techniques invariably show that, over time, internal compressive stresses are built into Sn films (matte, satin bright, or bright) deposited directly onto Cu or Cu-alloy substrates. These compressive stresses are usually attributed to buildup of the intermetallic compound Cu6Sn5 at grain boundaries and the film/substrate interface. Tin (matte, satin bright, or bright) deposited over a Ni underlay [48] exhibits a built-in tensile stress immediately after deposition that remains tensile after aging up to (at the time of this writing) 6 months. A study conducted by Vo used lead-frame Alloy 42 (Fe–Ni) substrates and deposited 10 Am of matte Sn over a 1-Am underlay of Ni [49]. The test samples were temperature cycled between 55 and +85jC for several hundred cycles and then examined using FIB. Some relatively short whisker growths were found on these samples. Fig. 22 is an FIB xsectional micrograph showing a well-defined matte-Sn columnar-grain structure with Cu6Sn5 precipitates on the vertically oriented grain boundaries and NiSn intermetallic at the boundary between the Sn film and Ni underlay. The whisker appears to be growing out of a surface grain and there is no evidence of precipitates in the boundary of the whisker grain. To date, there have been no reported instances of whiskers growing under ambient conditions for Sn films with a Ni underlay. To grow whiskers on Sn films with Ni underlay, it has been necessary to impose a stress test of some sort (e.g., temperature cycling, clamping pressure, or elevated temperature/ humidity). These facts are the basis for the pervasively high degree of confidence in Ni underlays as a whisker mitigation strategy.

VI.

WHISKER FORMATION AND GROWTH THEORY BASED ON RECRYSTALLIZATION

Material scientists recognize recrystallization as a phenomenon where a crystalline solid transforms its internal structure from one with high internal stresses and relatively small grains to a structure with low internal stresses and relatively large grains. The recrystallization transformation reduces the total grain-boundary area and defect density (e.g., dislocations, stacking faults, point defects) per unit volume. However, normal recrystallization does not involve net mass transport of atoms from one region to another. Early on, whisker growth was considered to be a stress-relieving process and a particular type of grain growth, so it was natural that some connection between whisker growth and recrystallization processes would be considered. There

FIG. 21 An FIB cross-sectional micrograph of a whisker grown on satin-tin deposited on a copper substrate showing (a) overview of an intermetallic SnCu particle at the root of the whisker. (From Ref. 48.)

STRUCTURE AND KINETICS OF TIN-WHISKER 881

FIG. 22 (a) SEM micrograph of a tin whisker before FIB cross-sectioning. (b) FIB cross-sectional micrograph of a whisker grown by thermally cycling a 10-Am-thick matte-tin film over a 1-Am-thick nickel underlay on a copper substrate. The whisker was determined to be pure tin and appeared to form from the surface grain. Intermetallics were observed to form at the grain boundaries. (From Ref. 49.)

882 CHOI ET AL.


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FIG. 23 A schematic diagram of a recrystallized grain structure showing grains that are about the same size and uniform in shape. (From Ref. 50.)

were several studies that inferred the possibility of recrystallization based on experimental data, but none had direct metallographic evidence of recrystallization [10,37]. Kakeshita et al. surmised that ‘‘whiskers are considered to grow on recrystallized grains’’ and showed schematics of recrystallizing grains forming within a region of smaller, highly stressed grains [50] (Figs. 23 and 24). However, Kakeshita also had no experimental evidence to demonstrate the existence of recrystallized grains. Additional indirect evidence for recrystallization was reported in Ref. [35], where it was determined that the orientation of the whisker grain was different from the predominant grain orientation (or texture) as determined by powder X-ray diffraction analysis. It was speculated that the whisker-growing grain was different because it had recrystallized from a parent-crystalline matrix and, in so doing, had grown a whisker structure. Any whisker formation theory must rationalize the observed whisker data, particularly the kinetics of whisker growth and the cessation (sudden stoppage) of whisker growth. A theoretical hypothesis must be consistent with the direct and visually evident observations provided by microstructural analysis. That is, any whisker recrystallization hypothesis must demonstrate that

FIG. 24 A schematic diagram of a secondary recrystallization grain structure in which one grain has grown at the expense of its immediate neighbors. (From Ref. 50.)

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the microstructures of whisker structures are consistent with the expected morphology of recrystallized grain structures.

A.

Recrystallization and Tin-Whisker Growth Kinetics

Growth-rate data for Sn whiskers have been limited because of the difficulty of making such measurements. Tin-whisker formation and growth are well described as erratic. The delay time (incubation period) for whisker formation can be hours or years, and seemingly identical platings can give significantly different whisker densities and growth rate results. However, some whisker growth rate data is available, and there is also data on the speed with which grain boundaries move during recrystallization. These data will be examined in the following sections.

B.

Growth Kinetics of Spontaneously Grown Whiskers

Whisker growth rates for spontaneously grown whiskers are much lower than those measured in clamp-pressure experiments. The rates cited for all reported spontaneous, whisker growth-rate data are less than 1.5 A˚/sec. Table 6, which is a summary of the reported data, indicates that the observed growth rates range from 0.01 to 1.50 A˚/sec, with most of the reported data being between 0.01 and 0.10 A˚/sec. Growth rates for spontaneously grown Sn whiskers are relatively slow metallurgical reactions. For example, the growth rate of whiskers grown from the vapor phase has been reported to be 10,000–100,000 A˚/sec [51]. The rate of grain-boundary migration during recrystallization can be as high as 105 A˚/sec [52], and single crystals of materials such as germanium have been grown at rates of 2 104 A˚/sec [53]. A number of models, based on a variety of concepts, have been developed that readily rationalize observed whisker growth rates on the basis of grain boundary self-diffusion of Sn. These models make reasonable assumptions for the grain boundary self-diffusion coefficients of Sn because there are no direct measurements available. Generally, researchers assumed an activation energy for grain-boundary (or dislocation) diffusion to be some fraction of the well known activation energy for bulk lattice diffusion.

C.

Whisker Growth Rates and Cessation

1. Differences Among Whisker Growth Models All modern whisker models presume that the Sn atoms necessary to sustain whisker growth move to the vicinity of a whisker via grain boundaries or dislocation networks. Differences between the various models come after the Sn atoms arrive at the boundary of the whisker itself. Dislocation theorists presume that Sn atoms arrive at the whisker via grain-boundary diffusion, and then migrate into the whisker and supply dislocation-source mechanisms that continually generate dislocations, which then climb or glide into the whisker structure. The cracked oxide theory discussed in Section VII [53], presumes that local stress relief is achieved at defective (cracked) oxide locations that allow Sn atoms, that have migrated there through long-range diffusion, to

TABLE 6

Spontaneous Tin Whisker Growth Rates on Various Substrate Materials

Plating Type Sn Sn Sn Sn Sn Sn Sn

on on on on on on on

brass copper zinc copper-clad steel quartz steel Fe on brass

Whisker Growth Rates (A˚/sec)

Reference

0.100–1.500 0.080–0.840 0.240 0.008–0.012 0.010–0.340 0.032–0.075 0.002–0.008

Glazunova and Kudryavtsev [10] Glazunova and Kudryavtsev [10] Glazunova and Kudryavtsev [10] Zakraysek Ellis et al. [37] Ellis et al. [37] Ellis et al. [37]


885

build up a whisker structure. Similarly, the recrystallization theory also recognizes that Sn atoms must be delivered to the recrystallized region via long-range, grain-boundary diffusion, and that Sn atoms must then migrate from the grain boundaries into the recrystallized grains (or whisker) whose growing volume then causes the oxide layer to crack. A detailed model of the migration from a grain-boundary network into a whisker-nodule structure has been described [37]. Because whisker growth is a relatively slow metallurgical process, it is not difficult for any model to rationalize observed growth rates. It has been necessary to recognize that the self-diffusion of Sn atoms in the Sn matrix is not quite fast enough to rationalize model growth-rate projections with the observed data and, therefore, it is necessary to include the grain-boundary (or dislocation network) self-diffusion of Sn into any proposed model. A difficulty for all whisker models is the linear growth rate as a function of applied external pressure, and the sudden halt in growth (cessation) that have been reported [27,28] by some investigators. It seems more reasonable to many that whisker growth rates should continually decrease over time as the driving force internal stresses are gradually relieved. In point of fact, there have been some data reported [29] of whisker growth rates that did continually decrease over time. These seemingly contradictory results may simply reflect the complex and time varying stress states in electroplated Sn, but they do complicate modeling. Stress measurements in plated films are the focus of much experimental work, and the interpretation of stress measurement data has been the subject of considerable controversy. 2. Effect of Intermetallic Compounds on Stress State The generally accepted rationale for the varying state-of-stress in Sn electroplates involves the formation of intermetallic compounds (IMCs) within the Sn matrix and at the Sn/substrate interface. The metallic elements of Cu, Fe, Ni, and Bi form IMCs in combination with Sn. Typical commercial Sn-plating solutions contain small amounts (30–100 ppm) of Cu, Fe, Ni, and Bi in solution, and the substrate material itself can contaminate the plating bath and, therefore, the asplated structure. These IMC-forming elements readily come out of solution and create discrete IMCs that tend to locate at the grain boundaries of the as-plated structure. The state of stress caused by elements (such as Cu) coming out of solution and precipitating out in grain boundaries as an intermetallic compound is a matter of some controversy. The molar volume of the Cu intermetallic Cu6Sn5 is less than the molar volume of the equivalent number of Cu and Sn atoms in solution; therefore, the induced stress state from this particular IMC formation process should be tensile. However, if the Cu diffuses up from the substrate into the tin film and then forms an IMC at the grain boundary, or anywhere else in the film, the resultant stress will be compressive. Experimentally, it is not possible to determine where the Cu atoms of an IMC particle in a Sn film grain boundary originally came from so the experimental observation of IMCs in the plating cannot be unequivocally interpreted as an indicator of a compressive stress. Lee and Lee [35] reported time-zero internal stresses, as determined by a cantilever-beam technique, that were initially tensile, and then became increasingly compressive over time. Another study [55] reported that bright-tin cantilever beam specimens were initially compressively stressed, and then relaxed to a less-compressive level over time. In yet another study [48], it was observed that the average residual stress for Sn on Cu substrates was compressive, but for Sn with a Ni underlay over a Cu substrate the average residual stress was tensile. The consensus opinion on stresses in electroplated films is that Sn on Cu substrates forms an intermetallic compound at the interface over time, and that the intermetallic Cu6Sn5 is itself in a compressive stress state relative to the underlying Cu substrate. For the case of Sn over Ni, the intermetallic compound that forms over time is in a tensile stress state relative to the underlying Cu substrate. The state of stress in the Sn film is a much more controversial subject. It is generally recognized that the time zero (right after deposition) state of stress in a Sn film can vary from tensile to compressive depending on the plating process conditions, impurity concentrations, and the type and amount of additives in the plating solution. The Sn-film stress state over time is the real controversy. XRD work tends to indicate that the surface layers of the Sn film (i.e., the top 1–2 Am) develop very low stress states (3–10 MPa) that can be either tensile or compressive. However, the data is argumentative for several reasons. Stress states of 3–10 MPa are extremely low and right at the edge of the detection

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capability of XRD equipment. In other words, the measured strains are right at the background noise level of the equipment detection capability. Second, interpreting X-ray diffraction patterns requires certain assumptions about the nature of the stress field, which may not be true. X-ray data interpretation requires an assumption that the stress fields be biaxial (i.e., uniform in at least two perpendicular directions). On a large enough scale (>100-Am square), it is probably a good approximation that the stress fields are biaxial. But when trying to discriminate stress gradients on a very small scale (95%) Sn electroplating. The application conditions must be understood as regards to temperature, humidity, expected length of service, critical spacings between adjacent circuit features under electrical bias, and the required reliability performance. The use of high-Sn content platings will require a trade off between cost and reliability. It is true that Sn platings have been used in many applications with great success, and with no evident whisker problems. But it is also true that there have been some significant and costly failures involving high-Sn content platings that grew whiskers and caused shorts in electrical circuitry. Users should seek expert advice before utilizing high-Sn content platings in critical applications. A list of mitigation practices, with no guarantees, is presented below: 1. Use alternative platings such as nickel (Ni)–palladium (Pd)–gold (Au)—Lead-frame finishes of Ni–Pd–Au have been successfully utilized in large quantities since 1992. This plating will not pass the relatively severe mixed-flowing corrosive gas tests required for some applications, but users must determine if this is a reasonable or necessary requirement because many commercial applications (e.g., business office environments) do not require meeting the harsh conditions imposed by a mixed-flowing gas test. The Au surface finish may be eliminated if long-term storage in uncontrolled environments is not an issue although it is often advisable to retain the Au layer because it promotes solder wettability. The solderability of Ni–Pd–Au with Pb-free tin–silver–copper (SAC) solders seems not to be an issue. 2. Fused Sn—Fused Sn lead-frame finishes were used for many years without any reported whisker problems. However, fused Sn will grow whiskers if scratched. 3. Tin over a Ni underlay (for Cu and Cu alloy substrates only—i.e, Sn/Ni/Cu)— Bright Sn over a Ni underlay on Cu substrates has been used for many years by a major electronics manufacturer without any reported whiskering problems. However, experiments have demonstrated that whiskers can be grown on Sn/Ni/Cu when subjected to accelerated thermal cycling, or when stored at 60jC/95% RH [49]. These whiskers are generally shorter (100 Am) whiskers in less than 3 months. Use of satin-bright Sn platings should only be carried out done(with expert advice and great care. 7. Bright tin—Bright Sn without an underlay is prone to growing long (100–300 Am) whiskers in normal (i.e., ambient) service conditions, and is not recommended for any application sensitive to whisker growth. Bright Sn over a Ni underlay has been used successfully by a major electronics manufacturer as a bus-bar and heat-sink finish without any known evidence of whisker growth. 8. Immersion tin—Immersion Sn is a relatively unstudied entity. Data from a major electronics manufacturing firm on the use of immersion Sn platings is encouraging. There were no reported Sn whisker problems in field service where the environments were relatively controlled (e.g., a typical office or computer room). Some published reports have indicated that immersion Sn films on printed circuit boards have grown whiskers inside the barrels of the plated through holes, but not on the surface film. It has very recently been reported that 0.6-Am-thick immersion-Sn films on copperfoil substrates have grown whiskers 10–15 Am in length within 60 days [62]. Use of immersion-Sn films should only be carried out with expert advice and great care.

X.

ACCELERATED TESTING

A.

Accelerated Test Considerations

There is no industry standard acceleration test to characterize the whisker-formation capability of a given film. Accelerated tests are typically performed at elevated temperatures to increase the reaction kinetics. For Sn films, it is necessary to keep the temperature below about 80jC to avoid the creation of intermetallics compounds (Cu3Sn) not seen at lower temperatures. At these lower temperatures whisker growth rates are still quite slow. At temperatures above 80jC, diffusion is much faster but the microstructure is not equivalent to the room temperature microstructure as a result of the formation of the Cu3Sn intermetallic compound. In addition, at higher temperatures the stresses are relieved quicker due to creep mechanisms, which decreases the driving force necessary for whisker formation. Hence, there exists a tradeoff between driving force and kinetics in standard temperature-based acceleration testing.

B.

Electromigration as a Whisker Formation Acceleration Test

The use of electromigration for conducting an accelerated test of Sn whisker growth has been suggested [42] by utilizing the classic Blech electromigration test [58] with Al films on a titanium nitride (TiN) base. In this test, Al atoms are current driven from the cathode to anode, which results in a compressive stress at the anode accompanied by nodule growth. The reader is directed to Chapter 20 of this book for a detailed discussion of electromigration. The potential advantage of electromigration as a means to determine sensitivity for Sn-whisker growth is that the driving force (current density) and the kinetic parameter (test temperatures) can be controlled independently. 1. Test and Model for Whisker Formation and Growth Fig. 48(a)–(d) shows the growth of a Sn whisker at the anode of a pure Sn Blech test sample under electromigration conditions. Measuring the growth rate and diameter of a whisker over time provides the volume change per unit time, V ¼ JAdtX;

ð8Þ


909

FIG. 48 Growth of a Sn whisker at the anode of a test sample of pure Sn under electromigration at 150jC and 105 A/cm2 for: (a) 20 hr, (b) 40 hr, (c) 60 hr, (d) 80 hr. (Courtesy of W. Choi, UCLA.)

where J is the electromigration flux, A is the cross section of the whisker, dt is unit time, and X is the atomic volume. The stress driving the growth of a whisker can then be determined from Eq. (6). Knowing J, Z* can thus be determined because J¼C

D kT

drX þ Z*ejq dx

ð9Þ

where C=1/X in pure Sn, D is diffusivity, kT is thermal energy, r is the stress at the anode, dr/dx is the stress gradient along the short strip of Sn of length of dx, Z* is the effective charge number of the diffusing Sn atoms in electromigration, e is the electron charge, j is the current density, r is the resistivity of Sn at the test temperature, and it is assumed that the stress at the cathode is zero. The stress, r, can be determined from Eq. (5), and the value of Z* can be determined from Eq. (9) because all other parameters are known. It is difficult to measure the size and volume of an irregularly shaped nodule emerging from the surface of a deposited film. According to Choi et al. [42], regularly shaped nodules of constant cross section can be obtained from the Blech test by depositing a thin coating of quartz over the entire Sn strip and then etching holes of given diameter at the anode, as depicted in Fig. 49(a) and (b). An applied current then causes whiskers to be pushed out from the holes etched at the anode end of the film. Using this technique, the growth rate of a whisker can be measured as a function of current density, temperature, time, and distance from the anode. The authors [42] also noted that it was possible to replace the Sn strip with eutectic Sn–Cu, or any other composition strip, for a comparison of whisker formation characteristics. However, this electromigration test may not be meaningful until it can be confirmed that a whisker

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FIG. 49 (a) A schematic diagram of the cross section of a Sn strip deposited on a TiN baseline, coated with a sputtered quartz film and with holes etched at the anode. (b) Under electromigration conditions, Sn whiskers push out of the holes. (From Ref. 42.)

driven by electromigration has the same basic atomistic growth mechanism as a spontaneously grown whisker.

XI.

SUMMARY AND FUTURE OUTLOOK

A.

Dislocation-Based Theories

As yet, there is no consensus within the scientific community on a model for whisker formation and growth. Dislocation theories should probably be regarded as nonessential, as was admitted by F.C. Frank, one of the original proponents of dislocation-based whisker growth [36]. The main objection to a dislocation-based whisker theory is that not all of the whisker growth directions are parallel to glide planes.

B.

Recrystallization and Cracked Oxide Theories

The recrystallization and cracked-oxide concepts are not necessarily contradictory. Recrystallization postulates a localized recrystallization event that causes the Sn oxide layer to crack from pressure exerted by the growing whisker nodule. The cracked oxide theory (COT) states that a locally weak oxide layer allows a whisker nodule to form and grow, thereby relieving the local stresses. However, the COT concept require the metal to be an oxide former. Oxide-free metal surfaces, such as gold, act as sinks for interstitial Sn atoms and lattice vacancies, and all internal stresses would tend to be uniformly relieved across the entire surface layer. Local relief mechanisms, such as whisker-nodule formation, are not be necessary. Recrystallization and cracked-oxide concepts both recognize that heavily oxidized surfaces act to prevent whiskernodule growth because of the strength of the oxide layer. One of the authors (G. Galyon) favors the recrystallization concept because experimental data shows that grains from which the whiskers grow always appear ‘‘different’’ than as-plated microstructures, and in some cases (bright Sn) the whiskers grow from grain structures that exhibit the classical appearance of recrystallized structures. There is, as yet, no direct experimental evidence for the existence of weak-oxide layers prior to whisker formation and growth. Most theoretical proposals for whisker formation recognize compressive stresses as the driving force behind whisker growth. The


911

Galyon–Palmer Integrated Recrystallization Theory is somewhat different because it recognizes a positive stress gradient, rather than a compressive stress, as the driving force. The stresses in a film could be entirely tensile, but if a localized positive stress gradient exists across a microstructurally recognizable boundary separating a surface grain from the surrounding region, whiskers or nodules can grow.

C.

Observations Differ Depending on Type of Tin Deposit

It seems clear from focused ion beam (FIB) microstructural analysis, energy-dispersive analysis, and cantilever-beam stress data that integration of multiple concepts is necessary to explain the various phenomena observed in whisker formation and growth. Recrystallization appears to play a role in bright-Sn whisker formation [57], whereas in matte Sn it remains to be seen if whiskers growth involves a recrystallization event. Matte-Sn microstructures often show large variations in as-plated morphologies, and some matte-Sn microstructures are quite similar to columnar bright-Sn microstructures; while other matte-Sn structures have been observed to consist of large, well-polygonized structures. Some preliminary data appears to indicate that even well-polygonized structures can grow whiskers under certain test conditions (e.g., temperature cycling or storage at 60jC/95% RH). An FIB cross section of a Sn whisker grown from well-polygonized matte-Sn under accelerated thermal cycle testing (45 to +65jC) is shown in Fig. 35. The appearance of the grain from which the whisker has grown is different than the grains of the immediately surrounding region. However, it is not clear whether the different appearing grain formed as a result of recrystallization subsequent to plating or formed during the plating process. Additional investigation should be directed at these large-grained, matte-Sn plated structures. It is not unreasonable that these different appearing matte-Sn grains from which the whiskers grew were caused by a localized recrystallization event, either during or after plating. This could occur in a manner similar to the formation of the nodule-whisker structures observed in bright-Sn films. Analyses of fused and immersion-Sn microstructures should be a subject of future research. Data on fused Sn indicates it is necessary to cold work the as-plated structure before whiskers will grow. A microstructural assessment of fused and immersion-Sn structures would be helpful in clarifying the modeling issues.

ACKNOWLEDGMENTS The authors Choi, Lee, and Tu would like to thank the Semiconductor Research Corporation contract #NJ-853 (Dr. Harold Hosack), and National Semiconductor Corporation (Dr. Luu Nguyen) for the financial support for the research described in this chapter. The authors would also like to thank George T.T. Sheng and C.F. Hu at Macronix International Company, Hsinchu, Taiwan, for some of the focus ion beam images of Sn whiskers. They would also like to thank N. Tamura at Advanced Light Source, Lawrence Berkeley National Laboratory for X-ray microdiffraction study of Sn whiskers. George Galyon would like to acknowledge the advice and counsel of a number of people in constructing portions of this chapter. To Dr. Karl Puttlitz of IBM for his leadership in lead elimination within IBM, and his encouragement of Sn-whisker research as an active member of NEMI (National Electrical Manufacturer’s Initiative Consortia). To Ron Gedney for his leadership as the NEMI coordinator of three Sn-whisker committees, one on modeling, one on User concerns, and one on the development of accelerated testing for Sn-whisker formation and growth. To my NEMI colleagues on the Whisker Modeling and USER GROUP Committees; Irina Boguslavsky, Peter Bush at SUNY, S. Lal at FCI, Chen Xu at Cookson Electronics, Maureen Williams at NIST, Valeska Schroeder at HP, Joe Smetana at Alcatel, Jack McCullen at Intel, Nick Vo at Motorola, Ben Huang at Indium Corp., Jim Martin and Neil Brown at Shipley, B. Radharkrishnan at Oak Ridge National Laboratories, Don Abbott and Doug Romm at TI, Ben Huang at Indium Corp., Rick Coyle at Lucent Corp., Dave Love at Sun Microsystems, Rick Charboneau at Storage Tek, Lynda Anderson at IBM, and Rich Parker at

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Delphi Electronics. . .thanks for the many stimulating discussions and insights. To Larry Palmer, my colleague at IBM Server Group, whose skill and initiatives produced many of the FIB microstructural cross section so useful in advancing the knowledge base on Sn-whisker formation and growth.

REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13. 14. 15. 16. 17. 18. 19. 20. 21. 22. 23. 24. 25. 26. 27. 28. 29. 30. 31. 32. 33.

Arnold, S.M. The growth and properties of metal whiskers. Proc. Am. Electroplaters Soc. 1956, 43, 26–31. Arnold, S.M. The growth of metal whiskers on electrical components. Proc. Elec. Components Conference; 1959; 75–82. Hedlund, L. Private communication: IBM Materials and Process Engineering Dept., eServer Group, Rochester, MN. Gutierrez, S. Private communication, IBM ECAT line, Austin, TX. Ranadive, N. Private communication, IBM LPAT line, Poughkeepsie, NY. Galyon, G.T. Private communication, IBM Corp., Systems Development Group, Poughkeepsie, NY. Britton, S.C. Spontaneous growth of whiskers on tin coatings: 20 years of observation. Trans. Inst. Met. Finish. April 1974, 52, 95–102. Lee, B.Z.; Lee, D.N. Spontaneous growth mechanism of tin whiskers. Acta Metall. 1998, 46 (10), 3701–3714. Arnold, S.M. Repressing the growth of tin whiskers. Plat. Mag. Jan. 1966, 53 (1), 96–99. Glazunova, V.K.; Kudryavtsev, N.T. An investigation of the conditions of spontaneous growth of filiform crystals on electrolytic coatings. (Translated from) Z. Prikl. Khim. March 1963, 36, 543– 550. Cobb, H.L. Cadmium whisker. Mon. Rev. Am. Electroplaters Soc., Jan. 1946, 33 (94), 28–30. Compton, K.G.; Mendizza, A.; Arnold, S.M. Filamentary growths on metal surfaces—‘whiskers’. Corrosion 1951, 7, 327–334. Dunn, B.D. Whisker formation in electronic materials. Circuit World July 1976, 2 (4), 32–40. McDowell, M.E. Tin Whiskers: A Case Study. Aerospace App. Conf., 1993; 207–215. Rozen, M. Practical whisker growth control methods. Plat. Mag., November 1968, 55 (11), 1155– 1168. M.Rozen, M. Effects of temperature on bright acid tin-plated deposits. Plating, October 1970, 57 (10), 1019–1024. Jafri, A. Fighting whisker growth in the communication industry. Plating, April 1973, 60 (4), 358– 359. Sabbagh, N.A.J.; McQueen, E.H.J. Tin whiskers: causes and remedies. Met. Finish., March 1975; 27–31. G.E. Web site: http://www.geindustrial.com/pm/support/dls/dlssb01.pdf. Galyon, G.T. Private communication, IBM Corp., Systems Development Group, Poughkeepsie, NY. Xu, C.; Fan, C.; Zhang, Y.J. Whisker prevention. Proceedings of the 2003 APEC Conf., April 2003; 1–8. Brusse, J. Tin whisker observations on pure tin-plated ceramic chip capacitors. AESF/SurFin Conf., June 2002; 45–60. Peach, M.O. Mechanism of growth of whiskers on cadmium. J. Appl. Phys. 1952, 23, 1401. Koonce, S.E.; Arnold, S.M. Growth of metal whiskers. J. Appl. Phys. 1953, 24, 134–135. letters to the editor. Frank, F.C. On tin whiskers. Philos. Mag. August 1953, 54, 854. Eshelby, J.D. A tentative theory of metallic whisker growth. Phys. Rev. 1953, 91, 774–775. letters to the editor. Fisher, R.M.; Darken, L.S.; Carroll, K.G. Accelerated growth of tin whiskers. Acta Metall. May 1954, 2, 368–373. Glazunova, V.K. A study of the influence of certain factors on the growth of filamentary tin crystals. (Translated from) Kristallografiya Sept.–Oct. 1962, 7 (5), 761–768. Pitt, C.H.; Henning, R.G. Pressure-induced growth of metal whiskers. J. Appl. Phys. 1964, 35, 459– 460. communications. Hasiguti, R.R. A tentative explanation of the accelerated growth of tin whiskers. Acta Metall. 1955, 3, 200–201. letters to the editor. Amelinckx, S.; Bontinck, W.; Dekeyser, W.; Seitz, F. On the formation and properties of helical dislocations. Philos. Mag., Ser. March 1957, 8, 2 (15), 353–377. Franks, J. Metal whiskers. Nature May 1956, 7, 984. letter-to-the-editors. Franks, J. Growth of whiskers in the solid phase. Acta Metall. Feb. 1958, 6, 103–109.

STRUCTURE AND KINETICS OF TIN-WHISKER 34. 35. 36. 37. 38. 39. 40. 41. 42. 43. 44. 45. 46. 47. 48. 49. 50. 51. 52. 53. 54. 55. 56. 57. 58. 59. 60. 61. 62.

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Lindborg, U. A model for the spontaneous growth of Zn, cadmium, and tin whiskers. Acta Metall. 1976, 24, 181–186. Lee, B.Z.; Lee, D.N. Spontaneous growth mechanism of tin whiskers. Acta Metall. 1998, 46 (10), 3701–3714. Selcuker, A.; Johnson, M. Microstructural characterization of electrodeposited tin layer in relation to whisker growth. Capacitor and Resistor Technology Symposium Proceedings, Oct. 1990; 19–22. Ellis, W.C.; Gibbons, D.F.; Treuting, R.G. Growth of Metal Whiskers from the Solid. Growth and Perfection of Crystals; John Wiley & Sons: New York, NY, 1958; 102–120. Koonce, S.E.; Arnold, S.M. Growth of metal whiskers. J. Appl. Phys. 1953, 24, 365. letters to the editor. Fujiwara, K.; Kawanaka, R. Observations of the tin whisker by micro-Auger electron spectroscopy. J. Appl. Phys. Dec. 1980, 51 (12), 6231–6233. Tu, K.N. Cu/Sn interfacial reactions: thin-film versus bulk case. Mater. Chem. Phys. 1996, 46, 217– 223. Baudry, I.; Kerros, G. Focused ion beam in microelectronics packaging applications—lead-free plating analysis. Solder. Assem. Technol., Issue 3, 2001. www.lead-free.org. Choi, W.J.; Lee, T.Y.; Tu, K.N.; Tamura, N.; Celestre, R.S.; MacDowell, A.; Bong, Y.; Nguyen, L.; Shen, G.T.T. Structure and kinetics of tin whisker growth on Pb-free solder finish. IEEE—Proc. of 52nd Elec. Components and Technology Conf., San Diego, CA, May 2002; 628–633. Gorbunova, K.M.; Glazunova, V.K. Present state of the problem of spontaneous growth of whisker crystals on electrolytic coatings. (Translated from) Zas. Met. May–June 1984, 20 (3), 342–358. Stevens, W. Private Communication, Brookhaven National Laboratory: Internal Report to NEMI Modeling Committee, 2002. Boguslavsky, Irina. Private Communication, NEMI Modeling Committee, 2002. Tu, K.N. Interdiffusion and reaction in bimetallic Cu–Sn thin films. Acta Metall. April 1973, 21 (4), 347–354. Sheng, G.T.T.; Hu, C.F.; Choi, W.J.; Tu, K.N.; Bong, Y.Y.; Nguyen, L. Tin whiskers studied by focused ion beam imaging and transmission electron microscopy. J. Appl. Phys. July 2002, 92 (1), 64– 69. Xu, C.; Zhang, Y.; Fan, C.; Abys, J.; Hopkins, L.; Stevie, F. Understanding whisker phenomenon— driving forces for the whisker formation. IPC APEX, S-06-2-1–S06-2-6. Vo, V. NEMI Workshop—Fall 2002; IPC/APEX workshop—Nov. 2002. Kakeshita, T.; Shimizu, K.; Kawanaka, R.; Hasegawa, T. Grain size effect on electroplated tin coatings on whisker growth. J. Mater. Sci. 1982, 17, 2560–2566. Sears, G.W. A mechanism of whisker growth. Acta Metall. July 1955, 3, 367–369. McLean, D. Grain Boundaries in Metals. Clarendon Press: Milltown, NJ, 1957. Chaps. 8 and 9. Bennet, D.C.; Sawyer, B. Single crystals of exceptional perfection and uniformity by zone refining. Bell Syst. Tech. J. 1956, 35 (3), 637. Tu, K.N. Irreversible processes of spontaneous whisker growth in bimetallic Cu–Sn thin-film reactions. Phys. Rev. B, January 1994; 2030–2034. Williams, M. Private Communication, NEMI Modeling Committee, December 2002. Ke, T. Experimental evidence of the viscous behavior of grain boundaries in metals. Phys. Rev. 71, 533. Galyon, G.; Palmer, L. Integrated Recrystallization Theory for Whisker Formation. NEMI Workshop, TMS, March 2002. Blech, I.A.; Petroff, P.M.; Tai, K.L.; Kumar, V. Whisker growth in Al thin-films. J. Cryst. Growth Feb. 1976, 32 (2), 161–169. Chang, C.Y.; Vook, R.W. The effect of surface aluminum oxide films on thermally induced hillock formation. Thin Solid Films May 1993, 32 (2), 205–209. Kehrer, H.P.; Kadereit, H.G. Tracer experiments on the growth of tin whiskers. Appl. Phys. Lett. June 1970, 16 (11), 411–412. Boguslavsky, I.; Bush, P. Analysis Report for NEMI; May 16, 2002. Galyon, G. Private communication, IBM eSystems Group.

22 Degradation Phenomena Michael J. Sullivan Lloyd Technology Associates, Inc., Katonah, New York, U.S.A.

Stephen J. Kilpatrick U.S. Army Research Laboratory, Adelphi, Maryland, U.S.A.

I. A.

INTRODUCTION General

The move to lead-free technology is ironically juxtaposed against the unparalleled virtues of lead (Pb)-based solders. Tin–lead solder is virtually an ideal material in nearly every regard. As will be discussed in this chapter, the features of that binary phase diagram, the mechanical properties of that alloy, and its thermomechanical fatigue behavior in service are extremely well suited to its utilization in microelectronic packages, providing a family of solders with a range of melting temperatures for a variety of applications. Lead-free systems, on the other hand, offer fewer options.

B.

Temperature Hierarchy

The challenge of identifying a replacement for lead-bearing solders requires satisfying the desirable attributes listed in Table 1 [1]. A key feature of the Pb–Sn solder system is that it provides a family of solders that melt over a range of temperatures which is very useful during assembly operations. For example, eutectic Sn–Pb (63–37) melts at 183jC, and a whole range of hypoeutectic (low Sn) alloys has been utilized for a variety of applications, up to the 97Pb–3Sn alloy (MP=325jC) utilized by IBM for flip chip solder bumps. A high melting point alloy (i.e., high Pb content) is utilized for chip solder bumps because they are attached early in the process. A lower melting point solder (i.e., lower Pb content) is utilized to attach ball grid array (BGA) components to cards, which occurs later in the process. In other words, the highest melting temperature solder is reflowed first and it does not remelt during a subsequent reflow operation, often referred to as a solder hierarchy. Use of a solder hierarchy can be beneficial in other ways also, such as cases where an underfill is used with a flip chip die to avoid the solid-to-liquid volume expansion or concomitant hydrostatic pressure that would result if the solder melted. Ideally, in migrating to a Pb-free system, it would also be desirable that it, too, can provide a series of alloys that melt over a broad range of temperatures to preserve the assembly benefits provided by a solder hierarchy.

915

916

SULLIVAN AND KILPATRICK

TABLE 1

Requirements for Suitable Alternative Pb-Free Alloys

Requirements Low cost Relatively nonhazardous No potential environmental problems; ranking in decreasing order of safety: Bi > Zn>In>Sn>Cu>Sb>Ag>Pb Capable of forming reliable joints Compatible with relatively low-temperature processing Corrosion-resistant Compatible with copper substrates, immersion gold-over-nickel, and a variety of nonlead substrates, as well as leaded substrates Available in bar, wire, and paste forms Available in sufficient quantities Low melting temperatures (1 1016 1 1014 >1 1018 2 1016 2 1015 6 1011 1.4 1010 7.0 108 4.5 109

Ce143 Nd144 Nd147 Sm147 Sm148 Sm149 Gd152 Dy156 Hf174 Os186 Pt190 Th232 U235 U238 Source: Ref. 67.

Relative specific activity 2 108 8 107 7109 0.01 8 108 8 108 2 107 4 1012 3 108 5 108 1 106 0.34 0.005 1.0

DEGRADATION PHENOMENA TABLE 10 Packages

943

The Most Common Sources of Alpha Particles in Microelectronic

Sources Pb–Sn solder Alumina substrates BEOL metallizations Fillers in plastics, encapsulants, underfills, mold compounds, and solder masks Flux Lead frame alloys Gold used for wirebonding and lid plating Particulates from plastic ball grid array (PBGA) trimming/handling operations Source: Ref. 67.

from lead-containing solders: U238 (uranium, 4.5 109 years) and Po210 (polonium, 138 days), corresponding to a Po210 specific activity about 1.2 1010 greater compared to U238. This implies that 1.2 ppm of U238 in the solder has the same SER effect as 1010 ppm of Po210. The startling reality is that just a few Po210 atoms located on the die active surface pose as an unacceptable SER risk [84,87].

C.

Elimination of Alpha Emitters from Lead Solders

Soft errors in microelectronic packages are a difficult, troublesome problem to address. Eliminating them is virtually impossible. Soft error rate effects can be reduced in a number of ways, including the use of shielding, software correction methods (although this reduces performance), alpha particle-tolerant structures (e.g., SOI), and utilization of low-alpha-emitting materials. Let us now focus on the reduction of alpha particle-induced SER from traditional leadbased solders.

TABLE 11 Alpha Radiation Activity of Some Common Materials Utilized in Microelectronic Packages Material Processed wafers Cu metal (thick) Al metal (thick) Mold compound Underfill Pb solders ‘‘Hot’’ Pb LC II Pb (HEM) LC I Pb (HEM) Alloy 42 (Hitachi) Au-plated alloy 42 (HEM) ELJ125/Au-plated alloy 42 (HEM) Sn (HEM) AlSiC (Lanxide) LC6 Al (HEM) Source: Refs. 68 and 69.

Alpha radiation flux (a/khr cm2) 0.9 1.9 1.4 24 to

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