Manufacturing Processes 1: Cutting (RWTHedition)

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Author: Fritz Klocke

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RWTH edition RWTH Aachen

Fritz Klocke

Manufacturing Processes 1 Cutting

Translated by Aaron Kuchle

123

Professor Dr.-Ing. Dr.-Ing. E.h. Dr. h.c. Dr. h.c. Fritz Klocke Laboratory for Machine Tools and Production Engineering RWTH Aachen University Steinbachstr. 19 52074 Aachen Germany [email protected]

Translated by Aaron Kuchle Yeungnam University Korean Language Institute 214-1 Dae-dong Gyeongsan Gyeongbuk 712-749 Korea [email protected]

ISSN 1865-0899 e-ISSN 1865-0902 ISBN 978-3-642-11978-1 e-ISBN 978-3-642-11979-8 DOI 10.1007/978-3-642-11979-8 Springer Heidelberg Dordrecht London New York Library of Congress Control Number: 2011925556 © Springer-Verlag Berlin Heidelberg 2011 This work is subject to copyright. All rights are reserved, whether the whole or part of the material is concerned, specifically the rights of translation, reprinting, reuse of illustrations, recitation, broadcasting, reproduction on microfilm or in any other way, and storage in data banks. Duplication of this publication or parts thereof is permitted only under the provisions of the German Copyright Law of September 9, 1965, in its current version, and permission for use must always be obtained from Springer. Violations are liable to prosecution under the German Copyright Law. The use of general descriptive names, registered names, trademarks, etc. in this publication does not imply, even in the absence of a specific statement, that such names are exempt from the relevant protective laws and regulations and therefore free for general use. Cover design: deblik, Berlin, Germany Printed on acid-free paper Springer is part of Springer Science+Business Media (www.springer.com)

Preamble About Compendium “Manufacturing Processes”

Key factors for quality and economic efficiency of industrial production are the choice of the manufacturing processes and their design. Manufacturing Technology is an elemental part of the fundamental knowledge of machining engineers. Also design engineers have to gain knowledge in this field, since they have high responsibility for the manufacturing costs. However, the students as well as practising experts who are willing to enhance their knowledge have the problem to collect information. To the current day there is no extensive, but still clear description of manufacturing processes focussing on the technology itself. In order to counter this necessity the compendium at hand is supposed to present an overall picture of the most common machining and non-machining manufacturing processes. Additional to the description of the techniques these volumes are desired to deliver an insight in the underlying physical principles whenever it is necessary for the understanding of the processes. The apportionment of the compendium “Manufacturing Processes” into Volume 1: Cutting Volume 2: Grinding, Honing, Lapping Volume 3: Electrical Erosion and Hybride Processes Volume 4: Forming Volume 5: Casting, Sintering, Rapid Prototyping groups techniques with similar active principles together. In front of the first volume is placed a technique-spanning section to the tolerances and questions of the workpiece measuring techniques used in manufacturing. Within the individual volumes was tried to avoid an encyclopaedic listing of the techniques. The book series are primarily intended for junior scientists in the fields of manufacturing technology and construction. In addition, the practitioner will be able to refurbish or extend his knowledge. The variety of manufacturing problems

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Preamble About Compendium “Manufacturing Processes”

is as large as the multiplicity of the products, and alone with text book wise sayings manufacturing questions are not to be solved. We wish that this book offers starting points and ways to its readers, on which they can come up with successful solutions by engineering thinking. Aachen, Germany September 2010

Fritz Klocke

Preamble for Volume 1 “Cutting”

The available volume treats machining with geometrically defined cutting edges. It approaches both the practical engineer and the student of engineering sciences. This book is based on the lecture “Manufacturing Technologies I and II” and the pertinent exercises, which are held at RWTH Aachen University. The arrangement of the book results from the experiences, which were gained when lecturing in the chronology of the course. The structure of the book is oriented to a large extent at didactical criteria. The first section deals with the metrology and the workpiece quality, followed by a chapter about the fundamentals of cutting. Based on the common technological principle of the cutting processes and their variants, the second section deals with the mechanisms occurring at the cutting edge and the resulting loads on the tools. The next section deals with the demands on the cutting materials as well as their manufacturing processes and application. The text presents topics like the use of the finite element method for cutting processes as well as cutting fluids. A detailed chapter about the machinability of the most important work piece materials, such as steel and cast iron, light metal alloy and high temperature resistant materials imparts a broad knowledge of the alternating coactions of work piece material, the cutting material and the process parameters. The necessity for process control, used sensor systems and application possibilities are regarded. Due to the increasing relevance and application in industrial manufacturing plants, monitoring systems are discussed in their own chapter. Finally, the different cutting processes and their kinematic characteristics are presented in detail. For their cooperation with the compilation of the available edition I would like to thank my co-workers, Dr.-Ing. M. Abouridouane, Dr.-Ing. K. Gerschwiler, Dipl.-Ing. M. Arft, Dipl.-Ing. C. Essig, Dipl.-Ing. P. Frank, Dipl.-Ing. A. Krämer, Dipl.-Ing. D. Lung, Dipl.-Ing. H. Sangermann, Dipl.-Ing. R. Schlosser, Dipl.-Ing.

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Preamble for Volume 1 “Cutting”

P. Vogtel, as well as Dipl.-Ing. S. E. Cordes, who was additionally responsible for the coordination of the work on this book. Further I would like to thank the many former co-workers, who participated in the past German editions and now have leading positions in industry and research companies. Aachen, Germany September 2010

Fritz Klocke

Contents

1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 Metrology and Workpiece Quality . . . . . . . . . . . . . . 2.1 Manufacturing Disturbances and Manufacturing History 2.2 Measuring and Testing . . . . . . . . . . . . . . . . . . 2.2.1 Measurement Errors . . . . . . . . . . . . . . 2.2.2 Macro- and Microgeometry of Components . . 2.3 Length Testing Devices . . . . . . . . . . . . . . . . . . 2.3.1 Material Measures . . . . . . . . . . . . . . . 2.3.2 Gauges . . . . . . . . . . . . . . . . . . . . . 2.3.3 Indicating Measuring Instruments . . . . . . . 2.4 Surface Inspection . . . . . . . . . . . . . . . . . . . . 2.4.1 Surface Parameters . . . . . . . . . . . . . . . 2.4.2 Subjective Surface Inspection . . . . . . . . . 2.4.3 Surface Measurement . . . . . . . . . . . . . . 2.5 Inspection of the Workpiece Rim . . . . . . . . . . . . 2.5.1 Surface Layers . . . . . . . . . . . . . . . . . 2.5.2 Inspection of the Surface Rim Zone . . . . . .

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3 3 8 9 10 13 13 13 14 21 22 26 27 30 31 33

3 Fundamentals of Cutting . . . . . . . . . . . . . . . . . . . . . 3.1 The Cutting Part – Concepts and Terms . . . . . . . . . . 3.2 Reference Systems . . . . . . . . . . . . . . . . . . . . . 3.3 Basic Process Variants . . . . . . . . . . . . . . . . . . . 3.4 Chip Formation . . . . . . . . . . . . . . . . . . . . . . . 3.4.1 The Cutting Process . . . . . . . . . . . . . . . 3.4.2 Different Types of Chip Formation . . . . . . . . 3.5 Kinematic Surface Roughness . . . . . . . . . . . . . . . 3.6 Mechanical and Thermal Strain on the Cutting Section . . 3.6.1 Influence of the Geometry of the Cutting Section 3.7 Wear . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.7.1 Wear Mechanisms . . . . . . . . . . . . . . . . 3.7.2 Causes of Wear . . . . . . . . . . . . . . . . . . 3.7.3 Forms and Dimensions of Wear . . . . . . . . .

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39 39 42 45 47 47 49 54 55 68 71 71 74 85

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Contents

3.8

Cutting Theory . . . . . . . . . . . . . . . . . . . . . . . . . . 3.8.1 Shear Plane Theory . . . . . . . . . . . . . . . . . . . 3.8.2 Application of Plasticity Theory to Cutting . . . . . .

4 Cutting Tool Materials and Tools . . . . . . . . . . . . . . . 4.1 Overview of Cutting Tool Materials . . . . . . . . . . . 4.1.1 Classification of Hard Cutting Tool Materials . 4.2 Tool Steels . . . . . . . . . . . . . . . . . . . . . . . . 4.2.1 Cold Work Steels . . . . . . . . . . . . . . . . 4.2.2 High Speed Steels . . . . . . . . . . . . . . . 4.3 Cemented Carbides . . . . . . . . . . . . . . . . . . . . 4.3.1 Historical Development . . . . . . . . . . . . 4.3.2 Cemented Carbide Production . . . . . . . . . 4.3.3 Components of Cemented Carbides and Their Properties . . . . . . . . . . . . . . . . . . . . 4.3.4 Microstructure . . . . . . . . . . . . . . . . . 4.3.5 Classification of Cemented Carbides . . . . . . 4.3.6 Uncoated Cemented Carbides and Cermets . . 4.3.7 Function Gradient Cemented Carbides . . . . . 4.3.8 Coated Cemented Carbides and Cermets . . . . 4.4 Coatings . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.1 Coating Methods . . . . . . . . . . . . . . . . 4.4.2 Specific Properties of Hard Material Coatings . 4.4.3 Substrate Pretreatment . . . . . . . . . . . . . 4.5 Ceramic Cutting Tool Materials . . . . . . . . . . . . . 4.5.1 Cutting Ceramics . . . . . . . . . . . . . . . . 4.5.2 Superhard Non-metallic Cutting Tool Materials 4.6 Tool Designs . . . . . . . . . . . . . . . . . . . . . . . 4.6.1 Solid Tools . . . . . . . . . . . . . . . . . . . 4.6.2 Tools with Inserts . . . . . . . . . . . . . . . . 4.7 Tool Preparation . . . . . . . . . . . . . . . . . . . . .

87 87 92

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114 116 117 123 123 126 127 128 149 161 161 162 173 184 185 187 193

5 Finite Element Method (FEM) . . . . . . . . . . . . . . . . . 5.1 Basic Concepts of FEM . . . . . . . . . . . . . . . . . . 5.2 Lagrangian and Eulerian Considerations of the Continuum 5.3 Explicit and Implicit Methods of Solution . . . . . . . . . 5.4 Combined Thermal and Mechanical FEA . . . . . . . . . 5.5 Nonlinearities . . . . . . . . . . . . . . . . . . . . . . . . 5.6 Material Laws . . . . . . . . . . . . . . . . . . . . . . . 5.7 Software . . . . . . . . . . . . . . . . . . . . . . . . . . 5.8 Hardware . . . . . . . . . . . . . . . . . . . . . . . . . . 5.9 Phases of a Finite Element Analysis (FEA) . . . . . . . . 5.10 The Use of FEM in Cutting Technology . . . . . . . . . . 5.10.1 Continuous Chip Simulation . . . . . . . . . . . 5.10.2 Segmented Chip Simulation . . . . . . . . . . . 5.10.3 Simulation of the Cutting Process . . . . . . . .

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197 197 199 199 200 200 201 204 205 205 206 207 209 212

Contents

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6 Cutting Fluids . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1 The Functions of Cutting Fluids . . . . . . . . . . . . . . 6.2 Types of Cutting Fluids . . . . . . . . . . . . . . . . . . . 6.2.1 Non Water-Miscible Cutting Fluids . . . . . . . 6.2.2 Water-Based Cutting Fluids . . . . . . . . . . . 6.3 Guidelines on the Use of Coolant Emulsions . . . . . . . 6.4 Effects of the Cutting Fluid on the Machining Process . . 6.5 Selection of Cutting Fluids . . . . . . . . . . . . . . . . . 6.6 Reducing or Avoiding the Use of Cutting Fluids . . . . . . 6.6.1 Reducing Cutting Fluids . . . . . . . . . . . . . 6.6.2 Minimum Quantity Cooling Lubrication (MQCL) 6.6.3 Avoiding Cutting Fluids . . . . . . . . . . . . .

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219 219 219 220 221 222 223 226 227 228 229 233

7 Tool Life Behaviour . . . . . . . . . . . . . . . 7.1 Determining Tool Life Parameters . . . . 7.2 Machinability . . . . . . . . . . . . . . . 7.2.1 Tool Life . . . . . . . . . . . . 7.2.2 Resultant Force . . . . . . . . . 7.2.3 Surface Quality . . . . . . . . . 7.2.4 Chip Form . . . . . . . . . . . 7.2.5 Cutting Speed . . . . . . . . . . 7.3 The Machinability of Steel Materials . . 7.3.1 Metallographic Constituents . . 7.3.2 Carbon Content . . . . . . . . . 7.3.3 Alloying Elements . . . . . . . 7.3.4 Types of Heat Treatment . . . . 7.4 Machinability of Various Steel Materials 7.4.1 Machining Steels . . . . . . . . 7.4.2 Case-Hardened Steels . . . . . 7.4.3 Heat-Treatable Steels . . . . . . 7.4.4 Nitrided Steels . . . . . . . . . 7.4.5 Tool Steels . . . . . . . . . . . 7.4.6 Hardened Steels . . . . . . . . 7.4.7 Non-rusting Steels . . . . . . . 7.5 Machinability of Cast Iron . . . . . . . . 7.5.1 White Cast Iron . . . . . . . . . 7.5.2 Grey Cast Iron . . . . . . . . . 7.6 Machinability of Non-ferrous Metals . . 7.6.1 Aluminium Alloys . . . . . . . 7.6.2 Magnesium Alloys . . . . . . . 7.6.3 Titanium Alloys . . . . . . . . 7.6.4 Copper Alloys . . . . . . . . . 7.6.5 Nickel Alloys . . . . . . . . . . 7.7 Machinability of Non-metals . . . . . . . 7.7.1 Graphite . . . . . . . . . . . . . 7.7.2 Fibre-Reinforced Plastics . . . .

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237 239 239 240 243 245 248 248 249 250 252 254 256 263 263 266 267 268 269 271 274 280 280 283 290 290 293 296 311 314 327 327 331

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Contents

8 Process Design and Process Monitoring . . . . . . . . . . 8.1 Finding Economical Cutting Parameters . . . . . . . . 8.1.1 Cutting Parameter Limits . . . . . . . . . . . 8.1.2 Optimizing the Cutting Parameters . . . . . . 8.1.3 Calculating the Machine Hour-Rate . . . . . 8.1.4 Planning Methods and Tools . . . . . . . . . 8.2 Process Monitoring . . . . . . . . . . . . . . . . . . . 8.2.1 Sensors for Process Monitoring . . . . . . . 8.2.2 Signal Processing and Monitoring Strategies

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339 339 339 342 350 351 352 354 371

9 Processes with Rotational Primary Movement . . . . . . . . . . 9.1 Turning . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.1 Face Turning . . . . . . . . . . . . . . . . . . . . 9.1.2 Cylindrical Turning . . . . . . . . . . . . . . . . . 9.1.3 Helical Turning . . . . . . . . . . . . . . . . . . . 9.1.4 Profile Turning . . . . . . . . . . . . . . . . . . . 9.1.5 Form Turning . . . . . . . . . . . . . . . . . . . . 9.1.6 Further Process Variants . . . . . . . . . . . . . . 9.2 Milling . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2.1 Process Variants, Specific Characteristics and Tools 9.3 Drilling . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.1 Profile Counterboring . . . . . . . . . . . . . . . . 9.3.2 Rotary Drilling . . . . . . . . . . . . . . . . . . . 9.4 Sawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.4.1 Bandsawing . . . . . . . . . . . . . . . . . . . . . 9.4.2 Hacksawing . . . . . . . . . . . . . . . . . . . . . 9.4.3 Circular Sawing . . . . . . . . . . . . . . . . . .

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383 383 385 386 387 388 389 390 390 395 416 416 418 439 439 442 442

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445 445 446 449 451 454 457 458 459 466

References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

469

Subject Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

497

10

Processes with Translatory Primary Movement 10.1 Broaching . . . . . . . . . . . . . . . . . . 10.1.1 Face and Circular Broaching . . . 10.1.2 Profile Broaching . . . . . . . . . 10.1.3 Form Broaching . . . . . . . . . 10.2 Shaving . . . . . . . . . . . . . . . . . . . 10.3 Planing and Shaping . . . . . . . . . . . . 10.3.1 Face Shaping and Face Planing . 10.3.2 Gear Shaping . . . . . . . . . . . 10.3.3 Gear Planing . . . . . . . . . . .

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Symbols and Abbreviations

Capital Letters A A Ac1 Ac3 Aα Aα Aγ A AFK AIP AV BAZ Cf

mm2 %

Cv

min

CBN CFK CNC CVD CWT D DC DLC DMS DP DWT E EB-PVD

mm2 mm2 mm2 mm2

mm min

mm

Cross section Stretch ratio Transformation temperature (Start of austenite formation) Transformation temperature (End of austenite formation) Flank face Minor flank face Rake face Shear face Aramid Faser reinforced plastics Arc Ion Plating Offset Machining center Axis intercept of the expanded TAYLOR equation, f = 1 mm Axis intercept of the expanded TAYLOR equation, vc = 1 m/min Cubic boron nitride Carbon fibre reinforced plastics (CFRP) Computerised Numerical Control Chemical Vapour Deposition Continuous wavelet transform Tool diameter Direct current Diamond-like carbon Strain gauge Polycrystalline diamond Discrete wavelet transform Young’s modulus Electron beam-PVD

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xiv

ECM EDX F Fa Fc FD Ff Fp F FFT FVK GFK H HB HC HF HF-CVD HM HRC HSC HSS HT-CVD HV KA KE KI KF,min KMH KML KR KZ KB KM KSS KT L LARC MQCL MQC MQS MQL MSIP MT-CVD NC


N N N N N N N

mm

C C C C/St. C/h C C C mm mm mm C/h

Electro chemical machining Energy-dispersive X-ray spectroscopy Resultant force Active force Cutting force Thrust force Feed force Passive force Shear force Fast Fourier transform Fibre-reinforced plastic (FRP) Glass-reinforced plastic (GRP) Stroke length Brinell hardness Diamond coated cemented carbide Fine grain cemented carbide Hot-Filament-Process-CVD Cemented carbide Rockwell hardness High Speed Cutting High Speed Steel High temperature-CVD Vicker hardness Imputed amortizations Energy costs Maintenance costs Minimal manufacturing costs per workpiece Machine hour-rate Machine and labour cost hourly rate Room costs Imputed interest Crater width Crater center distance Cutting fluid Crater depth Hourly wage Lateral Rotating Arc Quantity Cooling Lubrication Minimum quantity cooling Minimum quantity lubrication Minimum quantity lubrication Magnetron Sputter Ion Plating Moderate Temperature-CVD Numerical control


NE-Metalle PA-CVD Pc Pe Pf Pf Pn Pne Po Pp Pr Ps PW PKD PM PM-HSS PVD Q R Ra Re REM Rkin Rm Rp0,2 Rt Rz RF S S SKV STFT SVα

W W W

W

C μm N/mm2 μm N/mm2 N/mm2 μm μm

μm mm

SVγ

mm

T T Tm Tr Tsp U U0 VB

min ◦C ◦C ◦C ◦C V V mm

xv

Non-ferrous metals Plasmaactivated-CVD Cutting power Effective power Assumed working plane Feed power Tool cutting edge normal plane Effective cutting edge normal plane Tool orthogonal plane Tool back plane Tool reference plane Cutting edge plane Effective power Polycrystalline diamond Powder metallurgy Powder Metallurgy High Speed Steel Physical Vapour Deposition Charge Electrical resistance Average roughness Elastic limit Scanning electron microscope (SEM) Surface quality Tensile strength 0.2%–yield strength (Depth of) Surface roughness Averaged surface roughness Radio Frequency Major cutting edge Minor cutting edge Rounded cutting edge Short-Time Fourier Analysis Displacement of the cutting edge in the direction of the flank face Displacement of the cutting edge in the direction of the rake face Tool life Temperature Melting temperature Room temperature Temperature on the chip upper side Voltage Feeding voltage Width of flank wear land

xvi

VBmax VHM Wc We Wf WZ-D WSP Z ZTU


mm J J J mm

Maximal width of flank wear land Cemented Carbide Cutting work Effective work Feed work Tool diameter Insert Atomic number Time Temperature Transformation (TTT)

Small Letters ae ap apmax apok a-C a-C:H at.-% B bfα C cp cpk dij dW/dt f fa fok fmax fmin fr fz f0 H hch hcu,max hmin K kc1.1 kf1.1 ki kp1.1

mm mm mm mm

mm mm m/min

pC/N m/min mm mm mm mm mm mm mm 1/s mm mm μm mm N/mm2 N/mm2 N/mm2 N/mm2

Width of cut Depth of cut Maximum depth of cut Optimal-cost depth of cut Hydrogenfree amorphous carbon film Hydrogenated amorphous carbon film Atomic ratio Width of undeformed chip Width of chamfer Speed of sound Process capability index Rectified process capability index Piezoelectric coefficient Growth of wear Feed Axial feed Optimal-cost feed Maximal feed Minimal feed Radial feed Feed per tooth/cutting edge Eigenfrequency Undeformed chip thickness Chip thickness Maximal chip thickness Minimal chip thickness Strain sensitivity Specific cutting force, b = h = 1 mm Specific feed force, b = h = 1 mm Specific resultant force Specific passive force, b = h = 1 mm


lc lf mc mn pKSS rε rß sd T T te,min tg tv te ter th tn tr ta-C vc vch vf vfax ve x z

m m mm bar mm mm mm s min min min min min min min min m/min m/min m/min mm/min m/min mm

xvii

Cutting path Feed path Exponent of the specific cutting force Normal module Cutting fluid pressure Corner radius Rounded cutting edge radius Coating thickness Pitch Time Minimal time per unit Basic time Additional time Time per unit Recovery time Main process time Auxiliary process time Setting-up time Tetrahedral hydrogenated amorphous carbon film Cutting velocity Chip velocity Feed velocity Axiale feed velocity Effective velocity Profile offset Number of teeth

Greek Letters α α eff αf αF αo α oe β βo β oe γo γn γ ne γ oe γ eff

◦ ◦ ◦ ◦ ◦ ◦ ◦ ◦ ◦ ◦ ◦ ◦ ◦ ◦

Tool clearance angle Effective tool clearance angle First tool orthogonal clearance (Broaching) Clearance angle at the pitch diameter (Gear shaping) Tool clearance angle Working clearance angle Wedge angle Wedge angle Working wedge angle Tool orthogonal rake angle Tool normal rake angle Working rake angle Wirk-Orthogonalspanwinkel Effective rake angle

xviii

ε ε¯ ˙¯ε ε˙ εk εr η ϑ κr κ re λs λse μ φ φ φf ϕ ρ σ σn ω


s–1 1/s ◦ ◦ ◦C ◦ ◦ ◦ ◦ ◦ ◦ ◦ ◦

kg/m3 N/mm2 N/mm2 rad/s

Strain Shear strain Shear strain speed Strain rate Critical strain Tool included angle Effective cutting speed angle Temperature Tool cutting edge angle Working cutting edge angle Tool cutting edge inclination Working cutting edge inclination angle Coefficient of friction Phase Shear angle Slide angle Feed motion angle Density Stress Normal stress Angular speed

Chapter 1

Introduction

From about 12,000 to 50,000 years ago, human beings could already adapt stone tools with intentionally produced edges to specific machining tasks by varying the geometry of the cutting edge, as is shown in early tool findings from the Palaeolithic Age (Fig. 1.1). The discovery of how to extract metals such as copper, tin and iron was a huge milestone in the history of material and manufacturing technology. Since about 700 B.C., tools were almost exclusively made of iron. At the beginning of the 17th century, constant improvements in iron smelting led to the preferred use of iron and steel as construction materials instead of other metals known then and in place of wood, which had predominated until then. Inspired by the growth of the textile industry and the discovery of the steam engine, there was increased exploration into manufacturing technology at the beginning of the 19th century, leading by the second half of the 19th century to the first systematic investigations into cutting methods and initiating a completely new research area. At the end of the 19th century further research led to the discovery of new cutting tool materials and, at the turn of the century, to the development of high speed steel by F.W. TAYLOR, a significant contribution to the history of manufacturing technology [Tayl07]. In light of these developments, SCHLESINGER said: “The dividends may lie on the cutting edge of steel, but the speed of these cutting edges is a function of the machine moving them, so as wages increase, the cutting machine is a trump card.” [Schl11]. Subsequent research followed in this direction, leading to the development of cemented carbide in 1923 by SCHRÖTER [Schr23, DRP25] at the SOCIETY FOR ELECTRIC LIGHTING as well as its use in machining by FRIEDR. K RUPP AG as WIDIA cemented carbide. Following this was the invention of oxide-ceramic cutting tool materials and their application in cutting after 1938 by OSENBERG as well as WENTORF’s development and synthesis of the superhard cutting tool material cBN (cubic-crystalline boron nitride) in 1956 [Went57]. The refinement of cutting tools with wear-resistant cemented carbide coats starting in 1968 by the companies Sandvik Coromant and Friedr. Krupp AG was a major contribution to improving productivity and economy [Sche88]. In order to fulfill the constantly increasing requirements made on workpiece quality and to make machining processes more economical, all the influencing F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_1,

1

2

1

Introduction

Drill Saw

End scraper Gravettian point

Double tines

Round scraper

Side scraper

Notched point

Stemmed-base hafted scraper

Fig. 1.1 Flint tools, acc. to KERND’L [Kern72]

parameters relevant for productivity and workpiece quality must be considered: cutting edge geometry, process kinematics, cutting parameters, materials, cutting tool materials and their coatings, wear behaviour as well as auxiliary materials. In this pursuit, the following quotation by SCHLESINGER can act as a guideline: “The machine tool must be designed with a view to the place it is used”. As an introduction to the subject matter, the compendium will be preceded by a general chapter providing a survey of the most important attributes of workpieces and their test methods.

Chapter 2

Metrology and Workpiece Quality

2.1 Manufacturing Disturbances and Manufacturing History The basic task of manufacturing is to provide workpieces with specified quality characteristics in the required quantity in the most time- and cost-efficient way possible. Every manufacturing process is affected by variable disturbances, which can be both external and internal (occurring within the process itself). For this reason, the functionally determinative properties of the components are provided with tolerances. If a characteristic value lies outside of the permissible tolerance, it is defective. Thus important functional characteristics must be tested either already during manufacturing or at the end of manufacturing. Important disturbance factors which must be taken into consideration as possible causes of defects are: disturbances caused by static forces, such as deflections effected by the workpiece weight or clamping errors; disturbances caused by dynamic forces which lead to either self-starting or forced oscillations; disturbances caused by thermal influences, such as process heat or internal sources of heat in the machine tool; and disturbances caused by tool wear (Fig. 2.1). This group also includes disturbances resulting from the engagement kinematics between the tool and the workpiece, such as generated cut deviations occurring during hobbing or diffraction effects on optical surfaces caused by systematically applied tool marks. Functionally determinative quality characteristics can be defined through macrogeometrical properties. Macro-geometrical properties include, for example, the accuracy of dimension, form and position of geometrical elements. Microgeometrical parameters include, for example, characteristic surface values, such as roughness, mean roughness value and polishing depth. However, the functional integrity of the process requires that certain properties are tolerated although they lie below the surface in the peripheral surface zone. These may include hardness values, residual stresses or certain structural properties. The fluctuations of characteristics occurring in every manufacturing process can basically be derived from the influence of either systematically or randomly occurring disturbances. Systematic errors are caused by the system. They are reproducible under identical boundary conditions and change the characteristic in a single direction; they cause manufacturing trends. Systematically acting disturbances cause deviations in characteristics. If the causes of systematic errors are known, they can F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_2,

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4

2 Metrology and Workpiece Quality Statical disturbances

Thermal disturbances

• Clamping fault

• Internal heat sources

• Deflection

• Strain

•

• Process heat

•••

• Environmental effect •

•••

Prozess

Dynamic disturbances

Tool wear

• Self-starting oscillation

• Displacement of cutting edge

• Forced oscillation

• Flank wear

e.g. triggered over basement •

•••

• Flank roughness • Grinding wheel wear •

•••

Fig. 2.1 Variable disturbances in manufacturing

be corrected and compensated. Examples of systematic errors are geometric errors of machine guiding elements and the development of cutting edge offset through the continuous wear of turning tools used in turning processes. Changes of the cutting force caused by sudden cutting depth fluctuations and the workpiece deflections associated with them may also be of systematic nature (Fig. 2.2). Randomly occurring disturbances lead to changes to characteristics, the effects of which cannot be predicted in a strictly deterministic way. They influence the result in both a positive and a negative direction. Thereby, the work result becomes unstable. The influence of random disturbances on the work result can, however, be statistically recorded. Examples of this include fluctuations in the structure of a material, random temperature fluctuations or sudden self-starting oscillations (rattling oscillations). Within real processes, it is not always easy to clearly decide between systematic and random influences. An important parameter describing the effect of random disturbances on a quality characteristic (i.e. on the variation of a quality characteristic) is process capability or the process capability index (Fig. 2.3). These process capability indices can only be applied if the variation of the characteristics are normally distributed. In practice, it is frequently required that cp ≥ 1.33. The process capability index is always taken as the dominant process

Manufacturing Disturbances and Manufacturing History

5

Dimensional error Passive force FP / N / µm

Nominal diameter

2.1

Feed direction

1000 800 600 400 200 Cutting path 10 5 0

Workpiece material Cutting tool material Cutting speed Feed

Cutting path Ck55 N HW P10 vc = 160 m/min f = 0.25 mm

Cutting part geometry: κr = 70° γ0 = 6° εr = 90° α0 = 5° rε = 0.8 mm lS = 0°

Fig. 2.2 Influence of passive force on dimensional accuracy

design criterion when large numbers of pieces must be manufactured with the highest possible level of productivity. Variation in manufacturing batches occur, when it can not be guaranteed that all active elements work securely during the process being at its performance limits. If, for example, cutting edges are randomly failed during a cutting process because of cutting edge fractures or layer failures, this can have an immediate, significant on the process capability. Another example of critical process conditions is when the quantities of heat entering the workpiece at high process performances cause thermal strain to induce an increased variation of geometrical characteristics. These examples reveal that measuring and testing procedures in manufacturing are necessary both to guarantee the quality of the manufactured part and to start and develop processes. In general, several manufacturing steps are necessary to manufacture a part. Manufacturing progresses result from a manufacturing chain or a manufacturing sequence. With every manufacturing step, characteristic changes (geometry, surface, peripheral surface zone) are left behind on the part. Thus a manufacturing history develops (Fig. 2.4). The output properties of the part following a certain manufacturing step in turn become the input factors for the processing steps which follow. There can also be a significant interaction between the current step in the process and previous changes.

6

2 Metrology and Workpiece Quality

UTL 20

_

3s^

s

_

15

NV LTL

_

_

_

x

x - LTL

10

cp cpk

= 0,71 = 0,71

cp cpk

x

_

x

x

= 2,50 = 1,00

cp cpk

= 1,67 = 1,33

cp cpk

= 5,00 = 5,00

Process capability indices

cp

=

cpk

=

Tolerance band

=

Process spread

UTL - LTL 2 ( 3·σ^ )

Minimum proximity to process limit

=

=

Half process spread

_ x : Mean value of the random sample ^ : Estimation of the spread of the totality σ _ s : Mean deviation of the random sample

T _ 6s _ |x - NV|min 3 σ^

UTL: Upper tolerance limit LTL: Lower tolerance limit NV: Nominal Value

Fig. 2.3 Process capability indices

Lu Ll

Lu

Enveloping cut deviation

Anisotropism Ll

Forging, cold extrusion

k1a

Turning milling

c1a k2b

Machining center Lu,l Upper and lower limit

Lu Ll

Machining center k, c Damping and stiffness

Fig. 2.4 Manufacturing history (Source: Getrag Ford)

Ll

Hard machining

Hardening

c2b

Lu

Texture

knx

cnX

Grinding machine tool

Geometry

Assembling

2.1

Manufacturing Disturbances and Manufacturing History

7

Manufacturing process:

Primary shaping, machining

Manufacturing dispersion:

Affected by residual stress, chucking

0.01 0.008 Requested cpk = 1.67

Box part:

Automobile-gearbox made of Mg-alloy AZ 91hp Standard tolerance IT 4

Fig. 2.5 Representative case study – box part

The hole in a transmission part shown in Fig. 2.5 could not be manufactured in a procedurally secure way with the required process capability index and specified. There can also be a significant interaction between the current step in the process and previous changes. The hole in a transmission part shown in Fig. 2.5 could not Manufacturing process:

Forming, 5-axis-milling


Process instability, chatter marks, long and slander tools

Work piece: Blisk (Blade integrated Disk) made of TiAl6V4

Tool:

Depth of cut /mm

End milling cutter (D = 8 mm)

Source: MTU

Fig. 2.6 Stability bounds when milling a BLISK

stabil instabil stability bound

Speed / min–1

8


Manufacturing process:

Forming, turning, drilling, broaching


Process anomaly, tool life decrease of the part Source: Rolls-Royce

50 µm

Without anomaly Ktxσappl. / MPa

Smeared material

1000

Low thermal influencing High thermal influencing

10000

100000

1000000

Work piece:

Turbine disk made of Inconel718

Tool:

VHM-drill (D = 8 mm)

Numbers of change of load

Fig. 2.7 Surface damages when drilling with a worn drill

be manufactured in a procedurally secure way with the required process capability index and the demanded productivity. This is due to released residual stresses which arised in the previous die-casting process. Figure 2.6 shows a part manufactured by milling with long, slender end mills. The milling process is only stable for certain combinations for cutting speed and depth of cut. When these areas are exceeded, rattling oscillations occur. These characteristic stability areas must be known; otherwise the process cannot be carried out at the performance limit. However, not only impermissible macro-geometrically or micro-geometrically recognizable changes may limit the process, but also impermissible changes in the peripheral surface zone. Figure 2.7 shows a turbine disk which was exposed to high mechanical and thermal stresses caused by machining the drill holes. These stresses damaged the peripheral surface zone in a way that significantly diminished the fatigue strength of the part.

2.2 Measuring and Testing Quality characteristics must be tested during the manufacturing due to the boundary manufacturing conditions mentioned above. This can be done by means either of samples or of a 100% test. A 100% characteristic check is prescribed for many safety-related parts. Though, a 100% test does not assure freedom of errors.

2.2

Measuring and Testing

9

Testing means determining whether the test sample exhibits the required characteristics, such as dimension, form and surface quality. This determination can be made in different ways, thus the distinction between subjective testing and objective testing. Subjective testing is performed by means of the sensory perception of the tester, e.g. visual and tactile tests. Examples of this are the evaluation of burr formations on the edges of parts and of shadings on polished surfaces. In objective testing, the tester is supported by testing equipment. There are two further methods which fall under this category; measuring and gauging. Measurement is the comparison of a characteristic with a measurement standard. The result is a measured value. Gauging is the comparison of the test object with a dimension or a form. The result is information regarding whether the specified requirements were satisfied or not (go or not-go). Testing equipment is subdivided into measuring instruments, gauges and aids. Measuring instruments and gauges are distinguished by material measures corresponding to the measurand. Material measures can be of mechanical, electric, optical and electronic nature. Mechanical material measures may be, for example, distances between surfaces or angular positions of surfaces. Indicating measuring instruments have movable markers, divisions or counters. When using these devices, the measured value reading can be either direct analogue or digital. In contrast to this, the gauge body corresponds merely to the dimension or the dimension and form of the required characteristic. Intermediate values cannot be defined. Table 2.1 provides a survey of the attainable accuracy values when using different manufacturing processes. Table 2.1 Survey of manufacturing qualities Standard tolerance IT 1

2

3

4

5

6

7

8

9 10 11 12 13 14 15 16

Primary shaping Die forming (warm) Extrusion (cold) Deep drawing Turning Drilling Reaming Milling Grinding Erosion

2.2.1 Measurement Errors As a matter of principle, every measurement result is subject to error. Possible reasons for measurement errors are shown in the ISHIKAWA diagram in Fig. 2.8. The causes of measurement errors may lie in the measuring method, the measuring

10


Measuring method

Preparation

Unconsidered disturbances Aptitude Abbé principle error

Preliminary arrangements Cleanliness Protocol Calibration

Environment Temperature Air pressure Humidity Vibration Extraneous light Interference field Measurement error

Aim of the measurement Testing deviation

Measured object

Measuring forces Evaluation Aptitude Routine Assembling Visual acuity Measuring uncertainty Aptitude Attention Mistake Measuring device

Measuring person

Fig. 2.8 Causes for measuring errors by the measuring procedure

instrument, the measured object, the person measuring, the preparation involved and the environment. In order to achieve a high measurement accuracy, it is necessary to take the abovementioned influences into consideration and, where applicable, to minimize them, compensate them, or at least to estimate them in terms of their magnitude. It can principally be distinguished between systematic and random disturbances in measurement procedures, as well [DIN1319a].

2.2.2 Macro- and Microgeometry of Components 2.2.2.1 Structural Deviations With regards to components, the distinction is often made between macrogeometrical parameters and the surface quality. Macro-geometrical parameters refer to deviations of dimension, form and position. The surface quality is defined by roughness parameters. The transitions between these categories are not always clearly definable. DIN 4760 offers a general system for organizing structural deviations (Fig. 2.9). Therefore it is necessary to define the term “surfaces” at the beginning. The real surface (primary surface; see DIN EN ISO 4287) is the surface actually present on the part. The actual surface is the metrologically registered surface. It may differ from the real surface, since every measuring method can only approximate the real surface. A geometrically ideal surface is assumed in designs

2.2

Measuring and Testing

11

Structural deviations (in superelevated representation) 1st Order: form deviations 2nd Order: waviness 3rd Order: grooves (roughness) 4th Order: scores, scales (roughness) Not easily representable

5th Order: textural structure 6th Order: lattice structures of the material Superposition of structural deviations 1st up to 4th Order

Fig. 2.9 Structural deviations, acc. to DIN 4760

and forms the basis of tolerances. In Fig. 2.9 six orders of structural deviations are defined on the basis of these observations. Structural deviations of the 1st order (see also the following sections) are frequently the result of systematic errors. With regards to waviness, i.e. the structural deviations of the 2nd order, one cannot clearly define whether they are caused by systematic or random influences. The unbalance of a rotating tool and any periodical oscillations caused by it are forced, while sudden rattling oscillations are self-starting. In general, fundamentally different actions must be implemented in order to exclude any systematic or random causes of error. Structural deviations of the 3rd order also occur regularly. They are to be attributed to the penetration between tool and workpiece and are often determined by means of penetration calculations. Examples of these are kinematic roughness associated with turning, surface marks created in peripheral milling and generated cut deviations created in hobbing. In such cases, the structural deviations can be influenced in a targeted way by means of generation kinematics and tool design. The higher orders of structural deviation are primarily random in their occurrence. Examples of structural deviations of the 4th order include chip formation processes and removal processes. Roughness of the 5th order is rendered visible by structural properties on the surface. This can play a significant role in the high-precision machining of metallic optical mirrors. Thus in high-precision turning of multicrystalline metals, grain boundaries may become visible because the individual crystals exhibit varying orientations and therefore varying stiffnesses. In this case, anisotropism of the grains becomes visible on the surface.

12


In general, all the structural deviations on a real surface are superposed. Filters are employed to separate roughness and waviness in a measurement process [DIN EN ISO 4287]. The following sections will treat macrogeometrical deviations of structure, form and position, as well as the corresponding measurement technology. Section 2.4 and the following sections describe the metrological recording of higherorder structural deviations. 2.2.2.2 Form Deviations Form deviations refer to deviations from a specified ideal geometrical property, such as straightness, evenness, roundness or cylindric form [DIN EN ISO1101, VDI2601]. The following will introduce some examples of form errors and their causes (Fig. 2.10). • A cause for deviations from the cylindric form of a workpiece can be the incorrect alignment of the workpiece on the machine tool with respect to the tool. • Another cause for form deviations is the continuous alteration over time of the tool geometry caused by wear. • Thermal displacements of the workpieces, tools and machine tools can also lead to form deviations. • Deviations from cylinder form can also arise when the workpiece deflects because of a radial strain. This can happen, for example, e.g. when turning long, slender parts when the workpiece is not supported. • Roundness errors can develop through the incorrect clamping of the workpieces on the machine tool.

2.2.2.3 Position Deviations Position deviations are deviations of the position of a geometrical element with respect to a reference, such as an edge, circulation line or axis of the predetermined position. In general, the position of two surfaces or axes in relation to each other is indicated by simple length or angle specifications.

Fig. 2.10 Deviations from circularity and cylinder form

2.3

Length Testing Devices

13

Fig. 2.11 Deviations from concentricity, straightness and parallelism

Position deviations as shown in Fig. 2.11 may arise, for example, through incorrect clamping or defective clamping devices.

2.3 Length Testing Devices 2.3.1 Material Measures A material measure in length measurement technology represents lengths or angles by means of fixed distances or angles between surfaces or lines. Setting standards are material measures. The material measures used most frequently are gauges. The most important and most precise material measures in length measurement technology are parallel gauge blocks. They consist of two plane-parallel measuring surfaces made of hardened steel or cemented carbide. Parallel gauge blocks are used in prismatic and cylindric forms. Gauge blocks are composed of grades. A standard set contains 5 grades each containing 9 blocks, from which almost any dimensions can be assembled (see Fig. 2.12). The measuring surfaces are superfinished and even enough to allow, after careful cleaning of the measuring surfaces, individual gauge blocks to be joined together so that they adhere.

Fig. 2.12 Gauge blocks (Source: Mahr)

2.3.2 Gauges A gauge measures dimensions or forms generally according to limit dimensions [DIN2257]. It can be distinguished between inspection gauges, limit gauges and form gauges (Fig. 2.13).

Go


Gu

14

21,5

Fig. 2.13 Types of gauges: inspection gauges, form gauges, limit gauges Not go side

Not go side

Not go side

Not go side

Go side Go side

Go side

Go side

Drill hole

Shaft

Square profile

Multitooth profile

Fig. 2.14 Different types of limit gauges

Inspection gauges are parts of a gauge set for which the measure is assembled from the combination of gauges. Examples are the combination of parallel gauge blocks and the setting of feeler gauges. Form gauges allow for the testing of profiles using the light-slit method. They include angles, radius gauges and thread gauges. Limit gauges function according to Taylor’s principle: The go gauge must be designed in such a way that the dimension and form of the workpiece can be inspected when combined with the gauge. Not-go gauges are only used to inspect single dimensions. Limit gauges used most frequently are calliper gauges, limit plug gauges, gauging rings and thread gauges (Fig. 2.14). Gauges are the simplest testing equipment used in industrial manufacturing.

2.3.3 Indicating Measuring Instruments Indicating measuring instruments consist of a material measure and a display mechanism. The indicating range is the range of measured values that can be displayed on a measuring instrument. The measuring range is the part of the indicating range for which display deviations lie within specified or agreed limits. The suppression range of a measuring instrument is the range which must be exceeded in order that

2.3


15

the instrument can begin to display values. The hysteresis of a measuring instrument equals the difference of the display values determined by measuring from different directions. The hysteresis of a measuring instrument is often not constant, which means that only a certain lower limit is indicated. The reversal error is principally a systematic error. It can be compensated. The sensitivity of a measuring instrument is the ratio of an observable change on the measuring instrument display to the causative parameter. Using length measuring instruments, the sensitivity is defined as the ratio of the range of the indicating element (e.g. the indicator) and the range of the measuring element (e.g. the spindle or measuring arm). Conversion between the measurement range and the display may be based on different physical principles. The most frequently used conversion elements are mechanical, electrical, pneumatic, optical and electronic conversion elements. 2.3.3.1 Instruments with Mechanical Converters Instruments with mechanical converters utilize mechanical or optical indicating elements. In length measurement technology, the most often used measuring instruments function on the basis of mechanical converters, e.g. threads, racks, gearwheels and blade segments (Fig. 2.15). With mechanical callipers with sliding glideways as the conversion element, the material measure is mounted on the bar on which the display, i.e. the slider, is moved. Mechanical callipers have two different scales which can be read off against each other by the observer. The scale on the slider is called a vernier (Fig. 2.15). In the case of micrometers, highly precise threads are used as converters and as a material measure. The system comprises a measuring spindle running in a sleeve, whose forward face is designed as the measuring surface and whose back end carries a ground thread as material measure. Typical thread pitches are 0.5 and 1 mm. As seen with callipers, two different scales are read off against each other. One scale is mounted on the sleeve, the other on the scale drum. In order to avoid influencing

Fixed measuring jaw Measuring ambos Moveable measuring jaw Vernier

Scale sleeve Thimble

Clamp Bar Linear scale

Fig. 2.15 Setup of callipers and micrometers

Measuring spindle Coupler

16


Prestressed spring

Rack Arm with blade segment

Pinion Plug gauge

Fig. 2.16 Setup of dial gauges and dial comparator

the measurements through large fluctuations in the measuring force, the latter is limited to 5–10 N by means of a coupler (ratchet coupling). Micrometers are designed both as outside micrometers for external dimensions and as inside micrometers for internal dimensions (Fig. 2.15). Dial gauges are length measuring instruments equipped with racks and gearwheels as conversion elements which allow a larger view of the path of the plug gauge. The material measure lies in the gear mechanism, the conversion element which leads to measurement reversal errors due to anisotropic friction and potential tolerance. Needle positions greater than 360◦ are possible (Fig. 2.16). Dial comparators are the most precise mechanical length measuring instruments. They are equipped, unlike dial gauges, with a gear mechanism as lever arm system, gear segment and pinion as conversion elements. Thereby, the movement of the plug gauge is transmitted to the needle. The design only allows needle positions smaller than 360◦ (Fig. 2.16). In principle, all the instruments with mechanical conversion elements described here can be equipped with frictionless optical indicator elements in order to improve precision. Given constant conversion behaviour, the more precise indications the measuring instrument should display, the longer the indicator must be. However, since there is only a limited amount of space in the instruments and, above all, since the inertia of the indicator can lead to distorted measurement results when the measurement range is small, optical indicator elements are used. The principle is basically this: a beam of light is directed to the material measure and, according to the position of the material measure, this beam is reflected in a different direction. The optical indicator then displays the measurement range of the material measure without friction and inertia. Further potential for increasing precision can be exploited by means of a fine graduation of the scale which can no longer be observed with the naked eye, thus necessitating optical magnifying aids in the form of microscopes. By means of the microscope, a mechanism consisting of an objective and an ocular, a magnification of a real image by the objective can be viewed with the ocular. In this way, a remagnified virtual image of the object is created. The total magnification through

2.3


17

a microscope is the product of the objective and ocular magnifications. The scale graduation can be applied in the ocular directly or by means of an ocular plate. Microscopes including measuring oculars are called measuring microscopes.

2.3.3.2 Instruments with Electrical Converters These instruments are used to determine changes in the measurand caused by alterations of electrical properties, such as resistance, inductance and capacity. The turns of a coil, the distance between plates of a condensator and the resistance of an electrical conductor can all serve as material measure. Inductance, for example, is altered by changing the immersion depth of an iron core; the capacity of a plate capacitor is a function of the plate distance (Fig. 2.17). Position encoder Sensor

Output to measuring amplifier

Position encoder

L1

R4

L2

R3

s

UA

UB UA UB C s

= Bridge-output voltage = Bridge-supply voltage = Capacity = Measurement range ε o = Electric constant ε r = Dielectric constant A = Surface area d = Distance

Fig. 2.17 Capacitive and inductive measuring principles

C=

ε o˙ εr˙ A d

UA =C ˙ s UB

18


Inductive position encoder

Fig. 2.18 Inductive working sensor in real execution

An alteration of the ohmic resistance can be effected by changing the conductor length and/or by a cross-sectional variation [Hoff04, Grot05]. Figure 2.18 shows a real execution for multi-point measuring with inductive working sensors. 2.3.3.3 Instruments with Optical Converters These length measuring instruments exploit the wave property of light. The wavelength of light is used as the material measure. Figure 2.19 shows the functional principle of a laser interferometer.

Reference path

Wedge prism of the reference beam

Measuring path

Displacement

Laser Mirror

Polarizing beam splitter

Interference of measuring und reference beam Detector

Fig. 2.19 Longimetry with laser interferometer

Wedge prism of the measuring beam


λ1 λ3

λ2

ϕ3

Time

Ampl.

ϕ2 ϕ3

19

Ampl.

Ampl.

2.3

λ Time

Time

ϕ2 l

Light mixture of different wavelengths

Monochromatic and coherent light

Monochromatic light

λ ≠ const. ϕ ≠ 0

λ = const. ϕ = 0

λ = const. ϕ ≠ 0

Ampl.

Ampl.

λ

Time

Time

Interference of two coherent light waves with the same phase length (Addition of illuminance)

Interference of two coherent light waves out of phase (destructive interference)

λ: Wavelength ϕ : Phase difference

Fig. 2.20 Optical measuring principle, interference

A light source emits coherent, monochromatic light whose beam is split into a reference beam and a measuring beam at a polarizing beam splitter. These beams are reflected on a respective wedge prism and combined again through superposition. Finally, the combined beam is detected and evaluated. If the wedge prism of the measuring beam is shifted, interferences occur because of the difference of the optical paths of the two partial beams after combining them. These are fluctuations in light intensity via eliminating and amplifying light. Integral multiples of the wavelengths lead to an amplification, phase differences of 180 degrees to an elimination of light (Fig. 2.20). The number of interference lines is directly related to the path being measured. Since the wavelength of the light is a function of temperature and also air pressure and humidity, these influences are compensated within the device. The advantages of such systems lie in their large measuring range and especially in their contactless measurements. 2.3.3.4 Instruments with Pneumatic Converters These instruments use pressure and flow rates as material measures (Figs. 2.21 and 2.22). This measuring principle can be used for both tactile and contactless measurement (Fig. 2.22). There are two types of such measuring instruments: high-pressure and lowpressure instruments. High-pressure instruments work with an operating pressure greater than 0.5 bar, low-pressure instruments with an operating pressure lower than

20


Volume measuring method 1. Air filter 2. Pressure control 3. Flowmeter 4. Measuring nozzle

Velocity measuring method 1. Air filter 2. Pressure control 3. Venturi-nozzle 4. Flow of valve 5. Measuring nozzle 6. Differential pressure manometer

1

3

2

4 s

1

2 3 4

h 6

5 s

Fig. 2.21 Measuring principle of pneumatic distance measurement

Air inlet Valve Guide body

Measuring nozzles

Follower

Drill hole

Work piece

Tactile measurement

Contactless measurement

Fig. 2.22 Pneumatic sensor

0.1 bar. The pressure range between 0.1 and 0.5 bar is outside the operating range of the instrument [DIN2271a]. Figure 2.23 shows the functional principle of contactless pneumatic measurement. The left side shows an external measurement, the right side an internal measurement. If the measuring rod is designed in a way that both measuring nozzles

2.4

Surface Inspection

21

ps ∅dm pa

Annulus

Measuring nozzle pa

Test sample

h Work piece surface

Measuring nozzle

Fig. 2.23 Functional principles and examples for pneumatic measurement

work within the linear range of the characteristic diagram, then an exact centring in the hole is unnecessary, since the air currents of the two nozzles are added and the sum represent the measure for the total gap width. Since the type of hole measurement described is a two-point measurement, deviations from hole roundness and diameter can be determined by turning the rod or the workpiece. The measuring rods are often specially adjusted to the measurement of a specific workpiece. Therefore, for economical reasons, pneumatic measurement is used predominately in serial production, especially if a 100% test is required. The advantages of contactless pneumatic measurement lie in the self-cleansing effect of the measuring device (i.e. the workpieces do not generally have to be cleansed of oil, dirt or micro-chips) and in the quickness of the measuring process. 2.3.3.5 Electronic Measuring Instruments These devices work with photoelectric sensors, by which light-dark fields of a gauge are converted into electric signals. The material measure is realised through the light-dark fields. Impulse gauges and code gauges are preferably used as path measurement systems. To be precise, it is a mechanical guidance system whose positions are detected and indicated optoelectronically. This path measurement system is also executed with callipers, dial gauges and dial comparators. It is also realized in coordinate measuring devices and machine tools. To control measuring axes and to evaluate data, an exact control system and effective software are required which can influence the quality of the measuring result and the applicability of the data [Pfei01]. The measurement of freeform surfaces and tooth-flank topographies on spur gears and bevel gears with coordinate measuring devices is currently state of the art (Fig. 2.24).

2.4 Surface Inspection The task of characterizing a technical surface consists in attaining a complete topological record of its three-dimensional geometry and describing it. Often, in order to simplify the measuring process, only parameters for surface roughness along a single measured length are registered (one-dimensional parameters). This

22


Fig. 2.24 3-D coordinate measuring machines

procedure is frequently followed in practice and is sufficient in many cases. For certain applications, e.g. in optics or for characterizing grinding wheel surfaces, more comprehensive, multi-dimensional surface descriptions must be used. However, the following will exclusively treat one-dimensional parameters.

2.4.1 Surface Parameters In most of the procedures used in industrial surface inspection technology, only structural deviations of the second or higher orders are analyzed and measured on surface sections. These sections must be statistically representative for the entire surface [DIN2257]. The surface can be detected by means of surface sections or the bearing surface (Fig. 2.25) [DIN4760]. Profile sections are sections which are oriented to the surface normally, tangentially or at other angles. First, a profile section will be used to illustrate some of the basic terms of surface inspection technology (Fig. 2.26). With respect to a profile, the distinction is made between the test length Lt of the surface section being metrologically detected and Profile

Tangential cross section

Profile

α

Vertical cross sections

Diagonal cross section

Fig. 2.25 Registration of structural deviations by surface cuts

Cross section parallel to the tangential cross section

Surface Inspection L1

L2

L3

L4

L5

L6

Zp1

Z (x)

23

Zp6

2.4

Ra

Zv2

Zv1

Zt1

x

Measured length L

Fig. 2.26 Fundamental terms of surface inspection technology

the measured length L, which is used for evaluation (L < Lt ) [Grot05]. The detected profile (actual profile) of a surface depends on the measuring procedure and the filter used, thus it only represents an approximate image of the actual surface. The reference profile shifted within the measured length perpendicularly to the geometrically ideal profile is defined as the middle profile. This is oriented in such a way that the surface areas above and below the middle profile line are equally large (Fig. 2.26). By using different measuring procedures and depending on which filter is used, different profiles can be determined [DIN EN ISO 4287]. These are the P-Profile (primary profile), the R-Profile (roughness profile) and the W-Profile (waviness profile) (Fig. 2.27). Manufacturing technology: Short-stroke honing Average roughness: Ra = 0.75

Scanning system: HAT 25/6 Scanning path: 10 mm

P-Profile (waviness and roughness)

µm 1 1µm

Without wave filter

250 µm

250 µm

W-Profil (Welligkeit) W-Profile (waviness)

µm 11µm

Low-pass filter 250 µm

High-pass filter

µm 1 1µm

250 µm

0,75 mm Grenzwellenlänge: Cutoff wavelength: λλ== 0,75 mm (Rauheit) R-Profil(roughness) R-Profile

250 µm

250 µm

Fig. 2.27 Separation of waviness and roughness by wave filter

24


The examples given below refer to the R-Profile (roughness) (Fig. 2.26). According to DIN EN ISO 4287, the following roughness parameters can be distinguished: • The height of the highest profile point Rp : value of the y-coordinate Z(x) of the highest profile point of the middle profile line within the sampling length Li . • The depth of the deepest profile valley Rv : value of the y-coordinate Z(x) of the deepest point of the profile of the middle profile line within the sampling length Li . • The total height of the profile Rt : the sum of the highest profile point and the depth of the deepest profile valley within the measured length L. • The greatest height of the profile Rz : the sum of the height of the highest profile point Rp and the depth of the deepest profile valley Rv within a sampling length Li . • The mean roughness value Ra : the arithmetic mean of the values of the y-coordinates Z(x) within a sampling length Li .

1 Ra = L

L |Z(x)| dx

(2.1)

0

Horizontal parameters (distance parameters) are also designated as bearing lengths. They are determined by means of tangential sections. Forming a ratio of of the summed single bearing lengths and dividing by the measured length yields the relative bearing length (material ratio) in a specified section depth c. By creating sections at different depths c, the bearing ratio curve – also called the ABBOTT-FIRESTONE curve – can be determined (Fig. 2.28). The bearing ratio of the roughness profile Rm r(c) is calculated as follows:

Xi L

Rm r(c) =

(2.2)

c

Z(x) x1

x2

x3

x4 x5

x6

x7

x8

x9

x10

x

Measuring length L

Fig. 2.28 Bearing ratio curve of the profile, acc. to DIN EN ISO 4287

0

20 60 100 Rmr (c ) / %

2.4

Surface Inspection

25

The bearing ratio can be calculated on the basis of the primary profile (Pm ), the roughness profile (Rm ) and the waviness profile (Wm ). Using the average roughness, Table 2.2 provides an overview of the surface roughness values which can be achieved with different manufacturing processes. With respect to general specifications on achievable surface values, one must consider the fact that it is not necessarily possible to deduce the manufacturing process when using one-dimensional parameters for describing the surface [Abou76]. Due to the characteristic engagement conditions between the workpiece and the tool one would also have to specify, for each surface parameter, which manufacturing process is to be used to create that parameter. This problem is somewhat alleviated if multiple one-dimensional surface parameters are used instead of just one when describing the manufactured surface or if surface reference standards are available (Figs. 2.29 and 2.30).

Table 2.2 Achievable average roughness

Primary shaping Die forming Extrusion Turning Drilling Reaming Milling Grinding Erosion

Rough and finish turned

Honed

Planed

Grinded

Fig. 2.29 Surface quality for several manufacturing operations

630

1000

400

300

250

25

160

16

10

4

2.5

1

0.25

0.4

0.1

0.04

Achievable average roughness Rz / µm

Face milling

Cylindrical turning

2.5

0.4 3.2 1.6 0.8 Roughness Rt (µm) 12.5 6.3

Horizontal milling

Arithm. average 50 roughness Ra (µm)

32

16

8.0 4.0

N6 N5 ISO R 1302–1971

N3 N2

N8

N4 Reaming

Flat grinding

N7

N5 N6

N10 N9

N7 Flat lapping

Roughness markl Surface mark DIN 140

0.05 0.4 0.2 0.1

Roughness Rt (µm) 10 6,0 Arithm. average 1.6 0.8 roughness Ra (µm)

0.55


3.0 1.6 1.0

26

Fig. 2.30 Surface reference standard (Source: Rupert)

2.4.2 Subjective Surface Inspection In some cases, a visual inspection can determine whether the processed surface meets the specified requirements. A testing device frequently used in the workshop is the straight edge, which is used to check the evenness of a surface using the light-slit method. In workshop operations, spotting straight edges or spotting plates are used to detect form deviations or contact patterns on the workpiece surface (Fig. 2.31). Impression methods using lacquers have been developed for surfaces which are difficult to access (e.g. drill holes, internal surfaces). The lacquer is applied onto the

Spotting medium

Black areas Light areas (spotting medium) (lowest points) Work piece Spotting plate Spotting straight edge

Fig. 2.31 Surface inspection by spotting

Grey gleaming areas (bearing points)

2.4

Surface Inspection

27

test surface and pulled off after drying. The contact patterns attained on the lacquer film thus attained can then be evaluated. A frequent use of spotting is detecting contact patterns in quality testing and in the assembly of gear teeth (especially bevel gear teeth). Applying spotting paste and then rolling the teeth under light strain causes the bearing points on the tooth flanks to become visible. These bearing points can then be compared to the reference contact pattern. Nowadays in individual cases machine-tool guiding elements also become scraped, therefore, spotting is used to judge the surface.

2.4.3 Surface Measurement All the physical principles mentioned in the context of length measurement technology are also fundamentally applicable to surface measurements. Since mechanical and optical measuring methods have proven to be the most effective in this field, the following discussion will limit itself to these methods. 2.4.3.1 Mechanical Measuring Methods Devices for surface measurement which function on the basis of a mechanical functional principle are generally referred to as stylus instruments. Mechanical measuring methods are classified as either scanning methods or sensing methods. In the scanning method, a contact stylus descends upon the surface to be tested at a specified frequency. The surface is guided under the needle with a constant feed rate. The path of the contact stylus can be visualized mechanically, optically, electrically or electronically. The contact stylus can either be raised to a fixed level (the WOXEN principle) or raised by a fixed amount from the respective point of impact to the surface (differential tactile procedure) (Fig. 2.32). In the differential tactile procedure, the impact energy of the contact stylus is lower in comparison to the WOXEN principle and exhibits marginal dispersion. As a result, the penetration depth of the needle remains constant, the measuring accuracy being thus higher than in the case of the W OXEN principle. In the case of instruments functioning according to the sensing method, the contact stylus is guided continuously over the surface. The needle rises and falls in line with the profile pattern Scanning method

Woxen principle

Differential tactile procedure

Fig. 2.32 Mechanical measuring methods

Sensing method

Differential tactile procedure

28

2 Metrology and Workpiece Quality Oscillating scanning system

Plane of reference system Caliper slide

Semi-rigid System Scanning arm

Caliper slide

Caliper

Caliper

Table Skid Caliper

Skid

Fig. 2.33 Different surface scanning systems

(Fig. 2.32). The lifting motion is indicated relative to a reference point defined in the device or to a reference level. Here too, mechanical, electric, optical and electronic converters are used. Sensing profile methods are the most widespread in practice. Independently of the design type of the contact stylus instruments used, three system designs are distinguished (Fig. 2.33). Plane of Reference System In this system, the scanning unit is guided on a reference surface (plane, cylinder) which corresponds to the ideally geometrical surface of the test sample and is oriented along the surface to be measured (Fig. 2.33). Aside from errors arising due to the calliper geometry, this scanning system provides a faithful transmission of the roughness and waviness values of the test sample. When measuring small or very large surfaces, however, the handling of this scanning system can frequently become unwieldy. An alternative is the reference surface contact system. Here, the workpiece is conveyed on very precisely guided slides in a horizontal direction beneath the firmly anchored scanning system. Besides roughness, the macrostructure of a surface can also be detected in certain areas. Semi-Rigid System In semi-rigid systems, the scanning unit, which contains the contact stylus, is guided on the skid gliding on the surface to be measured (Fig. 2.33). This system has the advantage that it requires little space and is thus suitable for measuring small or hard-to-access surfaces. A disadvantage is that the system requires an orientation to the surface to be measured which can cause parts of the profile to be transmitted in a distorted manner. Oscillating Scanning System Two glide skids guide the scanning unit of the oscillating scanning system (Fig. 2.33). The latter is oriented towards the surface to be measured and is thus

2.4

Surface Inspection

29

comfortable to operate. However, it requires more space than the single skid system and can therefore not be used for small or hard-to-access surfaces. The distortions caused by long-wave profile sections are smaller than those in half-rigid systems, because the skids are flat and farther apart. Nevertheless the waviness must be filtered out in many cases [Henz68]. Usually, the magnification gauges for x- and y-coordinates are selected in a highly varying way in profile records. This is necessary because the measuring lengths (abscissa) lie in the mm region and roughness parameters are represented on the ordinate which lie in the μm region. The optical impression of the profile record is thus strongly distorted in comparison to reality. The contact styluses used in contact stylus instruments often have an apex radius of 2 μm. They function with a bearing strength of 0.5 N, which can result in considerable surface pressures (up to 6000 N/mm2 ), which possibly cause alteration of the test surface. On the other hand, excessively large contact stylus radii distort the result; they act like mechanical filters. Thus the optimal conditions must be determined on the basis of the material of the test sample and documented in their entirety in the measurement report. 2.4.3.2 Optical Measuring Methods White-Light Interferometer This technology differs from length measuring technology both in that it employs white light, i.e. light with the entire wavelength spectrum, and in that it does not just use one beam, as with laser interferometers, but rather an entire bundle. A reflectedlight microscope is used to display an image of a section of the test object on a detector (e.g. CCD camera). By using different interference lenses, a beam splitter can be used to superimpose a highly accurate reference surface with the image of the test object on the same scale. The interferences which are then created can be detected and evaluated (Fig. 2.34). The topography of the test object creates a spatial modulation of the light intensity in the interference image. Depending on the surface to be measured, two different measuring modes are used: one to characterize very flat surfaces with an average roughness value Ra < 1 nm and one for all other surfaces. Steps and roughnesses of up to several millimetres high can be displayed using this technology. Fringe Projection The fringe projection method functions according to the triangulation procedure, in which equidistant stripe patterns are observed and evaluated at a certain angle, i.e. the triangulation angle. The stripe patterns on the test object are detected and evaluated from a certain position, with the projected stripes following the arbitrary form of the test surface. These stripes appear from the observer’s standpoint to be “interferences”, though they only appear this way because the location of the projection


x Displacement

Reference mirror

Test object

Displacement

30

Light

Lenses

z

Interference of measuring and reference beam Detector

Fig. 2.34 Principle of a white-light interferometer (two-beam interferometer)

surface points varies in relation to that of the projector. The stripe patterns then are evaluated interferometrically on the basis of fluctuations in light intensity. The most important geometrical values on which this method is based are: • the real distance of the projected stripes • the stripe distance registered from the position of observation • the triangulation angle Fringe projection is used, for example, to judge surfaces bent over a large area, such as those found in deep-drawing tools and on deep-drawn parts. The advantages of this method are a high measurement speed and a spatial or laminar scanning. A disadvantage is that the test objects may not be transparent or reflective.

2.5 Inspection of the Workpiece Rim The functional behaviour and applicability of a component depend not only on its macrogeometry and surface roughness, but also from the physical properties of the material both on the interior and near the surface. While the inspection of material properties properly belongs to the field of materials science and materials testing and thus cannot be treated here in further detail, the following will discuss some of the properties of technical surfaces and of layers near the surface (i.e. the rim zone) and methods used to measure rim zone properties. Technical surfaces can be categorized according to types of load into the following groups:

2.5

Inspection of the Workpiece Rim

31

• external surfaces of technical products of all types, i.e. visible surfaces, covering surfaces, indicating surfaces, etc. They are generally mechanically unstrained, but are exposed to climatic or environmental stresses. • surfaces subjected to heat, radiation or electrical currents, such as insulating surfaces, electrical contacts, or the like. Such surfaces are referred to as thermally stressed, radiation stressed or electrically stressed surfaces. • surfaces which come into contact with fluids or gases. On the one hand, a corrosive stress may be predominant, or a flow stress with cavitation and erosion processes may occur in the case of flowing media. Also, surface boundary currents may be influenced via microstructures (ribblets). • surfaces in mechanical contact with moved counter bodies. This strain is referred to as tribological stress. These stresses are found, for example, in typical machine elements, such as bearings, couplings, brakes, gear-teeth, etc. With this type of stress, different kinds of wear may occur. • optical surfaces used to form and conduct electromagnetic waves. Optical surfaces are produced with mirrors and transparent components. The basic beam-conducting and forming principles are reflection, refraction and diffraction. • biologically stressed surfaces exposed to the effect of microorganisms.

2.5.1 Surface Layers The properties of technical surfaces relevant to component behaviour are determined through the entirety of the physical and chemical properties of the surface layer. These properties include textural structure, hardness, strength and residual stresses in the rim zone near the surface. The surface rim zones of technical bodies are created through machining processes. A part of the energy used to create the surface always flows into the workpiece and is either stored or causes change processes in the base material. Thus every machining process also causes a change of the surface rim zone vis-à-vis the base material. Whether these changes affect the functionality of the workpiece must be tested and confirmed in individual cases. The distinction is frequently made between the external and the internal boundary layer. Figure 2.35 shows the structure of both boundary layers. The layers referred to here cannot be clearly defined; there are no fixed boundaries between them. For further reference, please refer to the following sources: [Schm36, Schl51, Czic03]. 2.5.1.1 The External Boundary Layer The external boundary layer is located between the surrounding atmosphere and the atoms of the base material embedded in the crystal lattice. It is generated via the reaction between the material and the atmosphere during and after machining. The external boundary layer encompasses the following individual layers:

32

2 Metrology and Workpiece Quality External boundary layer Grease-/ oil film

10–100 µm

Adsorption layer

0.3–0.5 nm

Oxide film

1–10 nm

Internal boundary layer > 5 μm Plastic deformed and hardened Elastic und plastic deformed Elastic deformed

Matrix

Fig. 2.35 Layer structure of metallic surfaces, acc. to SCHMALTZ [Schm36]

Reaction Layers/Oxide Film On the surface of metals which are either freshly machined or created through breakage can form thin reaction layers/oxide films. These layers may only have a thickness of 10–20 molecular layers. They may, for example, influence wetting capability or adhesive behaviour. Adsorption Layer After a short duration of time of being exposed to the air, the oxide layer is covered with an adsorbed layer of water and gas. This adsorption layer becomes considerably important, for example, in the context of electrical resistance when metallic contacts are present. Grease/Oil Film Coolants in particular cause a grease and oil film to deposit itself after machining on the adsorption layer. This film ranges from 10 μm to several 100 μm in thickness. A grease and oil film can still be detected even after cleansing. These films can exert considerable influence interface formation in further galvanic, adhesion, or PVD/CVD processing. 2.5.1.2 The Internal Boundary Layer The internal boundary layer is the layer of the machined workpiece bordering on the base material and thus possesses practically the same chemical composition as the

2.5


33

Sharp-edged tool

Worn tool

Cutting tool material: HC Material: C15 vc = 200 m/min f = 0.1 mm 50 µm

50 µm

Fig. 2.36 Plastic deformation of the workpiece rim by drilling

base material. The physical structure and the expansion of this layer in the direction of the workpiece interior depend both on the material used and the manufacturing operation used. The transition from the affected rim structure to the unaffected base structure is continuous. Figure 2.36 shows a model of the plastic deformation of the structure of case-hardened steel C15 after a drilling operation. Larger mechanical and thermal stresses caused by worn tools lead to clear modifications of the textural structure. In steel material processing, for example, given high enough temperatures and high cooling rates, a hardening zone may develop in the outermost surface layer. In the layers below this, annealing processes are possible due to the lower temperatures and cooling rates in these layers (Fig. 2.37). As shown in Fig. 2.37, a hardening zone has developed which has an extremely negative effect on the flank bearing strength.

2.5.2 Inspection of the Surface Rim Zone Since, in addition to surface geometry, the physical properties of the areas near the surface play an essential role, these properties must be tested, too. Therefore methods familiar in the field of materials testing are applied: • • • • •

material analysis structure and texture investigation hardness testing fracture mechanical testing residual stress measurement

Structural changes caused by increased thermal stresses of the rim zones near the surface can be determined in structure and texture investigations which are standardly used in materials science. This is done by metallographic preparations. Crack detection in rim zones of workpieces can be either destructive or nondestructive. A crack is a locally limited detachment of the material structure of small width but often considerable length and depth. The crack can arise through internal

34

2 Metrology and Workpiece Quality Pitting at the tip of a tooth

RF

200 µm

Grinding burn with new hardening zone left flank

LF

200 µm

Vickers hardness / HV0,05

1000 900 Right flank

800

RF

700

LF

600 500

df dw

Left flank

Left flank

400 300

0

2 4 6 8 Distance from edge d / mm

10

Fig. 2.37 Grinding burn during gear tooth grinding

tensions or through external force effects. For further reference, please refer to the specified literature [Schu04]. 2.5.2.1 Residual Stresses Residual stresses are characteristic for the internal tension of a load-free workpiece. This tension results from inhomogeneous elastic and elastic-plastic deformations which are present without the influence of external forces and torques. For every component, there is a balance of all the internal forces and torques created by residual stress. Residual stresses are subdivided into macro-residual stresses (type 1) and micro-residual stresses (types 2 and 3). Residual stresses caused by mechanical and thermal influences is treated in the chapters devoted to the relevant manufacturing operations. The basic formation mechanisms of thermally and mechanically induced residual stress are shown in Fig. 2.38. Residual stresses become significant in the context of dynamic stresses. The attempt is therefore often made to use manufacturing processes in finishing which create compressive residual stresses in the surface rim zone in order to minimize the probability of crack formation and crack growth. Such operations include, for

Compressive residual stress (mechanical caused) Plastic deformation Yielding point

Compression(–)

sE–

Tensile residual stress (thermal caused)

High-temperature limit of elasticity

sE+

ain

Str

35

Tensile (+)


e

s lea

Re

Deformation path

Yielding point

Compression (–) Stress

Stress

Tensile(+)

2.5

Co oli n He

g

Temperature

ati

ng

High-temperature limit of elasticity Plastic deformation

Fig. 2.38 Formation of residual stress

example, finish and surface rolling as well as shot peening. In contrast, tensile residual stress is probable in electrical discharge machining, because of its thermal operating principle; in grinding and all manufacturing operations which use geometrically defined cutting edges, both mechanical and thermal load spectrums occur during processing, a prediction of the stress state is not easily possible. Methods of Analysis For all processes, the metrological definition of residual stresses occur indirectly through the measurement of strains, electromagnetic parameters or speed of sound. Below, the basic methods used in manufacturing technology will be explained. For further information on this topic, please refer to the specified sources [Peit92, Hauk87]. Mechanical Methods The mechanical methods of analysis include the borehole method and the toroidal core method. These are destructive methods which release residual stresses in the rim zone near the surface through the insertion of very small holes or ring grooves in this zone, which causes an altered stress state and thus strains, which can be measured with wire strain gauges, for example. The average residual surface tension can then be calculated from the strains in consideration of the material behaviour. By removing layers of the surface rim zone at different surface positions, distributions of residual stresses can be defined up to the depth corresponding to the drill’s diameter. Since drilling or milling processes only cause reproducible strain changes at a depth of several hundredths of millimetres, this method cannot be used to determine residual surface stress with strong rim zone gradients.

36


X-ray and Neutron Diffraction In research and practice methods are used which utilize the diffraction of monochromatic X- or neutron radiation on the crystalline lattice structure of the material phases (Fig. 2.39). The BRAGG condition applies here: n · λ = 2 · a · sin

(2.3)

Altered lattice distances a of the material phases can be calculated very accurately from the displacement of the diffraction lines, which in turn allow to draw conclusions about residual stresses. The chosen measuring direction defines the stress components to be detected. This allows the definition of the entire stress tensor by measuring in multiple directions. In order to calculate residual stresses from strain measurements, the elastic properties of the material phases are used. In the case of many materials, their anisotropic behaviour can even be taken into consideration when measuring at the respective lattice levels. Figure 2.40 shows the customary device for measuring residual stresses (goniometer).

λ

Source of radiation

Detector

ϑ

ϑ ϑ

a

Fig. 2.39 Principle of X-ray diffraction on crystalline lattice

Swivel table

Specimen

Rotating table Detector

Fig. 2.40 Goniometer

X-ray tube

2.5


37

Diffraction methods demand high requirements on measurement technology and on occupational safety (radiation protection). X-ray diffraction is used in practice to define residual stresses. In research, neutron sources can also be accessed at selected locations, with which it is possible to detect depth structures up to some millimeters without removing surface layers. The X-ray method has a depth of penetration which is limited to an extent depending on the absorption properties of the material and the wavelength used for the X-radiation. Given the usual measuring parameters, the average penetration depths vary between approximately 1 μm for tungsten carbide and 5 μm for steel; significantly greater depths of penetration are possible for light metals. Residual stress depth profiles with great rim depths can be produced through the electrochemical removal of layers of the surface rim zone. Magnetic Methods Magnetic analysis methods use the influence of mechanical stresses on resetting in ferromagnetic materials. These changes can be measured by means of an externally applied alternating magnetic field. The measurands are coercive field strength, superposition permeability, dynamic magnetostriction and magnetic and acoustic BARKHAUSEN noise. The simultaneous analysis of several magnetic parameters improves the evaluation of the quantitative measuring data. Magnetic methods are characterized by compact and easily manageable test set-ups. Depending on the material and the range of frequencies analyzed, surface rim zones can be gauged ranging from a few micrometers to over 1 mm. Disadvantegous is that these methods they are limited to ferromagnetic materials and have to be calibrated to known material states. Acoustic Methods Residual stresses are generated through lattice strains of the material which result in an altered propagation of structure-borne sound. This acoustic-elastic effect can be used to analyze stresses by means of ultrasonic measurements. Such methods are based on the assumption that the sound velocity is proportional to the lattice strain. The applicational potential of these methods are limited above all by the material state. This is because microstructural errors and textures exert as great an influence on sound velocity as structural gradients.

Chapter 3

Fundamentals of Cutting

DIN 8580 defines machining as all process variants of the third main group “Cutting”, in which form is altered by means of reducing material cohesion. Deformation is achieved by means of a relative motion between the tool and the workpiece that brings about a transfer of energy [DIN8580]. In the standard, this basic classification is further refined in order to categorise machining procedures [DIN8589]. Machining is defined as follows: cutting, in which layers of material are mechanically separated from a workpiece in the form of chips by means of a cutting tool. According to DIN 8580, machining comprises Groups 3.2 (machining with geometrically defined cutting edges) and 3.3 (machining with geometrically undefined cutting edges) in the manufacturing classification system. The first part of this compendium on manufacturing processes deals exclusively with process variants of Group 3.2. For this reason, the term “machining” will only be used in the sense of cutting with geometrically defined cutting edges. What all processes in the group of geometrically defined cutting edges have in common is that they use a tool, of which the cutting edge number, geometry and position to the workpiece are determined. Several concepts and terms are necessary to describe cutting part, which will be described in the following.

3.1 The Cutting Part – Concepts and Terms The concepts, designations and terms used to describe the geometry of the cutting part are set down in DIN 6581. The cutting part is the active part of the tool where the cutting wedges are located with the cutting edges. The idealized cutting wedge is made up of two faces: A rake face and a flank face, which cut in a line, cutting edge S. The angle between these two faces is designated as the wedge angle β. Figure 3.1 shows an idealized cutting wedge [DIN6581]. The rake face Aγ is the face of the cutting edge where the chip runs off. The flank face Aα is the face on the cutting wedge, which is turned towards the new workpiece surface (the cut surface). These terms make it clear that the cutting wedge F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_3,

39

40

3 Fundamentals of Cutting Trace of plane of rake face Aγ

Chip Tool Trace of plane of flank face Aα

β Cutting speed vc

η

ϕ Workpiece

Effective cutting speed ve

Feed velocity vf

Trace of plane of cutsurface

Fig. 3.1 Description of the idealized cutting wedge

(tool) should always be regarded in connection with the workpiece. This means that considerable importance should be attached to process kinematics. One model concept is often used for the sake of a simplified description of process kinematics, that of the selected cutting point. This model simplifies the actual kinematics by summarizing the spatial velocity fields in one point, the selected cutting point. At the selected cutting point, the velocity fields can be represented in a summarizing fashion by means of vectors. These vectors can be summarized in turn by vector addition in one total vector. Usually, the workpiece is assumed to be fixed; all motions are carried out by the tool. The resulting velocity vector is designated as the effective cutting speed ve . It can be divided into two components; the cutting velocity vc in the cutting direction and the feed velocity vf in the feed direction. To position the components of effective cutting speed clearly, two angles are defined: • The effective cutting speed angle η as the angle between the effective cutting direction and the direction of primary motion (see Fig. 3.1) • The feed motion angle ϕ as the angle between the feed direction and the direction of primary motion (see Fig. 3.1) Since there are no ideally sharp tools in practice, cutting edge rounding is taken into consideration (Fig. 3.2). In almost all cases the transition between flank and rake face is curved. This curvature is described by the cutting edge radius rß . Up to this point, the concepts have been explained using a simple cutting wedge formed by two faces. Generally, more complex tools are used, composed of several cutting wedges, in the simplest case of one major cutting wedge and one minor cutting face (Fig. 3.3).

3.1

The Cutting Part – Concepts and Terms

41 Trace of plane of rake face Aγ

Chip

Tool

Trace of plane of flank face Aα

β Cutting edge radius vc

rβ η ve

ϕ vf

Workpiece

Fig. 3.2 Cutting edge with cutting edge radius

Direction of primary motion Tool shank

Rake face Aγ

Direction of feed

Major cutting edge S Major flank face Aα Minor cutting edge S′

Cutting edge Minor flank face A'α

Fig. 3.3 Cutting edges and faces of the wedge, acc. to DIN 6581

Correspondingly, we speak of major and minor cutting edges. The major cutting edge S is always turned towards the cut surface, the minor cutting edge S towards the machined face [DIN6580]. If the selected cutting point is on the minor cutting edge, the concepts defined in the former are named accordingly and furnished with an apostrophe ( ). The flat areas of the major and minor cutting edges are joined by the corner radius r (Fig. 3.6).

42

3 Fundamentals of Cutting

3.2 Reference Systems In order to describe the location, position and direction of motion of a cutting wedge, reference systems are utilized in which characteristic planes are defined that are valid for all process variants. The two standardized reference systems are the tool-in-hand system and the toolin-use system (Fig. 3.4). The tool-in-hand system was developed for tool design as well as for the production and testing of cutting tools. In the tool-in-hand system, the tool angle is measured without considering process kinematics. In actual cutting processes however, the effective angle deviates from the nominal tool angle under certain circumstances due to process kinematics (exception: the wedge angle). For this reason, a tool-in-use system is also defined (Fig. 3.4, right). The tool-in-hand system is a system, the reference plane of which is oriented orthogonally to the assumed cutting direction. In contrast, the reference plane of the tool-in-use system is oriented orthogonally to the effective direction. In both systems, all planes contain the selected cutting point. The reference systems are in agreement when the cutting direction corresponds to the effective direction. First, the tool-in-hand system will be considered (Figs. 3.4, left, and 3.5): • The basic plane, upon which all other planes are based, is the tool reference plane Pr . It contains the rotation axis (if present) and lies perpendicularly to the assumed cutting direction.

Assumed direction of primary motion

Effective cutting direction

Pfe Pp

Pf

Direction of primary motion

ve

vc

Direction of feed

Assumed direction of feed

vf

Ppe

Pr

Fig. 3.4 Tool frame of reference (left) and tool-in-use system, acc. to DIN 6581

Pre

3.2

Reference Systems

43 Pn Po


Assumed direction of feed Ps

Pr

Fig. 3.5 Orientation in tool frame of reference Trace of tool back plane Pp Trace of tool cutting edge plane Ps Assumed direction of feed re κr

Trace of assumed working plane Pf

Tool

Tool reference plane Pr

Fig. 3.6 Position of tool cutting edge angle κ r

• The cutting edge plane Ps runs tangentially to the cutting edge S and perpendicularly to the tool reference plane Pr . • The tool orthogonal plane Po is perpendicular to the tool cutting edge plane Ps . • The assumed working plane Pf is perpendicular to the tool reference plane Pr and parallel to the assumed feed direction. • The tool back plane Pp stands perpendicularly on the tool reference plane Pr and perpendicularly on the assumed working plane Pf . • The tool cutting edge normal plane Pn is perpendicular to the cutting edge S. The tool cutting edge normal plane Pn is identical to the effective cutting edge normal plane Pne , since it is not oriented to the tool reference plane but rather to the major cutting edge. The tool-in-use system is rotated towards the tool-in-hand system by the effective cutting speed angle η. In tool-in-use system, the same signs are used as in tool-inhand system; they are followed however by an e, which stands for “effective”.

44


In order to designate all tool and effective angles clearly, the same index is assigned to them that signifies the plane in which these angles are measured. For example, the tool orthogonal wedge angle β o is measured in the tool orthogonal plane Po , or the effective side rake angle γ fe is measured in the working plane Pfe . Information about three angles around the three rotatory axes in space is required in order to determine the orientation of the cutting edges or the positions in space of the rake and flank faces. For an optimal cutting process, certain orientations are responsible for three rotatory axes in space. In the following, these angles will be explained exemplarily in the tool-in-hand system: • the orientation angle around a rotation axis that stands orthogonally on the tool reference plane Pr : The tool cutting edge angle κ r between the tool cutting edge plane Ps and the assumed working plane Pf , measured in the tool reference plane Pr , is defined as the angle that determines the position of the major cutting edge S in the tool reference plane Pr . It is measured mathematically positively on the basis of the assumed working plane Pf (Fig. 3.6). • the orientation angle around a rotation axis that stands orthogonally on the tool cutting edge plane Ps : The tool cutting edge inclination λs between the major cutting edge S and the tool reference plane Pr , measured in the tool cutting edge plane Ps , is defined as the angle that determines the position of the major cutting edge S in the tool cutting edge plane Ps (Fig. 3.7). • the orientation angle around a rotation axis that stands orthogonally on the tool cutting edge normal plane Pn : This rotation is executed by one of the tool angles α n or γ n , which are measured in the tool cutting edge normal plane Pn . The tool normal rake angle γ n is customarily used as the orientation angle, as it has a large effect on the cutting process (Fig. 3.8).

Trace of assumed working plane Pf

Major cutting edge S

λs Tool

Trace of Tool reference plane Pr


Tool-cutting edge plane Ps

Fig. 3.7 Position of cutting edge inclination λs

3.3

Basic Process Variants

45

Cutting line of planes Pn and Pr Trace of plane of rake face Ag γn Tool Trace of plane of flank face Aa βn


αn

Assumed direction of feed

Trace of tool cutting edge plane Ps

Tool cutting edge normal plane Pn

Fig. 3.8 Position of wedge in tool frame of reference

It is measured mathematically positively between the tool rake face Aγ and the tool reference plane Pr in the tool cutting edge normal plane Pn . The tool normal wedge angle β n is always fixed on the tool and cannot be changed by tool rotation. The tool normal orthogonal clearance angle α n lies between the tool flank face Aα and the tool cutting edge plane Ps , measured in the tool cutting edge normal plane Pn . The angle between the flank face Aα and the tool cutting edge plane Ps , measured in the tool orthogonal plane Po , is designated as the tool orthogonal clearance angle α o . The angle between the rake face Aγ and the tool reference plane Pr is defined as the tool orthogonal rake angle γ o . The angle between the rake face Aγ and the flank face Aα is called the tool orthogonal angle β o . As rule, the clearance angle, wedge angle and rake angle – defined in the tool orthogonal plane Po , the assumed working plane Pf and the tool cutting edge normal plane Pn – must in total make a 90◦ angle. The kinematic factors of tool cutting edge angle κ r , tool cutting edge inclination λs and tool normal rake angle γ n are important influencing parameters on the cutting process. For further information, see DIN 6582.

3.3 Basic Process Variants Despite the large number of cutting process variants, they can still all be subdivided into three main categories: • free, orthogonal cuts, • free, diagonal cuts and • bound, diagonal cuts.

46


vch

Tool

Tool

vch Chi

p

Chip

λs vc

vc Workpiece

Workpiece

Fig. 3.9 Free, orthogonal and free, diagonal cut

The free, orthogonal cut is a special case, which can be brought about by the following marginal conditions: • Only the major cutting edge is being engaged (free). • The tool cutting edge angle κ r is 90◦ (orthogonal). • Tool cutting edge inclination λs is equal to 0◦ (orthogonal). Practically speaking, this cut can be executed, for example, by means of longitudinal face turning or cross cylindrical turning (see Sect. 9.1) allowing for the above marginal conditions (Fig. 3.9). The free, diagonal cut is more general and contains the free, orthogonal cut. The following marginal conditions must be realized in the case of the free, diagonal cut: • Only the major cutting edge is engaged (free). • The tool cutting edge angle κ r can take on values that are not equal to 90◦ (diagonal). • Arbitrary tool cutting edge inclinations λs are permissible (diagonal). As soon as one of the last two marginal conditions is fulfilled, it is a free, diagonal cut. Both conditions don’t have to be met simultaneously, although this can occur. The general case, the bound, diagonal cut (Fig. 3.10), contains all abovementioned special cases and, in addition to the marginal conditions of the free, diagonal cut, also permits the engagement of the minor cutting edge (bound). All the basic process variants can be further extended by adding the following categories: • The uninterrupted cut and • the interrupted cut. The interrupted cut is the general variant among them, in which the cut takes place only intermittently. In the case of the uninterrupted cut, temporal interruption is infinitely small (the cut is continuous). Both process possibilities will be described

3.4

Chip Formation

47

Fig. 3.10 Bound diagonal cut

vch

Chip

Too

l

νc

ce

Workpie

in great detail in later chapters (see Chaps. 9 and 10). In total, there are therefore six basic process variants.

3.4 Chip Formation At the beginning of the chip formation process, the cutting section penetrates the material, causing it to deform elastically and plastically. After the maximum permissible material-dependent shear stress is exceeded, the material begins to flow. Contingent on a given cutting section geometry, the deformed material forms a chip, which runs off the rake face of the cutting section. The property of plastic deformability is not solely related to the material; it can also be brought about in a targeted way by altering the stress. The amount of stresses is influenced by the feed velocity vf , the cutting speed vc and the depth of cut ap . As far as process kinematics is concerned, the direction of cutting section stress is determined by defining the tool normal rake angle γ n , the tool cutting edge angle κ r and tool cutting edge inclination λs . In order to ensure chip formation, a minimum chip thickness and depth of cut must be exceeded (see Chap. 7 [Moll39, Djat52, Soko55, Bram61, Bram60]).

3.4.1 The Cutting Process By varying the direction and amount of stress, either tough or brittle material behaviour can be realized [Karm11, Böke14]. This has a large effect on chip formation. During cutting, the direction of a particular stress can be set by the process parameters tool normal rake angle, tool cutting edge angle and tool cutting edge inclination. The amount of the particular stress is influenced by the cutting parameters cutting speed, feed velocity and depth of cut. To put it another way, the same material can be tough or brittle depending on the direction and amount of stress,

48


Fig. 3.11 Model representation of deformation, acc. to Leopold [Leop00]

independently of how it behaves at room temperature and under single-axis tensile loads (tension test). When cutting steel, one must bear in mind that brittle behaviour only occurs at very low temperatures. This is not the case when machining cast iron, glass and ceramics. There is brittle fracture behaviour for these materials even at high temperatures due to their material structures. The particular stress must therefore be optimally adjusted to the given material properties with process kinematics in both direction and amount. To clarify plastomechanical processes during chip formation, a grid is superimposed on the workpiece (Fig. 3.11) so that material deformation during cutting can be observed (the visioplasticiy method) [Hast67, Chil71, Leop80, Leop00]. With this method, the chip formation process can be made visible, and with the help of this, chip formation models can be made. Finite element analysis is another possibility for modelling chip formation. The visioplasticity method can help to verify FE modellings, since it allows for a temporal and local definition of chip formation on the superimposed grid [Leop80]. Special marginal conditions exist during chip formation: both high rake face temperatures (about 770–1700◦ C) as well as high deformation speeds (of an order of magnitude of 104 /s). Figure 3.12 is a schematic representation of the chip formation process as sketched with the help of a photograph of chip initiation (right). In this representation, we can see a continuous plastic deformation that can be subdivided into four zones. The transition from the workpiece structure (a) to the chip structure (b) is made by simple shearing (shear zone). When cutting brittle materials, minor deformation on the shear plane can already lead to material detachment. If however the material has higher deformability, detachment first occurs in front of the cutting edge in zone (e). The tensile load under simultaneous perpendicularly

3.4

Chip Formation

49

Structure of the workpiece Shear plane

vc

Structure of the chip a

Shear zone 0,1 mm

e

b

d

c

vch

Flank face rake face Cut surface Tool

Tool Cut surface

Fig. 3.12 Location of chip initiation

active pressure leads, together with the high temperatures prevalent here, to strong deformations on the peripheries of the rake face (c) and cut surface (d). Sliding over the tool surfaces causes further plastic deformations to arise in the boundary layers. The “flow zone” (the non-etched white zone on the bottom of the chip), the deformation texture of which forms parallel to the rake face, gives the impression of a viscous flow process with an extremely high degree of deformation. The chip resulting from the described chip formation process is designated as a continuous chip. Other chip types include lamellar chips, segmented chips and discontinuous chips. Figure 3.13 shows a quantitative profile of normal and tangential stresses resulting from the resultant force components acting on the rake face. These stresses – in conjunction with temperatures prevalent in the contact zone, which can amount to over 1000◦ C in the continuous chip formation zone – lead to deformations with sheer strains ε¯ between 0.8 and 4.0 and sheer strain speeds ε˙¯ of up to 106 /s. For the sake of comparison, Fig. 3.13 provides corresponding figures from the tension test. Cutting conditions under which cemented carbide tools operate result in deformation and material heating durations in the order of magnitude of milliseconds; the heating velocities are theoretically around 106◦ C/s [Opit70].

3.4.2 Different Types of Chip Formation When chip formation has been ensured, it can occur in different ways. ERNST [Erns38, Erns41] has made a phenomenological classification of chip formation types. Continuous chip formation is characterized by an evenly deformed material

50


Workpiece Material: C45E; Tool Material: HW-P20; ap = 2 mm; f = 0.25 mm; vc = 160 m/min Cutting temperature: Chip flow velocity: Heating velocity: Average normal stress: Average shear stress: Shear strain Shear strain rate σn

T1 ~ 1030 °C vch = 67 m/min 106 °C/s σ−n = 350 N/mm2 − τ = 250 N/mm2 Tension test: ε ~ 0.2 Chipping: 0,8 < ε < 4.0 Tension test : ε ~10–3 s–1 Chipping: ε ~ 104 s–1

Distribution τ of stress w

σ

vch

τ

d

T3 vc

T4

Flow area

T5

Tool

vch

Detail A Fig. 3.13 Conditions of the cutting process, acc. to KÖNIG [Köni67]

chip structure, the cause of which is assumed to be in the temporally highly uniform friction conditions between the chip and the tool. Lamellar chip formation is characterized by an unevenly deformed material structure between the chip and the tool, the cause of which is explained by temporally highly altered friction conditions between the chip and the tool (stick-slip) or by dynamic stress transfer [Scha64]. Locally enhanced structural deformations, shear bands, can be recognized in the chip structure, that are characteristic of this type of chip formation. The frequencies of the vibrations caused by stick-slip are in the range of kilohertz and have small amplitudes. The high level of local structural deformation is

3.4

Chip Formation

51

explained by the fact that the thermally caused material softening is more dominant than its mechanical hardening. Research into these phenomena relies on the work of ZENER and HOLLOMON, who have described them and conceived the term “shear band”. Shear bands appear when machining high-strength materials with high levels of deformability. They are also especially frequent when machining with high cutting speeds [Hopp03]. If the stress condition in the shear zone exceeds the deformability of the material (shear strength), there is a detachment of material areas, which then fuse with each other again. This leads to the formation of segmented chips. This can be conceived of as a special case of lamellar chip formation in which highly localized deformations (shear bands) arise as well. In addition to vibrations, the entire system – consisting of the machine tool, tool, workpiece and fixtures – is influenced by further dynamic effects such as the regenenerative effect or directional coupling [Weck77]. The frequencies of these effects are much lower, in the area of a thousand hertz. In contrast to the higherfrequency chip formation dynamics, these vibrations do not so much influence chip formation in the sense of material deformation as much as chip formation from outside, since they are the cause of dynamic alterations in the nominal feed. SCHWERD has made microcinematographic experiments and suggested the concept of “discontinuous chip” [Schw36]. This process basically distinguishes itself from the other processes of chip formation by the fact that no plastic deformation occurs before fracture, but rather fracture takes place without plastic deformation. Discontinuous chips can be observed in the case of materials with very brittle properties, e.g. cast iron, stone, fibre-reinforced plastic or titanium alumnides. Figure 3.14 summarizes the principal chip types [Vier70]. • Continuous chips form when the material has sufficient deformability (εB > ε0 ), the microstructure is uniform in the cutting area, deformation does not cause embrittlement and chip formation is not impaired by vibrations. • Lamellar chips form when εB < ε0 < εF or the microstructure is not uniform or vibrations lead to variations in chip thickness. Lamellar chips can form with high feeds as well as with high cutting speeds. • Segmented chips consist of chip segments that are separated in the shear plane and fuse together again. The form when εF < ε0 , whereby this is not only the case for brittle materials like cast iron but also can come about if deformation causes embrittlement in the microstructure. Segmented chips can also be formed at extremely low cutting speeds (vc = 1 − 3 m/ min). • Discontinuous chips most form when cutting brittle materials with uneven microstructures such as certain types of cast iron and stone. The chips are not detached, but are torn off the surface, often causing damage due to small breakings from the workpiece surface. Figure 3.15 gives a tool-oriented overview of the dynamic system of spheres of influence, all complexly overlapping and mutually retroactive, which brings about

52

3 Fundamentals of Cutting 2 Lamellar chip

3 Segmented chip

Range of lamellar-, Range of segmented- and continous discontinous chip chip

4 Discontious chip

Elastic region Plastic region Yield range

1 2 3 4

Shear strength τ

Shear strength τ

1 Continous chip

ε 0: E

B

F

Strain ε

E: B: F:

Degree of deformation in the shear plane Elastic limit Breaking limit Fraction point

ε0

Degree of deformation ε Fig. 3.14 Types of chip depending on material properties, acc. to VIEREGGE [Vier70]

Range of temperature

Material properties Tool wear

Friction and Deformation

Range of velocity

Process parameter

Tool geometry

Mechanism of chip formation: • Continous chip formation • Lamellar chip formation • Segmented chip formation • Discontionous chip formation

Fig. 3.15 Influence on chip formation

the various types of chip formation. This system is heavily influenced by the velocity fields, the materials to be machined, tool geometry and process kinematics. This section will not deal with low-frequency dynamic effects on the part of the machine tool. See the relevant literature for further information [Kron54, Opit70, Weck77, Alti00].

3.4

Chip Formation

53 Feed motion

Tool

Primary motion Buffed and etched Workpiece Quarz glass pane Hollow spindle Tool reference plane Pr

Fig. 3.16 Microcinematograpy

One method with which chip formation is directly observable, is microcinematography, the first application of which on the cutting process was already carried out by KURREIN and then later by K LOPSTOCK [Kurr05, Klop23, Klop26] (see Fig. 3.16). The development of this process has been advanced by several researchers in the course of time [Schw36, Merc45, Oxle59, Spaa71, Warn74]. In this process, chip formation is observed through a quartz glass pane while photos are taken or films are recorded successively in short intervals. One disadvantage of this procedure is the limitation to very low cutting speeds (vc = 1 cm/min). Nevertheless, the process has helped to explicate basic phenomena. Today there are high-speed cameras available that can capture 150,000 images per second. With such cameras, the phenomena can be made visible at the lower threshold at higher speeds or especially during micro-cutting. In order to be able to study chip formation more effectively, various methods have been developed for cut interruption. During the cutting process, chip roots are held “in statu nascendi”, providing information about chip formation after subsequent metallographic preparation. In the case of the equipment used today, the tool is suddenly accelerated out of the cut. The acceleration is brought about by a predetermined breaking point on the swivel device that breaks during the cutting process due to an impulse caused by an explosion [Lola49] (Fig. 3.17). Gente has recently developed a method in which the delay time and the masses moved can be reduced in comparison with the “explosion method” [Gent02] (see also Fig. 3.17). In this way, investigations into chip roots are also possible with higher cutting speeds. Other methods have been developed in which a chip root is produced in the workpiece by means of a predetermined breaking point (brittle fracture, Fig. 3.18) [Buda68]. The advantage is that that chip root is detached from the workpiece so that the effective area can be spared during the following metallographic preparation.

54

3 Fundamentals of Cutting Gun Moment 1

Moment 1

Sample Moment 2

Tool Moment 2 Direction of primary motion

Tool Baffle

Sam

Buffed sheet steel with high Young’s modulus

ple

Fig. 3.17 Interruption of cutting process, acc. to LOLADSE [Lola49] (left) and GENTE [Gent02] (right)

Tool

Predetermined breaking point Fig. 3.18 Interruption of cutting process by predetermined breaking point, acc. to BUDA [Buda68]

3.5 Kinematic Surface Roughness Calculations of interpenetration between the tool and the workpiece can provide key information about the cutting process, including the associated undeformed chip cross-sections in the case of changing engagement parameters (see also gear manufacture) or micro-geometrical surface parameters. The following shows by way of example the generation of surface quality for a simple external turning process. Since chip formation processes are ignored, we speak in this case of the generation of kinematic surface roughness. For this, the penetration of the workpiece by the tool is geometrically evaluated taking into consideration the kinematics in the tool reference plane Pr (Fig. 3.19).

3.6

Mechanical and Thermal Strain on the Cutting Section

55

z

x rε

f

Rt

rε –Rt M1

κr

f/2 f

M2 Rt = rε − rε2 −

f2 4

or Rt =

f2 8 ⋅ rε

Area of validity: f < 2 · rε cos(κ r + ε r – 90°) and κ r + ε r < 180°

Fig. 3.19 Geometric ratio of engagement in the cutting process

3.6 Mechanical and Thermal Strain on the Cutting Section The resultant force F, represented here in the turning process as an example, can be decomposed into the components cutting force Fc , feed force Ff and passive force Fp (Fig. 3.20). Currently, these resultant force components are usually detected metrologically with the help of piezoelectric force sensors.

ve

vc

Primary motion (Workpiece)

Ff

vf Fp

Fc

Fa Fz FD

Direction of feed (Tool)

F : Resultant force Fc : Cutting force Ff : Feed force Fp : Passive force Fa : Active force FD : Thrust force

vc : Cutting speed vf : Feed velocity ve : Effective cutting speed

Fig. 3.20 Resultant force and its components in the cutting process, acc. to DIN 6584

56


A

h

vf Ff

Fc

Cutting force

Ff

Feed force

Fp

Passive force

Ffp

Resultant of Ff and Fp Normal force of major cutting edge Tangential force of major cutting edge Normal force of rake face Tangential force of rake face Normal force of shear plane Tangential force of shear plane Resultant of FγT and FγN, or FΦT and FΦN

FκN FκN

Fκ T A Fγ N

Ffp

Fp

κre

FκT Fγ T

λ se = 0°

FΦ N

Sectional view A-A vc

FΦ T Shear plane

h

Φ

Fκ N

βoe

FΦT

Fc Workpiece

hch vch

γoe

Fγ T

Fcγ N

Fγ N

FΦN

αoe

FcγN

Tool

h

Undeformed chip thickness

hch

Chip thickness

vc

Cutting speed

vch

Chip speed

α oe β oe γ oe κ re λ se Φ

Tool orthogonal clearance Wedge angle Tool orthogonal rake angle Working cutting edge angle Working cutting edge inclination Shear angle

Fig. 3.21 Components of resultant force in working plane of reference (top) and working plane (below)

The analysis of forces derived for the orthogonal cut by Merchant [Merc45, Merc45a] is the basis for determining the forces acting on the tool cutting edge (Fig. 3.21). Assuming an ideally sharp cutting edge and neglecting flank face wear, the tangential force Fγ T and the normal force Fγ N acting on the tool can be calculated from the force components. The following applies as long as the inclination is λs = 0◦ and the influence of the minor cutting edge is slight: Fγ N = Fc cos γo − (Ff sin κr + Fp cos κr ) sin γo

(3.1)

Fγ T = Fc sin γo + (Ff sin κr + Fp cos κr ) cos γo

(3.2)

3.6


57

where the bracket term in Eqs. (3.1) and (3.2) corresponds to the normal force of the major cutting edge FκN (Fig. 3.21). Because of the (in most cases) small difference between the angles in the tool-in-hand system and the tool-in-use system, calculations are made within the tool-in-hand system for the sake of simplicity. For the case γo = 0◦ , the cutting force Fc and the feed force Ff act perpendicularly/tangentially to the rake face. In contrast to Merchant’s original work, a neutral tool orthogonal rake angle (γo = 0◦ ) was selected, and the forces were shown with the effective cutting direction on the tool cutting edge. Designation of the acting forces and their signs is derived from the standards ISO 3002/4 and DIN 6584. As Fig. 3.22 clarifies for the simplest case of an orthogonal cross section with a neutral tool orthogonal rake angle, the cutting and feed forces on the worn tool are composed not only of the normal force of the rake face Fγ N and the tangential force of the rake face Fγ T but also of the tangential force of the flank face FαT and the normal force of the flank face FαN . The friction coefficient, which can be calculated from the cut and feed forces, thus contains not only rake angle friction but also forces acting on the flank face caused for example by elastic deformation of the workpiece surface and by tool wear [Klau65]. As research has shown, the normal force of the flank face amounts to about 40% of the feed force and thus to

h Φ F

ΦN

Fa

Chip F

Fz

ΦT

=

Fc

Ff vf

Tool Fγ N

vc

Fγ T acc to: Merchant

Fig. 3.22 In the cutting process the mechanical strain of the cutting edge results from normal and tangential forces and their components acting on rake and flank face, acc. to MERCHANT [Merc45, Merc45a]

Fα N

VB Fα T

|Fc | = |Fγ N |+ |Fα T | |Ff | = |Fγ T |+ |Fα N | True for: γ o = 0°, κ r = 90°, VB > 0

Cutting edge

58


Feed force Ff

Passive force Fp

Fig. 3.23 Components of resultant force depending on feed, cutting velocity, tool cutting edge angle, and depth of cut (qualitative)

Cutting force components

Feed f

Fc

Ff Fp

Tool cutting edge angle κr


Cutting force Fc



about 66% of the rake face component in the feed direction [Spaa67, Lutz68]. A percentage of approximately 10% of the cutting force Fc was ascertained for the tangential force of the flank face. The average normal and tangential stresses originating from the resultant force components acting on the rake face are between 350 and 400 N/mm2 or between 250 and 350 N/mm2 when machining construction steel [Köni72]. Materials that are difficult to cut result in values of 1100 N/mm2 . Their profile is qualitatively reproduced in Fig. 3.13. The size and direction of the resultant force are strongly influenced by the cutting parameters and cutting section geometries used. Figure 3.23 shows the dependence of the static components of the resultant cutting force Fc , Ff and Fp on the feed f, cutting speed vc , depth of cut ap and the tool cutting edge angle κ r qualitatively in a linear coordinate system. The extremes in the profiles of the resultant force components over cutting speed can be ascribed to growth of built up-edge. The reduction of forces with increasing cutting speed is caused by the reduction of material strength at higher temperatures. The components of the resultant force increase proportionally over the depth of cut ap . Yet this is only valid if the depth of cut is larger than the corner radius of the tool. The profile of feed force Ff and passive force Fp over the tool cutting edge angle κ r results from the geometric position of the cutting edge with respect to the workpiece axis, since with a larger cutting edge angle the resultant force component aimed in the feed direction increases, and its maximum is reached at κr = 90◦ . If the tool cutting edge angle is increased, the undeformed chip thickness h increases proportionally to the reduction of the width of undeformed chip b. Since cutting

Fc

Ff Fp

Cutting speed vc Fc

Ff Fp

Depth of cut ap

3.6


Influencing variables

59

Change of cutting force components per degree angle variation passive force Fp

Decreasing

Feed force Ff

Tool orthogonal rake angle

1.5 %

5.0 %

4.0 %

Tool cutting edge inclination

1.5 %

1.5 %

10.0 %

Increasing

Cutting force Fc

Tool orthogonal rake angle

1.5 %

5.0 %

4.0 %

Tool cutting edge inclination

1.5 %

1.5 %

10.0 %

Fig. 3.24 Influence of cutting edge and rake angle on the components of the resultant force

force Fc is proportional over depth of cut ap (= ˆ width of undeformed chip b) but increases degressively over feed (= ˆ undeformed chip thickness h), a light reduction of Fc with increasing κ r is the outcome of both changes. Figure 3.24 provides some standard values for how the components of the resultant force change if the tool orthogonal rake angle or the cutting edge inclination is varied. This information can vary greatly however and should only be seen as reference values. Changing the tool orthogonal clearance angle in the range of 3 ≤ αo ≤ 12◦ has no mentionable effects on the resultant force components. Changing the corner radius also does not have a significant effect on the forces as long as the condition 2r ≤ ap is met. Tool wear (see Sect. 3.7) is another variable that influences resultant forces. Varying effects on the components of the resultant force can be observed depending on the type of wear. Crater wear, resulting in a larger positive tool orthogonal rake angle, generally leads of a reduction of cutting forces. On the other hand, these forces increase when flank face wear is dominant since the friction surface between the workpiece and the flank face becomes larger. Quantitative determinations concerning force increase with increasing tool wear can only be approximated due to the large number of influencing parameters. The following reference values can be assumed approximately for force increase up to a width of flank wear land of VB = 0.5 mm: About 90% for the feed force Ff , about 100% for the passive force Fp and about 20% for the cutting force Fc . Several specific characteristic values can be distinguished for calculating the components of the resultant force. These component forces are directly proportional to the width of undeformed chip b. To identify empirical laws, it is usually advisable to relate the measurement to already known, linearly dependent magnitudes in

60


Cutting force Fc = Fc’ / (N/mm) Width of undeformed chip b

2000 kc1.1 = 1740 N/mm2

mc

1000 1

800

1 –mc = 0.7265

600 45° 400 1 Fc' = kc1.1 ⋅h1–mc 200 0.1

0.2 0.4 0.6 0.8 1.0 Undeformed chip thickness

2.0

h / mm

Fig. 3.25 Graphical determination of characteristic values kc1.1 and (1–mc )

order to minimize the number of variables going into the physical law. In this case, we form the quotient Fc from the cutting force Fc and the width of undeformed chip b. If we now plot the values thus found over undeformed chip thickness h in a double logarithmic plot, the measurement points arrange themselves in a straight line (Fig. 3.25). The corresponding linear equation log(Fc /b) = log(kc1.1 ) + (1 − mc ) · log h

(3.3)

can be converted into the KIENZLE Equation Fc = kc1.1 · h(1−mc )

(3.4)

The specific cutting force kc1.1 is the cutting force required to detach a chip of undeformed chip width b = 1 mm and undeformed chip thickness h = 1 mm. The exponent (1–mc ) designates the gradient of the straight line Fc = f (h) in the double logarithmic system. To determine kc1.1 and (1–mc ), cutting experiments are carried out for the combination of workpiece material and cutting tool material under investigation. In these experiments, the relevant cutting forces are measured with constant cutting speed, depth of cut and cutting section geometry and plotted in accordance with Fig. 3.25. The required specific cutting force characteristic parameter kc1.1 is determined by

3.6


61

extrapolating the undeformed chip thickness to h = 1 mm. The tangent of the angle between the straight line and the x-axis is the desired gradient value (1–mc ). Corresponding equations and characteristic values can be defined for resultant force components Ff and Fp : Ff’ = kf1.1 · h(1−mf )

(3.5)

Fp = kp1.1 · h(1−mp )

(3.6)

The values determined in this way are however only valid for undeformed chip thicknesses of h > 0.1 mm. Values for kc and (1–m) can be found in [Köni82]. According to DIN 6584, the energy for cutting is the result of the product of the paths to be travelled or paths travelled and the components of the resultant force acting in their direction. Correspondingly the powers arising during the cutting process result from the product of the speed components and the components of the result force acting in their direction. Cutting energy Wc and cutting power Pc : Wc = lc · Fc

(3.7)

Pc = vc · Fc

(3.8)

Feed energy Wf and feed power Pf : Wf = lf · Ff

(3.9)

Pf = vf · Ff

(3.10)

Effective energy We and effective power Pe are understood as the sum of all corresponding cut and feed amounts: We = Wc + Wf

(3.11)

Pe = Pc + Pf

(3.12)

Due to the relatively low feed speeds and feed paths, feed energy/power during turning amounts to only about 0.03–3% of the corresponding cutting energy or cutting power. For this reason, we can assume We ≈ Wc and Pe ≈ Pc for most cases. Figure 3.26 provides an overview of the segmentation of the total active energy into shear, cutting and friction energy as a function of the undeformed chip thickness [Vier70]. The illustration shows that the amounts of the different types of energy depend on the undeformed chip thickness, whereby shear energy has the largest share for most undeformed chip thicknesses. The mechanical work used in cutting is almost completely converted into thermal energy. Since the heat centres are identical with the deformation centres, the shear

62

3 Fundamentals of Cutting Shear energy Effective energy We

Deformation energy Cutting energy Friction of flank face

We = Fe ·Ie

Latent energy and heat

Friction energy Friction of rake face

700

Material:

/ (m·daN/m)

600

Tensile strength: Rm = 800 N/mm2

Total energy

Cutting speed:

500

Shear energy

300

b = 4.25 mm Tool orthogonal clearance: αο = 5°

200

Effective energy We

vc = 100 m/min

Width of undeformed chip:

400

Path length due to primary motion

55NiCrMoV6

Friction of rakeface

100

Friction of flank face and cutting energy

0 0

0.2

0.4

0.6

0.8

Tool orthogonal rake angle: γο = 10°

1.0

Width of undeformed chip: h / mm

Fig. 3.26 Segmentation of effective energy in cutting process depending on chip thickness, acc. to VIEREGGE [Vier59, Vier70]

zone and the friction zones on the tool come into consideration as heat sources. As Fig. 3.13 shows, the degree of deformation on the bottom of the chip is much higher in the flow zone than in the shear zone, so one can expect the highest temperatures between the chip and the tool. Since the thickness of the flow zone is very small in comparison with the shear zone however, these higher temperatures should not also be equated with a high energy conversion. The illustration in Fig. 3.27 left gives information about heat that is absorbed/dissipated by the workpiece, chip and tool. Most of the heat is dissipated by the chip. Most of the mechanical energy (in this case 75% and generally more than 50%) is converted in the shear zone. The heat arising in the individual development locations is dissipated by thermal conduction, radiation and convection to the environment. As a result of this heat balance, corresponding temperature fields form in the workpiece and tool that change until equilibrium between added and removed heat is achieved. The right side of the figure shows such a temperature field. If we consider one material particle in the cutting zone, its temperautre will be at least equal to that of a particle in the shear zone. As it glides further into the contact zone, the material is intensively heated on the chip bottom side and the tool on the rake face, because the energy required to overcome friction between the chip

3.6


63

Workpiece 300

310

Chip

o

5% 75% 2% 18%

380 C 130 80 500 30

600 700 650 600

For steel

400 450 500

Tool

Material: Steel Flow stress: kf = 850 N/mm2 Tool material: HW-P20 Cutting speed: vc = 60 m/min Width of undeformed chip: h = 0.32 mm Tool orthogonal rake angle: γo = 10o

Fig. 3.27 Distribution of heat and temperature in workpiece, chip, and tool in the process of steel cutting, acc. to KRONENBERG [Kron54] and VIEREGGE [Vier59]

and the rake face is almost completely converted into heat. Since this process only takes place in boundary layers of the chip and cutting tool material it heats up the rake and chip bottom side all the more strongly the less time is available for heat dissipation due to the higher cutting speeds. The maximum temperature does not appear directly on the cutting edge, but rather at a certain distance from it on the rake face depending on the cutting conditions. Determining the temperatures arising during cutting has long been the subject of intensive research [Gott25, Lang49, Schm53, Küst54, Vier55, Küst56, Axer55]. Already in 1956, KÜSTERS demonstrated temperatures of over 1000◦ C on the rake face when cutting steel with cemented carbide [Küst56]. Later investigations have qualitatively confirmed these results again and again [Cass94, Denk90, Ehme70a, Lenz66, Laus88, Beye72, Dama90, LoCa94]. The level of the temperatures acting on the cutting section depends on the machined material, the cutting tool material, the selected cutting conditions, tool wear and the cooling medium. Figure 3.28 provides an impression of the order of magnitude of the mean cutting temperatures to be expected on the rake face as a function of cutting speed for various cutting tool materials. In the vc = 20–50 m/min range, the temperature profile is non-linear in the double logarithmic coordinate system. The reason for this is the the growth of built up-edge arising in this cutting speed range (see Sect. 3.7.2), which interferes with direct heat conduction. In comparison to the rake face, temperatures on the flank face are generally low (Fig. 3.29) [Denk90, Ehme70a, Axer55, Laus88, Beye72, Dama90]. The size of the temperature difference between the rake face and the flank face, about 200–300◦ C

64

3 Fundamentals of Cutting 1150 1100 1050 1000 950

HW-P1 x

900

Temperature of rake face ϑ / °C

850 800 750

x x

End of growth of– build–up edge

HW-P30

x x

700

x

650 HS12–1–4–5

600 550

Tool material: HW–P10, P30 High speed steel HS12–1–4–5 Material: C53E Geometry of cutting section:

500 450 400

κr

εr

αo

γo

λs

6°

6°

0° 70° 84° 0.8 mm

rε

Cross section of undeformed chip: ap · f = 3 x 0.25 mm2 Cutting time: tc = 15 s

350

300 10

16

25

40

63

100

250

Cutting speed vc / (m/min)

Fig. 3.28 Mean rake face temperature

when cutting with cemented carbide tools according to [Ehme70a], is strongly codetermined by the position of the contact zone on the rake face. The closer the rake face moves towards the cutting edge, the more the temperatures prevalent on the rake and flank faces conform. This means that temperatures are very high on the flank face as well, especially during finishing operations with high cutting speeds and small feeds. According to the information provided by the literature [Ehme70a, Vier53, Schm53, Küst54, Vier55, Küst56, Axer55, Beye72], temperatures of over 800◦ C can be found on the flank face depending on the selected marginal conditions when cutting with cemented carbide tools. There are various methods and devices available for temperature measurement during the cutting process (Fig. 3.30) [Lowa67], which vary according to technique, experimental setup and measurement position. The only metrological techniques that come into question for measuring temperature during the cutting process are time-resolution measurement techniques, since the cutting process duration (and thus the achievement of a thermally stationary state) is too short for other methods

3.6


65

Weld

1200°C

3

2

1

0

1000°C 4

800°C 1

600°C 500°C

Compensating resistance Material: Cutting tool material: Cutting speed: Cross section of undeformed chip:

2

Isolation of silicon varnish

Ni-tube ∅ 0.13−0.20 mV

400 °C

Example for the distribution of temperature inside the cutting edge

CrNi-Wire ∅ 0.07 mm

30Mn4 HW-P10 vc = 180 m/min

Cutting wedge geometry γo

αo λs

0°

Bevel width: 0.5 mm 4° 0° 45° 90° 45° to cutting edge

2

κγ

εr

rε

ap·f = 3 ·0.74 mm

Fig. 3.29 Measurement of temperature with embedded thermocouple on cemented carbide, acc. to Küsters [Küst56]

Temperature measurment during machining

1. Caloric heat measurement

2.1. Coat of paint at an uninvolved surface of the tool

2. Measurement with thermo-colours

3. Thermo-electric measurement

2.2. Coat of paint at the inner surface of a splitted tool

4.1. Radiation pyrometer

4.2. Infrared photothermometry

3.1. One-chisel method

3.2. Two-chiselmethod

4. Radiationmeasurement

3.3. Assembly of a complete thermocouple

Fig. 3.30 Temperature measurement in cutting processes

4.3. Measurement of infrared radiation with PDS-cell

3.4. Assembly of an isolated wire as themocouple

66


(e.g. coating with temperature-sensitive colours or powders with constant melting points). Of the processes introduced in Fig. 3.30, the one and two-chisel methods, temperature measurement with thermocouples, pyrometers and photo thermometry are of technical interest. GOTTWEIN developed the one-chisel method as the first direct method for measuring temperature and then, in cooperation with REICHEL, the two-chisel method [Gott25, Herb26, Vier70]. Both methods are based upon the principle of a thermocouple. Both tool and workpiece form the heat soldering joints, while the tool clamp is the cold soldering joint. The workpiece and the tool must be clamped in isolation, since the thermoelectric voltage arising in the contact zone between the cutting insert and the chip would otherwise be dismantled through the machine in a short circuit. Contact between the passing chip and the workpiece outside of the contact zone must therefore also be avoided. Both methods require electrically conductive materials and cutting tool materials. In the case of the two-chisel method, cutting tool materials with varying thermoelectric properties must be used. The greatest disadvantage of the single-chisel method is the protracted and elaborate process needed to calibrate the thermocouple which must be carried out again for every cutting tool material and workpiece material combination. This process can detect temperatures of up to 1200◦ C. Temperature measurement with thermocouples is among the most common techniques today. Installing a thermocouple into the tool or workpiece allows for a point-wise determination of the temperature field (Fig. 3.29). The thermocouples included are classified as encapsulated thermocouples, sheathed thermocouples and single filament thermocouples. Thermocouples can be applied in blind holes in the tool [Küst56, Qure66, Abra97] or in the workpiece [Osul02]. Temporal resolution is influenced by the response time of the thermocouple and heat transfer between the thermocouple and the device under test. These techniques generally have low temporal resolution. Also problematic is the contact heat transfer resistance between the surface under test and the thermocouple due to the roughness of the bore. This causes a difference in temperature between the measurement surface and the thermocouple. In the case of sheathed thermocouples with isolated measuring points, there is also the distance between the thermocouple surface and the internal measurement point. Due to the extremely high temperature gradients with short test times characteristic of cutting processes, this can lead to much lower measurements. The maximum values dominant in the friction zones cannot be measured in this way. Thermal compounds are used to improve heat transfer between the thermocouple and the surface. Another disadvantage is that direct contact between the thermocouple and the test object is necessary and that the holes used to position the thermocouples can significantly affect the distribution of temperature. The most important techniques in radiation measurement, which determines temperature by measuring the heat radiation emitted from a surface, are pyrometry and thermography. Pyrometry is the contact-free determination of absolute temperature by measuring the inherent radiation of a body without spatial scanning of the object field. Thermography provides a pictorial representation of temperature distribution. Radiation techniques have decisive advantages compared with thermoelectric

3.6


67

methods: the time resolution is much higher (whereby pyrometers are principally faster than infrared cameras), and they are also contact-free. One significant problem when measuring for an exact absolute temperature with a radiation method is the dependence of the radiation emitted on the grade of emission of the surface. Since the emission grade is the function of many factors like temperature, wavelength, angular position, material and surface condition, calibrating the measurement device for a particular surface is very difficult. The precision of total radiation and broadband partial radiation pyrometers are especially influenced by factors that alter the spectral grade of emission of the surface. In cutting, effects such as surface roughness and oxidation influence the grade of emission of different surfaces greatly. To limit the influence of the grade of emission on measured temperatures, narrow-band partial radiation, two-colour and multi-colour pyrometers have been developed. The two-colour pyrometer (Fig. 3.31) has the advantage that the spectral grades of emission ε 1 and ε2 of the surface need not be known. Since the two selected wavelengths lie directly next to each other, ελ1 ∼ ελ2 . An error in measurement will only result if both wavelengths λ1 and λ2 differ greatly. Further advantages of this principle are that the measured temperature is independent of signal dampening, due to dust for example, so long as both signals are dampened equally. Moreover, the temperature of objects that are smaller than the optical field of vision can be measured without error [Müll85]. Thermography with infrared cameras can be used as an alternative to pyrometrical measurements. The advantage to this is the pictoral representation of temperature information. Commercial cameras usually work with long wavelengths and large broad-bands, which makes it possible to measure lower temperatures but which also has a negative effect on the attainable accuracy. Scanning cameras that work with a single detector are too slow for fast processes. Some high-speed infrared cameras

Workpiece

Chip Lens

Dichroic beam-splitter

Point of measurement

IR-filter Insert

Glass fiber (∅ = 0.42 mm)

Photodiode

Amplifier

Characteristics • Measurement of two discrete wavelength bands λ 1 = 1.7 µm, and λ 2 = 2 µm which are set into ratio • High temporal resolution (approx. 2 µs) • Independence from grade of emission (no calibration required) • Large range of temperature measurement (250–1200 °C) • Contact free measurement

Fig. 3.31 Build-up of a two-colour pyrometer in principle, acc. to MÜLLER [Müll85]

68


offer integration times in the range of microseconds, making it possible to capture fast processes without motion blur. One major problem in the metrological determination of cutting temperatures, besides poor accessibility, is the limited local resolution of the available measurement techniques as well as the small range (only a few micrometers) within which the maximum temperatures arise. Local temperatures in the boundary layer between the cutting tool material and the sliding chip also cannot be captured with existing measurement methods. Promising attempts to determine thermal tool load more accurately than previously have been made with sensors in thin-film technology, which are applied to tool contact zones and determine temperature distribution within the contact zones during the machining process [Zieb95, Töns95, Kloc97]. This technology is however not yet suitable for cutting processes with high wear. Another highly promising method for determining temperatures in the boundary layers between the material and the cutting tool material is the implantation of a window made of diamond in the tool contact zone. Heat radiation emitted from the chip bottom side passes through the window via a mirror integrated into the chisel holder and reaches an infrared camera [Müll96]. A direct comparison of the temperatures measured by various authors is only rarely possible, since experimental conditions generally deviate from each other. There are numerous parameters that may have an effect on temperature: the type of cutting process, materials, cutting tool materials, possible coatings of the cutting tool materials, the use of cutting fluids and cutting parameters such as feed, cutting speed, depth of cut and cutting section geometry. Moreover, the measurement position is also crucial as well as metrological parameters like local or temporal resolution. The duration of the cutting process and the measurement time must also be taken into consideration. Cutting speed clearly has the greatest influence on the maximum temperatures that arise in the contact zone between the chip and the tool. In the case of the basic process, most measurements indicate an apparent rise in temperature at low to average cutting speeds, while at high speeds there is an approximately constant temperature level.

3.6.1 Influence of the Geometry of the Cutting Section Depending on the cutting task at hand, we can select from a diverse array of cutting section geometries. Geometry selection depends on • • • •

the cutting tool material, material, cutting parameters and tool geometry.

Typical tool angles in steel cutting are given in Fig. 3.32. Determining the tool angle always involves a compromise that can only do approximate justice to various requirements. Figure 3.33 shows how a change in cutting section geometry influences cutting parameters.

3.6

Mechanical and Thermal Strain on the Cutting Section Geometry of cutting section

Tool Tool Cutting Cutting edge orthogonal orthogonal edge rake angle clearance inclination angle

Cutting tool material

γο

αο

High speed steel (HSS)

–6° to +20°

6° to 8°

Cemented carbide

69

Tool included angle

Corner radius

λs

κr

εr

rε

–6°

10°

60°

0.4

to

to

to

to

+6°

100°

120°

2 mm

–6° to +15° 6° to 12°

Fig. 3.32 Tool angle for machining of steel Lower minimum chip thickness

Increasing stability of cutting section – γ o = –10° to +20° + Improved chip formation Improved surface Reduced cutting force Decreasing wear

α o = 6° to 12° Increased stability of cutting section

Chamfer of the rake face

κ r = 10° to 100° Decreasing wear

Decreasing chatter, Decreasing cutting force

Decreasing wear Increasing stability of cutting section Increasing of passive forces, Increasing stability of cutting section – + λ s = –6° to +6° rε = 0.4 to 2 mm Decreasing chatter, Decreasing cutting force Increasing surface quality

Controlled chip flow

Fig. 3.33 Influence of cutting edge geometry on characteristic machining values

3.6.1.1 Tool Orthogonal Clearance α o Wear on the flank face (denoted by the width of flank wear land VB) is mostly determined by the size of the tool orthogonal clearance. If it is large, the cutting section is weakened in two respects: a heat build-up can ensue in the tool that can possibly lead to impaired or loss of hot hardness; also, if the wedge angle is too small, this increases the danger of breaking on the cutting edge. If α o → 0◦ , surface

70


wear increases because pressure weldings appear on the contact areas of the friction partners to a greater extent. 3.6.1.2 Tool Orthogonal Rake Angle γ o , Wedge Angle β o The tool orthogonal rake angle γo can, in contrast to α o , be either positive or negative. It is responsible for the detachment of the material to be cut. The size of the tool orthogonal rake angle γo affects the stability of the wedge; highly positive tool orthogonal rake angles can thus lead to tool breakage as a result of cutting section weakening. The primary advantages of a positive tool orthogonal rake angle are the low cutting and feed forces as well as a usually improved workpiece surface quality. However, a chip flow that is supported by a positive tool orthogonal rake angle often has insufficient chip breakage (tendency towards the formation of continuous chips). Negative tool orthogonal rake angles increase cutting edge stability (used for example for planing and machining workpieces with apertures, rolling skin or cast skin). The thereby increased deformation of the passing chip and the large cutting forces result in a high temperature load on the cutting section. Increased crater wear appears on the rake face, which can lead to a decreased tool service life. Together with the wedge angle β o , the tool orthogonal rake angle γo and the tool orthogonal clearance form a right angle α o (Fig. 3.8). 3.6.1.3 Tool Included Angle εr Due to the tool stability that is desired under extreme cutting conditions, the tool included angle εr should be as large as possible. Small tool included angles are required especially for copier and NC machining. The possible range is limited by specifying the position of the major cutting edge and by making sure that the angle between the minor cutting edge and the feed direction is at least 2◦ in order to avoid subsequent shaving of the minor cutting edge on the workpiece. 3.6.1.4 Cutting Edge Angle κ r At constant feed and depth of cut, the width of undeformed chip b increases with decreasing κ r . This causes the specific cutting edge wear to sink, so small cutting edge angles are used especially to machine high strength materials in order to keep tool load and wear low. On the other hand, the passive force Fp rises with decreasing κ r , increasing the danger of clattering vibrations due to growing instability of the cutting process. 3.6.1.5 Cutting Edge Inclination λs A negative cutting edge inclination can stabilize the cutting process to a large extent since the lead of the tool does not take place on the cutting edge but rather towards the middle of the cutting edge. This results in an improved load profile, so that the danger of cutting edge breakage due to local overloading is reduced. First cuts that have minimal loads are of particular importance especially in the case of interrupted

3.7

Wear

71

cross sections (e.g. in milling or planing) as well as in the machining of cast iron and forged parts (workpieces with transverse drill holes, shrinkage cavities). Negative cutting edge inclinations induce large passive forces, which must be absorbed by the machine tools (stiffness perpendicular to the main spindle!). The cutting edge inclination also influences the direction of chip flow. A negative cutting edge inclination can result in the chip getting diverted to the workpiece surface, decreasing the surface quality. 3.6.1.6 Corner Radius rε The corner radius to be selected depends on the feed f and the depth of cut ap . Together with the selected feed, it influences the attainable workpiece surface quality to a great extent (see Sect. 7.2.3), whereby the following relation is approximately true: Rt =

f2 8rε

(3.13)

Large corner radii improve the surface quality and cutting stability. Small corner radii have the advantage of a smaller clattering tendency due to smaller passive forces.

3.7 Wear During the cutting process, deformation, separation and friction processes take place in the area of the cutting edge. The cutting tool materials used are subject to an extremely complex load collective characterized by high compressive stresses, high cutting speeds and high temperatures. Using cutting parameters common in practice, cutting tools reach the end of their service life because of continuously increasing wear on both rake and flank faces. This is explained as the progressive loss of material from the surface of a solid body, brought about by mechanical causes, i.e. contact and relative motion of a solid, liquid or gaseous counter body.

3.7.1 Wear Mechanisms The main mechanisms that cause wear are adhesion, abrasion, tribochemical reactions and surface disruption (Fig. 3.34). 3.7.1.1 Abrasion Abrasion occurs when rough areas of the counter body or particles present as intermediate material or also as counter bodies penetrate into the surface of the cutting tool material and simultaneously make a tangential motion, producing scores and

72


Adhesion

Abrasion

Tribochemical reaction

Surface damage

Fig. 3.34 Schematic illustration of the four main mechanisms of wear, acc. to ZUM GAHR [ZumG92]

Micro-plowing

Micro-cutting

Microcracks

Fig. 3.35 Types of material damages caused by abrasive parts, acc. to ZUM GAHR [ZumG87]

micro-cutting. The various forms of abrasion are designated as groove, fluid erosion, mill, notch or blast wear. Often, the term “groove wear” is used synonymously with the term “abrasive wear” [Habi80, Czic06]. In the case of groove wear, hard areas of roughness or hard particles penetrate into the surface of the stressed material, creating scores or grooves by a sliding movement. Wear can be caused by a rough, compact counter body (counter body grooving) or by loose particles in the case of sliding load (particle grooving). The process of material damage can be subdivided into micro-plowing, microcutting and microcracks (Fig. 3.35). Micro-plowing and micro-cutting are the predominant wear processes in the case of ductile materials. In the case of microplowing, the material is plastically deformed within the wear groove and pushed towards the groove edges. In the ideal case, no material is removed. This can however occur if the material is repeatedly pushed to the groove edges due to the simultaneous effect of many abrasive particles or due to multiple effect of a single particle, and the material finally fails due to fatigue. Ideal micro-cutting leads to material removal in the form of a chip, the volume of which is equal to the volume of the wear groove created. In the case of brittle materials, microcracks also arise. Material particles are hereby generated by means of cracking and crack growth in the stressed surface. The volume of the wear particle created in micro-cracking is much larger than that of the continuous wear groove [ZumG87, ZumG92, Habi80].

3.7

Wear

73

3.7.1.2 Adhesion Adhesion is defined as the formation of bonds between certain molecules [Erin90, ZumG87]. As a wear mechanism this process is understood as one in which atomic bonds (e.g. in the form of micro-welds) are formed in the contact zone between the material and the cutting tool material. These bonds are then deformed, reinforced and sheared off during a tangential motion of the friction partners. The process of adhesion (i.e. the formation of micro-welds between two friction partners) can come about by atomic interaction (chemical adhesion) between the partners, such as thermally induced diffusion processes, electron exchange or electric polarisation. It can also result from mechanical snagging (mechanical adhesion) of the workpiece material, which has plastically become extremely deformable under high temperatures, with the cutting tool material [Tell93]. By means of adhesion, particles can be transferred from one wear partner to the other. If this transfer is the primary cause of wear, it is sometimes referred to as adhesive wear, although other wear mechanisms are usually also involved in the formation of loose wear particles [Habi80]. Adhesion is a pairing property that depends on the characteristics of the base and counter bodies. Whether a material has the propensity to adhesion can only be answered with reference to the material of the counter body, with which it may either form strong, weak or no adhesion bonds at all [Habi80]. The strength of an adhesion bond is described by the adhesion coefficient. This is defined as the quotient of the normal force FN with which two solid bodies in relative motion are pressed against each other and the opposing force FA that must be applied to undo the bond formed by adhesion [Habi80, Andr59, Siko63]. The propensity of solid bodies to form adhesion bonds can be evaluated with the help of their surface energy. Contact angle measurement is one method for ascertaining the surface energy of solid bodies. [Bobz00] describes a measurement process that makes it possible to compare the surface energies of coating systems with respect to their polar and dispersive parts. Since it is the polar energy share that is primarily responsible for the propensity of a solid body to adhesion, the surface energy of cutting tools (i.e. especially that of the coating system) should have as little polar energy as possible in order to minimize that propensity [Bobz00, Kloc05]. During chip formation, new material surfaces are generated that, without adsorption or reaction boundary layers in statu nascendi, come in a chemically highly active state into contact with the tool surface under high pressures and temperatures, approaching that surface on the atomic level. Plastic deformation processes lead to the formation of large real contact areas. The cutting process thus provides very good opportunity for the appearance of adhesively caused interactions [Bömc89, Neis94, Erin90]. 3.7.1.3 Tribooxidation The term “tribooxidation” refers to chemical reactions of the cutting material and the material with components of the intermediate material or the surrounding medium as a result of activation caused by friction. Tribooxidation changes the properties

74


of the external boundary layer. Reaction products can be formed that are either removed with the chip or remain stuck to the cutting tool material as a coating. Wear can be increased or reduced by this. Whether tribooxidation increases the amount of wear essentially depends on the hardness of the reaction products formed in comparison with the hardness of the cutting material. A decrease is especially possible when the reaction layers prevent direct metallic contact between the base and counter bodies, limiting the effect of adhesion [Erin90, Habi80, Tell93]. 3.7.1.4 Diffusion Diffusion is the thermally activated change of position of individual atoms [Horn67]. This involves a temperature-dependent physicochemical process in which the wear resistance of the cutting tool material can be reduced by foreign substances diffusing with it or its own components diffusing away from it. The diffusion of atoms from the tribologically loaded areas of the tribopartners leads on the one hand to a direct loss of material, which is usually very small but can still certainly be measured as a quantity of wear; however, much more serious is the potential reduction of the wear resistance of cutting tools by the diffusion of certain alloying elements. Diffusion of essential alloying elements can lead to decreased hardness and thus to reduced resistance of the cutting tool material to abrasion [Neis94, Erin90, Habi80]. The high pressures and temperatures in the contact zones characteristic of the cutting process provide an ideal setting for diffusion processes between the material and cutting tool material. Diffusion processes appear especially when high cutting speeds are used, thereby leading to high temperatures in the contact zone. In the area of the contact zone, the cutting tool material and the material approach each other on the atomic level. Furthermore, there is a difference in concentration due to the differing compositions of the material and the cutting tool material which also is preserved in the material because of the constantly added chip [Neis94, Erin90]. 3.7.1.5 Surface Damage Surface damage occurs as a result of tribological alternating stresses. In the stresses surface areas, alternating mechanical stresses lead to structural changes, fatigue, cracking, crack growth and even to separation of wear particles [Habi80]. As opposed to abrasion, in which wear particles can be formed by a single stress process, surface damage is usually preceded by a longer incubation period during which no measurable wear occurs. In this period, the formation of wear particles by means of structural changes as well as cranking and crack growth is prepared [Habi80].

3.7.2 Causes of Wear Friction processes in tool contact zones are comparable to those of dry friction in a vacuum. Together with extraordinarily high mechanical and thermal stresses, the tool is generally worn out quickly [Rabi65, Krag71, Opit70a, Opit70b].

Wear

75

Diffusive wear

Total wear

3.7

Abrasion Adhesion High temp. oxidation

Cutting temperature (cutting velocity; feed et al.)

Fig. 3.36 Causes of wear in cutting processes

According to current information, the following separate causes can be given for the collective term “wear” (Fig. 3.36): • • • • •

mechanical wear (abrasion), the shearing off of adhered material damage to the cutting edge due to mechanical and thermal overstress, diffusion, scaling.

These processes overlap to a large extent and are only partially separable from each other with respect both to their cause and to their effect on wear [Opit67, Köni65, Ehme70b, Ehme70c]. 3.7.2.1 Mechanical Wear (Abrasion) Mechanical wear or abrasive wear occurs at both low and high cutting speeds. Groove wear in the form of counter body or particle grooving can be seen as the dominant form of wear in this category. In the case of counter body grooving, tribological stress on the cutting tool material is based on the abrasive effect of hard particles fixed on the contact surface of the workpiece or the chip. These particles can originate from the workpiece material (oxides, carbides, nitrides) or be transferred by adhesion from the cutting tool material to the workpiece or chip bottom side. In the case of particle grooving, loose particles are the cause of wear. These can arise directly from abrasion (micro-cutting, microcracks) or from surface disruption. However, these may also be adhesion particles or products of tribooxidation removed by abrasion or surface fatigue. Due to the high pressures and temperatures that predominate in the contact zones on both rake and flank faces, we must presume that the loose particles generated are pressed into the softer counter bodies flowing by them, contributing to further wear by counter body grooving.

76


Due to abrasion or surface disruption on the cutting edge or in the contact zones on the rake and flank faces, cutting tool material particles that have broken away from the cutting tool material flow over the rake or flank face under high pressure. This can cause further wear by micro-cutting or microcracks. This process, also called “self-wear” [Ehme70a], is of particular importance especially with respect to the formation and development of flank face wear. 3.7.2.2 Shearing-Off of Adhered Material Particles (Adhesion) In the case of micro-welds being sheared off, material separation can occur in the boundary layer, within one or within both bodies. The term adhesive wear is used as soon as material is separated in the cutting tool material. Adhesion is also responsible for the formation and growth of built-up edges, in which material is transferred from the material to be cut to the cutting tool [Erin90, Habi80, ZumG87]. Ferritic and austenitic steel materials have a high propensity to adhesion with the cutting tool material. The reason for this is above all the high plastic deformability of these materials. The high ductility of ferritic materials is based above all on their relatively low strength, in the case of austenitic steel materials on their face-centred cubic crystal lattice. Tungsten carbide, the basis of hardness and wear resistance in conventional WCCo cemented carbides, has a hexagonal crystal structure. On the other hand, the crystal lattice of the binding metal cobalt is face-centred cubic above 690 K, which is favourable for adhesive processes. The titanium-based coating systems commonly deposited on cemented carbides also have face-centred cubic structures, resulting in a strong tendency to adhesion when machining austenitic steels. The types of wear resulting from this strong adhesive tendency can range from material bonding on the rake and flank faces to de-coating of coated tools in the area of the contact zone. Built-up edges are highly reinforced layers of the machined material that take over the function of the tool cutting edge as bondings on the tool. This is made possible by the property of certain materials to harden during plastic deformation. The material adhering to the cutting edge is deformed by chip pressure, making it very hard. This makes it possible for it to take over the function of a chip-removing tool. Depending on the cutting conditions, built-up edge particles slip periodically between the flank face and the cutting surface. In the case of high hardness and removal frequencies up to about 1.5 kHz, these particles lead to increased flank face wear and considerably deteriorate the surface quality of the workpiece (Fig. 3.37). Since the chip is diverted with the built-up edge and not the rake face, crater wear is usually negligible. Figure 3.38 shows a wear-cutting speed function (VB-vc graph). According to it, flank face wear does not increase with cutting speed continuously, but has at least two distinct extremes [Opit64]. Wear first reaches a maximum at the cutting speed at which the built-up edges reach their largest dimensions. A wear minimum appears at the cutting speed at which no more built-up edges form. Flank face wear decreases after exceeding the maximum, despite higher cutting speeds. This can be ascribed to the fact that reinforcement of the built-up edge

3.7

Wear

77

1

2

3

4

Width of flank wear land VB / mm

Fig. 3.37 Scheme of the periodic growth of build-up edge 0,24 Material: C53E Cutting mat.: HW-P30 Cross section of undeformed chip: ap· f = 2 · 0.315 mm2

0,20 0,16 0,12

vc = 20 m/min

0,08 0,04 0

1

2

3

5

10

20 30 50

100

vc = 30 m/min

Cutting speed vc / (m/min) Geometry of cutting wedge:

αo

γo λs κr

εr

re

8° 10° –4° 90° 60° 1mm Cutting time tc = 30 min

Fig. 3.38 Wear on flank face and growth of build-up edge

begins to decrease as a result of recrystallization processes. It becomes unstable and no longer migrates partially between the cut surface and the flank face, but as a whole across the rake face. The position of the maxima and minima of the VB-vc graph is contingent on temperature. It is shifted to lower cutting speeds by any kind of measures taken

78


to increase the cutting temperature (e.g. higher feed, smaller tool orthogonal rake angle, higher material strength). Measures to reduce the cutting temperature (e.g. cooling) shift the extremes accordingly to higher cutting speeds [Opit69, Peke74]. 3.7.2.3 Mechanical and Thermal Overstress Damage to the cutting edge, such as breaking, parallel cracks, comb cracks or plastic deformation, is the consequence of mechanical or thermal overstress. Cutting Edge Chipping Large cutting forces lead easily to breaking on the cutting edge (cutting edge chipping) or corner if the wedge angle or tool included angles of the tool are too small or the cutting tool material used is too brittle. In the case of such breaking, the profile of the fracture surface is determined by the direction of the cutting force [Köni75]. Cut interruptions can also lead to breaking, especially when tough materials are being cut, the chips of which tend to stick. Small amounts of breaking occur when the workpieces contain hard non-metallic inclusions that originate during deoxidation of the steel [Opit64a, Opit66, Opit62]. Sintered oxides and more wear-resistant types of cemented carbide are sensitive to this type of local overstress, especially in the case of manufacturing processes with relatively small cross-sections of undeformed chip (e.g. reaming or shaving). Distinct breaking of cutting tool material on the major and/or minor cutting edge can also be caused by chips striking against the cutting edge or during the rotation of shaft sections because of chip jamming between the cutting insert and the workpiece. Breakage of the cutting edge can occur both on the top and bottom sides of the insert. Parallel Cracks In the case of interrupted cuts (e.g. milling), the cutting tool material is subject to strong mechanical alternating stress. This dynamic pressure threshold load can lead to fatigue failure. Quick consecutive cutting force changes lead, especially in the case of milling with cemented carbide tools, to parallel cracks (Fig. 3.39). The quickly alternating stress in the formation of lamellar chips can also lead to parallel crack formation if a critical stress cycle is exceeded [Domk74, Beck69], e.g. when cutting titanium materials. Comb Cracks Comb cracks are a form of damage to the cutting edge as a result of thermal alternating stresses (Fig. 3.39). Such stresses originate mainly while working with interrupted cuts. During tool engagement, the cutting edge quickly heats up to high temperatures. It cools down after the tool exits the workpiece. The difference between the highest and the lowest temperature depends, among other things, on the material, cutting

3.7 Di

Wear

79

cti

on

of

co

m

b

cr

ac

ks


re

Comb and parallel cracks

ce

e fa

Rak

KM vc = 200 m/min

GJS70

llel

Comb cracks

f para tion o Direc cracks

VB

e

k fac

Flan

42CrMo4+QT

Temperature

Stress Tensile + 0 – Compressive

vc = 275 m/min

Stress Temperature Tensile + 0 – Compressive

–y

Heating during cutting process

Cooling

Fig. 3.39 Comb and parallel crack formation in milling, acc. to VIEREGGE [Vier59]

conditions and the ratio of the paths covered in the material and in the air. In the case of interrupted cuts, the use of cutting fluids is particularly important with respect to the size of the temperature difference because it has a much larger quenching effect than air. Cooling is advantageous to the formation of comb cracks in the case of cemented carbides and ceramic cutting tool materials. The profile of the comb cracks is aligned with that of the isotherms of the temperature field in the cutting part. The comb cracks arising in the cutting part during milling or also in short-cycle turning are generally temperature change cracks. These should be distinguished from thermoshock cracks such as develop during turning or milling while using cutting fluid. Thermoshock cracks are formed by a nonrecurring steep temperature change, while temperature change cracks form only gradually in the cutting part in the course of the cutting edge engagement. While in the case of thermoshock the thermally caused tensile strain arising in the contact zone exceeds the strength of the cutting tool material and leads directly to cracking, the stresses leading to the formation of temperature change cracks are below the strength of the cutting tool material. The development of temperature change cracks thus requires several periodic temperature changes as a rule.

80


Due to the high energy level thermoshock cracks propagate with greater speed. Crack growth is largely transcrystalline. Depending on the stress level, they can acquire very large depths in the cutting tool material (> 1 mm), thus leading to considerable weakening of the cutting edge [Gers98]. Temperature change cracks can be observed in both wet and dry cuts. Besides the amount of thermal load, the cutting cycle and the number of cutting cycles play a decisive role in their formation. The use of cutting fluids leads to a strong reduction of the amount of cutting cycles required for cracking [Gers98]. Depending on the bending stress which the cutting edge is subject to during the cutting process and which are superimposed over thermally induced stresses, the cracks become larger with increasing numbers of cutting cycles. With a corresponding output length, they can relatively quickly achieve a critical size leading to fracture. As a result of the notching effect coming from the crack, they can also become the source of further cracks [Gers98]. Plastic Deformation Plastic deformation of the cutting edge arises when the thermomechanical stress effecting the tool cutting edge exceeds the deformation resistance of the cutting tool material. The influencing parameters are the strength of the material to be cut, the cutting parameters, the geometry of the insert and, on the part of the cutting tool material, its high temperature properties like hardness, compressive strength and creep behaviour. Cutting edges made of tool steel or high speed steel deform the stronger the smaller the difference is between the temperature of the cutting edge and the annealing temperature of the cutting tool material. Plastic deformations also appear in the case of cemented carbides and cermet, however only at higher temperatures (cutting speeds) and under higher forces as is the case for tool or high speed steels. Cemented carbides deform more the more binder phase there is, usually cobalt. Plastic deformation of the cutting edge results in a considerable increase in wear and can lead to sudden disruption of the tool by cracking and shearing off of the cutting edge. It thus limits above all the cross-section of undeformed chip and cutting speeds applicable in rough turning. 3.7.2.4 Diffusion In the case of heat wear resistant cemented carbide tools, diffusion wear must be expected at high cutting speeds and mutual solubility of the partners. Tool steel and high speed steel already become soft at temperatures at which diffusion can hardly manifest itself (e.g. about 600◦ C for high speed steel). When cutting with uncoated cemented carbide substrates, the following reactions can occur (Fig. 3.40): • diffusion of Fe into the binder phase Co, • diffusion of Co into the steel, whereby Fe and Co form a gapless series of mixed crystals,

3.7

Wear

81

(Ti, Ta, Nb, W) C

Co-W-C-mixedcrystal Fe WC: Dissolution into: Fe3W3C; (FeW)6C; (FeW)23C6

C Co

Tool

Chip

Cutting material Cemented carbide WC-Co W, Co

50 Mass percent %

Material X6CrNiMoTi17−12−2

W

40

Fe, Cr, Ni

30

Fe

20 Co 10 0 µm

5

0

5

Fig. 3.40 Presentation of the diffusion phenomena in cemented carbide tools

• dissolution of tungsten carbide to form mixed and double carbides in the form of Fe3 W3 C, (FeW)6 C and (FeW)23 C6 . The carbon released during the dissolution of tungsten carbide migrates in the direction of lesser concentration, i.e. into the steel. The diffusion of carbon passes into the cobalt phase. The maximum solubility of carbon in cobalt is about 0.7% at 1200◦ C. In the presence of Fe the solubility is increased to 1.5–2%. The diffusing iron thus introduces two reactions, both of which accelerate dissolution. It lends itself to the formation of iron mixed carbides and increases the reception of carbon in cobalt, which is in turn the prerequisite for the dissolution of tungsten carbide. The diffusion processes that are possible between cemented carbide and steel materials can be shown very effectively with the help of annealing tests. To this purpose, cemented carbide-material pairs were annealed at a temperature of 1100◦ C, a pressure of 12.6 MPa and a holding time of 2 h in a protective gas atmosphere of argon. Then the samples that were fused together on the contact surface were separated on an abrasive cut-off machine perpendicularly to the contact level, and

82


after these sample halves were grinded and buffed, the distribution of elements was analysed with a microprobe perpendicularly to the contact surface cutting tool material/material. As the quantitative line scans in Fig. 3.40 show, diffusion phenomena take place under the selected annealing parameters between the uncoated cemented carbide and the austenitic steel material here selected. On the side of the steel material, there is a diffusion of iron, chrome and nickel into the cutting tool material. In the opposite direction (i.e. from the cutting material into the material) tungsten and cobalt diffuse. The dissolution of tungsten carbide, the bearer of hardness and wear resistance, caused by the diffusion of iron into the cemented carbide leads to a weakening of the structure and lowers the resistance of the cemented carbide substrate against abrasion. The consequence of this is the formation of distinct crater wear (Fig. 3.41). Since diffusion phenomena depend on temperature, the cutting speed influences on an elementary level the speed with which crater wear develops. Uncoated cemented carbides can thus only be used at relatively low cutting speeds (vc < 100 m/ min). The cemented carbide substrate also has a serious effect on crater formation. Mixed carbides on the basis of titanium and tantalum/niobium have, in comparison to WC, a much higher chemical stability towards iron. Mixed carbides reduce the diffusion of iron into the cobalt binder phase, thus increasing the resistance of the cutting tool material to crater wear and increasing its heat wear resistance quite substantially. The most effective way to reduce crater wear is coating the cemented carbide substrate. Cemented carbide layers seal the substrate surface. This prevents direct contact between the substrate and the material, thus reducing or suppressing diffusion phenomena. As the line scans of annealing tests in Fig. 3.42 show, diffusion processes between the cemented carbide substrate and the material can by all means still occur in the case of TiN or TiCN layers. In contrast, no more diffusion takes place in the case of the coated cemented carbide with Al2 O3 as an intermediate layer. The aluminium oxide layer thus functions as a diffusion barrier. The elements iron, nickel and chrome are found in the ZrN surface layer, while tungsten and cobalt are in the TiN boundary layer (Fig. 3.42). In the aluminium oxide intermediate layer Uncoated cemented carbide, tc = 40s Coated cemented carbide, tc = 20 min

vc = 120 m/min, f = 0.32 mm, ap = 2 mm vc = 200 m/min, f = 0.2 mm, ap = 2 mm Material: X6CrNiMoTi17–12–2

Fig. 3.41 Formation of crater wear on an uncoated and a coated cemented carbide, acc. to GERSCHWILER [Gers04]

3.7

Wear

83

WC-7,5Co- 6,8MK

TiN

Ti (C,N)

TiN

X6CrNiMoTi17-12-2

W Co Fe Cr Ni WC-7,5Co6,8MK

W Co

TiN

Ti (C,N)

Al2O3 ZrN

X6CrNiMoTi 17-12-2

Fe Cr Ni

Fig. 3.42 Verification of the diffusion phenomena between coated cemented carbide and steel dependant on the coating system on the basis of annealing tests

the traceable amount of these elements is reduced to 0 mass-%. There is above all no indication that iron, chrome, tungsten or cobalt have diffused through the Al2 O3 intermediate layer on the opposite side of that layer. The observation that, excluding air or water, electrically isolated oxide layers function as diffusion barriers on cemented carbides as opposed to semiconducting nitride or carbon nitride layers can be explained as follows. In electrically isolating oxides like Al2 O3 or HfO2 , only ion diffusion is possible, no migration of electrons or holes. The result of this is that “component diffusion” becomes impossible. “Component” refers to an atom or iron, which theoretically can be split into an ion and an electron. For reasons of electron neutrality, the iron atom cannot “release” its electrons outside the layer and then migrate through the layer as an ion. The consequence is that these oxide layers act as barriers as long as there is no way for electrons to get around it, and the transport of oxygen ions is also slow enough (in oxides this transport is empirically slower than cation transport by several orders of magnitude.) In semiconductive or conductive layers however, electrons can migrate through the layer together with the ions, leading finally to the iron element migrating through the layer to alloy with the substrate. Conversely, conveyance of Co or W into the chip is also possible. Diffusive transport is already possible at the temperatures prevalent on the tool during machining of > 800◦ C. The diffusion coefficients are indeed still small, but the necessary diffusion paths are very short. While diffusion of iron through the cemented carbide layer into the cemented carbide substrate is indeed possible in the case of cemented carbides coated with TiN and TiCN, this diffusion is not as significant as is the case for uncoated substrates. As opposed to uncoated cemented carbides, the coat provides the substrate

84


with additional protection from the abrasive action of the flowing chip. Especially because they are much harder than the substrate, cemented carbide layers help to reduce abrasive wear to a significant extent. Cemented carbide layers thereby ensure the high performance level of coated tools and allow for much higher cutting speeds when cutting steel materials in comparison with uncoated cemented carbides (Fig. 3.41). But as soon as the hard material layer in the contact zone of the rake face has been rubbed off and the substrate exposed, diffusion processes now once again taking place between the substrate and the chip contribute, together with the abrasive effect of the chip on the substrate, to an acceleration of crater wear. 3.7.2.5 Scaling If we examine a tool after cutting, we can recognize several annealing colours near the contact zones, indicating scaling (oxidation) of the cutting tool material. Scaling varies in importance depending on the cutting tool material alloy and cutting temperature (Fig. 3.43). Cemented carbides already begin scaling at 700–800◦ C, whereby

Increase in mass / %

20

15

HW-K10 Atmosphere: air Temperature: 1000 °C

HW-P25

Annealing temperature: 500 °C Annealing time: 3h exp. to air

10 HW-P10 5 HT-P05 HT-P10 HT-P20

0 1

2

3

4

HW-K10

5

Annealing time / h Annealing temperature: 1000 °C; Annealing time: 3 h exp. to air

HW-K10

HW-P25

HW-P10

HT-P10

Fig. 3.43 Gain of weight during the annealing process in air dependant on the used cutting tool material

3.7

Wear

85

Oxidation

Material: C53E Cutting tool material: HW-P30 Cutting velocity: vc = 125 m/min tc = 20 min Cutting time: Feed: f = 0.25 mm Depth of cut: ap = 3 mm

Rake face

Oxidation zones

Flank face

Fig. 3.44 Oxidation zones at a cemented carbide turning tool

those made of pure tungsten carbide and cobalt oxidize more than those alloyed with titanium carbide or other carbides [Kief65]. Even under the usual cutting conditions, an oxide film is formed on WC tools in the area of the cutting edge under the effect of the cutting temperatures and atmospheric oxygen. This film covers the areas to which atmospheric oxygen has free access, i.e. the ends of the contact zones on the flank face, minor flank face and rake face (Fig. 3.44). The destructive effect of oxidation on the structure of cemented carbide can be observed especially clearly on the minor flank face. A complex tungstencobalt-iron oxide is formed that, as a result of its larger molar volume compared with cemented carbide, develops in a wart-like fashion and can lead to corner breaking [Köni75]. Scaling is of no practical importance for tool steels and high speed steels since their heat resistance is exceeded before their surfaces are oxidized more strongly.

3.7.3 Forms and Dimensions of Wear Wear phenomena that are generated on the cutting part during cutting vary greatly depending on the type and duration of stress. Figure 3.45 shows wear forms that occur primarily on turning tools. The cutting part is worn on the rake face (crater wear) as well as on the major and minor flank faces (flank face wear) (Fig. 3.46). Depending on the cutting parameters and the combination of workpiece and cutting tool material, flank face wear tends to be strongest on the edge of the contact zone on the major and minor cutting edge of the tool, which continues on the rake face. Such “notch wear” is caused by the abrupt transitions of mechanical and thermal stress prevalent on the end of the contact area, the abrasive character of the sharp

86


Cratering on rake face Smeared material, partly overlapping the cutting edge in a palisade-like manner Spalling of cutting edge along major cutting edge Notch wear on minor cutting edge Notch wear on rake face Plastic deformation of corner of cutting edge to the point of breaking

Spalling of cutting edge due to chip impact or clamped chips

Layer flaking due to plastic deformation of cutting edge

Layer flaking beneath width of flank wear land

Notch wear on flank face

Wear on flank and minor flank face

Delamination on cutting edge beyond contact zone

Fig. 3.45 Characteristic wear forms at the cutting part during the turning process Cutting tool material: Cermet Material: 42CrMo4+QT Cutting parameters: vc = 400 m/min f = 0.1 mm tc = 12 min ap = 0.5 mm Insert geometry: SPGN120308 Cutting tool material: PCBN Material: Nickel-based alloy Cutting parameters: vc = 300 m/min f = 0.16 mm tc = 1.6 min ap = 0.3 mm Insert geometry: VBMW160412 Notch wear on: Minor cutting edge

Major cutting edge

Cutting tool material: Ceramic Material: Nickel-based alloy Cutting parameters: vc = 220 m/min f = 0.16 mm tc = 2.3 min ap = 0.3 mm Insert geometry: RCGX090700T00515

Fig. 3.46 Examples for the formation of crater wear, flank face wear and notch wear dependent on the used cutting tool material and the machined material

3.8

Cutting Theory

87

SVα: Displacement of cutting edge toward flank face SVγ: Displacement of cutting edge toward rake face

rε C

VBN b/4 B

A

N Wear notches on major cutting edge

b

Sectional view A-A KT

KB

VBB

VBB max.

SVα

KB: Crater width KM: Distance betw. crater centers KT: Crater depth

VBC

Width of flank wear land VB

Wear notches on minor cutting edge A

KM

crater

SV

A

γ

Plane Ps SV

α

Fig. 3.47 Wear forms and measured quantities at the cutting part, acc. to DIN ISO 3685

chip edge or the feed comb and the tool edge as well as direct contact with the atmosphere (Fig. 3.46). Figure 3.47 is a schematic representation of the dimensions of wear. In particular, we distinguish the width of flank wear land VB, the displacement of cutting edge toward flank face SVα and rake face SVγ , crater depth KT and the crater centre distance KM from which the crater ratio K = KT/KM is formed.

3.8 Cutting Theory 3.8.1 Shear Plane Theory The model representation of the shear plane is based on ideal cutting conditions. Required are: • • • •

ideally sharp cutting edges, cutting edges that are not subject to wear, the representation of mechanical loads as substitute forces and ideal plastic material behaviour.

88

3 Fundamentals of Cutting Trace of tool reference plane Pr γn

Trace of shear plane PΦ

hch

h

Tool

Φ

Trace of cutting edge plane Ps

Workpiece Selected point of cutting edge

z

Po ≡ Pf ≡ Pn

y Fig. 3.48 Model representation of the shear plane

The chip formation theory derived from this model (shear plane theory) assumes that deformation takes place on only one plane, the shear plane. By definition, the shear plane passes through the selected point of cutting edge just like the tool-inhand and tool-in-use reference system (Fig. 3.48). In the following the shear plane P will be considered in the tool-in-hand reference system. This is a plane which is inclined toward the cutting edge plane Ps by the shear angle F. The requirements made on process kinematics are fulfilled when: • tool cutting edge angle κr = 90◦ • tool inclination λs = 0◦ We are thus assuming a free, orthogonal cut. The shear plane model has the advantage of making complicated cutting processes highly comprehensible. If the effects of the corner radius and minor cutting edge are negligible in comparison with that of the major cutting edge, the theory of the shear plane can also be used for the bound cut. With the shear angle we can macroscopically consider material and cutting tool material properties as well as thermal loads and friction conditions which predominate during contact of the chip with the rake face. Moreover, the shear angle establishes the relation between undeformed chip thickness h and chip thickness hch . hch =

cos ( − γn ) ·h sin

(3.14)

3.8

Cutting Theory

89

By means of the shear angle we can now calculate the surface area of the shear area A in the shear plane: h A = b · (3.15) sin The ratio of undeformed to deformed magnitudes is designated in forming technology as deformation. If the material is compressed and only the change in height is described, the term degree of compression is also used. This concerns volume consistency. In cutting theory, the term chip compression λh is used analogously [Krys39]. hch cos ( − γn ) λh = = (3.16) h sin The shear angle is a purely theoretical quantity and should not be confused with the actual slip direction of the material, even if the shear surface and slip surface are nearly identical under certain marginal conditions. In conclusion, to show that the theory of the shear plane makes it possible to calculate various quantities with complicated relations, three classical applications of shear plane theory will be described. 3.8.1.1 Calculation of Chip Speed From simple trigonometric relations (law of sine) in the velocity plan, MERCHANT calculated the chip speed vch with the help of the shear angle F (Fig. 3.49) [Merc45]. | vch | =

sin · | vc | cos (γn − )

(3.17)

Trace of tool refernce plane Pr

γn Trace of shear plane PΦ

hch

h

Tool

Φ Trace of cutting edge plane Ps

vch

Workpiece z

vΦ

vc Selected point of cutting edge

y

Fig. 3.49 Velocity plan, acc. to MERCHANT [Merc45]

Po ≡ Pf ≡ Pn

90


The continuity equation can also be used to calculate chip speed vch . The idea behind this is that a mass flow flows through a control surface, in this case the shear surface. Assuming that the density of the material is not altered by mechanical and thermal load, we obtain: m ˙ = const. ⇒ | vc | · A = | vch | · Ach

(3.18)

From which follows: | vch | =

h 1 · | vc | = · | vc | hch λh

(3.19)

3.8.1.2 Calculation of Resultant Force It is assumed that the shear angle is a function of the shear stress in the shear plane, which will be shown in the following steps. In the first place we have: τ =

| |F A

(3.20)

The component of the resultant force in the shear plane, the shear force F , can be represented by the resulting resultant force Fz . With the help of the angle relations in the circle of THALES we obtain according to Fig. 3.50:

z | · cos ( + ρ − γn )

| = |F |F

(3.21)

Trace of tool reference plane Pr γn

Trace of shear plane PΦ Tool

Φ


FΦ

Fc

Ff

z

ρ − γn

FΦn

Fz

Fγn

Fγ Po ≡ Pf ≡ Pn

y

Fig. 3.50 Circle of THALES for the force diagram, acc. to MERCHANT

3.8

Cutting Theory

91

The amount of resultant force can now be calculated from Eqs. (3.20) and (3.21):

z| = |F

τ ·b·h sin · cos ( + ρ − γn )

(3.22)

The functional relations of the values of the resultant force components cutting force and feed force can be taken from the circle of Thales (Fig. 3.50).

c | = |F

z | · cos (ρ − γn ) |F

(3.23)

z | · sin (ρ − γn )

f | = |F |F

(3.24)

If we now insert Eq. (3.22) into Eqs. (3.23) and (3.24), the following is valid for the values of the resultant force components: cos (ρ − γn ) · τ · b · h sin · cos ( + ρ − γn ) sin (ρ − γn ) · τ

f| = ·b ·h |F sin · cos ( + ρ − γn )

c| = |F

(3.25) (3.26)

3.8.1.3 Calculation of the Shear Angle KRYSTOF developed a model for the simple calculation of the shear angle [Krys39]. He made the following assumptions: • The theory of the shear plane is valid. • The maximum shear stress leads to material collapse. In the main axial system, maximum shear stresses occur at an angle of 45◦ . For calculations of the cutting process, the position of the main axial system is assumed to be approximately in the direction of the resulting contact stress (Fig. 3.51). For purposes of simplicity, the principle stress is not represented by the stress tensor but simply by vector addition. In this case, the friction angle approximates the orientation of the main axial system. The function relation between the shear angle, the tool orthogonal rake angle and the friction angle can be derived from Fig. 3.51. The following is valid for the shear angle: π (3.27) = − (ρ − γn ) 4 MERCHANT has suggested a possibility of calculating the shear angle that makes the following assumptions [Merc45, Merc45a]: • The theory of the shear plane is valid. • The position of the shear plane is determined by the minimum of cutting energy. The value of cutting force can be read off the circle of T HALES in Fig. 3.50:

92

3 Fundamentals of Cutting Trace of tool reference plane Pr

1

γn Trace of shear plane PΦ Tool

Φ

τΦ


−

π 4 Fz

ρ

Fγ

3 Fγn

Po ≡ Pf ≡ Pn

Fig. 3.51 Principle of the maximum shear stress, acc. to KRYSTOF [Krys39]

c| = |F

b·h cos (ρ − γn ) · · τ cos ( + ρ − γn ) sin

(3.28)

The necessary and the sufficient conditions for calculating the position of the energy minimum are nec.:

∂Ec =0 ∂

suff.:

∂ 2 Ec = 0 ∂

(3.29)

Since the cutting energy is calculated as the product of the cutting force and the cutting path, and the cutting path is not a function of the shear angle, we can simplify to: nec.:

c| lc · ∂|F =0 ∂

suff.:

c| ∂ 2 |F = 0 ∂

(3.30)

From which follows for the shear angle: =

π 1 − · (ρ − γn ) 4 2

(3.31)

3.8.2 Application of Plasticity Theory to Cutting HUCKS and OPITZ developed a shear angle equation by applying MOHR’s slip theory to the orthogonal cutting process [Huck51, Opit53, Mohr06]. As opposed to the models of K RYSTOF and M ERCHANT, this theory makes it possible to take a material-specific slip angle into consideration instead of an idealized 45◦ . It should

3.8

Cutting Theory

93 Trace of tool reference plane Pr 1

Trace of shear plane PΦ

Sli

γn

pd

zMT yMT

zFE yFE

ω

cti

on

ϕ

Φ1

τzy

τyz

σy

Tool τyz

σρ y y

τzy

3

Φ3

ire

Po ≡ Pf ≡ Pn

Sli

pd

1,3: Principle normal stress directions

cti o

n

Trace of cutting edge plane Ps ≡ Pp

ire

Fig. 3.52 Surface element at the cutting part, acc. to HUCKS [Huck51]

again be pointed out that the previously known shear angle need not necessarily be equal to the angle given by the actual position of the slip surface, since the direction of the shear plane was previously only purely geometrically determined by ideal considerations. In this case the directional angle of the slip surface towards the tool cutting edge plane is designated as the shear angle . First a square surface element is considered on the rake face that experiences both shear and compressive stresses from the tool (Fig. 3.52). The shear stresses also appear on all remaining faces of the quadratic element as a result of the law of equality of shear stress. Due to the unhindered flow of the chip over the face there is a very small amount of compressive stress in this direction, which will be neglected in the following. There are therefore two compressive stresses and four shear stresses on the surface element under consideration. If we now determine the principle normal stress with the help of MOHR’s stress circle, we establish the validity of the following equation: ω=

1 arctan(2 · μ) 2

(3.32)

The angle between the rake face and the principle stress direction is thus determined by the friction coefficient between the rake face and the chip. This relation makes it clear, for example, why different cutting speeds yield different shear angles. The friction coefficient is dependent on the cutting speed and consequently alters the principle stress direction. The slip direction on the other hand is inclined towards the principle stress directions by a materially dependent angle ϕ. This angle can be

94


calculated from the compressive flow stress σ D and the shear flow stress τ F with the help of flow tests and MOHR’s stress circle: 1 σD − 2τF ϕ = 45 − arcsin 2 σD

(3.33)

For an ideal plastic body which is approximately realized by steel at the yielding point, the flow plane is under 45◦ in relation to the principle stress direction, since there the shear stress is at a maximum [Huck51]. With this, we obtain the principle stress and slip directions shown in Fig. 3.52 relative to the rake face. With the help of the angles γ , ϕ, and ω, the slip angle 1 can now be calculated: 1 1 = ϕ − arctan(2μ) + γn (3.34) 2 Slip angle 3 is in the still undeformed area, which is why only slip angle 1 (which is identical with the shear angle) is significant. The shear angle calculated here presumes an extrapolation of the stress conditions via the shear plane and the width of flank wear land as well as the formation of a continuous chip. This approximation was proven to be in accordance with experiments made on materials with envelope lines that are inclined and parallel to the σ -axis. HUCKS did not experiment with materials with bent envelope curve [Huck51]. With the help of the shear angle, the resultant forces can now also be calculated. To do this, the stresses that are active in the shear plane are multiplied with shear plane surface, resulting in a normal and tangential force in the shear plane. We will in this context proceed without the derivation of HUCKS’s formulae. For steel, which has an envelope line that is parallel to the σ -axis, the following formulae for cutting and feed force result with width of undeformed chip b and chip thickness h according to HUCKS: ⎡ ⎤ 1 Fc = τ 0 · b · h · ⎣ + cot g(φ)⎦ (3.35) 2 μ2 + 14 ⎡

⎤ cot g(φ) − 1⎦ Ff = τ0 · b · h · ⎣ 2 μ2 + 14

(3.36)

To sum up, we can say that the highly numerous experimental relations in chip formation depend only on a few numbers of materials and the friction coefficient. HUCKS’s work made the fundamentally new discovery that the stress field in the chip on the cutting edge depends, besides its intensity, only on the friction coefficient and the normal direction of the envelope line [Huck51].

Chapter 4

Cutting Tool Materials and Tools

Tool change times, and with them both manufacturing times and tool, machine and labour costs, are affected by wear. Wear is affected in turn by the properties of the cutting tool materials. Development in the cutting tool material sector is therefore far from finished, but is constantly aiming both to improve cutting tool materials that are already established as well as to discover new materials for use in the manufacture of cutting tools. Cutting tool materials should have the following properties in order to do justice to the stresses placed on them: • • • • • • • • •

hardness and pressure resistance, bending strength and toughness, edge strength, inner bonding strength, high temperature strength, oxidation resistance, small propensity to diffusion and adhesion, abrasion resistance, reproducible wear behaviour.

Should we consider all these characteristics simultaneously, we are then faced with the requirement for the “ideal” cutting tool material (Fig. 4.1). However, there is no one cutting tool material that unifies all the required properties. One reason for this, for example, is the physical opposition of hardness and toughness. Developments in the cutting tool material sector are concentrating on optimizing and modifying chemical composition, manufacturing methods, geometry and coating in order to broaden the areas of application of cutting tool materials and tools in accordance with the requirements of modern production.

4.1 Overview of Cutting Tool Materials The cutting tool materials used in cutting can be summarized as (Fig. 4.2): • tool steels, • cemented carbides, F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_4,

95

4 Cutting Tool Materials and Tools

m

at

er

1

BN

tin

g

to

ol

DP

ea

lc

ut

2 Id

Cutting speed Wear-resistance and heat-resistance

ia l

96

Al2O3 ceramic Coated cermet Al2O3 + TiC

Si3N4 ceramic

Cermet

Coated cementedcarbide

Cemented-carbide on tungsten-carbide-basis

Fine grain and submicron grain cemented carbide Coated HSS

DP: 1 Considering the hardness 2 Considering the hardness and the temperature

HSS

Durability and bending-strength Feed

Fig. 4.1 Schematic classification of several cutting tool materials

Cutting tool materials for chipping with geometrical defined cutting edge Cold work steels Tool steels High speed steels

Cemented carbide

WC-Co WC-(Ti,Ta,Nb)C-Co TiC/TiN - Co,Ni

Cutting - ceramics

Oxide-ceramic Mixed-ceramic Silicon-nitride-ceramic

Super-hard cutting tool materials

Boron-nitride Diamond

Fig. 4.2 Classification of cutting tool materials for machining

4.1

Overview of Cutting Tool Materials

97

• cutting ceramics and • super-hard cutting tool materials made of boron nitride and diamond. The hardness and wear resistance of the cutting tool materials increases in this order. On the other hand, their bending strength decreases in this order. Tool steels include unalloyed and alloyed cold work tool steels as well as high speed steel. The term cemented carbide refers to conventional cemented carbides based on tungsten carbide as well as “cermets” based on titanium boron nitride. Ceramics is the generic term for cutting tool materials made of oxide, mixed and non-oxide ceramic.

4.1.1 Classification of Hard Cutting Tool Materials According to DIN ISO 513, the notation and application of hard cutting tool materials of cemented carbide, ceramic, diamond and boron nitride are classified in accordance with Fig. 4.3. According to it, uncoated cemented carbides based on tungsten carbide with a WC grain size of ≥ 1 μm take the abbreviation HW and with grain sizes of < 1 μm the notation HF. “Cermets” – cemented carbides based on titanium boron nitride – are notated with HT, coated cemented carbides and coated cermets with HC. Corresponding abbreviations from Fig. 4.3 are valid for other hard cutting tool materials based on ceramic, diamond or boron nitride. The goal of DIN ISO 513, besides that of identifying cutting tool materials, is above all their assignment to materials, for which they are most suitable for cutting. Expanding earlier norms, DIN ISO 513 provides six main application groups and thus six classes of workpiece-materials, which are classified with the code letters P, M, K, N, S and H as well as by colour (Fig. 4.4). DIN ISO 513 thus maintains the established code letters P, M and K but now only designate with them the main application groups that comprise the material groups steel (P), stainless steel (M), and cast iron (K). The major machining group K included according to the old standard not only cast iron materials but also a number of other materials. In DIN ISO 513, these now new major application groups were provided with the labels N for nonferrous metals, S for special alloys and H for hard materials. Every main application group is subdivided into particular application groups (Fig. 4.4). These are notated with a code letter for the main application group they belong to and with an index. The index refers to the toughness and wear resistance of the cutting tool material. The higher it is within each application group, the lower is the material’s wear resistance and the higher its toughness. The indices are merely reference numbers indicating a certain sequence. They provide no information about the extent of the wear resistance or the toughness of a cutting tool material. Manufacturers of cutting tool materials should assign their cutting tool materials to the appropriate application group, depending on wear resistance and toughness. Examples are HW-P10, HC-K20, and CA-K10. An application group thus contains comparable cutting tool materials of different manufacturers, although they may differ in wear behaviour and performance. It is also possible for a cutting tool material from one manufacturer to be categorized under multiple application groups if suitable to them.

98

4 Cutting Tool Materials and Tools Cemented carbide Code-letter Group of materials

HW

Uncoated cemented carbide, predominantly made from tungsten-carbide (WC) with a grain size = 1 µm

HF

Uncoated cemented carbide, predominantly made from tungsten-carbide (WC) with a grain size < 1 µm

HT

1)

HC

Uncoated cemented carbide, predominant titan-carbide (TiC) or titan-nitride (TiN) or both Cemented carbide like above, but coated

1)

These cemented carbides are also called "Cermet"

Ceramic Code-letter CA

Oxide-ceramic, predominantly made from aluminium-oxide (Al2O3)

CM

Mixed-ceramic, based on aluminium-oxide (Al2O3), but also with other components than oxides

CN

Silicon-nitride-ceramic, predominantly made from silicon-nitride (Si3N4)

CR

Cutting-ceramic, predominantly made of Aluminium-oxide (Al2O3), reinforced

CC

Cutting-ceramic like above, but coated

Diamond Code-letter

Group of materials

DP

Polycrystalline diamond

DM

Monocrystalline diamond

Boron-nitride Code-letter Group of materials

BL

Cubiccrystalline boron-nitride with a low content of boron-nitride

BH

Cubiccrystalline boron-nitride with a high content of boron-nitride

BC

Cubiccrystalline boron-nitride like above, but coated

Diamond Code-letter

Diamond category

DP

Polycrystalline diamond

DM

Monocrystalline diamond

Group of materials

Boron-nitride Code-letter

Boron-nitride category

BL

Cubiccrystalline boron-nitride with a low content of boron-nitride

BH

Cubiccrystalline boron-nitride with a high content of boron-nitride

BC

Cubiccrystalline boron-nitride like above, but coated

Fig. 4.3 Notation of hard cutting tool materials for machining, acc. to DIN ISO 513

4.2

Tool Steels

99

Main application group Code-letter

P

Code-color

Workpiece-material

blue

Steel: All kinds of steel and cast-steel,exept stainless steel witch austenitic structure

Application group Hard cutting tool materials P01 P10 P20 P30 P40 P50 M01 M10 M20 M30 M40

P05 P15 P25 P35 P45

yellow

Stainless steel: Stainless austenitic and austenitic-ferritic steel and cast-steel

red

Cast-iron: Cast-iron with flake– graphite, cast-iron with ductile graphite, annealed cast-iron

K01 K10 K20 K30 K40

K05 K15 K25 K35

green

Nonferrous metals: Aluminium and other nonferrous metals, nonmetal materials

N01 N10 N20 N30

N05 N15 N25

S

brown

Specializations and titanium: Highly heat resisting special-alloys based on iron, nickel and cobalt, titanium and titanium based alloys

S01 S10 S20 S30

S05 S15 S25

H

grey

Hard materials: Hardened steel, hardened cast-iron-materials, castiron for chill-casting

H01 H10 H20 H30

H05 H15 H25

M

K

N

M05 M15 M25 M35

Increasing cutting speed, increasing wear-resistance of the cutting tool materials Increasing feed, increasing durability of the cutting tool materials

Fig. 4.4 Application and classification of hard cutting tool materials, acc. to DIN ISO 513

4.2 Tool Steels DIN EN ISO 4957 defines tool steels as high-grade steels suited to machining and processing materials as well as handling and measuring workpieces. We distinguish between cold work steels, hot work steels and high speed steels. Cold and hot work steels are suited to purposes in which the surface temperature during engagement

100


is generally under 200◦ C. High speed steels on the other hand can be applied at temperatures of up to 600◦ C. Tools for both cutting and forming are made out of cold work steels and high speed steels. Hot work steels are used primarily for forming tools such as forging dies, pressure casting dies or tools for forging machines and extrusion presses.

4.2.1 Cold Work Steels Cold work steels obtain their wear and toughness properties from heat treatment, consisting of heating to austenitizing temperature, quenching in an oil or water bath (a high cooling speed is necessary, martensite hardness) and tempering (with the goal of a partial loss of hardness in order to increase toughness). Cold work steels can be classified as unalloyed and alloyed steels (Fig. 4.5) [DIN EN ISO 4957, Wegs95]. Unalloyed tool steels (carbon steels) contain up to 1.25% C

Unalloyed cold work steels Material grade Material number

C

Si

Composition Mn Cr Mo

V

W

Examples

C45W 1.1730

0.40 0.15 0.60 – – – 0.50 0.40 0.80

Hammers, hatchets, axes, scissors, screwdrivers, bits

C85W 1.1830

0.80 0.25 0.50 – – – 0.90 0.40 0.70

Wood saws, hand saws, steel blades for segmentbuzz saws

C125W 1.1563

1.20 0.10 0.10 – – – 1.35 0.30 0.35

Files, ductors, chisels, paper knifes

Alloyed cold work steels 45CrMoV7 1.2328

0.42 0.20 0.85 – – – 0.47 0.30 1.0

1.7 – 1.9

115CrV3 1.2210

1.10 0.15 0.20 – – – 1.25 0.30 0.40

0.5 – 0.8

X210CrW12 1.2436

2.0 0.10 0.15 11.0 – – – – 2.25 0.40 0.45 12.0

Chisel of all kinds

0.25 – 0.30 0.05 – –

0.07 – 0.12

–

–

–

–

Files, ductors, chisels, Paper knifes

0.6 – 0.8

Fig. 4.5 Examples for alloyed and unalloyed cold work steels

Cutting-tools, shear-blades for cutting of steel, broaches, woodworking-tools

4.2

Tool Steels

101

and small amounts of Si and Mn. Alloyed tool steels have about 1.25% C as well as up to 1.5% Cr, 1.2% W, 0.5% Mo and 1.2%V. The hardness and wear resistance of unalloyed tool steels depend on their martensitic structure. Wear resistance increases with hardness and carbon content; yet toughness decreases simultaneously and thus the material’s sensitivity during heat treatment and tool use becomes greater. All unalloyed tool steels are shell hardeners, i.e. they do not harden all the way through along the entire cross-section but only on the surface of the workpiece. The advantages of alloyed tool steels in contrast to unalloyed are the increase in wear resistance (addition of carbide-forming elements), retention of hardness and high temperature strength (chrome, tungsten, molybdenum, vanadium alloys) and in their higher hardness (carbon in solution). Moreover, the critical cooling speed is lower, allowing for improved hardenability. They can be used at cutting temperatures of up to 200◦ C. They are used above all in steel machining with low cutting parameters (reaming, thread die cutting) and to manufacture tools for repair work, since their cost is lower than high speed steel (HSS) tools because they have fewer alloy elements. Due to their low hot hardness, which limits the cutting speeds with which they can be used, cold work steels are used only rarely for metalworking on machine tools. Their area of application extends mainly to hand tools such as files, gouges, reamers or on saw blades for woodworking (Fig. 4.5).

4.2.2 High Speed Steels A new cutting tool material was introduced for the first time in the year 1900 at the world’s fair in Paris, with which TAYLOR could realize significantly high cutting speeds. The productivity of machining processes could thus be increased considerably. This new group of cutting tool materials was called “high speed steels” (HSS). High speed steels are high-alloyed steels containing tungsten, molybdenum, vanadium, cobalt and chrome as their main alloy elements. They have relatively high bending fracture strength and thus have favourable toughness properties. In contrast to cold work steels, their matrix is characterized by improved retention of hardness, and they have higher hardness. Their hardness of about 60–67 HRC is preserved up to 600◦ C in temperature. Due to both this and their machinability, they continue to have a broad range of application, especially for tools with sharp cutting edges and small wedge angles such as broaching tools, twist drills, threadcutting tools, reamers, milling cutters and turning tools for grooving and parting-off operations as well as for finishing. While the hardness of high speed steels are affected by the hardness of the base material (martensite) and the number and distribution of the carbides, the alloy elements dissolved in the matrix W, Mo, V and Co, which, partially precipitated as stable special carbides, are responsible for retention of hardness. Hardness and wear resistance are increased by the martensite tempered in the matrix and the embedded

102


carbides (especially Mo-W double carbides, Cr and V carbides). Carbide formation and hardening is promoted by additional alloying of chrome. 4.2.2.1 Classification of High Speed Steels High speed steels are denoted with the letters “HS” and the percents of alloy elements in the order W-Mo-V-Co (e.g. HS10-4-3-10). High speed steels are classified according to their W and Mo content into four alloy and performance groups (Fig. 4.6) [DIN EN 10027a]. Group I includes the steels that contain high amounts of tungsten (18% W). The type containing Co HS18-1-2-5 has good retention of hardness. Tools made of this alloy are used for roughing high strength steels, cast irons that are difficult to machine, nonferrous metals and non-metallic materials. The material HS12-1-4-5 of the group with 12% W has excellent wear resistance due to its high V content. Its Co content also gives it high hot hardness and retention of hardness. This alloy is used in the production of turning tools and profile steels of all kinds, for finishing tools, high performance milling cutters and automatic tool machines. They are suitable for machining highly heat-treated Cr-Ni steels and non-ferrous metals.

Composition

W - Mo - V - Co

I

II

III

IV

For conditioning of steel under... ...average load ...highest load

Symbol

< 850 N/mm2 > 850 N/mm2

Roughing

Finishing

HS18-0-1

+

–

–

–

HS18-1-2- 5

–

–

+

–

HS12-1- 4 - 5

–

–

(+)

+

HS10-4 - 3 - 10

–

–

(+)

+

HS6-5-2

–

+

–

–

–

–

(+)

+

HS6-5-2- 5

–

–

+

–

HS2-9-1

+

–

–

–

HS2-9-2

–

+

–

–

HS2-10-1-8

–

–

+

–

18% W

12% W

6% W + 5% Mo

2% W + 9% Mo

HS6-5- 3

Fig. 4.6 Alloy and performance groups of high speed steels

4.2

Tool Steels

103

The last two groups are represented primarily by steels containing tungsten and molybdenum. Molybdenum replaces tungsten from metallurgical points of view and is more effective at equal masses because it has an approximately twice as large volume percentage due to its lower density. Steels containing molybdenum are especially tough. Types containing cobalt are used for machining with simple tools when robust stresses are expected (drills, turning, milling, planing and broaching tools, hobs). They are characterized by high toughness, temperature resistance and retention of hardness. They are universally applicable for roughing and finishing tasks, especially for highly stressed milling tools. Steels of both of these groups that have low amounts of or no cobalt are used to manufacture tools of all kinds. The most important group are high speed steels with 6% tungsten and 5% molybdenum. Usually, the universal or standard type HS6-5-2 is used [Habe88]. As a result of its balanced alloy structure, its high toughness and good wear resistance, this alloy is very diverse. It is suitable for manufacturing machining tools for roughing or finishing such as twist drills, milling cutters of all types, tappers, reamers, broaches etc. The basic effect of alloy elements in high speed steels is summarized once more in abbreviated form in the following: • Tungsten: forms carbides; increases hot hardness, retention of hardness and wear resistance. • Molybdenum: forms carbides; improves hardening and toughness, increases hot hardness, retention of hardness and wear resistance, molybdenum can replace tungsten (has half its density!). • Vanadium: exists as the primary carbide VC and increases wear resistance (finishing). • Cobalt: shifts the limit of overheating sensitivity to higher temperatures, making it possible to achieve higher hardness temperatures. Most carbides are dissolved and the hot hardness increases. • Chrome: improves hardenability, participates in carbide formation. • Carbon: the source of hardness in the base material. Increases wear resistance because it forms carbides. 4.2.2.2 Areas of Application Figure 4.7 provides a list of the main applications of high speed steels [DIN EN ISO 4957, Wegs95]. As the amount of alloy elements increases, so does the effectiveness of these cutting tool materials with respect to improved wear resistance and lifetime. At the same time however, machinability becomes more difficult, which has negative consequences particularly when manufacturing complicated moulding tools. Generally, a higher amount of alloy elements means higher tool costs. The cost effectiveness of a manufacturing process is thus also determined by the choice of high speed steel. This use of high-alloyed high speed steels lends itself especially to the solution of machining problems in which an increase in high temperature strength or

104


Steel grade Material name according no. DIN EN ISO 4957

Main use

HS6-5-2

1.3343

Standard high speed steel for all cutting tools for rough turning or finish turning, taps and twist drills, milling cutters of all kinds, pull broaches, reamers, threading dies, counterboring drills, planing-tools, buzz saws

HS6-5-3

1.3344

Highly loaded taps and reamers, high speed milling cutters, broaching-tools, twist drills, cutting and shaving wheels

HS6-5-2-5

1.3243

High speed milling cutters, planing- and turning-tools of all kinds, highly loaded twist drills and taps, broaches, woodworking- and cold working-tools, roughing-tools with high durability

HS10-4-3-10

1.3207

All-purpose usage for roughing- and finishing-works, turning- and highly loaded milling cutters, automated works, woodworking-tools

HS2-9-2

1.3348

Milling cutter, reamers, broaching-tools

HS2-9-1-8

1.3247

End-milling cutter, turning-tools for automated works, twist drills, taps

Fig. 4.7 Main applications of the most important high speed steels, acc. to DIN EN ISO 4957

toughness would have a great effect. High speed steels alloyed with cobalt (e.g. HS6-5-2-5, HS18-1-2-5) are ideal for machining tasks that place increased requirements on the high temperature strength of the tools. Steels that contain vanadium in addition to cobalt, such as the qualities HS12-1-4-5 and HS10-4-3-10, are characterized by improved wear properties at increased temperatures stresses and are suited to machining tasks that place the highest requirements on the wear resistance of the tools. 4.2.2.3 High Speed Steel Fabrication High speed steels can be fabricated using cast-metallurgical [Habe88] or powdermetallurgical methods [Duda86, Beis82]. A comparison of both fabrication methods is provided in Fig. 4.8. Despite attempts to develop alternative fabrication techniques, the cast-metallurgical fabrication of high speed steels remains the most economically important.

4.2

Tool Steels

Casting

105

Recast (ESU)

HotMachining deformation

Cast-metallurgical fabrication Melting

Powder-metallurgical fabrication Gas spray aerated Housing powder

Water spray aerated powder

Hot isostatic pressing

Heattreatment

Grinding

Hot Machining deformation

Cold Sintering HotMachining isostatic ρ < 100 % deformation pressing Softening Cavitypressing

Sintering ρ = 100 %

Hot isostatic repressing

Fig. 4.8 Production-sequence of the powder- and cast-metallurgical fabrication of high speed steels

Cast-Metallurgical Fabrication After melting and casting (1550◦ C) the steel in moulds, it is then subjected to block tempering (900◦ C) (homogenization, Fig. 4.9). This is followed by a forging process (to break up the ledeburite and carbides) with a possible intermediate heating of the workpiece, as well as by rolling (1200◦ C). Softening improves machinability. The softened condition is in most cases the post practical both for cutting and cold forming and offers the most favourable output structure for hardening. In order to avoid cracks or warpage during hardening, it is advisable to anneal the tools with a minimal amount of stress in order to reduce machining stresses before the last mechanical processing, especially in the case of tools that are irregular or difficult to mould. High speed steels require a highly complex hardening sequence due to the high amounts of different alloy elements. Conventionally produced high speed steels thus tend to segregate (demix) during the hardening phase, which is especially pronounced in the case of large cast blocks and high-alloyed steels. Such phenomena are usually causes of lifetime deviations in later tool use. Independently of previously applied measures such as improved melting process or seeding as well as improved design of the block format, electro slag remelting (ESR) can further improve the degree of purity and toughness of high speed steels. In this process, only a small part of the block is fluid during remelting, not the whole block. In this way, macroscopic demixings such as carbide agglomerations can be reduced and slag inclusions removed. The disadvantage is the higher fabrication cost.

106

4 Cutting Tool Materials and Tools HS10-4-3-10

HS12-1-4-5

HS12-1-4

HS6-5-2

1280

1240

1240

1240

1230

3× 560 °C

2 × 570 °C 1 × 550 °C

2× 570 °C

2× 560 °C

2× 540 °C

Code = Grade

HS18-1-2-5

Hardening temp. [°C] Annealing, each 0.5–1 h

1600

1000 800 600 400 200

Block turning

1200 Melting and casting

Temperature T / °C

1400

Block tempering

Forging

Rolling

Softening

0

Slow oven cool-down

F B

If heated in a vacuum furnace

Air

Air

g ne An 3.

An

ne

al in

al in g

g al in

air

2.

A

ne

oil/air

C

Annealing duration each approx 1–2 h

An

D

Annealing

1.

E

Balancing-temperature

Finishing works

Preparatory work

Temperature T / °C

St re s tre s re at lie m f en he t at W ar m in g up Au st en iti sa tio n Q ue nc hi ng

Time t /h

Air

G

Time t / h A 600 to 650 °C B 1st preheating stage aprrox. 400 °C (in salt bath) C 2nd preheating stage 850 °C

D E F G

3rd preheating stage 1050 °C Hardening temperature ~1200 °C Hot bath 500 bis 600 °C 50–80 °C

Fig. 4.9 Cast-metallurgical fabrication and heat treatment of high speed steels

4.2

Tool Steels

107

The advantage of the ESR process is that the end structure is more even, has a small amount of segregations and a high degree of purity. The practical consequences of this are reduced warpage in both transverse and longitudinal directions during hardening and annealing due to the more homogeneous microstructure – an especially important aspect in the case of tools that are long or are premachined with a small stock allowance – as well as improved toughness resulting from the higher degree of purity and the generally more homogeneous structure. Powder-Metallurgical Fabrication The starting materials for powder-metallurgical fabrication of high speed steels is powder obtained by means of gas or water spray aeration of the molten mass (Fig. 4.8). Gas spray aerated powder is housed and undergoes hot isostatic pressing. After soft deformation, the semi-finished products are conventionally machined into tools. Water spray aerated powder undergoes a cold isostatic pressing after softening, and then the semi-finished products are sintered in a vacuum furnace. Sintering is not performed until the full theoretical density is reached. Final pressing takes place later in a forging process downstream. On the other hand, mouldings such as indexable inserts are hardened in matrix presses. During sintering, the aim is to reach the full theoretical density. The sintering process can sometimes be followed by a hot isostatic repressing. Powder-metallurgically obtained high speed steels have become increasingly important in recent years. As opposed to cast-metallurgically fabricated high speed steels, they generally have a somewhat higher alloy content. PM steels are characterized by a homogeneous structure (no carbide segregation) with an even distribution of fine carbides. Due to their structural composition, PM steels have better grindability and higher toughness. With respect to their efficiency as cutting tools, powder-metallurgically fabricated high speed steels vary to some degree. Numerous cutting experiments have shown that they are at least equal to conventional high speed steels with the same nominal composition. PM steels become advantageous under high mechanical stress resulting from greater feeds, especially when cutting tool materials that are difficult to machine, such as nickel-based and titanium alloys. Until now, powder-metallurgically fabricated high speed steels have no material indexes of their own. They are designated according to the company that manufactures them. To simplify their classification, the American nomenclature is often used, or the standard designation of the corresponding cast-metallurgically fabricated high speed steel is applied. 4.2.2.4 Heat Treatment of High Speed Steels HSS tools are subjected to heat treatment in order to give them their required hardness [Habe88]. First they are hardened (heating and maintaining an austenitisation temperature followed by high-speed cooling) and then annealed several times. Figure 4.9 shows a time-temperature chart for the heat treatment of high speed steel HS6-5-2.

108


Due to their poor heat conductivity, high speed steels must be heated very slowly in order to avoid heating stresses and cracks. Warming up thus generally takes place in two or three preheating stages at 400, 850 and 1050◦ C. After reaching the level of austenitisation, a temperature of 1150–1250◦ C is held for hardening. The goal of austenitisation is to dissolve as many carbides as possible, whereby the dissolution of the secondary carbides increase with a rising hardening temperature and extended holding times. On the other hand, coarseness must also be avoided. Depending on the critical quenching speed and on the shape and size of the workpieces, quenching can take place in different media, including water, oil, hot baths (salt or metal molten baths) or gas. Gaseous media include air that is still or in motion, nitrogen or other gasses. Quenching intensity, which in the case of gasses is much lower than with liquid media, can be increased by raising the flow speed and pressure. After quenching, the microstructure of high speed steels consists of martensite, residual austenite and carbides. The hardness of this structure is unsuitable for cutting tools; due to the high amount of residual austenite and the instability of the structure, it has insufficient wear resistance, is not dimensionally stable during use and exhibits an increased fracture tendency. The tools must therefore undergo annealing after hardening. High speed steels are annealed at temperatures of 540–580◦ C, the temperature range of secondary hardness. This is defined as the increase in hardness at higher annealing temperatures after an initial decline in hardness from annealing at low temperatures. Under normal hardness and annealing conditions, this second increase in hardness leads to hardness values that are clearly higher than those existing after quenching. Figure 4.10 is a schematic depiction of the processes that occur during the annealing of high speed steels. The cutting tool material hardness resulting

M

~ 66

M2 S A2

Hardness/HRC

M1

A1

A

Annealing temperatureT /°C

~ 600

Fig. 4.10 Schematic illustration of cutting tool material hardness resulting from the superposition of various effects during the annealing of high speed steel

4.2

Tool Steels

109

from the superimposition of various effects is plotted as the cumulative curve S as a function of the annealing temperature. As the annealing temperature increases, the hardness of martensite (M) decreases continuously (M1). The cause of this is precipitation of the carbon trapped in solution and the associated reduction of crystal lattice tension. The precipitation of the carbides acts against this process, which leads to the decline in hardness observable during annealing at up to about 350◦ C. Extremely finely distributed carbides precipitate from the martensite (M) above 350◦ C and from the residual austenite (A) above around 450◦ C, causing another increase in hardness (M2/A1) due to precipitation hardening. As a result of carbide formation, the matrix becomes low in carbon and alloy elements, so that part of the residual austenite is converted to secondary martensite (A2) during cooling from annealing temperature. All three processes – carbide precipitation from the martensite, carbide precipitation from the residual austenite and the conversion of residual austenite into secondary martensite – lead to the increase in hardness known as the secondary hardness effect. The maximum of the hardness profile is shifted to higher temperatures (560◦ C), and, if austenitisation occurs correctly, hardness can be higher after annealing than in the hardened state [Habe88]. Until now, heat treatment usually took place in salt baths. Their advantages are optimal heat transfer and the possibility of executing partial heat and surface treatments. The important disadvantages are the high costs involved in detoxifying the salts and the necessity of cleansing the workpieces of salt residue after treatment. The last few years have seen increasing amounts of heat treatments taking place in vacuum furnaces. The advantages of this method are above all its environmental friendliness as well as the clean and shiny workpiece surfaces it yields. Quenching the workpieces from hardening temperature is achieved by flooding them with nitrogen at a pressure of 10 bar. The gas is circulated such that it flows at high speed through the batch and then through a heat exchanger for recooling. Components are quenched in this manner to 50◦ C. 4.2.2.5 Surface Treatment The wear properties of HSS tools can be further improved by treating the surface by nitrating (enriching the surface layer with nitrogen by thermochemical treatment at temperatures of 500–580◦ C), steam annealing (formation of a thin iron oxide layer on the tool surface in a steam atmosphere at about 500◦ C), chrome-plating (deposit of a hard chrome layer with a thickness of 5–50 μm at temperatures of 50–70◦ C) or coating (PVD coating with thin, highly wear-resistant hard material coats such as titanium nitride (TiN) or titanium aluminium nitride (Ti,Al)N. 4.2.2.6 HSS Indexable Inserts Indexable inserts made of high speed steel can be manufactured by precision casting, by machining pre-products or directly using powder-metallurgical methods.


Cutting speed vc

70 m/min

Material: rpm: Feed velocity.:

Profile milling cutter for rotors

X5CrNi13–4 n = 65 min–1 v f = 35 mm/min

Requirements for the cutting tool material Sufficiant High ductility temperature wear at low vc resistance

HSS

98 m/min

110

Cemented carbide

345

480 Diameter D/mm

Indexable inserts HS10-4-3-10 P20 - P30

Fig. 4.11 Fit-for-purpose design of indexable inserts composed of high speed steel and cemented carbide (Source: Fette)

While pre-products still require further machining, powder-metallurgical fabrication of HSS indexable inserts for example exploit the cost-intensive alloy elements much more fully. Besides lower manufacturing costs and more efficient utilization of material, the most important advantage of sintered HSS indexable inserts is their increased flexibility in tool use due to the use of indexable insert technology. In the case of manufacturing processes such as milling or drilling, this technology allows us to fit tools with various fit-for purpose cutting tool materials (Fig. 4.11). Indexable inserts made of precision casting have so far gained no importance for steel machining [Bong91, Wähl86].

4.3 Cemented Carbides Cemented carbides are composite materials. They are composed of the carbides of transition metals (fourth to sixth subgroup of the periodic system), which are embedded in a soft metallic binder phase made of cobalt and/or nickel. The carbides are on the boundary between metals and ceramics. They still have metallic properties (e.g. electrical conductivity), but they assigned to the non-oxide ceramics as “metallic hard materials” [Horn06, Salm83]. Additionally, cermets contain titanium nitride. Hard materials are the bearers of hardness and wear resistance. The binder phase is responsible for binder the brittle carbides and nitrides to a relatively solid body.

4.3

Cemented Carbides

111

The advantages of cemented carbides include good structural uniformity due to their powder-metallurgical fabrication, high hardness, pressure resistance and hightemperature wear resistance. At 1000◦ C, cemented carbides have the same hardness as high speed steel at room temperature. It is also possible to manufacture types of cemented carbides with various properties by means of intentionally changing the amounts of hard material and binder [Sche88, Kola92].

4.3.1 Historical Development In 1927, cemented carbides were first introduced as new high performance cutting tool materials at the Leipzig Trade Fair under the name WIDIA (Fig. 4.12). This was a revolutionary development at the time, opening up completely new dimensions in cutting technology. In contrast to the high speed steel tools used hitherto, cemented carbides allowed the use of twice or three times the cutting speed with the same cross-section of undeformed chip. The capacities of the machine tools of the time were far beneath the potential of the new cutting tools. Materials such as chilled cast iron, which had been very difficult to cut with HSS tools, could be machined easily with the new cutting tool material. Initially, WC-Co cemented carbides were used exclusively to machine cast iron materials. Due to the high amount of crater wear, these cemented carbides were not suited to machining long-chipping steel materials. This changed with the introduction of titanium carbide as an alloy component. With cemented carbides containing TiC, similar sensational cutting speed increases were made possible in steel machining as previously was the case when machining cast iron with WC-Co cemented carbides. This period also saw the development of the first cermet. Cemented carbides made for patent reasons without tungsten based on TiC and Mo2 C with Nickel as the binder proved in many cases to be 1923

Patent for the fabrication of WC-Co-HM, K. Schröter

1927

Leipzig Trade Fair: 1st WC-Co-HM, producer Krupp, labeling WIDIA

1928

Carboloy (General Electric)

1931

WC-TiC-Co cemented carbide

1934

TiC-Mo2C-Ni (1st generation of Cermets) producer Plansee, labeling Titanit S

1937

Titanit (Plansee), Böhlerit (Böhler)

1942

Coromant (Sandvik)

68/69

Implemention of coated cemented carbide

1970

Fine grain cemented carbide

1973 since 73/74

Spinodal - cemented carbide (1st Cermet with TiN)

1993

Submicron grain and nano crystalline cemented carbide

Accelerated developing of Cermets

Fig. 4.12 Milestones in the development of cemented carbides

112


excessively fragile for the machinery of the time [Sche88, Kola92, Kola93, DRP23, ÖP31, Häus90, Spri95]. Further development of cemented carbides in the following years led to continuous improvement of their composition, production and cutting performance. The influence of carbide grain size on the properties of cemented carbides was recognized early on. The relation, whereby cemented carbides can be increased in hardness only with a reduction in toughness, could be overcome with the development of fine grain cemented carbide. By reducing WC crystallite size to under 1 μm, both hardness and bending strength could be increased with the same amount of binder [Sche88, Kola93, Spri95]. The introduction of coated cemented carbides at the beginning of the 1970s was another great innovation. The combination of tough cemented carbide substrates with highly wear-resistant hard material coats led to an enormous increase in possible cutting speeds and tool standing times. The CVD and PVD methods are the most important coating process variants today. Multi-layer coats on substrates with binder-rich rim zones or gradient structures are new developments [Sche88, Kola92, Gill95]. With the development in 1973 of “spinodal” cemented carbides, the first Cermet, which contained titanium nitride as a further hard material component, the basic form of today’s highly efficient cermet was created. Cermets are today among the high performance cutting tool materials that meet the demands of modern cutting technology excellently by allowing for the use of high cutting speeds with moderate feeds and realizing long standing times with a high level of reliability. This is due to their high chemical stability and high-temperature wear resistance, making these cutting tool materials especially interesting for cutting operations with high thermal stress on the cutting edge.

4.3.2 Cemented Carbide Production Cemented carbides are fabricated using various powder-metallurgical means (Fig. 4.13). Due to the great variety of shapes that cemented carbide components can assume, very different forming methods can be used. The most general distinction is between direct and indirect fabrication, combined fabrication and special methods (injection moulding). The fabrication method used depends mainly on the geometry and quantity of the product to be manufactured [Sche88, Kola92]. About two thirds of all cemented carbide products are manufactured by direct fabrication, primarily indexable inserts. Complicated moulding in small quantities, such as pistons, screws, rolling rings or matrices are made indirectly, i.e. using additional machining steps such as cutting, drilling, turning or milling. The starting material is a cemented carbide in a pre-sintered or cold-isostatically pressed condition, the consistency of which is still chalky [Sche88, Kola92]. The individual components of the cemented carbide are weighed out as powder to a batch and homogenized in mixers. In the wet grinding following this, the grinding fluid (alcohols, acetone, hexan) should protect the powder from oxidation during

4.3

Cemented Carbides

113

Tungstencarbide WC

Auxiliary metal Co, Ni

Alloying elements TiC, (Ta,Nb)C, Mo2C, VC, Cr3C2

Blending, wet grinding, wet-sieving Pressing additives

Vacuum drying, screening Spray drying

HM-powder

Granulation

HM - Granulate

Cold isostatic compression

Cold isostatic compression

Machining

Debindering Sintering

Plasticising Vacuum drying dispersing

HM-modelling material

Powder compaction

Debindering Sintering/ HIP

Semi-finished products beams, profiles

Extrusion

Vacuum drying dispersing HM-injection moulding material Injection moulding

Debindering Sintering HIP

Grinding, separating

Coating

Finished products Grinded mouldings for wear protection

Thick beams Profiles Drills, milling cutter

Indexable inserts Mining tools Small bearing parts Short beams

Beams, profiles d < 22 mm Drills with integrated coolant bore

Small mass products with complex geometry

Fig. 4.13 Fabrication sequences for the production of workpieces made of cemented carbide (Source: Widia)

grinding and guarantee optimal dispersal of all the components in the suspension. After the grinding process is finished, the powder mix is prepared in accordance with the subsequent shaping method. To produce indexable inserts, the powder is transformed into granulate in order to assure good flow properties and a suitable granulate size for compression into matrices. The granulate is produced with the help of spray drying or granulation processes. Shaping the indexable inserts is achieved by compressing the granulate in matrix presses. The matrix press method allows for short cycle times and is thus especially suited to producing large quantities [Kola92, Daub95]. Within the fabrication sequence, sintering process is probably the most important operation, since it is here that the component obtains the mechanical and technological properties that are essential for its functional capability. In principle, sintering is defined as a thermally activated material transport in which the compressed, loosely bonded powder material (pressed part, green compact) is increasingly compressed due to diffusion-controlled place-shifting processes (surface diffusion, grain

114


boundary diffusion, volume diffusion). It is decisive for fabricating cemented carbides that not the entire alloy system is converted to a molten state, but rather (as it is called in professional jargon) sintered with liquid phase [Sche88, Kola92]. Molten baths first appear at about 1300◦ C. Tungsten carbide dissolves increasingly in the molten bath. At a sintering temperature of about 1400◦ C, the entire binder phase, composed of cobalt and tungsten carbide, is molten. The liquid phase wets the carbide, penetrates into all the pores and causes the hard material particles to glide together under the influence of surface tension into the smallest space. This particle rearrangement and re-precipitation process leads to tight packing with minimal surface energy. Depending on the green density and sintering parameters, linear shrinkage can reach 20%. As the mass cools from sintering temperatures, tungsten carbide precipitates again from the liquid binder phase. Up to a temperature of 700◦ C, a further thermally activated material transport of tungsten and carbon takes place in accordance with the solubility of WC in Co. Atoms of the hard material phase still dissolved after solidification in the binder phase stabilize the cobalt-rich binder phase in the cubic crystalline crystal lattice structure until room temperature; otherwise the cobalt converts into hexagonal form below 417◦ C [Sche88, Kola92].

4.3.3 Components of Cemented Carbides and Their Properties • WC: Mono tungsten carbide is the most important hard material phase in technical sintered cemented carbides. WC is soluble in Co, resulting in the high inner bonding and edge strength of WC-Co cemented carbides. WC is also even more wear-resistant than TiC and TaC. The applicable cutting speed is limited at higher temperatures due to its tendency to dissolution and diffusion. • TiC: Titanium carbide has a low tendency to diffusion. The consequence of this is that TiC cemented carbides have considerable high-temperature wear resistance but little bonding and edge strength. Cemented carbides high in TiC are therefore brittle and fragile. Their use is preferred when cutting steel materials with high cutting speeds. TiC can join with WC to make a composite carbide. • TaC: In small amounts, tantalum carbide has a grain-refining effect, thus improving toughness and edge strength; the inner bonding strength does not decrease as sharply as is the case for TiC. • NbC: NbC has a similar effect as TaC. Both carbides appear as a mixed crystal (Ta, Nb)C in cemented carbides. • TiN: Titanium nitride is the property-determining component in all modern cermets. TiN in steel is even less soluble and is thus more resistant to diffusion than titanium carbide. Nitrogen causes an increase in wear resistance. In addition, grain growth is inhibited. Cermets containing nitrogen have as a rule a very fine-grained structure. In a solid state, TiC and TiN are completely mixable. The physical properties of cermets are based on those of titanium carbon nitride. • Co: Cobalt is still unrivalled as a binder metal for cemented carbides based on tungsten carbide. This is due to the high level of solubility of WC in cobalt and

4.3

Cemented Carbides

115

to the good wettability of tungsten carbide crystals by the molten WC-Co binder phase [Sche88]. • Ni: due to the improved wettability of hard materials, nickel is used as a binder for cermets. But since nickel is more easy to deform than cobalt, cobalt is also added with nickel as a binder in cermets today in order to improve their high temperature properties.

Hardness Pressure resistance

The properties of the cemented carbide substrate are of key importance with respect to the wear resistance and efficiency of uncoated and coated cemented carbides. Cemented carbides should have high hot hardness and pressure resistance for the sake of increasing the resistance of the cutting edge to plastic deformation. They should however also have high bending strength and therefore sufficiently high toughness and resistance to cracking, crumbling and fracture. In general, tough cemented carbides are low in hardness and pressure resistance. These opposing cutting tool material properties are influenced decisively by the microstructure of the base material. The tendency is that fracture toughness increases with the cobalt content and average grain size, while hardness and pressure resistance are decreased (Fig. 4.14). As the concentration of composite carbides goes up, fracture toughness is reduced. Tantalus carbide has a favourable effect on temperature change strength.

TiC TaC

Bending strength

(Ti,Ta)C TaC

Wear resistance

TiC

Co-Concentration

WC-Grain size

Composite carbide concentration

Cutting tool material attributes High wear resistance

High ductility

Co-Concentration:

Co-Concentration:

WC-Grain size:

WC-Grain size:

Composite carbide concentration:

Composite carbide concentration:

Fig. 4.14 Influencing factors for wear resistance of cemented carbides

116


This is exploited in many cases and, in the case of substrates for milling, the ratio TiC/TaC is modified in favour of tantalum carbide [Sche88, Schi89, Köni90]. The high material costs put a tight limit on the use TaC alloys however.

4.3.4 Microstructure In conventional cemented carbides based on WC, the tungsten carbide is usually in the form of prisms with triangular bases. As opposed to cemented carbides containing larger amounts of cobalt, in which crystal growth is less hindered, this crystal form is less well developed in types that are fine-grained and low in cobalt. Cubic mixed crystals exist both in both cubical (with rounded edges) and nearly spherical shape (Fig. 4.15). The carbide skeleton is filled with the binder phase. The structure of cermet contains only rounded composite carbon nitrides. The core/shell structure of the hard materials is characteristic of their microstructure. The causes of this are demixing phenomena in the hard materials (spinodal demixing), mainly however the dissolution and re-precipitation processes taking place during liquid phase sintering. By selective solution, the Ni-Co molten bath is enriched with carbide components. Composite carbides precipitate as a rim zone around the remaining hard material grains when cooling down from sintering temperature [Mosk66, Rudy73, Kola89, Kief71, Leop87, Kola89a, Kola93a].

Cermet HT-P15

Cemented carbide HW-P10

Chart of reflected electrons

Chart of reflected electrons

Hard material particles:

Hard material particles: TiN (Ti, Ta, W) (C, N) (Ti, Ta, W, Mo) (C, N) (Ti, Ta, W, Mo) (C, N)

WC (Ti, Ta, Nb, W) C Binder phase: Co-W-CMixed crystall

Binder phase: Co, Ni (Ti, W, Mo) Ni3 (Ti, Al) Main component

Schematic

Schematic

Fig. 4.15 Microstructure of conventional cemented carbides and cermets

4.3

Cemented Carbides

117

The microstructure of the hard material particles is determined by the components contained in the original powder. The hard materials can exhibit widely varied textural structures depending on composition and grain size. If the initial powder consists of unalloyed binary hard material components, usually titanium nitride or titanium carbide are in the cores (dark cores). When prealloyed powders are used consisting of composite carbides or quaternary carbon nitrides, higher concentrations of molybdenum and/or tungsten are also found in the cores (light cores). The back scatter electron image in Fig. 4.15 shows this varying core/shell structure very clearly. The initial powder consists in this case of single carbides, composite carbides and carbon nitrides. The hard material particles with the dark cores have Ti(C,N) in the centre, which is surrounded by a relatively titanium-rich composite carbon nitride (Ti,Ta,W)(C,N). The lighter (Ti,Ta,W,Mo)(C,N) shell consists primarily of a titanium-tungsten-molybdenum composite carbide containing a relatively large amount of tungsten and/or molybdenum. The hard material particles in the light core contain this composite carbide inside, surrounded by a titanium and nitrogen-rich shell.

4.3.5 Classification of Cemented Carbides Cemented carbide cutting tool materials can be subdivided into three groups. These are cemented carbides based on: • WC-Co, • WC-(Ti,Ta,Nb)C-Co and • TiC/TiN-Co,Ni. The term “cermet” is now used for cemented carbides based on TiC/TiN-Co,Ni. 4.3.5.1 WC-Co Cemented Carbides The cemented carbides of this group consist almost exclusively of hexagonal tungsten monocarbide and the binder phase cobalt. They can contain up to 0.8 mass % VC and/or Cr3 C2 and/or up to 2 mass % (Ta,Nb)C as doping additives to control structural fineness and consistency (Fig. 4.16). WC-Co cemented carbides are characterized by high abrasion resistance. Due to the strong diffusion tendency of tungsten carbide, they are not suitable for machining soft steel materials. They are mostly used for short-chipping materials, cast iron materials, non-ferrous metals and non-metals, high temperature materials as well as in stone and wood processing (Fig. 4.4). According to previous terminology, WC-Co cemented carbides were subdivided in accordance with the average WC grain size in the sintered microstructure into fine (0.8–1.3 μm), submicron (0.5–0.8 μm) and ultrafine (0.2–0.5 μm). Deviating from this, the standard DIN ISO 513 distinguishes only between cemented carbides with grain sizes ≥ 1 μm (HW) and those with grain sizes of < 1 μm (HF).

118


Cemented carbide Application group (DIN ISO 513)

HW-K05

HW-K10

HW-K25

HW-K40

WC-4Co

WC-6Co

WC-9Co

WC-12Co

15.1

14.9

14.6

14.2

1730

1580

1420

1290

5700

5400

5000

4500

1600

2000

2350

2450

650

630

590

580

Fracture toughness (N⋅m1/2/mm2)

6.9

9.6

12.3

12.7

Poisson-constant

0.21

0.22

0.22

0.22

Heat conductivity (W⋅m–1⋅K–1)

80

80

70

65

Coefficient of thermal expansion (293–1073 K) (10–6⋅K–1)

5.0

5.5

5.6

5.9

Grade Density (g/cm3) Hardness HV 30 Pressure resistance (cylinder test) (N/mm2)

(ISO 3369) (ISO 3878)

(ISO 4506)

Bending strength (N/mm2)

(ISO 3327)

Young’s modulus (103⋅N/mm2)

(ISO 3312)

Fig. 4.16 Composition and attributes of conventional WC-Co cemented carbides, acc. to Kolaska [Kola92]

Conventional Fine-Grain Cemented Carbides (HW) Conventional uncoated WC-Co fine-grain cemented carbides (fine cemented carbides) with an average grain diameter of 0.8–1.3 μm (acc. to DIN ISO 513 grain size ≥ 1 μm) are still widely used in cutting technology applications where high demands are placed on cutting edge sharpness and toughness, e.g. in steel milling, finishing, grooving and parting-off or in the manufacture of threads [Kola92]. Submicron and Ultrafine Cemented Carbides Submicron and ultrafine cemented carbides have until now been the terms used in literature and practice for WC-Co cemented carbides with an average WC grain diameter of 0.5–0.8 and 0.2–0.5 μm in the sintered microstructure. The small grain diameter gives there cemented carbides a special combination of attributes: reducing the WC crystallite size below 1 μm with equal binder content leads both to increased hardness and bending strength (Fig. 4.17). This property makes submicron and ultrafine cemented carbides applicable to a wide variety of tasks [Kola92, Daub95, Köni90a, Köni93, Drey01, Gill01]. Valuable submicron and ultrafine cemented carbides are superior to conventional fine-grain cemented carbides in hardness, edge strength and toughness. Moreover, they have a small inclination to adhesion and to wear by diffusion. These attributes

Tool life T / min

(B): Bending strength N/mm2

(H): Hardness HV30

14 12

Fine grain WC-6Co-Cemented carbide

Submicron grain

Turning of white iron (80 Shore)

vc = 16 m/min

4300(B)

3000

1780

Standard

2000(H)

119

2000

Cemented Carbides

1580

4.3

f = 0.1 mm

10 8

ap = 1.0 mm Indexable insert: SPGN 120308 Sharp-edged

6 4 2

κ r = 75°

0 Standard

Fine Submicron grain grain

Fig. 4.17 Microstructure and attributes of fine and submicron grain cemented carbides in comparison to standard K-class cemented carbides (Source: Widia)

are necessary when hardened materials must be machined with finishing quality and the smallest machining allowances. Submicron and ultrafine cemented carbides are used whenever high toughness, high wear resistance and the highest edge strength is required of the cutting edge, e.g. for broaching, milling and shaping of heat treated and hardened steels, in cast iron machining and for machining fibre-reinforced plastics and non-ferrous metals [Kola92, Daub95, Köni90a, Köni93, Drey01, Gill01]. 4.3.5.2 WC-(Ti,Ta,Nb)C-Co Cemented Carbides The cemented carbides of this group contain, besides tungsten carbide, mixed carbides (MC) made of titanium, tantalum, niobium and/or zirconium carbide (Fig. 4.18). Compared with WC-Co cemented carbides, they are characterized by improved high-temperature properties. This is especially the case for hot hardness and high temperature strength, oxidation resistance and diffusion resistance to iron materials. Their main area of application is in machining long-chipping steels (Fig. 4.4). Due to the mixed carbide content, the cemented carbides of this group can be classified according to their use into two subgroups. Group A: Mixed carbide content >10–11 mass %. Due to their mixed carbide content, the cemented carbides of this group are characterized by a high level of

120


HW

HW

HW

HW

HW

-

-

-

-

-

-

P10

P15

P25

P30

M10

M15

Composition (mass percentage) WC (Ti, Ta, Nb)C Co

60.0 31.0 9.0

64.5 25.5 10.0

72.7 17.3 10.0

78.5 10.0 11.5

84.5 9.5 6.0

82.5 11.0 6.5

Density (g/cm3)

10.6

11.7

12.6

13.0

13.1

13.3

(ISO 3369)

Hardness HV 30

(ISO 3878) 1560

1500

1490

1380

1700

1550

Pressure resistance (cylinder test) (N/mm2)

4500

5200

4600

4450

5950

5500

(ISO 4506)


(ISO 3327)

1700

2000

2200

2250

1750

1900


(ISO 3312)

520

500

550

560

580

570


8.1

9.5

10.0

10.9

9.0

10.5

Poisson constant

0.22

0.23

0.22

0.23

0.22

0.22

Heat conductivity (W⋅m–1⋅K–1)

25

20

45

60

83

90

Coefficient of thermal expansion (293 K–1073 K) (10–6⋅K–1)

7.2

7.9

6.7

6.4

6.0

6.0

Cemented carbide Application group (DIN ISO 513)

HW

Fig. 4.18 Composition and attributes of WC-(Ti, Ta, Nb)C-Co cemented carbides, acc. to KOLASKA [Kola92]

high-temperature wear resistance, a low tendency to diffusion with iron materials and low abrasion. They are utilized above all in steel and cast steel cutting, with the exception of rust and acid-resistant steels with austenitic microstructures. Group B: Mixed carbide content < 10–11 mass %. The cemented carbides assigned to this group have relatively good high-temperature wear resistance and abrasion resistance. They are especially suitable for cutting rust, acid and heat-resistant steels with austenitic structures as well as for alloyed or hard austenitic/ferritic cast iron materials. 4.3.5.3 TiC/TiN-Co,Ni Cemented Carbides (Cermets) Cemented carbides based on titanium carbide and titanium nitride with a Ni,Co binder phase are designated as cermets (formed from ceramic + metal). Present-day cermets are complex multi-material systems that can contain a number of additional elements such as tungsten, tantalum, niobium, molybdenum or complex carbides, from which intermetallic phases are formed during sintering.

4.3

Cemented Carbides

121

Cermet Application group (DIN ISO 513)

HT - P05

HT - P10

HT - P20

89.0 0.6 10.4

85.7 0.8 13.5

82.3 1.0 16.7

6.1

7.0

7.0

1650

1600

1450

5000

4700

4600

2000

2300

2500

460

450

440


7.2

7.9

10.0

Poisson constant

0.21

0.22

0.21

Heat conductivity (W⋅m−1⋅K−1)

9.8

11.0

15.7

Coefficient of thermal expansion (293 K–1073 K) (10−6⋅K−1)

9.5

9.4

9.1

Composition (mass percentage) Carbon nitrides Additive nitrides Co/Ni Density (g/cm−3)

(ISO 3369)

Hardness HV 30

(ISO 3878)

Pressure resistance (cylinder test) (N/mm2)

(ISO 4506)


(ISO 3327)


(ISO 3312)

Fig. 4.19 Composition and attributes of cermets, acc. to KOLASKA [Kola92]

The composition and some characteristic attributes of three typical cermet types is given in Fig. 4.19 [Kola92]. Compared to conventional cemented carbides, cermets are characterized by their lower density. Some essential differences to WCbased cemented carbides are their significantly smaller thermal conductivity coupled with their greater thermal expansion. Due to their lower heat conductivity, a larger portion of the cutting heat is dissipated with the chip, resulting in a lower total heating of the cutting insert. Near the contact zones however, the temperature gradient increases inside the cutting tool material. In association with the large thermal expansion coefficient, this leads to high tensile and pressure stresses in the cutting tool material. The result of this is that cermets react much more sensitively to temperature change than conventional cemented carbides. They thus have a strong tendency to form comb cracks, especially in interrupted-cut machining [Köni90a, Köni93, Kloc96]. Cermets are very hard, have a low tendency to diffusion and adhesion as well as a high level of high-temperature wear resistance. Due to their high edge strength, high resistance against abrasive wear and low adhesion, cermets are especially good for planing steels. Until now, cermets have been predominantly utilized to machine

122


500

Fine finishing ap < 0.5 mm

Finishing ap < 2 mm

200

Rough finishing ap < 3 mm

100

0.50 0.40 Cemented carbide HW−P10

0.30 0.20

0.10

Cermet

0.05

Coated cemented carbide

Cutting speed vc /(m/min)

400

300

C45+N, Rm = 700 N/mm2 vc = 400 m/min, f = 0.1 mm, ap = 1 mm Width of flank wear land VB/mm

Material: Cutting conditions:

3

20 10 5 Cutting time tc /min

30

vc = 220 m/min f = 0–25 mm ap = 1–1.5 mm Tool life quantity: 150

0 0

0.1

0.2 Feed f /mm

0.3

0.4

Cermet

C45E

Tool life quantity: 240 vc = 400 m/min f = 0.3 mm ap = 1–1.5 mm

Fig. 4.20 Field of application of cermets during turning steel (Source: Kennametal Widia, CeramTec)

steels with high cutting speeds and small cross-sections of undeformed chip. The development of tougher cermet types has led to the expansion of their field of application to include more average roughing conditions (Fig. 4.20). Basic areas of application include both turning and milling. Cermets also are suitable for grooving and thread turning. The high wear resistance of the cutting edges, in association with the low diffusion tendency and high oxidation resistance, generally leads to better surface qualities than coated cemented carbides are capable of producing in planing and finishing. At the same time, due to these properties, they allow for higher cutting speeds in comparison with conventional cemented carbides. But they are also suitable for cutting speeds below 100 m/min. Tougher cermet types, corresponding to the range P15 to P25 of conventional cemented carbides based on WC-(Ti,Ta,Nb)C-Co, are used successfully in average rough turning operations and in milling (Fig. 4.21). Because of their clear tendency to

4.3

Cemented Carbides

123

Chamfering of gears

600 150 Cemented Carbide -TiN PVD 50 HSS -TiN Cutting speed (m/min)

300

25 % 70 %

200 Tool life quantity (pieces)

Tool life travel path lf / m

260* Cermet

Peripheral down-cut milling

*Limited by the nmax of the maschine

Cermet

Cemented carbide K25 0

100 %

200

300


Tool costs per workpiece

Material: 16MnCr5 Tools: End milling cutter, D = 16 mm, (HSS -TiN) and D = 18 mm, (Cemented Carbide, Cermet-inserts) Coolant: Emulsion

100

Material: Tool: Cutting cond.:

50CrV4, 250 – 300 HB End milling cutter, D = 8 mm, z = 3 fz = 0.025 mm ap = 8 mm, ae = 1 mm Dry cutting

Fig. 4.21 Milling with cermets (Source: VW-Kassel, Kennametal)

form comb-type fractures, cermets should always be dry-milled [Drey97, Köni90a, Köni93, Kloc96].

4.3.6 Uncoated Cemented Carbides and Cermets Uncoated conventional cemented carbides are still used consistently in cutting technology whenever there are high demands placed in cutting edge sharpness and toughness properties, e.g. in steel milling, finishing, grooving and parting off operations or in thread manufacture. They are however in direct competition with coated cemented carbides and cermets, which are replacing them in an increasing amount of applications. According to DIN ISO 513, uncoated cemented carbides based on tungsten carbide take the abbreviation HW and uncoated cermets the prefix HT (Fig. 4.3).

4.3.7 Function Gradient Cemented Carbides Function gradient cemented carbides are defined as materials, whose composition and/or microstructure differ in the rim zone and kernel area. The transition from the rim zone to the kernel area is continuous [Berg03].

124


The goal in the development of function gradient cemented carbides was, and still is, to improve even further the wear resistance of coated indexable inserts (especially in the case of dynamic stress) by modifying the rim zone. Such modifications, e.g. via cobalt enrichening and mixed carbide depletion in a approx. 50 μm thick rim zone, both cracking and fracture sensitivity of the insert are reduced [Berg97]. Such measures are, however, only practical when applied to indexable inserts intended for larger cross-sections of undeformed chip and interrupted cuts. In planing operations, the strength of cemented carbides is generally completely sufficient. In this case, a rim zone that takes over or supports the wear-reducing effect of the coating (mixed carbide enrichening) is advantageous [Berg97]. In the following, three types of function gradient cemented carbides will be introduced. 4.3.7.1 Cemented Carbides with Rim Zone Free of Mixed Carbides Gradient cemented carbides are understood as cemented carbides, the rim zone of which is free of hard and brittle (Ti, Ta, Nb)C mixed crystals (Fig. 4.22) and consists practically of nothing but tungsten carbide and cobalt up to a depth of about 50 μm. The cobalt content in the rim zone is higher than within. To reduce impact sensibility, a coarse WC grain is advantageous. Moreover, nitrides, which are less hard but more wear-resistant than cubic carbides or carbides of other transition metals (e.g. zirconium carbides) can also be contained in these rim zones. In this way, cemented carbide substrates with high hardness but very tough rim zones of lesser hardness and increased resistance against cracking can be produced. The fabrication of rim zones free of mixed carbides can take place in various ways. The basic principle behind it is the formation of a nitrogen gradient in the rim zone of the cemented carbides [Berg97]. In the most commonly used method, nitrogen is introduced into the initial cemented carbide mixture as TiN or Ti(C,N). The titanium nitride decomposes during the sintering process. The nitrogen dissolving in the liquid phase diffuses from the cutting tool material into the surrounding atmosphere. In this way, a nitrogen concentration gradient arises in the cemented carbide plate from the inside outwards. The result of this is the formation of a titanium activity gradient. Titanium diffuses into the interior and accumulates there on existing cubic mixed crystals. The same is valid for tantalum and niobium. The mixed carbides near the surface gradually dissolve. On the other hand, cobalt penetrates from inside towards the outside. The growth rate follows a parabolic temporal law. It increases with a rising nitrogen content and decreases with a rising titanium content. An edge effect is characteristic of these processes, the end effect of which is that the zone free of mixed carbides is thinner on sharp edges than on even surfaces [Berg97]. Other method variants include nitrogen pressure treatment, in which the cemented carbide is subjected to a nitrogen treatment under increased pressure after sintering, and the creation of zones free of mixed carbides without nitrogen, in which the cemented carbide contains a defined excess of carbon and a certain sintering regime is adhered to during cooling [Berg97, Köni90, Yohe93].

4.3

Cemented Carbides

125

Material = C25 vc = 200 m/min f = 0.15–0.3 mm ap = 0.5–1.5 mm

Multilayer coating

360 %

gradiented rim zone

Multilayer coating on gradiented rim zone

100% Tool life quantity

P25 Hardness gradient

10 µm

1000 1200 1400 1600 Hardness HV

Fig. 4.22 Influence of the rim zone on the tool life of CVD tools (Source: Sumitomo)

The thickness of mixed-carbide-free rim zones produced in this fashion is up to 50 μm. The methods described can therefore only be used for indexable inserts that are already in their final shape after the sintering process. At best a minor cutting edge rounding can still be undertaken. Therefore, in the case of highly accurate milling inserts that are ground, such toughness-increasing measures are not used, although they would be highly desirable. In the case of turning operations with tools of the application groups P20–P40, indexable inserts with rim zones free of mixed carbides are widely used [Berg97]. As has been shown in practice, thus designed cemented carbides have significant advantages in their wear behaviour compared with coated ones. Especially when turning with interrupted cut, tools coated according to this design achieve significantly longer standing times.

126


4.3.7.2 Gradient Cemented Carbides with External Ti(C,N) Enrichment The hallmark of these cemented carbides is an enrichment of nitrides and carbon nitrides on the surface. The substrates are first tightly sintered and then impinged with nitrogen. Diffusion of nitrogen into the rim zone leads to the formation and enrichment of carbon nitrides high in nitrogen on the surface, the composition of which becomes more rich in carbon with increasing depth. The thickness of the surface carbon nitride layer can be varied by changing the treatment time. This process of enriching the rim zone with Ti(C,N) is called coating the substrate during the sintering process. The thus “coated” cemented carbides are efficient alternatives to uncoated and conventionally coated tools. Cutting experiments with gradient cemented carbides with Ti(C,N)-rich rim zones showed clearly improved wear behaviour compared with uncoated cemented carbides. They can however also be conventionally coated. This guarantees that the coating adheres better, especially when one succeeds in designing the coating and the rim zone such that neither are clearly delineable from each other after the coating process [Leng06, Leng04]. 4.3.7.3 Gradient Cemented Carbides with Internal Ti(C,N) Enrichment In this substrate variant, first a rim zone free of mixed carbides is produced. Then nitrogen pressure is altered at a relatively high temperature so that nitrogen diffuses into the interior through the mixed-carbide-free WC-Co rim zone. Diffusion of Ti through the WC-Co layer is, however, minimal. The rim zone functions as a membrane that permits the diffusion within of nitrogen but prevents the diffusion without of titanium. By means of this one-sided diffusion of nitrogen, a transition of the WC-Co zone to the base cemented carbide develops, i.e. a zone enriched with Ti(C,N) a few micrometers beneath the surface in the interior of the substrate. In milling experiments, inserts with internal Ti(C,N) enrichment exhibited excellent wear behaviour in comparison with TiAlN-coated standard inserts [Leng06, Leng04].

4.3.8 Coated Cemented Carbides and Cermets Coating cemented carbides with thin, highly wear-resistant hard material coatings is one of the milestones in their continued development. Coated cemented carbides consist of a relative tough base body (e.g. P20, K20) on which is applied a 5–20 μm thick hard material coating made of carbides (e.g. titanium carbide, TiC), nitrides (e.g. titanium nitride, TiN), carbon nitrides (titanium carbon nitride, Ti(C,N)) and/or oxides (e.g. aluminium oxide, Al2 O3 ). Although uncoated cermets are already relatively high in wear resistance, hard material coating also can further improve their wear performance and efficiency just as with the conventional cemented carbides. According to DIN ISO 513, coated cemented carbides and cermets have the prefix HC (e.g. HC-K20).

4.4

Coatings

127

4.4 Coatings Since about 1968, cemented carbides have been coated with hard materials in order to improve their wear resistance and thus also their performance capability. Today, the use of coated cemented carbides and high speed steels for machining various materials is state-of-the-art. The high efficiency of coated cemented carbides is made even more clear by the fact that in the meantime more than 80% of all cemented carbides used in cutting have a highly wear-resistant hard material coating. In turning, the percentage is about 95% and in milling about 60%. The primary function of the hard material layer is to inhibit contact between the material, thereby reducing tool wear, which is caused by adhesion, abrasion, diffusion and oxidation phenomena (Fig. 4.23). These processes are superimposed to a large extent and are only partially separable in their causes and in their effects on wear. But in general, besides abrasion at low cutting speeds, it is above all adhesion and, at high temperatures, diffusion and oxidation that determine wear on the tool. With their high hardness and chemical stability, several hard materials possess the required prerequisites for improving resistance effectively against crater, flank face and notch wear at both low and high cutting speeds. While the hard material layer improves wear resistance, lessens adhesion between the tool and the workpiece and acts as a diffusion barrier, the task of the substrate is to act as an effective support for the hard material layer and to provide the composite body (consisting of substrate and coating) with sufficient hot hardness and, in particular, toughness. The toughness attributes of a substrate are of decisive importance for applications with interrupted cut. Due to input and output impacts and associated mechanical and thermal stress variations, the toughness of the cutting edge is much more responsible

Adhesion Delamination Abrasion Oxidation Diffusion Surface effects

Strains Crack formation Crumbling Fractures

Fig. 4.23 Wear phenomena at coated cutting tools

Volume effects

128


for its dimensional accuracy than wear resistance. For the interrupted cut, the composite material “coated cemented carbide” must first exhibit sufficient resistance to cracking, crumbling and fracture (Fig. 4.23). Only when these demands are met can, analogously to the interrupted cut, wear mechanisms like adhesion, abrasion, delaminations and – as long as the contact zone temperature and contact time are sufficient – oxidation and diffusion be effectively hindered. From this we can derive two requirements for coatings [Köni92]: • The hard material layer has to drastically reduce the effect of all (if possible) wear mechanisms involved in wear process. • The coating process may not reduce the inner bonding strength, i.e. especially the toughness of the substrate.

4.4.1 Coating Methods Coating cutting tools can be done chemically and physically. Method variants include • chemical vapour deposition (CVD) and • physical vapour deposition (PVD). Figure 4.24 shows an overview of technically established coating methods for the production of wear-resistant hard material layers. As can be seen, layer deposition lies within a specific pressure/temperature range for each method and includes some characteristic coating systems, which are given in bold in the illustration [Berg03]. (Ti,Zr)(C,N) (Ti,Hf)(C,N) Hf(C,N) Zr(C,N) Ti(C,N)

(Ti,Zr)(C,N) TiN Ti(C,N) 5 10 Al2O3

105

TiN Ti(C,N) (Ti,Al)N (Ti,Al)(O,N) MoS2 WC/C

Pressure p /Pa

104 MT-CVD

103

CVD

102 101

PA-CVD

100

10–1 10–2 400

Al2O3 / ZrO2 / TiOx TiC TiN Ti(C,N) Al2O3

PVD 600 800 1000 Coating temperature T /°C

1200

Fig. 4.24 Comparision of pressure and temperature conditions of coating methods with characteristic coating systems (Source: Kennametal Widia)

4.4

Coatings

129

4.4.1.1 CVD Processes CVD processes are chemical reactions that take place in gaseous phase under rough vacuum conditions (103 –105 Pa) and with the addition of thermal or radiant energy, forming technically useful solids (hard materials) as well as volatile products. The chemical reaction is determined by the laws of thermodynamics and thus by the partial pressure of the gaseous components and temperature. There are three types of CVD coating procedures: • HT-CVD (high temperature CVD, 900–1100◦ C) • MT-CVD (medium temperature CVD, 700–900◦ C) • PA-CVD (Plasma activated CVD, 450–650◦ C) A broad palette of hard material coating systems can be synthesized with CVD methods. The most common coating materials are based on hard materials with primarily metallic bonds, such as TiC, Ti(C,N), TiN, heteropolar (ionic) bonds such as Al2 O3 , but also covalent bonds such as diamond [Damm04]. Figure 4.25 provides an overview of the most common hard materials manufactured with the CVD process. Figure 4.25 provides a comparison of some essential physical characteristics with the properties of tungsten carbide, the main component of classical cemented carbides.

WC6Co Crystal system

TiC

TiN

ZrC

ZrN

HfC

HfN

α-Al2O3

ZrO2

HfO2

fcc

fcc

fcc

fcc

fcc

fcc

rhombohedral

monoclinic

monoclinic

4.93

5.44

6.46

Density (g/cm3)

14.9

7.35

12.30 13.94

3.99

5.7

9.7

Micro hardness HV0,05

1580 3100 2300 2800 1700 HV30

2600 1850

2300

1200

1100

352

530

380

200

3700 3380

2050

2700

8.0

7.7

3070 2950 3420 2980

7.7

9.4

6.7

7.3

6.6

6.9

80

33

29

38

21

29

22

Gold yellow

Metallic grey

Light yellow

Metallic grey

Green yellow

Colour

5.5

Metallic grey

Thermal conductivity (W⋅cm–1⋅K–1)

495

Grey

Coefficient of thermal expansion (10–6.K–1)

348

6 (at 0.6 (at 1000 °C) 1000 °C)

5.6

< 0.1

Grey white

Melting point (°C)

570

White

451

Uncolored

Young’s modulus 590 (103⋅N /mm2)

Fig. 4.25 Characteristics of hard materials for coatings in comparison to a cemented carbide of the application group HW-K10 (Source: Kennametal Widia)

130


Because the process is near equilibrium, thermally activated HT and MT-CVD methods can only be used to produce equilibrium phases. On the other hand, the low temperature PA-CVD process is also capable of producing metastable phases such as TiAlN. Among the most technically advanced methods is the deposition of coating systems with nanostructures. In this way, complexly doped hard material phases made of Al2 O3 or superlattice structures made of TiN/TiB2 can be produced with a single coating layer thickness of 5 nm [Damm04]. The HT-CVD Process The high temperature CVD process (HT-CVD) is the classic CVD method for coating cemented carbides. With it, HM tools can be coated at temperatures of 900–1100◦ C. High temperature coatings are characterized by a high level of adhesive strength on the substrate, which is produced by an easily reproducible material transfer between the surface of the substrate and the reactive periphery [Ever94]. In order to produce a TiC coating, titanium tetrachloride (TiCl4 ), for example, is vaporized and transferred with methane (CH4 ) to a reaction vessel, which can contain several thousand indexable inserts (Fig. 4.26). Titanium carbide is formed at temperatures of 900–1100◦ C and a pressure below atmospheric pressure in a hydrogenous environment (H2 ) in a chemical reaction following this equation:

Coating furnace CH4

H2 Gas inlet H2 H2+TiCl4

Fluid TiCl4

Evaporator

Fig. 4.26 Schematic illustration of a CVD coating unit (schematic)

Gas exhaust

4.4

Coatings

131

1000◦ C TiCl4 + CH4 + n H2 −−−−−−−−→ TiC + 4HCl + n H2 10 − 150 mbar

(4.1)

This process allows us to deposit hard materials like TiC, TiN, Ti(Cx Ny ), (Ti,Hf)(C,N), (Zr/C,N), Al2 O3 , AlON and others as single coats or in different combinations as multilayer coats. The coating material is formed by chemical reaction from the gaseous phase directly on the surface of the parts to be coated. The reaction products rinse the substrates so that no shading effects arise. Parts with complex geometries can thereby be coated thoroughly and consistently without difficulty. Cemented carbides that are coated with the classic high temperature CVD process are characterized by high wear resistance due to their relatively thick hard material coatings (up to 20 μm for turning and up to 6 μm in the case of milling). The high process temperature is problematic in the case of coating tool and heattreated steels. In this case, coating must be followed by rehardening, making the material vulnerable to unacceptable warpage. A further disadvantage to the HT-CVD coating process is that the toughness of the coated cemented carbide body is reduced in comparison to the uncoated substrate. The causes for this loss of toughness in HT-CVD coating are extraordinarily complicated. They include on the one hand the properties of the substrate to be coated (chemical composition, crystalline structure, grain size, thermal expansion, bending and pressure strength, pre-treatment) and process control on the other hand (gas atmosphere, pressure conditions, temperature/time cycle). Furthermore, processes taking place during coating in the rim zone (formation of eta phases), residual stresses present in the substrate, transition zone (interface) and hard material layer as well as the microstructure, texture, thickness and adhesion of the applied coating are all mutually responsible for toughness reduction in HT-CVD coating. We will now consider more closely processes in the rim zone (eta phase formation), the “temperature/time cycle” in the coating process and coating thickness based on all the above factors, which offer us many starting points for the improvement of toughness in coated cemented carbides. During the high temperature CVD coating process, there is a danger of forming brittle phases in the interface. Cemented carbides of the first generation exhibited an additional approx. 3–5 μm thick brittle zone with an “eta phase” (W6 Co6 C, W3 Co3 C), caused by decarburization processes in the rim zone [Schi89]. Embrittlement from the eta phase has a negative influence on the toughness of coated cemented carbides. Figure 4.27 shows the time/temperature transformation chart for a cobalt alloy with 5% tungsten and 0.23% carbon. It shows typical coating temperatures and times for the HT-CVD, MT-CVD, PA-CVD and PVD methods. From this chart, we can see that in the classic HT-CVD and medium temperature CVD process the area of eta phase precipitation is passed through/tangent. In contrast, no changes in constitution are to be expected in the case of low temperature coating methods, as in the PA-CVD or PVD processes [Köni90].

132

4 Cutting Tool Materials and Tools 1200 Co-5W-0.23 C

1000

HT-CVD

Temperature T /°C

WC 800

Eta'

Eta

MT-CVD Co3W

600

PA-CVD PVD

400

200 0.01

0.1

1.0

10 Time t /h

100

1000

Fig. 4.27 TTT-chart of a cobalt alloy with typical temperature sequence of several coating processes (Source: Widia)

The comparative examination of the bending strength of thin, variously coated WC-Co cemented carbides shows what effect coating temperature and coating thickness have on these toughness-relevant parameters (Fig. 4.28). With increasing coating thickness and coating temperature, bending strength is reduced. The effect is especially negative in cutting operations with interrupted cut, in which the tool is subject to dynamic stresses. In the case of milling therefore, cemented carbides with thinner hard material coatings (4–6 μm) than in turning are used. With these coating thicknesses, toughness loss is relatively low. Reduction of wear protection due to the reduced coating thickness must be expected. The logical consequence of these discoveries is that more and more coating methods have gained significance for coating cemented carbides in the past few years, which can be used at lower coating temperatures in comparison with the classic HT-CVD process. The MT-CVD Process By using acetonitrile (CH3 -CN) instead of a methane (CH4 )/ nitrogen (N2 ) gas mixture, it is possible to lower the coating temperature in comparison with the HT-CVD method by up to 200◦ C [Chat86]. In the 700–900◦ C temperature range, carbon

4.4

Coatings

133 10 TiN TiN/Ti(C,N)

Changes of bending strength / %

0

PA-CVD

–10

400 °C

550 °C

PVD –20 700 °C

550 °C

–30 CVD

700 °C

–40 F

–50

CVD 2,6 14,5

–60

0

2

4 6 Coating thickness / µm

8

10

Fig. 4.28 Influence of coating method and coating thickness on the bending strength of coated cemented carbides (Source: Widia)

nitride coatings of the elements Ti, Zr and Hf can be applied with high deposition rates. As opposed to the conventional HT-CVD process, the MT-CVD method has the following advantages: • Thermal stress on the cutting tool material is lower due to the reduced coating temperature at simultaneously higher deposition rates. • The danger of decarburization and thus of embrittling eta phases is lower for cemented carbide substrates. • The coatings applied with the MT-CVD method exhibit lower tensile residual stresses compared with HT-CVD coatings. Trial results from the longitudinal turning test show the same qualitative dependence on cutting speed for MT-CVD coated cemented carbides as for the uncoated substrate (Fig. 4.29) [Köni92, VDI3324]. Both MT-CVD coated and PVD coated cemented carbides are clearly superior to uncoated cemented carbides in performance. In higher cutting speed ranges, the performance of the 3 μm thick

134

4 Cutting Tool Materials and Tools Material: 42CrMo4+QT Rm = 980 N/mm2 Substrate: P30/40 SPUN120312 f = 0.2 mm ap = 2.5 mm

100000

Number of impacts n

50000

Arc-PVD

10000 5000

MT-CVD

1000 Uncoated

500

100

60

80

100

200

300

400


Fig. 4.29 Influence of coating methods and cutting speed on the toughness behaviour of carbides in an interrupted longitudinal turning test

Arc-PVD-TiN coating and the 6 μm thick MT-CVD-TiN-Ti(C,N)-TiN coating do not differ. The PVD coating tends to show advantages at low cutting speeds. It extends the range of applicable cutting speeds clearly to lower cutting speeds to a range that is also covered by coated HSS. At low cutting speeds, the cutting edges succumb to flank face wear and no longer to cracking and crumbling. The causes for the improved toughness are the unaffected properties of the substrate and the more favourable compressive residual stresses resulting from PVD coating of the cemented carbides [Köni90]. MT-CVD coated cemented carbides are used above all in milling. There are significant performance advantages compared with conventional coatings. The end of their standing times is no longer brought about by cutting edge crumbling, but in a reproducible way by a continuous wearing of the flank face. Their excellent toughness properties make these cemented carbides also ideal for milling with cutting fluids. Fewer comb cracks form and crack development is slower than it is the case with HT-CVD coatings. Transverse cracks and cutting tool material crumbling occur later, resulting in longer tool standing times and higher machining quality.

4.4

Coatings

135

Residual Stresses in CVD Coating Systems As a rule, coatings deposited by means of the HT or MT-CVD process exhibit tensile residual stresses. This is, however, no longer exclusively true. In the meantime, there are also CVD systems with compressive residual stresses available. Tensile residual stresses originate when cooling from coating temperature to room temperature and result primarily from the difference of the thermal expansion coefficients between the substrate and the coating or, in the case of multilayers, between the individual coatings. In addition, stress-free annealing in the substrate cemented carbide always occurs at the coating temperatures typical for the CVD process, leading (as opposed to the low temperature PVD process) to a reduction of the (compressive) residual stresses introduced, for example, during grinding [West00]. Coatings with compressive residual stresses, generally familiar in the case of PVD coatings, have the advantage that crack development is hindered in the hard material coating, and the fracture strength of the coating/substrate composite is largely maintained. Compressive residual stresses are therefore the most beneficial to performance in interrupted turning and in milling in general. Compressive residual stresses can also be produced when using the CVD method by a corresponding choice and arrangement of individual coating systems. One example is the coating sequence TiN-Ti(C,N)-Al2 O3 . The outside coating composed of Al2 O3 has the smallest thermal expansion coefficient, the inside TiN coating the largest. After cooling from coating to room temperature, the Al2 O3 coating exhibits a small amount of compressive residual stress. Considerable larger compressive residual stress is measured with the same coating sequence if it is supplemented externally with a further Zr(C,N) or Hf(C,N) coating (ratio C:N = 1:1) (Fig. 4.30). PVD-Coating system

CVD-Coating system Internal stress (N/mm2)

17 µm

–4000

–2000

–3000

0

–1000

–3000

–2000

0

Pressure

4 µm

Hf(C,N) (MT - CVD)

Internal stress (N/mm2)

Pressure –1000

1000

Strain

TiAlN

Al2O3

Ti(C,N) (MT - CVD)

Cemented carbide substrate

Cemented carbide substrate

Fig. 4.30 Influence of coating methods and coating structure on residual stresses in hard coating systems

136


To produce these multilayer systems, the classic high temperature CVD process is combined with the MT-CVD process, whereby the precipitation of the carbon nitride coatings takes place with the help of the medium temperature CVD process [Drey97, West00, Berg03, Berg05].

The Plasma Activated CVD Process In the PA-CVD process (“plasma assisted” or “plasma activated”) – also known in the literature as PE, “plasma enhanced” – precipitation of fine-grained hard material coatings takes place at the, in contrast to HT and MT-CVD methods, much lower coating temperatures of 450–650◦ C. At these temperatures, the thermal energy alone is insufficient for introducing the chemical reaction needed to form the hard materials from the gaseous phase. For this reason, additional energy is added to the process by means of the pulsed plasma of a low pressure glow discharge. In this way, chemical reactions are made possible that, at thermodynamic equilibrium, are only possible at much higher temperatures [Tabe89, Köni89]. As in the HT-CVD process, coatings made of titanium nitride, titanium carbide, titanium carbon nitride and aluminium oxide can be deposited individually in crystalline structure by changing the gas composition with individually or in an arbitrary sequence. Due to the non-equilibrium conditions during deposition, there is also a possibility of synthesizing metastable systems like (Ti,Al)N using PA-CVD [Lemm03]. The attributes of the cemented carbide remain largely unaffected by the lowering of temperature during coating. The residual stresses are in the compressive range after coating. PA-CVD coated cemented carbides measure about 30% higher bending strengths than HT-CVD coated ones. The properties of PA-CVD coated cemented carbides have a positive effect on their performance when cutting strong steel materials in interrupted cut. The sensitivity of the composite to comb cracking and failure due to crumbling is much lower than it is the case for HT-CVD coated cutting tool materials. The application example “side milling” proves this (Fig. 4.31). As opposed to the uncoated substrate, the tool life travel path is slightly diminished after HT-CVD coating due to thermal influence on the substrate as well as tensile residual stresses in the hard material coating, while it is over 300% larger after PA-CVD coating. Modifying the Layer Structure The layer structure of many multilayer systems usually begins with a thin titanium nitride layer coating the substrate, followed by one or more titanium carbon nitride intermediate layers. The TiN surface layer functions as a diffusion barrier during coating and prevents brittle eta phases forming in the interface [Köni90]. The Ti(C,N) layers building upon the TiN surface layer can consist of one layer or several layers differing in stoichiometry. They act as a kind of “buffer” between the relatively soft and tough substrate and the much harder and more brittle carbides, nitrides or oxides of the coating.

4.4

Coatings

137 16

Tool life travel path Lf / m

Side milling cutter

12

Material: 60WCrV8, 220 HB Cutting tool material: HW-P25

8

vc = 80 m/min fz = 0.08 mm ae = 42 mm ap = 6 mm

4

0 Cemented carbide P25 / TiN, PA-CVD-coated

Uncoated

HT-CVD

PA-CVD

Coating method

Fig. 4.31 Cutting edge durability of a PA-CVD-coated cemented carbide in side milling (Source: Widia)

With increasing coating thickness, the wear resistance of the CVD-coated cemented carbides increases, but their bending strength and toughness is reduced. The toughness of the coating/substrate composite is affected not only by the thickness of the entire coating, but also to a large extent by the coating material and the thickness of the individual coating components. Figure 4.32 illustrates this using the example of a coating made of TiC and Al2 O3 . With a constant total coating thickness of 10 μm, the failure rate of the coated cutting tool material due to fracture clearly increases with the thickness of the Al2 O3 layer [Naka88]. Due to its high chemical stability as a coating material, aluminium oxide does indeed have excellent resistance against crater wear, but it only has a small amount of thermal shock resistance and toughness. In order to reduce the brittleness of coating systems containing aluminium oxide, they are often designed in the form of fine lamellar, fine grained multilayer coatings (Fig. 4.32). Such multilayer coatings consist usually of a combination of TiN or Ti(C,N) with Al2 O3 or, as in Fig. 4.32, with aluminium oxynitride (AlON). Coatings such as these can also be composed of a large number of individual layers, whereby each layer can be thinner than 0.2 μm. Total coating thicknesses, depending on the intended use, are around 5 μm in the case of milling and between 12–20 μm for turning. Using PA-CVD coating technology, we are now also capable of depositing nanolayer coatings. These are multilayer coatings with individual layers whose thickness is given in nanometres, not in micrometers (Fig. 4.33). In the case of coating systems 1 and 2 shown in Fig. 4.33, sixty 80 nm thick layers and thirty-six 50 nm thick layers are deposited on a top layer composed of titanium nitride. In


Wear resistance

138

PVD CVD TiN AlON

Coating thickness Bending strength

PVD PA-CVD

TiN AlON

CVD TiN Coating thickness

Single layer thickness Al2O3 / TiC

CVD − coating Al2O3 / TiC Total coating thickness for all tools 10 µm

AlON TiN AlON

1 µm / 9 µm 4 µm / 6 µm

Transistion layers

6 µm / 4 µm Substrate

3 µm

0 10 20 30 40 50 60 70 80 90 100 Occurance of failure / (%) (cutting edge breakage)

Fig. 4.32 Fine lamellar multilayer coating (Source: Widia, Sumitomo)

the case of the coating system TiN-AlN, there are as much as 2000 single layers of titanium and aluminium nitride of 9 nm thickness that are applied alternately. Lowering the thickness of the coating increases hardness, toughness and wear resistance. This effect is based, among other things, on the fact that coating materials below a certain thickness take on completely new properties. For example, AlN is, with a hardness of about 1200 HV0.05, relatively soft in comparison to hard materials like TiC, however it has a hardness of over 3000 HV0.05 when the coating thickness is below 10 nm. At high cutting speeds or during dry machining, demands on thermal resistance of the tool cutting edge are increased. Due to their low thermal conductivity, oxide coatings are especially suitable to meet these demands. As can be seen in Fig. 4.25, in addition to aluminium oxide, both zirconium oxide ZrO2 and hafnium oxide HfO2 have very low thermal conductivity. Since both oxides are insufficiently hard by themselves, they are combined with Al2 O3 [West00]. This can be accomplished in the form of multilayer coating systems with single or multi-phase layers, whereby the latter – so-called composite coatings – are finely distributed in the Al2 O3 base matrix of a further oxide, e.g. HfO2 oder ZrO2 (Fig. 4.34). In the case of the Al2 O3 /ZrO2 /TiOx composite coatings shown in Fig. 4.34, the oxide layer, consisting of three components, is condensed from a gas mixture comprising AlCl3 , ZrCl4 , TiCl4 , CO2 , H2 , N2 and Ar. Al2 O3 then exists as a κ-phase and ZrO2 as a phase mixture in tetragonal/monoclinical (coating temperature > 990◦ C)

4.4

Coatings

Coating system: TiN-30x(TiCN-TiN) Coating thickness: 5.2 µm Single layer thickness: approx. 80 nm

139

Coating system: TiN-18x(Al2O3-TiN) Coating thickness: 2 µm Single layer thickness: approx. 50 nm

Coating system: 1000x(TiN-AlN) Coating thickness: 9 µm Single layer thickness: approx. 9 nm Hardness of coating: 3200HV0,05

Fig. 4.33 Nanolayer created by a PA-CVD coating process (Source: Kennametal Widia)

or monoclinical form (coating temperature < 990◦ C) depending on the coating temperature. In the fracture structure image, the ZrO2 particles can be recognized as finely distributed in the Al2 O3 matrix. The highly fine-grained coating structure is originated by means of a small addition of titanium as doping agent. Doping layers made of Al2 O3 with titanium or boron (or TiN layers with boron) leads to the formation of very fine-grained coating structures. In the CVD process, titanium is added to the gas mixture as TiCl4 and boron as BCl3 . The doping products TiOx und Bx Oy , well detectable in the EDX (energy dispersive X-ray) analysis but still not clearly identifiable as a phase, are effectively dissolved in the Al2 O3 lattice. Due to their nanodispersive distribution, they have a positive effect on coating growth and the grain size of the hard materials forming the coating system [West01, Kloc01, Berg03]. Doping brings about higher deposition rates, makes it possible to reduce deposition temperature and clearly improves wear resistance at high thermal cutting edge loads (Fig. 4.35) [Kath03]. Aftertreatment Aftertreatment helps to improve further the performance characteristics of coatings, especially in the case of interrupted cut. For example, smoothing the surface of CVD-coated tool bodies by means of brushing, jet treatment or polishing is

140


HfO2

1 µm 1.5 µm

Al2O3 /HfO2 /TiOx

0.5 µm

6 µm

Ti(C,N) (MT-CVD) TiN

(12–13 µm)

9.5 µm

Al2O3 with dispersed HfO2

9.5 µm

TiCN (MT-CVD) TiN

(16–18 µm)

Al2O3 /ZrO2 /TiOx Composite

7 µm

TiCN (MT-CVD)

5 µm

TiN (12–17 µm)

Fig. 4.34 Fracture structure of CVD-coated cemented carbides with modern coating-design (Source: Kennametal Widia) TiN A TiN

B Ti(C,N)

MK reduced rim zone

A

Single layers of Al2O3 and TiN endowed with titan and boron

B

Single layers of TiN and Ti(C,N) endowed with boron

Fig. 4.35 Formation of microcrystalline layer structures via doping with titanium and boron (Source: Ceratizit)

4.4

Coatings

141

the common practice today for removing potential cracking sources. Since brushing only covers the cutting edge area, all-round smoothing with jet treatment is being given increasingly more preference. It is of particular importance that the jet treatment, in addition smoothing, also causes an alteration of the tensile residual stresses as a function of the treatment parameters. For example, the Al2 O3 layer in the Al2 O3 /ZrO2 /TiOx coating system has, after jet treatment, compressive residual stresses (–3.500 N/mm2 ) in place of tensile residual stresses (1000 N/mm2 ). Jet treatment not only improves the residual stress condition, but also reduces the range within which the wear and standing time of coated hard material indexable inserts are scattered [Berg03].

Modifying the Substrate Rim Zone In comparison to uncoated cemented carbides, CVD coatings lengthen standing times enormously. On the other hand, they also lower the toughness of indexable inserts, thereby also reducing their capacity to bear high stress peaks during cut interruptions [Köni90, Köni92]. The reason for this is assumed to be cracks, which spread from the hard tensile stressed material coating into the cemented carbide substrate. Attempts to improve the toughness and resistance of cemented carbides substrates to crack development without reducing their hot hardness led to the development of substrates with graded rim zone structures [Sche88, Köni90, Schw89]. The goal was to modify the rim zone of the cemented carbide such that crack development towards the centre of the indexable insert was hindered or suppressed. One possible solution are cemented carbides, the rim zones of which are largely lacking hard and brittle (Ti, Ta, Nb)C mixed crystals and consist practically of only tungsten carbide and cobalt up to a depth of approx. 50 μm (mixed carbidefree or reduced rimzone, Fig. 4.35). The coating and application of such cemented carbides is presently state of the art. Further developments in this area are aimed at establishing the basic properties of a coated cemented carbide body – hard outside and tough inside – already during sintering [Leng04]. This would also have positive effects on the following coating process. The fabrication of function-graded cemented carbides with rim zones lacking mixed carbides and with external or internal Ti(C,N) enrichment is described in more detail in Sect. 4.3.5.

4.4.1.2 PVD Coating The early 1980s saw the succession of the CVD process by the PVD process in various process variants (vacuum evaporation, Arc-PVD, sputtering). The initial benefit of these methods was the possibility of coating HSS tools with complex geometries. In the meantime, these processes are also being used for cemented carbide and cermet coating. In addition, decorative layers, oxidation layers and corrosion layers can also be produced.

142


The fundamental differences between PVD coating and the classic HT-CVD process can be summarized as follows [Sche88, Schi89, Köni90, Rödh87, Knot89, Gühr89, VDI3198, VDI3824]: • The process temperature of 160–600◦ C means relatively low temperature stress on the substrate materials, so that temperature-sensitive substrate materials can also be coated. • The bending strength of the substrate remains largely unaffected by the low coating temperature. • PVD coatings exhibit residual compressive stresses, which currently limit the possible layer thicknesses to 3–6 μm. Residual compressive stresses reduce the risk of fracture formation under shock lead and thermal cycling. • In the PVD process, workpieces that are to be coated require a very careful pretreatment of the surface and process control in order to guarantee sufficient adhesion of the coating. CVD coatings have better adhesion due to the effects of evaporation and diffusion. • Because of shadowing effects, consistent layer thicknesses can only be partially realized by means of very costly devices for rotating the parts to be coated. Inner contours are as a rule only coatable up to a depth/diameter ratio of 1, since the layer thickness is diminished with increasing depth. • High number of potential coating systems and substrate materials. • Besides stable coating systems such as TiN and Ti(C,N), multi-component metastable phases can also be deposited under the non-equilibrium conditions of physical vapour deposition. These phases often exhibit structures that do not exist under equilibrium conditions. An example of this is the coating system TiAlN. We can distinguish several different methods of PVD coating: • methods with thermal evaporation (vacuum evaporation) • methods with arc evaporation (arc-PVD) • methods with cathode evaporation (sputtering) Thermal Evaporation (Vacuum Evaporation) In vacuum evaporation, usually the coating material is evaporated in a crucible with a resistance heater or an electron beam gun (EB-PVD) in high vacuum at a pressure of 10–3 –10–6 Pa (Fig. 4.36). The vapour atoms exhibit average free path lengths of up to several meters at these pressures. Therefore, they generally do not interact with each other and proceed straight to the substrate. Since the latter is considerably colder than the vapour, the particles condense on the substrate in accordance with the principle of water vapour. A columnar fine crystalline layer is formed [Möhw96]. Due to the straight path of the particles, the substrate material must be moved in the coating chamber in order to avoid shadowing effects and inconsistent coating thicknesses. Vacuum evaporation allows for high coating rates (coating thickness increase per unit time) and is used above all for optical applications, in mass production for large-scale metallization of films, plastic parts and paper as well as for coating turbine blades (EB-PVD) [Sche88, Rass96].

4.4

Coatings

143

Gas inlet

Target Watercooled recipient

High vacuum coating plant evaporation by: - resistance heating or by - electron beam

Electron beam

Evaporator

Vacuum pump

Energy supply

Fig. 4.36 PVD-Process – Vacuum evaporation, acc. to Möhwald [Möhw96]

In order to form hard material coatings composed of carbides, nitrides or oxides, the substrate is reactively vapour-deposited with the help of a reaction gas (N2 , CH2,. . . ). [Haef87]. For example, from 2Ti + C2 H2 the hard material layer 2TiC and the gas H2 are formed. In the case of simple reactive evaporation, bonding strength is relatively low due to the low particle energy of less than 2 eV. Arc Evaporation (Arc-PVD) In the arc-PVD method, the target material is vapour-deposited and ionised by means of an arc in high vacuum. The arc travels either randomly (random arc) or in a controlled fashion (steered arc) by means of the vaporizing source, usually connected as a cathode (Fig. 4.36). This method is distinguished by its high degree of ionisation of up to 90% and thus makes it possible to produce high quality coatings with thick structures and good adhesive properties. Due to the high energy density in the focal point of the arc, the material is explosively evaporated and propelled, but molten particles are also torn out at the same time. The liquid particles that are carried away are found again as dropshaped defects (“droplets”) in the layer [Bran05]. These are undesirable because they adhere to the substrate insufficiently [Möhw96]. The appearance of droplets is

144


the main disadvantage of arc evaporation, and since the droplets can lead to pores in the coating, this method is not suited to depositing corrosion-resistant coatings [Möhw96]. Droplet emission can be partially reduced by optimizing the process parameters. Since evaporation by means of cathode arcs requires that the coating material is electrically conductive, no oxide coatings can be deposited with this method. One process variant is the LARC method (Lateral Rotating Arc Cathodes). The basis of this method are rotating, water-cooled arc cathodes [Damm04]. The magnetic field is created by permanent magnets and coils that are shifted and steered vertically and radially. In contrast to conventional arc technology, fewer droplets are produced, resulting in smoother coatings [Csel03].

Cathode Evaporation (Sputtering) In a low-pressure plasma, an inert gas (e.g. argon) is ionized by applying high voltage. The positively charged inert gas ions are propelled onto the target switched as a cathode (coating material) and knock out atoms, atomic groups and molecules of the coating material via impulse exchange [Rass96] (Fig. 4.37). Gas inlet

Target Sputtering plant

- without additional magnetic field: Diodes-Sputter-Source - with additional magnetic field: Magnetron-Sputter-Source

Electrode as additional ion source

Plasma-room Evaporator (Target) Additional magnetic system Vacuum pump +

-1...5 kV

Fig. 4.37 PVD-Process – Cathode evaporation (sputtering)

4.4

Coatings

145

As opposed to vacuum evaporation, sputtering requires a process gas (usually argon) as well as a high voltage source. After evacuating the recipient, the process gas is admitted at up to 0.1–1 Pa pressure. The average free path length of the particles at this pressure amounts to a few millimetres. At a target voltage of –1 to –5 kV, a glow discharge plasma is created. The plasma burns between the target which acts as an anode and the substrate connected to the ground or the recipient wall, thus making it the anode. The plasma consists of ions, electrons and uncharged atoms [Möhw96]. This evaporation method is universally applicable, since the material is evaporated not thermally but by means of impulse transfer. Almost all substances [Haef87] and composites of very diverse melting substances can be evaporated, i.e. not only metallic, but also isolating layers as well as materials with high melting points can be deposited. If not only metals or metal alloys are to be deposited, but also carbides, nitrides, or oxides, this can be achieved with reactive sputtering. In this case, a reaction gas (e.g. N2, CH4 , O2 , H2 , . . .) is injected in addition to the process gas. From this reaction gas are obtained the non-metallic components of the coating (e.g. TiN, TaC, Al2 O3 , . . .). As in the arc method, sputtering also involves the possibility of depositing metastable coatings (i.e. coatings that do not exist in equilibrium). Because the process conditions are far from equilibrium conditions, it is possible to freeze metastable composite phases due to their rapid condensation from the gaseous phase, thus synthesizing a completely new material type that is often superior to stable boundary systems [Leye04]. We differentiate between DC sputtering, in which the sputtering material is electrically conductive, and RF sputtering, in which there is a high-frequency AC field that is used for non-conductive materials. The use of direct current (DC) plasma, high-frequency (RF) plasma and the combined use thereof as well as a nearly unlimited number of possible coating materials making sputtering enormously flexible with respect to the development of application-specific component coatings [Rass96]. Although RF sputtering has important advantages when atomizing insolating materials, this technology is of lesser importance in the creation of wear-protection coatings due to the appreciably higher capital costs in comparison with DC sputtering and the significantly lower deposition rates [Lemm03]. The disadvantages of RF sputtering can be surmounted by using pulsed DC sputtering (MF: medium frequency sputtering, up to 50 kHz). The coatings produced with sputtering at low vapour-deposition rates have good adhesive properties. Magnetron Sputtering In comparison to pure diode sputtering, magnetron sputtering involves the arrangement of permanent magnets behind the target (Fig. 4.37), the magnetic field lines of which vertically penetrate the cathode plate and the target affixed onto it. Due to the overlapping of magnetic and electric fields, the electrons in the plasma are

146


directed on extended spiral-shaped paths. In this way, the probability of ionisation is increased as well as the deposition rate as the discharge voltage is reduced. It was the invention of magnetron sputtering that has first made possible an economic utilization of the sputtering effect for coating deposition and thus the commercial deposition of wear-protective layers by means of sputtering technology [Lemm03, Bobz00].

High-Ionisation Pulsing Rising interest in isolating, mostly oxidic wear-protection coatings and the necessity of increasing ion density in the area of the substrate in order to deposit layers of dense, optimized morphology have led to the development of pulsed DC plasmas. With this technology, not only coating systems usually reserved for RF technology can be deposited, but completely new coating systems can also be synthesized due to the extremely enhanced ion currents [Lemm03, Leye04]. Pulse technology makes it possible to deposit conductive and electrically isolating layers in almost arbitrary stoichiometry as well as nano-structured single and multi-phased layers with new property combinations. Figure 4.38 shows a (Ti,Al)N coating created with conventional DC technology in comparison to one made with the help of pulse technology. One can clearly recognize the alteration of the coating micromorphology towards finer crystalline structure made possible with the increased plasma ionisation of pulse technology [Lemm03]. Pulse technology thus also creates the possibility of depositing crystalline γ-Al2 O3 at substrate temperatures of < 500◦ C or producing (Ti,Al)N layers with a AlN content of up to 66 mol% (Fig. 4.39). The coating process with pulsed plasmas is also known as high ionisation pulsing (H.I.PTM ). We distinguish between three process variants: unipolar, asymmetrical bipolar and bipolar two-cathode magnetron sputtering [Lemm03, Erke05].

Fig. 4.38 Fracture pattern of (Ti,Al)N-layers on cemented carbide substrates, deposited with conventional DC-technique (left) and using bipolar-pulse-technique (right) (Source: CemeCon)

4.4

Coatings

147 900

Fig. 4.39 Metastable phase diagram of the system TiN-AlN, acc. to Cremer [Crem98]

800

(Ti,Al)N+ (Al,Ti)N

700

T / °C

600 500 400

(Ti,Al)N (NaCl structure)

300 200 (Al,Ti)N (wurtzite structure)

100

TiN

0,2

0,4

0,6

0,8

AlN

xAlN

Ion Plating With respect to process control, ion plating is defined as a variant of PVD coating, in which the condensing layer is altered by means of an additional electric field applied to the substrate. This method variant is independent of the type of process and physical principle involved in creating the metal vapour. The prerequisite is a sufficiently high number of ionised steam and process gas atoms. In ion plating, the substrate is coated with a negative voltage or “bias voltage”. The metal vapour is ionized by means of electrodes arranged in a gas space and electromagnetic fields. Part of the ionised particles are accelerated towards the substrate (Fig. 4.40). Ion plating thus involves surrounding the substrate with gas ions. By bombarding the growing layer with ion, coating properties such as density and residual stress are modified and adhesion to the substrate is improved. The resultant coatings are very adhesive and tight. All PVD methods in which a bias voltage is used to modify the coating are termed “ion plating” [Bobz05]. In order to produce carbides, nitrides or oxides for coating purposes, here too a reactive gas is fed into the chamber. This method is called reactive ion plating [Möhw96, Haef87]. Ion plating can also be utilized for purifying and activating the component surface prior to coating (also called etching) or for activation during coating [Rass96]. A variety of ion plating plant designs have emerged as practicable coating variants. For tool coating, arc ion plating (AIP) and ion plating with low-voltage arc discharge are among the process variants that have become the most significant in tool coating.

148


-

+ 10...5000 V

Rotating arc plasma channel

Vaporising via vacuum arc vaporiser: Contact ignition device

Magnets for stabilisation rather leading of the arc Vacuum pump

Fig. 4.40 PVD-Process–Ion plating, acc. to MÖHWALD [Möhw96]

Arc Ion Plating (AIP) In the AIP process, the target material is vaporized and ionised by means of an arc in high vacuum. This method is characterized by the high degree of ionization of up to 90% and can thus be used to produce high quality coatings with dense structures and good adhesion. The pre-voltage applied to the substrate accelerates the ionised particles, which impact the substrate at high speed. Their energy is partially converted into heat, which can result in high coating temperatures (approx. 550◦ C) in the substrates.

Magnetron Sputter Ion Plating (MSIP) MSIP guarantees a stable plasma and high atomization rates [Bobz00]. Coatings deposited with MSIP exhibit very good mechanical and optical properties. Nearly all materials are depositable with this method. Since the ionisation degree of the atomized particles is under 5%, the substrates are heated to a much smaller degree than in the AIP process, so that plastics and materials with low annealing temperatures can also be successfully coated [Bobz00].

4.4

Coatings

149

Low-Voltage Arc Coating Ion plating with low voltage arc discharge is characterized by a low acceleration voltage, with which an electron jet is created from a low-voltage arc discharge. This jet is directed towards the target (anode) in a water-cooled crucible. The vapour made of the coating material arising from the electron bombardment is about 50% ionised and, since the substrate is applied with a voltage of –200 V, directed towards the coating target. This method is ideal for depositing nitride or carbide layers by introducing the corresponding reactive vapours [Haef87]. No droplets are formed when ion plating with low-voltage arc discharge is used for coating. This method thus makes it possible to deposit very smooth layers (Rz = 0.1 μm) and is thus suitable for coating mirror-polished surfaces without the need for post-polishing. The Use of PVD-Coated Cemented Carbides at Low Cutting Speeds Broaching soft and hardened steels offers an impressive example for the cutting edge durability of PVD-coated cemented carbides at low cutting speeds [Köni92a]. Characteristic of the broaching process are the generally low applicable cutting speeds, usually 1–25 m/min depending on the broaching machine, cutting tool material and the target material. Built up edge formation is characteristic of machining relatively soft or tough steels at low cutting speeds. These lead to dimensional and shape inaccuracies as well as to reduced workpiece surface quality. One measure for reducing built up edge formation among others is the use of broaching oils. By using PVD-TiN-coated cemented carbides in place of uncoated ones, built up edge formation is almost completely suppressed due to lowered adhesion between the flowing chip and the cutting tool material in the entire cutting speed range under investigation of vc = 10–63 m/min (Fig. 4.41). Furthermore, the use of coolants – a constant ecological problem whose maintenance and disposal are ever costlier – is superfluous.

4.4.2 Specific Properties of Hard Material Coatings The most commonly utilized wear-protection coatings in metal machining commercially are TiN, TiC, Ti(C,N), Ti(Al,N), CrN and Al2 O3 . Besides these coating systems, those that are characterized by a low friction coefficient such as DLC (diamond-like carbon) and MoS2 -coatings are successfully used for wearprotection. Due to the unsurpassed hardness of diamond, CVD diamond coatings are also being increasingly used [Lemm03]. Different coating systems can be deposited as monolayers (e.g. TiN, Ti(C,N), Ti(Al,N), (Ti,HF,Cr)N) or, to improve their wear-protective effect, combined as multilayers with single layer thicknesses in the micrometers (e.g. TiN-Ti(C,N)Al2 O3 -TiN, (Ti,Al)N-WC/C).

150 vc = 63 m/min broaching oil

vc = 10 m/min dry

4 Cutting Tool Materials and Tools Uncoated

vc = 63 m/min dry

PVD-TiN-coated

vc = 63 m/min dry

0.2 mm Material: 17CrNiMo6+FP Cutting tool material: HW6Co Chip thickness: h = 0.1 mm

Tool geometry:

αo

γo

λs

rβ

5°

9°

0°

35 μm

Fig. 4.41 Coating prevents formation of built up edges

The primary goals of multilayer technology are not only to improve the toughness, hardness and wear resistance of the coating but also to increase the total thickness of the coating. The transitions that are formed between layers of different materials place limits on shifting, which leads to a large increase in strength, especially in the case of coatings with very diverse Young’s modulus. For cracks that are growing in the coating, transitions between the individual layers signify barriers at which they are forced into energy-dissipating detours, slowing or even stopping further growth (Fig. 4.42). The multilayer structure leads, in comparison to an equally thick monolayer coating, to lower residual stresses. Multilayer coatings also make it possible to integrate layers with different properties within one coating system, thus combining the functionalities of diverse coating materials. Such specific coating functions include, for example, high thermal stability, high hot hardness or the low friction resistance of intermediate or cover layers [Krie01, Lugs03]. Larger layer thicknesses mean a larger coating volume that has to be worn. From this fact results the generally longer standing times as well. One further development in multilayer technology is the deposition of nanolayers or coatings with superlattice structures. These are coating systems consisting of a large number of extraordinarily thin layers, the thickness of which amounts to

4.4

Coatings Crack growth down to the substrate

151 Crack growth along the grain boundaries

Substrate

Substrate

Coating with low internal residual stresses

Nanocrystalline structures

Crack growth along the boundary surfaces

Crack growth is hindered

Substrate

Substrate

Multilayer

Coatings with high residual stresses

. Ductility of the coating is as important as the hardness of the coating . Balance between high residual compressive stresses (bad coating adhesion) and low blank (no obstruction of crack growth) ones

Fig. 4.42 Schematic diagram of the crack growth in hard coatings depending on structure and properties (Source: Oerlikon Balzers)

only a few nanometres (Fig. 4.43). In contrast to conventional, usually columnar layers, coating materials deposited as nanolayers are characterized by their significantly higher hardness. As has already been explained in the context of the PA-CVD coating process, this is based on the fact that the hardness of a material increases significantly below a certain layer thickness. This increase in hardness is explained by the phenomenon that diminishing layer thicknesses have altered crystal lattices and thus Young’s modulus as well. For example, AlN has a hexagonal lattice structure at layer thicknesses > 10 nm and a cubic structure when the layer thickness is < 10 nm [Csel03a]. Another way to improve a coating’s properties is to produce so-called nanocomposites. These are noncrystalline isotropic multiphase systems, in which two mutually insoluble phases (e.g. Al, Ti, Si) are deposited during the coating process on the tool surface. Examples of this are the embedment of nanocrystalline cubic TiN in an AlN matrix, of nanocrystalline TiAlN/AlCrN in an amorphous Si3 N4 matrix [Csel04] or of nanocrystalline (Ti,Al)N into a matrix composed of (Al,Ti)N. The nanocomposites can be deposited as monolayer or multilayer coating systems. The boundary layers/grain boundaries in the nanolayer/nanocomposite layers are energy-dissipating barriers to cracks. Cracking and the speed of crack growth are thereby reduced (Fig. 4.42). Both nanolayer and nanocomposite layers are thus characterized not only by extremely high hot hardness and high-temperature wear resistance but also by favorable toughness attributes. With respect to the structure of hard material coatings, we differentiate between monophase layers (e.g. TiN), multiphase layers (e.g. TiN + Ti2 N) and graded layers, i.e. layers the chemical composition of which exhibits a gradient (e.g. (Ti,Al)N layers with an increasing Al content) [VDI3824].

152

4 Cutting Tool Materials and Tools Monolayer (TiN, Ti(C,N), (Ti,Al)N, etc.)

Super hard layer (CVD-Diamant, Bornitrid)

Multilayer with functional intermediate layers (TiN-Ti(C,N)-Al2O3-TiN, TiN-Ti(C,N)-Al2O3-Zr(C,N), etc.)

Hard layer with shear-flexible cover (TiN + MoS2, (TiAl)N + WC/C, etc.)

Nanolayer (Superlattice) thickness of one layer: a few atom layers to 100 nm (Ti,Hf)N-CrN

Hard layer + low-friction amorphous carbon layer (DLC-layer) (a-C, a-C:H, a-C:H:Me)

Graded layer Ti(C,N)-Ti(C,N) (Ti,Al)N-(Al,Ti)N

Nanocomposite (nc-TiN/a-AlN, nc-(Ti,Al)N/a-Si3N4)

Fig. 4.43 Schematic diagram of the structure of current CVD and PVD layer systems

The properties of hard material coatings are determined by their chemical composition and structure, which depends on the conditions of their deposition. Figures 4.25 and 4.44 show some characteristic values for layer properties for some exemplary hard material coating systems produced by PVD and CVD processes. By changing the deposition conditions, such properties as chemical composition, morphology, structure, texture, residual stresses and thus microhardness, the thermal expansion coefficient and oxidation resistance can be altered within certain limits. The data presented can thus only be considered as guides for possible areas of application. Besides the hard material coatings here presented and explained in more detail below, there is a number of other hard material layers used for tool coating that are not more fully described here. 4.4.2.1 Titanium Carbide Coatings (TiC) Coating tools with hard materials began with the deposition of TiC. Due to the high hardness realizable with titanium carbide (3100–3400 HV0.05), it provides more effective protection against abrasive wear than TiN. On the other hand, its tendency to diffusion is a bit higher than that of TiN and Al2 O3 due to the relatively small enthalpy of formation. TiC s resistance to flank face wear is thus higher than that of TiN, while its resistance to crater wear is lower. Its oxidation wear is also the lowest of the layers shown in Fig. 4.44. TiC coatings are chiefly applied using the CVD process. For the most part, TiC is used as a monolayer when abrasion is predominant or in multilayer coatings together with TiN, Ti(C,N) or Al2 O3 [VDI 3824]. 4.4.2.2 Titanium Nitride (TiN) Since 1980, tools were already being coated with titanium nirtide. TiN is the most frequently employed hard material for coating cutting tools today. The coating material is an intercalation compound composed of titanium and nitrogen.

4.4

Coatings

153

TiN Production process

TiCN

PVD/CVD PVD/CVD

TiC

TiAlN

CrN

Al2O3

CVD

PVD

PVD

CVD/PVD

Coating thickness/µm

1 to 5

1 to 5

1 to 5

1 to 5

1 to 10

1 to 5

Microhardness/HV 0,05 1)

2300

3000

3100

3000

1900

2100 HV 0.1

Oxidation temperature/°C 2)

> 450

> 350

> 350

> 700

> 600

–3)

Thermal barrier effect 4)

+++

++

+

++++

+

+++++

Resistance to abrasion

++

+++

+++

+++

++

++

Resistance to wear due to adhesion (against steel)

++

++

+

++

++

+++

Resistance to wear due to diffusion (to steel)

++

+

+

+++

++

+++

Protection of basis material against corrosion 5)

+

+

+

+

++

+

1) For the microhardness, mean values are indicated. They are obtained from reported measured values resulting from different compositions, coating thicknesses and internal stresses. 2) The oxidation temperature is the temperature at which oxidation of the coating material begins, considerably affecting the characteristics of the coating. 3) Al2O3 is already an oxide. 4) Being poor conductors of heat, the coatings act as a thermal barrier to the heat produced during metal-cutting so that most of the heat can be removed through the chip. 5) As the aforementioned hard coatings themselves do not corrode, they protect the basis material from corrosion. (Leaks in hard coatings may lead to the development of local galvanic elements and to pitting corrosion).

Fig. 4.44 Quantities Characterising the coating and the coating performance, acc. to VDI 3824/1

Due to the high interaction between the metal and nitrogen atoms, this compound is highly stable. With an enthalpy of formation that is almost twice as high as that of TiC, TiN is thermodynamically more stable, and thus more diffusion resistant and less inclined to adhesion. Therefore, TiN’s resistance to crater wear is higher than that of TiC. TiN coatings are characterized by a high level of toughness. TiN can be deposited with both the CVD and PVD process. The characteristic colour of TiN is goldish yellow. Silver-hued TiN coatings consist of a phase composite of Ti2 N and TiN, whereby the Ti2 N phase is dominant. These coatings are harder but also more brittle than pure TiN layers. They are most commonly used when abrasive wear predominates [VDI 3824].

154


4.4.2.3 Titanium Carbonnitride Coatings (Ti(C,N)) TiC and TiN can be mixed at any ratio. The properties of titanium carbonnitride are adjustable by varying the C/N ratio. With increasing carbon content, the colour of the Ti(C,N) layers change from coppery to violet, bluish-grey to grey [VDI 3824]. Titanium carbonnitrides are used industrially both as hard materials in cemented carbides and as wear-resistant thin-films. Often Ti(C,N) coatings are multilayered, i.e. deposited with increasing carbon content in the direction of the coating surface. By integrating carbon atoms in place of nitrogen atoms into the titanium nitride crystal lattice, a considerable increase in hardness can be realized, which is positive for wear resistance but also increases brittleness [Sato78, Schi74, Berg90]. To compensate for this increase in brittleness, Ti(C,N) coatings are deposited as multilayers, so that residual stresses between the individual coating layers can be reduced. Ti(C,N) coatings are suited for machining steels with high tensile strength and thus for higher cutting temperatures. 4.4.2.4 Titanium Aluminium Nitride Coatings ((Ti,Al)N) The (Ti,Al)N coating system was developed in order to improve the oxidation resistance, hot hardness and wear-protection properties above the levels of previously used coatings [Quin87, Knot87]. In comparison to TiN and Ti(C,N) coatings, (Ti,Al)N coatings have the highest oxidation resistance with a comparably high level of hardness. Since Ti(Al,N) is a metastable coating system, it can only be deposited with the PVD, PA-CVD or MT-CVD process. (Ti,Al)N coatings are a further development of TiN, whereby titanium is substituted by aluminium by 20–60 at.-%. Depending on the composition, these layers range from brown (lower Al-content) to black-violet (higher Al-content) [VDI 3824]. Due to their high level of oxidation resistance and hot hardness, the preferred areas of application are dry machining, hard machining and HSC machining. The excellent wear resistance of (Ti,Al)N coatings is explained by the fact that, as opposed to stable composite phases, metastable layers decompose into stable boundary phases or form stable oxides in an oxidizing atmosphere as long as the energy required to convert into an equilibrium condition is added (e.g. as heat), as it is the case in dry machining or high-speed machining. In this context, the high oxidation resistance of (Ti,Al)N is derived from the fact that a thin aluminium oxide layer is formed on the coating surface which is constantly renewed during the cutting process, thus decelerating the progress of wear. With increasing Al-content, the oxidation resistance of the (Ti,Al)N coating is increased. Despite its much further improved oxidation resistance compared to TiN and Ti(C,N) coatings, this coating also fails beyond approx. 800◦ C. TiAlN coatings are deposited as monolayer, multilayer or gradient layers [Lemm03, Leye04, VDI 3824]. (Ti,Al)N is one of the most commonly used high-performance coating systems. Important measures being taken to further augment the cutting edge durability of (Ti,Al)N coatings include the development of nanolayers and increasing the content of aluminium. The primary goal of these developments is to improve their hot

4.4

Coatings

155

hardness, wear resistance and oxidation resistance. In order to differentiate them from conventional (Ti,Al)N coatings, those with over 50% aluminium content are designated as (Al,Ti)N coatings. 4.4.2.5 Al-Cr-N Coatings Besides increasing the aluminium content another approach to further improving the oxidation resistance and high-temperature properties of (Ti,Al)N coating systems is adding small amounts of oxide formers such as Cr, Y or Si. One example for this direction in coating development is the Al-Cr-N coating system. In comparison to the conventional (Ti,Al)N coating, the (Al,Cr)N coating system has proved to have a higher resistance to abrasive wear as well as higher hot hardness and oxidation resistance [Gey04]. Its areas of application include the dry, HSC and hard machining of steel and non-ferrous metals by turning, drilling and milling [Gey04, Denk04]. 4.4.2.6 Aluminium Oxide Coatings (Al2 O3 ) The exceptional resistance of Al2 O3 against both abrasion and diffusion wear and its simultaneous insensitivity to oxidative wear, known from its use in cutting ceramics, has made Al2 O3 an obvious candidate as a hard material for coating [Sche88]. Because it is very brittle, Al2 O3 is generally not used in monolayer form, but only in combination with other hard materials in multilayer coatings. The electrically nonconductive Al2 O3 layers could thusfar not be fabricated with the PVD, but only with the CVD process for technical reasons. The deposition of the thermodynamically stable high-temperature modification of aluminium oxide – α-Al2 O3 – by means of CVD process has been established industrially for over two decades. Due to its excellent wear protection and high performance potential, attempts are being made to produce Al2 O3 layers with the PVD process as well. A first step in this direction was the deposition of amorphous aluminium oxide (α-Al2 O3 ) with the help of RF sputtering technology. Because of the low coating rates and consequently long process durations, this technique is proven to be uneconomical for coating cutting tools. With the help of the pulsed magnetron sputtering process, it has become possible to coat cutting tools with crystalline γ-Al2 O3 layers at substrate temperatures of 500–600◦ C. γ-Al2 O3 can be applied either as a monolayer or multilayer coating in combination with carbidic or nitridic intermediate layers [Hauz05]. 4.4.2.7 Amorphous Carbon Coatings The concept “amorphous carbon coatings” includes a number of coatings and coating systems. They are often referred to as “diamond-like carbon” or “DLC” coatings. These are carbon-based, highly cross-linked amorphous layers with different amounts of sp2 (graphite bond) and sp3 bonds (diamond bonds) as well as various amounts of embedded hydrogen. In the case of Me-C:H layers, finely distributed metal carbides are also embedded. DLC coatings are manufactured by means of plasma-activated PVD and CVD techniques and deposited at coating temperatures ranging from room temperature to about 300◦ C. Their typical layer

156

Production process Coating thickness in µm Microhardness in HV 0,05 1) Internal stresses in GPa Graphitisation temperature in °C 2) Resistance to abrasion

Resistance to wear due to adhesion (against steel) Protection of bais material against corrosion 3)


MeC:H

a-C:H

a-C

Diamond

PVD

CVD

PVD

CVD

1 to 10

1 to 5

1 to 3

3 to 10

800 to 1800

1500 to 3500

3000 to 7000

10000

0.1 to 1.5

1 to 3

2 to 6

–

350

400

450

> 600

+

+++

++++

++++

+++

+++

+++

(+++) with good cooling

+

+++

+++

+++

1) For the microhardness, mean values are indicated. They are obtained from reported measured values resulting from different compositions, coating thicknesses and internal stresses. 2) The graphitising temperature is the temperature at which amorphous carbon begins to convert from a three-dimensional lattice into graphite, considerably affecting the characteristics of the coating. 3) As the aforementioned hard coatings themselves do not essentially corrode (except for MeC), they protect the basis material from corrosion. (Leaks in hard coatings may lead to the development of local galvanic elements and to pitting corrosion.)

Fig. 4.45 Quanties Characteristing the coating and the coating performance, acc. to VDI 3824/1

thicknesses are 1–5 μm. Figure 4.45 shows some characteristic values for the layer properties of the most important carbon-based hard material layers producible with PVD and CVD available today [VDI 3824, VDI 2840]. In order to improve their adhesive strength, carbon layers are as a rule deposited on a hard material layer (e.g. CrN) as a top layer. Graded intermediate layers made of CrCN have proved favourable for this, in which case hardness and the Young’s modulus can be adjusted to the mechanical properties of the carbon top layer by successive exchange of nitrogen and carbon. This procedure makes it possible to obtain a largely continuous transition from the hard material layer to the carbon layer in order to optimize the bond between both coating systems [Leye04]. Amorphous carbon layers are characterized by extremely low friction coefficients in the case of solid body friction, high wear resistance and low tendencies to adhesion. In cutting technology, these coatings are used above all on tools for dry machining NE metals (e.g. aluminium or magnesium) and in fine machining with great success. Further areas of application are in machining graphite and fibre-reinforced plastics.

4.4

Coatings

157

The property profile of amorphous carbon layers can be purposefully influenced by process control and by integrating additional chemical elements. This potential has led to a number of diverse types of carbon coatings in the last several years. In the following, the essential characteristics and properties of some selected amorphous carbon layers will be described in close detail. The classification, designation and characterization of the coatings are derived from VDI guideline 2840. According to VDI guideline 2840, amorphous carbon coatings can be subdivided into two groups, hydrogen-free and hydrogenous coatings. Since all carbon layers contain a certain amount of hydrogen even without adding hydrogen gas (e.g. from residual gasses), a limit of about 3 at.-% hydrogen is seen as the transition from hydrogen-free to hydrogenous carbon coatings. With increasing amounts of hydrogen, the amount of interconnection between the carbon atoms decreases, leading to softer layers [VDI 3824, VDI 2840]. As wear-protection coatings on cutting tools, hydrogen-free carbon layers are used above all in the form of the amorphous carbon layer a-C and the tertraedic amorphous carbon layer ta-C. The atoms are arranged randomly in amorphous solid bodies. The only bonds are those between a few isolated atoms. If the sp2 bond is predominant in the amorphous carbon layer, it is softer, while it is harder if the amount of sp3 is higher. Depending on the deposition energy, the carbon atoms in the layer are predominantly arranged in one of the two hybrid states. At low deposition energies, the amount of sp2 bonds is higher, making the layers softer. These have the abbreviation “a-C”. With high deposition energies, sp3 hybridizations with a tetraedic arrangement predominate. The layers exhibit a high hardness level, and the compressive residual stresses are increased. Tetraedic amorphous carbon coatings have the designation “ta-C” [VDI 3824, VDI 2840]. Of all amorphous carbon coatings containing hydrogen, it is above all the nonmodified layers that besides carbon contain only hydrogen, designated with “a-C:H” as well as modified carbon layers that are of any importance. The latter contain further elements in addition to hydrogen and are classified in accordance with VDI 2840 into the group of carbon layers containing metal and the group of hydrogenous amorphous carbon layers modified with non-metals [VDI 3824, VDI 2840]. The a-C:H coating system represent the origin of coating with amorphous hydrocarbon layers. In addition to carbon, it contains 10–30% hydrogen. The latter is derived from hydrocarbon gases like acetylene, which are employed in the fabrication process. If, during vacuum deposition, metals are additionally integrated, layers containing metal are produced. As opposed to a-C:H layers, these are electrically conductive and can be manufactured with a technically simpler and cost-efficient DC method. Of all amorphous carbon layers, these therefore have the broadest and deepest range of application. Amorphous carbon layers containing metal are also generally known as “a-C:Me” or “a-C:H:Me” (the abbreviation Me-C:H is also common), where “Me” stands for “metal”. Instead of the abbreviation “Me”, the integrated metals can also be named specifically, e.g. tungsten (a-C:W, a-C:H:W) oder titanium (a-C:Ti, a-C:H:Ti). The added metals form finely distributed carbides

158


with the carbon in the matrix. By integrating metal carbides, layer adhesion can be improved and the tribological properties of the coating can be affected [VDI 3824, VDI 2840]. Hydrogenous amorphous carbon layers modified with non-metals contain nonmetallic elements such as silicon (Si), oxygen (O), nitrogen (N), fluorine (F) or boron (B), which partially also form carbides. By incorporating different nonmetals, further improvements become possible. Silicon for example helps to increase temperature resistance. Moreover, different elements can be built into the layer at the same time. In this way, special layer properties like surface energy (adhesion tendency, wettability) can be altered.

4.4.2.8 Self-Lubricating Coatings Self-lubricating coatings include coating systems which include graphite or molybdenum disulfide (MoS2 ). On cutting tools, graphite and MoS2 coatings are applied as top layers on a supporting hard material layer, where they act as a solid lubricant. Graphite is deposited in combination with WC as a WC/C multilayer coating system consisting of several layers of WC and graphite. Graphite and MoS2 have similar crystal structures. It is characteristic of graphite that its C atoms are arranged in a plane resulting in a very high level of internal bond strength due to the small inter-atomic distance. There is a relatively large distance between the atom layers however, so the layers of atoms can be easily shifted against each other under the influence of external forces. MoS2 also has a lamellar structure. There is a very strong chemical bond between the Mo and S atoms forming each lamella. The bond between two neighbouring lamellae is made between the sulphur atoms. Due to the intermolecular van der Waals bonds between the sulphur atoms of two lamellae, these can be shifted towards one another very easily. These slip planes are responsible for the soft, lubricating effect of graphite and molybdenum sulphide. Both coating systems are used primarily for machining materials that tend to adhere strongly with the tool, for example in the case of dry machining aluminium wrought alloys. In the case of drilling, these layers decrease the friction between the chip and the tool because of their “lubricating” effect and thus help to improve chip removal. In the case of tapping, in which case the machining results are determined above all by crushing, friction and adhesion processes, these coating systems can bring about significant improvements in performance even without hard material intermediate layers. Because of their properties, both coating systems are gradually worn out during the cutting process (loss lubrication). Even if the layer appears optically to be worn, its effect is still partially demonstrable. This is explained by the fact that the coating material still exists in microscopically small recesses on the tool surface where it can still be effective. After a complete loss of the soft top layer, the further wear properties of the tools is then determined by the properties of the supporting hard material system.

4.4

Coatings

159

4.4.2.9 Diamond Coatings (Crystalline Carbon Coatings) Coating tools with polycrystalline diamond has been available only since 1990 and is as such the most recent technique in diamond tool production. Diamond coatings are applied to tools made of cemented carbide or ceramics in a low-pressure diamond synthesis (a CVD process). The diamond layer consists of pure diamonds and does not have a binder phase. Industrial deposition temperatures are around 600–1000◦ C. This coating technology makes it possible to manufacture even tools with complex geometries, such as curved surfaces, economically out of diamond [Leye95]. Application examples for this are drilling tools, end milling cutters and indexable inserts with chip form geometries. These tools are used in the machining of plastics, cemented carbide and ceramic green bodies, non-ferrous metals as well as in the machining of abrasive materials. Diamond-coated tools can not be used to machine steel and other ferrous materials since the diamond coating is easily worn due to the solubility of carbon in iron. We differentiate between thin diamond layers (sd = 1–40 μm) and thick diamond layers (sd = 0.3–2 mm). Thin diamond layers are used to coat components (tools, for example) directly. On the other hand, thick diamond layers are generally deposited on an auxiliary substrate and then detached from it again. Diamond plates produced in such manner are then used as free-standing diamond (e.g. as radiation windows) or mounted on supports, usually by means of vacuum soldering in order to produce tools or other components. Since this can no longer be referred to explicitly as a layer, we often no longer refer to it as a CVD diamond layer but as a CVD diamond [VDI 2840]. Several processes have been established for fabricating diamond coatings. The most important include the hot filament process (HF-CVD), the microwave process and the plasma jet process (DC arcjet). What all of these processes have in common is the decomposition of carbonic gases (e.g. methane, acetylene or carbon monoxide) followed by a deposition of carbon on the substrate surface. The formation of unwanted graphite must be avoided in the process. Atomic hydrogen plays a key role in the germ formation and growth of the diamond layers. In the case of the hot filament method, the best-known process used for depositing CVD thin diamond layers, the atomic hydrogen required for the process is produced by electrically heated filaments made of refractory metals (Ta, W or Re). Gas temperatures of over 2500◦ C are necessary to initiate the required chemical reactions [Lemm04, VDI 2840]. The rim zone is of particular importance when coating cemented carbides. Cobalt, which forms the binder phase of the cemented carbide, reduces the germ formation of diamond on the substrate surface, accelerates graphite formation, and affects crystal growth and layer adhesion. Cemented carbide substrates must therefore be subjected to pretreatment prior to coating, in which case we differentiate between physical and chemical pretreatment and the introduction of an intermediate layer. The common process presently is a multi-stage pretreatment consisting of irradiation, chemical etching, purifying and seeding. The substrate is first purified and homogenized with microbeam treatment. In the subsequent chemical etching

160


process, the cobalt binder phase is removed from the rim zone of the cemented carbide substrate. The depth to which the binder phase is removed depends on the back diffusion of cobalt during the coating process. Due to the high substrate temperature of 900◦ C and long processing times, cobalt diffuses through the etched rim zone to the substrate surface during the diamond coating process. If the etching depth is too small, the cobalt that has diffused to the surface reacts with the diamond layer and leads to the separation of diamonds on the boundary surface. This lowers the adhesive strength of the diamond coating. If the etching depth is too large, the carbide bonds in the binder phase are weakened, also lowering layer adhesion. The rough surface caused by etching serves mechanically to clamp the diamond layer to the cemented carbide substrate. Due to the extreme differences between the coating and substrate material (Young’s modulus and hardness) this mechanical clamping is of special importance with respect to the adhesive strength of the diamond layer. After etching, the substrates are again purified and seeded. In order that the diamond layer will develop, there must be diamond germs on the substrate surface that serve as starting crystals and from which the layer grows. Seeding can be performed, for example, by applying small diamond crystals with a diamond suspension [Gram04, VDI 2840, Lemm04]. Standard diamond layers have a microcrystalline structure (Fig. 4.46). The crystallites grow from the germs with varying speed depending on their crystal orientation. Slowly growing crystallites are covered by the faster-growing ones. With increasing layer thickness, the crystallites forming the surface become larger and larger, resulting in the typical appearance of a microcrystalline CVD diamond layer with sharp-edged crystal surfaces [VDI 2840]. If growth conditions are set such that new germs are constantly formed and largescale growth of individual crystallites becomes impossible, crystallite size remains in the range of nanometres within the entire layer (Figs. 4.46 and 4.47).This is how “nanocrystalline” CVD layers are formed, which are much smoother compared to crystalline layers [VDI 2840]. Both layer types can also be combined in multilayer structures, uniting the positive attributes of microcrystalline and nanocrystalline diamond layers [Uhlm05] (Figs. 4.46 and 4.47). This increases the fracture toughness of the entire system and contains the development of cracks, such that they can no longer easily reach the

Microcrystalline (standard)

Nanocrystalline (smooth)

Multilayer (smooth)

Microcrystallin Nanocrystalline

Fig. 4.46 Growth model of diamond coatings, acc. to VDI 2840 and LEMMER [Lemm04]

4.5

Ceramic Cutting Tool Materials

Microcrystalline (standard)

Nanocrystalline (smooth)

161 Multilayer (smooth)

Fig. 4.47 SEM photographs of CVD diamond coatings on a cemented carbide substrate [VDI 2840, Lemm04]

interface between the first diamond layer and the substrate, potentially causing the entire surface to chip [Lemm04].

4.4.3 Substrate Pretreatment Both CVD and particularly PVD coating processes place high demands on the surface conditions of the components that are to be coated. Even the smallest impurities on the surface of the substrate – be it residual cutting fluid or grease, water deposits, grinding and polishing agents, hardening salt residues or oxide deposits – can degas during the coating process, negatively influence that process and lead to coating adhesion problems. The pretreatment of the substrate therefore has a key role in creating high and reproducible coating quality [Lugs03, VDI3824]. In substrate pretreatment a distinction is drawn between mechanical pretreatment (grinding, deburring, microbeaming and polishing) and physicochemical pretreatment (coating-removal, ultrasound-supported purification). In order to optimize layer adhesion, the parts are subjected to a multi-staged ultrasound purification process before charging. In the coating plant, the parts are then first heated in a vacuum, whereby further impurities can evaporate. Then the tools are etched by means of argon ion bombardment in order to produce a purely metallic surface. Coating follows directly thereafter. If the substrates still need to be stored temporarily before charging, this takes place in a drying closet in order to protect the substrate surfaces from water deposits (from air moisture), rust film formation and skin contact [Lugs03, VDI3824, Gey03].

4.5 Ceramic Cutting Tool Materials Ceramic materials include all non-metallic, inorganic solid materials. Essentially these are chemical compounds of metals with non-metallic elements of group III A to group VII A of the periodic system of elements. A distinction is made between oxidic and non-oxidic ceramics. The largest group of ceramics are the oxides. Non-oxidic ceramics are carbides, borides, nitrides and silicons. Some authors differentiate again between metallic hard materials (compounds between C, B, N or

162


S and Ti, Zr, Nb, Ta, W among others) and non-metallic hard materials such as diamond, SiC, Si3 N4 , B4 C and BN. Characteristic properties of ceramic materials are compressive strength, high chemical resistance and high melting temperatures, which can be derived from the strong covalent and ionic bonds of the atoms [Salm83a, Horn06]. The following will describe more closely those ceramic materials that are used as cutting tool materials in cutting technology. Following the usage common in practice, we will differentiate between cutting ceramics (oxidic and non-oxidic ceramics) and superhard non-metallic cutting tool materials (diamond and boron nitride).

4.5.1 Cutting Ceramics Ceramics have applications in many areas of machining with geometrically defined cutting part geometries. Mentionable among these are the high-speed machining of cast iron materials, hard machining or nickel-based alloy machining. The use of ceramics in the mass production of brake disks, flywheels and like workpieces has become a matter of course. In the field of automotive technology, silicon nitride ceramics have been able to capture an important market share in the high-speed cutting of cast iron workpieces [Schn99]. Hard machining is a primary application area of mixed ceramics. “Whisker-strengthened” ceramics are often used in the area of engine construction for machining engine components made of nickel-based alloys. Whiskers are needle-shaped monocrystals with a small degree of misorientation in the lattice. Correspondingly, they have high mechanical strength (Rm up to 7000 N/mm2 ). Their length is about 20–30 μm, their diameter between 0.1 and 1 μm. The increased interest in ceramics in production is based not only on the excellent wear properties of ceramic cutting tool materials but most importantly on their toughness properties, which have been clearly increased in recent times. The brittle fracture attributes characteristic of ceramic materials, the wide variation of their strength properties and the resulting stochastically appearing tool fractures are as before the main causes limiting the use of these cutting tool materials in machining technology (in contrast to cemented carbides for example). New developments and improvements in ceramics have been concentrated therefore on further augmentation of the toughness and dependability of these cutting tool materials in production. In the case of pure oxidic ceramics, research is focused on improving these properties by increasing their zirconium oxide content, improving the distribution of these phases and by making the microstructure more consistent. In the case of mixed ceramics, finer-grain hard materials are used and titanium carbide is being partially substituted with titanium carbon nitride. Further measures include fibre and/or whisker reinforcement, reduction of the amount of glass phase at the grain boundaries of non-oxidic ceramics as well as the formation of special microstructures [Kola86, Clau77, Schn99].

4.5


163

Cutting ceramics

Oxidic ceramics

Oxide ceramics

Mixed ceramics

Al2O3

Al2O3 + TiC

Al2O3+ ZrO2

Al2O3 + ZrO2 + TiC

Non-oxidic ceramics

Whiskerreinforced ceramics Al2O3 + SiC-whisker

Silicon nitride ceramics β-Si3N4 β-Si3N4 + hard materials α-Sialon (α + β)-Sialon

Fig. 4.48 Classification of cutting ceramics

Ceramic cutting tool materials can be subdivided into oxidic and non-oxidic cutting ceramics (Fig. 4.48). Oxidic cutting ceramics include all cutting tool materials based on aluminium oxide (Al2 O3 ). A distinction is made between oxide ceramics that only contain oxides (e.g. ZrO2 ) in addition to Al2 O3 , mixed ceramics, which in addition to Al2 O3 also contains metallic hard materials (TiC/TiCN) and whisker-reinforced ceramics, in which SiC whiskers are integrated in the Al2 O3 matrix. Figure 4.49 shows images of fractures of different cutting ceramics taken with a scanning electron microscope. It is typical of oxidic cutting ceramics that they do not have a visible binder phase such as cemented carbides (cobalt) have. One can recognize the globular grains of oxide ceramics, the extremely fine grains of mixed and whisker-reinforced ceramics as well as the SiC whiskers, which are only a few μm large in the detail magnification. Characteristic of non-oxidic ceramics based on Si3 N4 is the needle-shaped formation of the crystals. In order to improve wear resistance, Si3 N4 cutting ceramics can be coated with Al2 O3 or with multilayer coating systems made of Al2 O3 and TiN [Schn99]. Figure 4.50 shows some physical properties of commercially available ceramic cutting tool materials. Depending on their chemical composition, these ceramic types can have considerably diverse properties. 4.5.1.1 Cutting Ceramics Based on Al2 O3 Oxide Ceramics White oxide ceramics are the traditional type of cutting ceramics. Ceramics based Al2 O3 on were already introduced as cutting tool materials at the end of the 1930s [Kola86]. Inserts were made for a long time out of pure aluminium oxide. Due

164


Oxide ceramic (Al2O3 + ZrO2)

Mixed ceramic (Al2O3 + TiC )

Silicon nitride ceramic (Si3N4 + MgO )

Whisker-reinforced ceramic (Al2O3 + SiC-Whisker )

Detail magnification (Al2O3 + SiC-Whisker )

Coated silicon nitride ceramic (Si3N4 + Al2O3)

Fig. 4.49 Fractures of cutting ceramics

Whisker-reinf. Silicon nitride Mixed ceramic Oxide ceramic Al O 2 3 ceramic Si3N4 Al2O3 +15%ZrO2 +3.5% +15% +20%SiC- +10%ZrO2 +30% +10%Y2O3 ZrO2 ZrO2 Whisker +5%TiC Ti(C,N) Oxide ceramic Al2O3

Properties

Density

ρ

g/cm3

4.0

4.2

3.7

4.1

4.3

3.3

Vicker hardness

−

-

1730

1750

1900

1730

1930

1750

Bending strength

σ

N/mm2

700

800

900

650

620

800

N/mm2

5000

4700

-

4800

4800

-

bB

Compressive σ dB strength Young‘s modulus

E

103⋅N/mm2

380

410

390

390

400

Fracture toughness

KIC

N⋅m1/2/mm2

4.5

5.1

8.0

4.2

4.5

Thermal conductivity

λ

W/m⋅K

16.4

15

32

14.7

20

Thermal expansion

α

10–6⋅K–1

8

8

-

8

8

3.4

Melting point

T

°C

-

-

-

-

-

-

7.0

Fig. 4.50 Physical and mechanical properties of different cutting ceramics and their main components

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to their brittleness and vulnerability to fracture, such inserts are no longer used in machining. The pure ceramics used today are dispersion material containing Al2 O3 as well as about 3–15% of finely distributed zirconium dioxide to improve toughness. The toughness-enhancing effect of dispersed ZrO2 particles in an Al2 O3 matrix is based on the phase transformation of zirconium dioxide. ZrO2 , which exists in the form of a tetragonal lattice modification in the sintering temperature range (1400–1600◦ C), transforms during cooling into its monoclinic low-temperature modification. The temperatures in which this transformation takes place depend on the size of the particles. The smaller the ZrO2 particles are, the lower the transformation temperature. Since the transformation from the tetragonal to the monoclinical modification is associated with a volume expansion, various specific mechanisms of action can assert themselves depending on the size of the particles. The common effect of all these mechanisms is that they ultimately absorb fracture energy. The speed of crack development is reduced by microcracking, crack branching, the stress-induced transformation of small ZrO2 particles as well as crack diversion. The result of this is that critical cracks only develop at a higher level of energy, which corresponds to an increase in fracture resistance and an improvement of ductility [Clau77, Clau84, Zieg86]. Mixed Ceramics Mixed ceramics (black ceramics) are dispersion materials based on Al2 O3 that contain between 5 and 40% of non-oxidic components in the form of TiC or TiCN. The hard materials in the matrix form finely distributed phases, which limit the growth of aluminium oxide grains. Correspondingly, these ceramics have a very fine-grained structure, improved toughness properties and a high level of edge strength and wear resistance. Compared to pure ceramics, they are harder and have more favourable thermoshock properties due to their high level of thermal conductivity (Fig. 4.50). The toughness of these ceramics can be further improved by adding ZrO2 . The development of mixed ceramics is progressing towards finer and finergrained cutting tool materials with extremely homogeneous textural structures. As opposed to conventional cutting ceramics based on aluminium oxide and titanium carbon nitride with an average grain size of < 2 μm, the more advanced mixed ceramics have a submicron structure with grain sizes of < 1 μm. The finer-grained structure increases hardness and bending strength and consequentially the mechanical and thermal loadability, wear resistance and edge strength of ceramic inserts. Fine-grained mixed ceramics are used in hard-fine machining, e.g. for machining hardened rolling bearing steels, for hard-fine turning case-hardened automotive components such as drive wheels, crown wheels, gearwheels or sliding sleeves with a Rockwell hardness of 54–62 HRC but also for planing and fine planing cast iron at very high speeds. In the case of hard turning, submicron mixed ceramics are competing with PCBN cutting tool materials in many areas of application as more economical alternatives due to their good cost-benefit ratio [Krel97, Schn99].

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Whisker-Reinforced Cutting Ceramics Whisker-reinforced cutting ceramics are cutting tool materials based on Al2 O3 with about 20–40% silicon carbide whiskers. The goal of whisker-reinforcement is to enhance the toughness properties of ceramic cutting tool materials. The increase in toughness gained by incorporating whiskers in oxidic ceramics is remarkable. Compared to mixed ceramics, whisker-reinforced types have up to 60% higher fracture toughness. The whiskers bring about a more consistent distribution of mechanical loads in the cutting tool material as well as a more rapid transport of heat from the thermally highly loaded cutting areas due to their improved heat conductivity. This results in improved fatigue limits and thermoshock resistance, so that whisker-reinforced cutting ceramics can also be used in wet cut processes.

Fabrication of Oxidic Cutting Ceramics The starting materials of cutting ceramics, including pressing and sintering auxiliaries, are metered according to an exact formula, homogenized in vibration mills and converted to a compressible powder with a high bulk material stability in a spray-dryer. Oxide ceramics, as well as mixed ceramics with small amounts of hard materials, are compressed at room temperature to inserts (cold pressing) and then sintered. The sintering temperature is about 1600◦ C. Mixed ceramics with a high percentage of hard material (> 10%) as well as whisker-reinforced ceramics have to be hot-pressed, i.e. pressing and sintering take place in one cycle. Due to the special equipment required for this (e.g. graphite moulds instead of cemented carbide ones), hot pressing is more cost-intensive than normal sintering. Furthermore, the number of manufacturable tool geometries is much more limited. Hot isostatic pressing can be used to subsequently compact sintered ceramic parts, thus reducing the porosity of the cutting tool material. One method variant is hot isostatic compacting of pre-sintered closed-pore mouldings in an installation. After sintering/hot pressing, the contact face, and in the case of precision inserts, the lateral surface and cutting edge chamfer are ground with diamond grinding discs [Kola86, Grew85].

Properties of Oxide Ceramic Cutting Tool Materials Because of their high hot hardness and chemical resistance, ceramic cutting tool materials based on Al2 O3 are characterized by excellent wear properties. However, the highly favourable wear properties of oxide ceramic cutting tool materials is accompanied by their sensitivity to tensile, bending, impact and thermal shock stresses. As mentioned at the beginning of this chapter, current further development of oxide ceramic cutting tool materials has for this reason been concentrated on improving toughness and thermoshock properties without diminishing wear resistance.

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The excellent wear resistance of oxidic cutting tool materials, particularly at high temperatures, is based among other things on their superior hardness and compressive strength compared to other cutting tool materials at high temperatures. For example, oxide ceramic cutting tool materials at 1000◦ C are still harder than high speed steel at room temperature (Fig. 4.51). While the compressive strength of Al2 O3 at room temperature roughly corresponds to that of cemented carbide, it is at 1100◦ C still as great as that of steel at room temperature, while neither steel nor cemented carbide are resilient to pressure at 1100◦ C. The high compressive strength of ceramics has been a primary reason for their being privileged as cutting tool materials. But because their banding fracture strength is relatively low, the use of oxide ceramics as the cutting tool material requires that the process kinematics be selected such that the cutting forces are effective as much as possible only in the form of compressive stresses [Kola86, Kola86a, Zieg86]. A further property which makes the use of ceramic cutting tool materials advantageous at high cutting speeds is Al2 O3 ’s low level of creeping. In the case of cemented carbides, high temperature strength is limited by the property-determining cobalt phase at temperatures of 800–900◦ C. At higher temperatures, creeping processes are introduced which are smaller by several orders of magnitude in the case of Al2 O3 .

5000

Diamond

Hardness HV 10

4000

Boron nitride 3000 Mixed ceramic Al2O3 + TiC/TiCN 2000 Carbide Oxide ceramic Al2O3

1000

0

250

Tool steel 500 750 Temperature T / K

Fig. 4.51 Hot hardness of cutting tool materials (Source: CeramTec)

HSS 1000

1250

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The high wear resistance of oxidic cutting tool materials can also be attributed to the good chemical resistance of Al2 O3 . Al2 O3 is resistant to oxidation at the cutting temperature used in practice and has only a small amount of affinity to metallic materials. The most crucial disadvantage of ceramic cutting tool materials is their brittleness, i.e. lacking capacity to reduce stress peaks by means of plastic deformation. The cause of this is the low number of sliding systems in the crystal structures of ceramic materials. This results in high sensitivity to tensile stresses and low resistance to mechanical and thermal shock compared with metals. In the case of a sudden mechanical overload due to impulse-like stress, the cutting tool material is destroyed by brittle fracture. Because of insufficient ductility, cracks begin to form when the inner material cohesion is exceeded by external stress. If the crack reaches a critical size, unstable crack growth appears leading to fracture of the ceramic tool. The stress intensity factor KI is a parameter for judging the stress condition at the peak of a crack. Failure of the cutting tool material due to fracture occurs if the stress intensity factor reaches a critical value, crack toughness (fracture toughness) KIC . In comparison to other materials, the values of KIC are very low for ceramics (Fig. 4.50). Another important disadvantage of oxide ceramics are their relatively low resistance to temperature change. In the case of temperature changes of more than 200◦ C, pure Al2 O3 is destroyed. This can only be improved by alloying a component of superior resistance to temperature change. Due to their sensitivity to thermoshock, coolants should not be used in conjunction with oxide ceramics when roughing and planing. Coolants should only be used to temper the workpiece, e.g. due to narrow workpiece tolerances. Low bending fracture strength and relatively high sensitivity to impact and temperature change stresses require a slanted entering and leaving cutting path, which is associated with delayed load of the cutting edge of the tool (Fig. 4.52). Due to their low edge strength, chamfered edges should stabilize the cutting edge. Areas of Application of Oxide Ceramic Cutting Tool Materials It must be stressed that, due to their extreme brittleness, a careful adjustment of cutting parameters to tool geometry is the prerequisite of a successful use of cutting ceramics. There is no room for error, since ceramics can be destroyed immediately in the form of fracture should the marginal conditions be unfavourable. Cutting ceramics are therefore almost exclusively used in mass production. This is justified by the reduction of processing times and the high cost of selecting and adjusting the cutting tool materials. The main area of application of oxide ceramic cutting tool materials is rough and finish turning of grey cast iron, case-hardened steels and heat-treated steels. However, pure oxide ceramics are being increasingly replaced by silicon nitride ceramics, especially in cast iron processing (Fig. 4.53). The increased popularity of dry machining however is opening up new areas of application to oxide ceramics in the case of turning extrusion parts with a small machining allowance.

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1. Chamfering

1. Chamfering and external turning

Finished part 1. Clamping (external)

2. Clamping (internal) 2. Chamfering and internal turning

3. Face turning

2. Face turning

Fig. 4.52 Entering and exit tool path when turning with cutting ceramics (Source: Ford)

Fine-grained mixed ceramics are used in hard-fine machining (e.g. for machining hardened rolling bearing steels) for fine turning case-hardened automotive components such as drive wheels, crown wheels, gearwheels or sliding sleeves with a Rockwell hardness of 54–62 HRC. They are also used for finish planing and fine planing cast iron at very high speeds. In the case of hard turning, submicron mixed ceramics are competing with PCBN cutting tool materials in many areas of application as more economical alternatives due to their good cost-benefit ratio [Krel97, Schn99]. Both types of cutting ceramics still however remain restricted with respect to the carbon content of the machined steels. For example, in the case of steels with a C content of under 0.35% adhesion occurs on the tool as well as chemical reactions which increase wear and generally make the process uneconomical. Whisker-reinforced oxide ceramics have been used with great success for turning high temperature nickel-based alloys (Fig. 4.54). In contrast to otherwise used cutting tool materials (cemented carbide or HSS), it is possible to increase the cutting speed by a factor of 10 or more. In the case of rough and finish milling Inconel 718 with milling head face cutters, cutting speeds of 800–1000 m/min can be realized with cutting ceramics based on silicon nitride or whisker-reinforced aluminium oxide. As opposed to machining with cemented carbide tools, this represents an increase in performance by factors of 25 or more [Momp93, Gers02, Krie02, Uhlm04].

170

4 Cutting Tool Materials and Tools Al2O3-ceramic 300

vc = 600 m/min f = 0.35 mm ap = 3.0 mm

250 Tool life quantity N

Si3N4-ceramic vc = 1000 m/min f = 0.5 mm ap = 3.0 mm

200 150 100 50 0 non-chamfered chamfered

non-chamfered chamfered

Fig. 4.53 Influence of cutting ceramics and entering path on the tool life quantity in brake disc manufacturing (Source: CeramTec)

Tool life travel path Lc / m

3000 2000 Cemented carbide K10/20

1000

Polycrystalline cubic boron nitride

500 400 300

SiC-whisker reinforced cutting ceramics and SiAlON

200

100

Criteria of tool life: VB = 0.3 mm, VBN = 0.3 mm 20

200 300 400 600 100 30 40 50 Cutting speed vc / (m/min)

VB

Notch wear VBN

VB Flank wear

vc = 400 m/min f = 0.12 mm ap = 0.30 mm Inconel 718 Wet cut

Determinant tool life wear criteria

Fig. 4.54 Achievable tool life travel paths and applicable cutting speeds in finish turning Inconel 718 with uncoated cemented carbide, whisker-reinforced cutting ceramic, silicon nitride ceramic and PCBN

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Chemical reactions and built-up edge formation when machining light metal alloys make Al2 O3 cutting ceramics unsuitable for machining Al, Mg and Ti alloys. Cutting ceramics composed of thin ceramic tubules represent a new generation of ceramic cutting tool materials. In the case of these ceramics, first introduced at EMO 2005, the tool body is formed of small thin tubes consisting of a tough ceramic material and filled with a hard, wear-resistant ceramic material. The diameter and thickness of the tubes, as well as their orientation and filling, are variable. The tubes are arranged in a plane. The tool is formed with a large number of planes built on top of each other, where the tubes forming the individual planes can be parallel from plane to plane or can be arranged alternately at an angle of 45◦ /90◦ to each other. In this way, the properties of the cutting tool material, such as toughness and wear resistance, can be modified and adjusted to the stresses of various machining tasks. Initial machining experiments in turning high-temperature nickel-based alloys as well as in machining hardened steel materials demonstrated the promising wear properties and performance of this new cutting ceramic [Haid05]. 4.5.1.2 Non-oxidic Cutting Ceramics Among the non-oxidic ceramics (carbides, nitrides, borides, silicides. . .), materials based on Si3 N4 have been especially successful as cutting tool materials for machining purposes. Si3 N4 cutting ceramics, in contrast to oxidic cutting ceramics, are characterized by increased toughness and improved thermoshock resistance (Fig. 4.50). Moreover, they have a higher hot hardness and high temperature strength. In machining processes involving grey cast iron, they make it possible to use the highest cutting values while maintaining the highest tool life spans and low failure rates. The high safety level of these cutting tool materials has been an especially important factor in the acceptance of Si3 N4 cutting tool materials by manufacturers. The significantly higher fracture resistance compared to oxide and mixed ceramics is based on the needle-like shape of the hexagonal β-Si3 N4 crystals as opposed to the globular Al2 O3 grains. The non-directional growth of the needle-shaped crystals leads to a microstructure made up of mechanically interlinked components which provides the cutting tool materials with excellent strength properties. The low thermal expansion of Si3 N4 ceramics compared with oxide and mixed ceramics is responsible among other things for its favourable thermoshock properties. However, in order to press silicon nitride ceramics completely, sintering auxiliaries (Y2 O3 , MgO, Al2 O3 ) are required that form a glass phase and fill in the gaps between the crystals. This glass/binder phase has a negative influence on the high-temperature properties of silicon nitride ceramics. Besides sintering auxiliaries, Si3 N4 cutting ceramics can also contain other additives that affect their crystal structure or texture and thus their properties as well. Corresponding to their chemical composition and crystallographic structure, the silicon nitride ceramics available today can be subdivided into three groups: • I: β-Si3 N4 + binder phase (Y2 O3 , MgO, Al2 O3 ) • II a: α sialon (α–Si3 N4 + Al2 O3 + AlN) + binder phase (Y2 O3 )

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• II b: (α + β) sialons (α Si3 N4 + Al2 O3 + AlN) + additive (e.g. ytterbium) + binder phase (Y2 O3 ) • III: β silicon nitride + hard materials (e.g. TiN, ZrO2 , SiC-whiskers) + binder phase The cutting tool materials from Groups I and II are fabricated by means of hot pressing, sintering, hot isostatic pressing or a combination of these methods. Most of the Si3 N4 cutting tool materials available on the market today belong to Group I. The cutting tool materials from Group II are usually designated as sialons. Silicon nitride can receive up to 60% aluminium oxide in solid solution. Some nitrogen atoms are replaced in the process by oxygen atoms and silicon atoms by aluminium atoms. While the α sialon mixed crystals have a globular form, β sialon mixed crystals are stem-shaped. Including special additives makes it possible to stabilize the β sialons, thereby creating a cutting tool material the structure of which consists of both α and β mixed crystals. The amount of additives determines the amount of β sialons in the cutting tool material’s microstructure. In comparison to the cutting tool materials from Group I, sialons are harder, more chemically resistant and have an increased resistance to oxidation. The manufacturing method in this case is sintering, which can sometimes be followed by hot isostatic pressing. Group III can include Si3 N4 cutting tool materials, the properties of which can be specifically altered by adding hard materials such as titanium nitride, titanium carbide, zirconium oxide or Sic whiskers. Areas of Application of Non-oxidic Cutting Ceramics The classic field of Si3 N4 cutting tool materials is grey cast iron machining. In this case, usually the tougher silicon nitride ceramics of Group I are preferred, especially in automated manufacture. Due to the high fracture toughness of these cutting tool materials, large feeds, high cutting speeds and thus large volume removal rates can be realized when machining cast iron materials in smooth and interrupted cut [Köni86, Köni87, Schn99]. For example, when turning automotive brake discs, using Si3 N4 ceramics improves tool life quantity considerably compared with oxide ceramics (Fig. 4.53). Improved resistance to fracture makes it possible, as shown in Fig. 4.52, to omit workpiece chamfering as long as one forgoes maximal utilization of the cutting tool material. The main area of application of sialons is in turning nickel-based alloys under finishing or average roughing conditions (Fig. 4.54). They are therefore in direct competition with whisker-reinforced oxide ceramics. Like the latter, sialons can be used for machining nickel-based alloys at much higher cutting speeds than cemented carbide tools. They are also good for high-speed milling of nickel-based alloys with tools mounted with indexable inserts at cutting speeds of 800–1000 m/min. The wear resistance of silicon nitride ceramics is slightly lower than that of oxide ceramics. Silicon nitride cutting tool materials exhibit a strong affinity to iron and oxygen under machining conditions. They wear very quickly in the case of steel machining, so the use of these materials groups currently remains uneconomical.

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Coating silicon nitride ceramics opens up further prospects for their use. The classic coating materials are TiN, TiC, TiCN and Al2 O3 , which are applied in different combinations and layer thicknesses. A multilayer coating system consisting of aluminium oxide and titanium nitride remains the most popular. The service life of Si3 N4 can be considerably increased by applying a wear and diffusion inhibiting coating, especially when machining cast iron with globular graphite (e.g. GJV40) [Schn99]. Coated silicon nitride ceramics are also suitable therefore for cases that had hitherto been reserved for cemented carbides. As opposed to cemented carbides, they allow for higher cutting speeds and thus for shorter machining times and in many cases also for dry machining [Schn99].

4.5.2 Superhard Non-metallic Cutting Tool Materials In machining technology, cutting tool materials based on diamond and boron nitride are designated as superhard and non-metallic. According to the definition, both are ceramic cutting tool materials, where diamond is assigned to monatomic, boron nitride to non-oxidic ceramics. 4.5.2.1 Diamond as a Cutting Tool Material Elemental carbon appears in the two crystal modifications graphite and diamond. Diamond solidifies in the cubic crystalline lattice system, in which the C atoms are bonded covalently and tetraedically. The extremely high bond and lattice energy is the reason why diamond is the hardest of all known materials. Diamond is equally superior to all other known hard materials with respect to thermal conductivity. Figure 4.55 shows the hardness and thermal conductivity of some materials that tend to be used as abrasives as well as different carbides, which are components of cemented carbide tools and permits a comparison with diamond. Classification of Diamond Cutting Tool Materials To classify diamond cutting tool materials, we differentiate between natural and synthetic diamond, which both can appear in monocrystalline or polycrystalline form. According to DIN ISO 513, the identification letters of polycrystalline diamond are DP while those of monocrystalline diamond are DM (Fig. 4.3). Following the usage common in practice and in the literature, we will also refer to polycrystalline diamond cutting tool materials with the abbreviation PCD. (a) Natural Diamond Natural diamond is only of importance for the purposes of cutting with geometrically defined cutting edges in its monocrystalline form. Although diamond also exists in nature in polycrystalline form (ballas, carbonado), these types of diamond are of lesser interest since the synthetic fabrication of polycrystalline diamond is more advantageous both economically as well as technologically. One significant property of monocrystalline diamond is the anisotropy (directional dependence) of mechanical properties such as hardness, strength

174

4 Cutting Tool Materials and Tools Thermal conductivity λ RT / (W/m⋅K)

Hardness HV

Diamond cBN B4C SiC TiC α -Al2O3 10000

5000

2500

100

1000

10000

Fig. 4.55 Hardness and thermal conductivity of diamonds in comparison to other hard materials

or elastic modulus. While this directional dependence is compensated in the case of polycrystalline materials by the completely random distribution of the individual crystals, anisotropy remains in monocrystalline materials due to the aligned orientation of the crystal lattice. This also the cause of the cleavability of monocrystalline diamonds in four select cleavage directions. From this we can see that the position of the lattice directions must be known both for grinding monocrystalline diamond as well as for using it as a tool. While grinding must always take place in the direction of lowest hardness, monocrystalline diamond tools have to be oriented in the tool holder such that the cutting force points in the direction of a hardness maximum. (b) Synthetic Diamond The fabrication of synthetic diamonds takes place using a catalyst solution in a pressure and temperature range in which diamond forms a stable phase. By a careful selection of pressure and temperature, the crystal growth rate and dimensions of the crystals can be controlled within a range of a few micrometers to several millimetres and specific physical properties like purity or porosity can be influenced. Diamond synthesis results in monocrystalline diamond particles that are utilized in machining with geometrically defined cutting edges as synthetic monocrystalline diamond tool bodies or for further processing into polycrystalline cutting parts. Fabrication and Designs of Diamond Tool Bodies (a) Monocrystalline Diamond Tools Monocrystalline diamond tool bodies are manufactured with both natural diamond and synthetic monocrystals [Heus96]. The most common cutting edge shapes of

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175

κ'r ≈ 2° κr ≈ 45°

Tool with one cutting edge

Tool with round cutting edge

κ'r ≈ 2°

κr ≈ 18°

160°

Tool with facette shape

Fig. 4.56 Shapes of monocrystalline diamond tools

monocrystalline diamond tools are shown in Fig. 4.56 [Wein69]. The tool with one cutting edge is used both for drilling as well as for external turning, whereby in the latter case a corner radius is ground onto it in order to improve the surface profile. By selecting a very small minor cutting edge lead angle κr < 2◦ , the minor cutting edge can be designed as a broad-tool cutting edge which takes on the function of further polishing the surface. Tool variants with round cutting edges have the advantage of having a very large useful cutting edge length, but due to the curvature of the cutting edge they have unfavourable chip formation conditions and relatively high passive forces. In the case of tool bodies with facette shapes, three to five cutting edges are ground on, whereby two neighbouring edges form an angle of about 160◦ . The tool is adjusted so that the minor cutting edge has a very small lead angle as acts as a broad polishing cutting edge. Surface quality can be affected considerably by varying κr slightly. The grinded tips are brazed on the support material or fixed with special clamps. (b) Polycrystalline Diamond Tools Tool bodies made of a synthetic polycrystalline diamond layer were first introduced in 1973 and have replaced monocrystalline diamond tools and cemented carbides in some areas. The starting materials are synthetic diamond particles of a very small defined granulation (grain diameter 2–25 μm) in order to obtain a maximum amount of homogeneity and packing density. The polycrystalline diamond coating is fabricated by means of a high pressure, high temperature process (60–70 kbar, 1400–2000◦ C), in which synthetic diamonds are sintered together into polycrystalline bodies in the presence of a metallic catalyst. To this end, usually cobalt, but also silicon, tungsten or tungsten carbide are

176


used [Neis94]. During the sintering process, “diamond bridges” are formed between the diamond grains which give the polycrystalline bodies their high strength. As the diamond grains consolidate, gaps arise between neighbouring diamond crystals that are filled in by the catalyst. In the literature, this phase is also called the binder, which gives the polycrystalline diamond bodies the required toughness [Pret06]. One way to reduce the size of the gaps between the diamond grains – especially in the case of coarse-grained DP types – is using diamond particles with varying grain sizes (2–30 μm). The small diamond grains accumulate in the gaps formed by the large crystals, they increase the number of diamond bridges and with it the impact resistance of the composite material as well as the quality of the cutting edge. The polycrystalline coating fabricated in the high pressure/high temperature process, the thickness of which is about 0.5 mm, is either directly applied on a presintered cemented carbide base or over a thin intermediate layer consisting of a metal with a low elastic modulus bonded with the cemented carbide in order to balance stresses between the diamond layer and the cemented carbide base. Due to its polycrystalline structure, the diamond coating is a statistically isotropic overall body in which the anisotropy of the individual monocrystalline diamond particles is balanced by the random distribution of the diamond grains. Polycrystalline diamond thus does not exhibit the hardness isotropy and cleavability of monocrystalline diamonds. On the other hand, it also does not reach the hardness level of a diamond monocrystal in its “hardest” direction, especially since hardness is also influenced by the degree of consolidation between the individual crystals and their bonding to the binder phase. Depending on the grain size of the diamond granulation used as well as the catalyst used, the properties of the polycrystalline composite material can be specifically influenced. Coarse-grained DP cutting tool materials (grain size about 25 μm) are characterized by higher hardness and thermal resistance, but their cutting edges are rounder and more jagged than those of finer-grained types. As a result, wear resistance increases in the case of primarily abrasive stress during the machining process (Fig. 4.57). They are the most advantageous with respect to wear, cutting edge quality and cutting edge strength in the machining of aluminium composite materials and alloys, which are not critical with respect to the workpiece surface. DP types of average grain size (approx. 10 μm) are very common as multi-purpose variants. They are used for rotating tools with high requirements on cutting edge quality and service life. Fine-grained types (about 2 μm grain size) have established themselves especially in the automobile industry in cases demanding the best possible cutting edge quality [Bail01]. Changing the composition of the “binder phase” can also increase the limits of thermal stability, for example by using SiC. However, this considerably reduces the thermal conductivity of the cutting tool material compared with that of materials with the usual Co catalyst [Neis94]. Cutting edge shaping of the cutting part blank (i.e. of the cemented carbide base with applied diamond coating) is accomplished by means of spark-erosive cutting and grinding. The tool bodies are either soldered onto the tool bearer or clamped into standardized tool holders, since indexable inserts made of polycrystalline diamond


Process:

External cylindrical turning

Material:

G-AlSi12

Cooling:

dry vc = 500 m/min

Cutting cond.:

f = 0.1 mm ap = 1 mm Cutting time:

tc = 40 min

177 Displacement of the cutting edge SV / μm

4.5

14 12 10 8 6 4 2 0 DP1

DP1: grain size: 10–15 µm, binder: β −SiC

DP3: grain size: 2–6 µm, binder: W (12%), Co (88%)

DP2

DP3

DP4

DP2: grain size: 6–10 µm, binder: W (9%), Co (91%)

DP4: grain size: 0.5–1 µm, binder: W (47%), Co (53%)

Fig. 4.57 Microstructure and wear of polycrystalline diamond cutting tool materials when turning AlSi-alloy G-AlSi12, acc. to NEISES [Neis94]

correspond in shape and dimensions with commercially available cemented carbide or ceramic indexable inserts. Application Areas of Diamond Cutting Tool Materials Cutting iron and steel materials with diamond tools is impossible due to the affinity of iron to carbon. Diamond turns into graphite and reacts with the iron in the contact zone between the tool and the workpiece due to the high temperatures arising there.

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As a result, the cutting edge wears quickly in the case of both monocrystalline and polycrystalline diamond cutting tool materials. Monocrystalline diamond tools are especially suited to cutting light, heavy and precious metals, hard and soft rubber as well as glass, plastics and stone. Their field of application is mainly in finishing, since large depths of cut and feeds are impossible due to limitations on cutting edge dimensions and relatively low flexural strength. The use of monocrystalline diamond cutting tool materials promises advantages when the demand for very high dimensional accuracy and surface quality is of foremost importance. For example, the use of nearly notch-free polished diamond cutting edges in ultra-precision cutting (turning, fly-cutting, planing) can result in surface finishes of between 3 and 6 nm. One classic area of application is machining hard and soft contact lenses. Further fields include the production of mirrors for lasers and other optical applications, of shaping tools for making impressions on blank CDs, of tools for making imprints on plastics or of tools for plastic injection moulding [Spen91, Ikaw91, Weck95, Heus96, Brin96, Kloc96, Take00]. Besides light, heavy and precious metals, the palette of materials machined with polycrystalline diamond tools comprises different plastics, coal, graphite and presintered cemented carbide. Their use is not restricted to finishing, but also includes roughing. It is in many cases possible to unite both pre-machining and finishing in one working cycle. In the case of machining workpieces made of aluminium and other non-ferrous metals, PCD tools have acquired a secure place in modern cutting processes as highperformance tools [Schü01, Halw04, Kass04, Vogt04, Wick04, Fall05, Hedr05, Hedr05a, Kasp05, Brun99]. Polycrystalline diamond tools are of special importance in the machining of aluminium alloys containing large amounts of silicon (Fig. 4.58). Since these alloys have a hard/soft structure, the cutting edge cuts in an alternating fashion through the soft aluminium phase and through the hard silicon particles. Due to the strong abrasive effect of the silicon particles, tools made of cemented carbide are subject to a lot of wear. Moreover, when using cemented carbide tools, the adhesive tendency of aluminium with the cutting tool material has a negative effect on the machining process. In comparison to cemented carbides, PCD cutting tool materials can be used with much higher cutting speeds when cutting aluminium alloys with high silicon content. They have extremely high tool lives of up to 80 times higher than those of cemented carbide tools and have excellent surface quality and precision. Their high level of manufacturing safety is another essential reason why tools made of polycrystalline diamond are preferred over those of cemented carbide for this machining task. This is especially the case in mass production on transfer lines, where short cycle times and high safety against unforeseen cutting edge fracture are a must. Further application examples include the milling of magnesium pressure cast alloys and the high-performance milling of aluminium integral parts in the airline industry with end milling cutters equipped with PCD [Wein69, Obel84, Köni82, Chry79, Spur84, Hoff88, Jäge89, Beck95, Wald92, Stie99, Zwah00]. Under certain conditions, it is possible to machine cast iron materials with PCD tools as well. One example of this is drilling finishing with reaming tools equipped

4.5


179 HW (ap = 0.5 mm; fz = 0.023 mm; z = 24)

36000

1600 1600

DP (ap = 0.5 mm; fz = 0.15 mm; z = 8) DP with optimal adjusted cutting edges (ap = 0.5 mm; fz = 0.1 mm; z = 8) - Reduction of tool change time - Improvement of surface quality - Reduction of rework (minimized burrs) - Complex cutting edge survey - Difficult regrinds - Special tools (settings; form fit) - Accurate application

1600 380 2500 100

360

880

100

1130

16 Cutting tool material costs per workpiece / %

Tool life quantity / workpieces

Tool costs per workpiece /%

Cutting speed / [m/min]

100

50

60

Cutting time per 1 m milling distance / %

Fig. 4.58 Comparison of cemented carbide and polycrystalline diamond when milling the aluminium-based alloy GK-AlSi17Cu4Mg

with DP. In the case of very small depths of cut, adjusted feeds and cutting speeds as well as an intensive cooling of the cutting edge, it is possible to obtain long tool lives as well as high surface quality and component precision in cast iron machining [NN02]. (c) CVD Diamond Tools CVD diamonds are diamond coatings deposited on a base body made of cemented carbide or ceramics in a CVD process. A distinction is drawn between thin and thick diamond coatings. As opposed to polycrystalline diamond, these diamond layers consist only of diamond crystals, containing no “binder phase”. The fabrication of tools coated with diamond is treated extensively in Sect. 4.4.2. Thin diamond coatings can be deposited on tools with complex geometries such as drills, end milling cutters and indexable inserts with chip form grooves. They are suited to machining graphite, copper, fibre-reinforced plastics and aluminium alloys high in silicon. In the field of superfinishing, thin diamond coatings have not been successful thus far. The reason for this is that sharp cutting edges are blunted by the relatively thick diamond coatings (about 20 μm). Recent developments are making it possible to produce very sharp cutting edges on diamond-coated tools as well by means of a sharpening process [Hage04]. Thick CVD diamond coatings are excellently suited to cutting highly abrasive materials such as aluminium-based metal composite materials. Examples of this are

180


the machining of brake drums made of AlSi9Mg into which 20 vol% SiC particles are embedded or of cylinder crankcases made of AlSi9Cu3, which are reinforced with 15 vol% Si particles and 5 vol% Al2 O3 short fibres [Suss01, Wein02, Uhlm00]. In the case of machining GFK/CFK materials as well, much longer tool lives can be expected compared with PCD cutting edges due to the lack of a binder phase [Feuc05]. 4.5.2.2 Boron Nitride as a Cutting Tool Material Boron nitride exists like carbon in a soft hexagonal modification, which crystallizes in the same lattice type, and in a hard cubic modification, which has an identical structure to the diamond lattice. There is also a third modification, which crystallizes in the wurzite structure. The wurzite lattice is a lattice type with a hexagonal symmetry but with a different atomic configuration than the graphite lattice. With respect to hardness, this form is somewhere between the other two modifications. In contrast to silicon nitride, naturally existing hexagonal boron nitride is soft and not suitable as a cutting tool material for machining with defined cutting part geometries. Only after transforming the hexagonal into the cubic crystalline lattice with the help of a high-pressure/high temperature process does boron nitride exhibit those qualities that distinguish it as a cutting tool material (Fig. 4.55). Cubic boron nitride is, after diamond, the second hardest material. Hexagonal boron nitride is synthesized by reacting boron halogenides with ammonia. It has a density of 2.27 g/cm3 and a melting point of 2730◦ C [Salm83]. Cubic boron nitride (ρ = 3.45 g/cm3 ) does not exist in nature. Its fabrication under the conditions of diamond synthesis was first successful in 1957. The transformation of hexagonal into cubic boron nitride is accomplished with pressures of 50–90 kbar and temperatures of 1800–2200 K under the catalytic influence of alkaline nitrides or alkaline-earth nitrides. Figure 4.59 shows the phase diagram of boron nitride system with the four phase areas molten, hexagonal boron nitride with graphite structure, hexagonal boron nitride with wurzite structure and cubic boron nitride with diamond structure. Despite their identical lattice structures, there are essential differences between diamond and cBN. cBN has six cleavage planes – two more than diamond. This property is insignificant for the use of cBN in machining with defined cutting edges, since in that case only polycrystalline tools are used. More important is the fact that boron nitride is not a chemical element like carbon, but a chemical compound. The boron nitride lattice contains boron and nitrogen atoms and can therefore not reach the same level of bonding force symmetry and hardness as diamond, the lattice of which consists exclusively of carbon atoms. But with respect to its chemical resistance, especially against oxidation, cBN is far superior to diamond. It is stable from atmospheric pressure to about 2000◦ C, whereas diamond begins to graphitize already at about 900◦ C. In the case of machining with geometrically defined cutting edges, cBn is mainly used as a polycrystalline cutting tool material (BN). Since the maximum crystal size in the fabrication of cBN is limited (1–50 μm), cBN grains are sintered with

4.5


181

180 160 Cast

120 100 80 60 40 20

0

Hexagonal BN (wurtzite-structure)

Pressure p / kbar

140

Cubic BN (diamond-structure)

Hexagonal BN (graphite-structure) 1000

2000 3000 Temperature T / K

4000

5000

Fig. 4.59 Phase diagram of boron nitride

the help of a binder phase in a high pressure/high temperature process to an approx. 0.5 mm thick polycrystalline BN layer and simultaneously applied to a cemented carbide base. A distinction is made between cutting edge blanks, which are connected by soldering to the tool holder and only receive their final cutting edge shape by grinding, and indexable inserts, which are clamped in common clamping holders [Töns81, Töll81, Wern81, Nott82, Clau85, Töll82, Wern84, Wein87, Buzd87, Krat06]. Aluminium, titanium, titanium aluminium nitride, titanium carbide or cobalt, nickel and tungsten are used as binders. Aluminium and titanium react chemically with the cBN, forming AlN and AlB2 / TiN and TiB2 . Cobalt forms complex borides with tungsten in the form of CoWB, W2 CoB2 or W2 Co21 B6 . The coW borides arise from grain boundary reactions during the sintering process. This leads to an extremely stable bond between the cBN grains and the Co-W binder phase [Krat06]. The currently most common cutting tool materials based on boron nitride can be subdivided into two groups, one containing high amounts of cBN (BH) and the other containing low amounts (BL) (Tables 4.1 and 4.2). Types with high amounts of cBN consist of 80–90% cBN and a metallic W-Co or ceramic binder phase based on titanium and aluminium. Their grain size varies within a range of 0.5–10 μm. Low cBN-containing types are generally 45–65% cBN and have a ceramic binder phase based on titanium carbide or titanium nitride [Krat06]. They are as a rule very fine-grained (grain diameter < 2 μm). As Fig. 4.60 shows, the thermal conductivity and resistance of BN cutting tool materials against abrasive wear increases with

182


Table 4.1 PCBN types (BH) available on the market for cutting sintered steel and sintered forged steel, acc. to KRATZ [Krat06] cBN content/ cBN grain vol% diameter/μm

Binder phase

Knoop hardness/GPa

Thermal conductivity/ Bending (W/mK) strength/GPa

80 85 90

Co-W Co-W Co-W

26–30 38–41 40–43

80–100 80–100 100–110

0.5–4 0.5–5 0.5–4

1.4 1.0 1.3

Table 4.2 PCBN types (BL) available on the market for cutting hardnened sintered steel and sintered forged steel in continuous cut, acc. to KRATZ [Krat06] cBN content/ cBN-grain vol% diameter/μm Binder phase

Knoop hardness/GPa

Thermal conductivity/ Bending (W/mK) strength/GPa

45 50 60

24 28 33

34 44 45

0.5 2 1

Ti ceramic Ti ceramic Ti ceramic

0.6 0.6 1.2

increasing cBN content and grain size. Fine-grained BL types have a higher cutting edge quality, an essential prerequisite for producing high-quality surfaces in the hard finishing of components. Tools made of cubic boron nitride are preferred for turning, drilling and milling hardened steel with a hardness of 55–68 HRC, high speed steel, grey cast iron, hard cast iron, sintered metals and high temperature alloys based on nickel and cobalt – materials which are very difficult to machine using cemented carbide tools. The main uses of BH types are generally roughing/pre-machining of hardened steel and

Cutting edge quality

80 ce

tan

50 45

ar

We

is res

Low cBNcontent 0

250 m/min) and notch wear on the major and/or minor cutting edge for low cutting speeds (vc < 250 m/min). Besides low cutting speeds, small corner radii also are advantageous for the formation of notch wear. An emulsion is used as a cooling lubricant during the cutting process – dry machining is generally impossible.

4.6 Tool Designs Various tool designs for separate processes have become available on the market, since different areas of use place particular demands on the design of the tools. In order to guarantee tool functionality, the following requirements should be considered already in the design phase: • mechanical tool stresses (cutting forces), • thermal tool stress (friction and deformation heat, coolant), • rapid changing and secure positioning of the tool cutting edge,

4.6

Tool Designs

185

• simple and fast replacement of worn tool parts, • versatility in use, • manufacturing and maintenance costs Two main tool groups have emerged based on these requirements. Tools can be designed either as • solid tools or as • tools with inserts. Included in the solid tool category are tools whose cutting part and shaft have been manufactured with the same material. In the case of tools with an insert, the shaft usually consists of a tougher and cheaper base material and the insert of a harder, more wear-resistant material (e.g. cemented carbide, cBN). Different designs have become established for different manufacturing processes. For example, in the case of milling and drilling, solid tools are more economical, whereas they are only used in special cases for turning. In the following introduction to tools designs, special attention will be paid to turning tools. Further specifications regarding tool systems of other chip-removing machining processes with defined cutting edges can be found in Chaps. 9 and 10.

4.6.1 Solid Tools Solid tools are turning tools whose tool body and shaft consist of one material (e.g. high speed steel). This tool type, also called a tool holder bit, is fabricated by grinding the cutting part geometry on to a wide variety of basic shapes. Parting tool holder bits already have a ready-made geometry. The angles required for the parting process are ground up to the tool orthogonal clearance. After grinding the tool orthogonal clearance, these tools are placed in a clamping as shown in Fig. 4.62. Final grinding occurs only on the flank face. Any necessary height correction is made by subsequent adjustment of the tool holder bit in the clamp. One advantage of these tool holder bit is that theoretically freely selectable geometries can be ground onto them for every regrind. Figure 4.63 shows two forming tools in both round and flat design. The latter is designated as a tangential tool because the tool can be tangentially shifted, as opposed to the radial tool, which must be radially adjusted. Round tools are turned after regrinding by a corresponding angle. These types of forming tools are reground only on the rake face, since any regrinding of the flank face would alter the shape, which is very expensive to produce. These tools are advantageous only for turning materials that do not require chip formers, since in these cases they can simply be reground. Round tools can be prepared until the circumference of the cutting part has been reduced to 25%.

186


Fig. 4.62 Parting off tool clamping Holder bit

Solid tools are being used increasingly less commonly for turning, whereas usually tools with inserts are used. If soldering the cemented carbide inserts poses problems (e.g. in the case of small tools), turning tools can be fabricated completely out of cemented carbide. In order to guarantee a small amount of shaft deflection, in the case of boring bars for example, complete cemented carbide shafts are utilized, since cemented carbide has a larger elastic modulus.

γ°

α°

e

Workpiece

γ° α° Round formed turning tool

Flat formed turning tool

Fig. 4.63 Forming tools

4.6

Tool Designs

187

4.6.2 Tools with Inserts In accordance with joining techniques, tools with inserts should be classified as being either • tools with fixed inserts (material connection) and • tools with detachable inserts (frictional/positive connection).

4.6.2.1 Tools with Fixed Inserts Soldering as a technique for joining the insert and tool shaft is now only used in exceptional cases for turning processes. It is more common in other areas, such as in milling and drilling processes or in the manufacture of woodworking tools. Today, inserts are most usually screwed or clamped. For turning tool shafts, unalloyed construction steels with a strength of 700–800 N/mm2 have proven successful. If tools for large chip cross-sectional areas are required, the strength should be increased to 800–1000 N/mm2 . Shaft cross sections are standardized in DIN 770-1 and DIN 770-2. In the case of soldered tools as in Fig. 4.64, usually inserts made of PCBN or cemented carbide are soldered onto the tool shaft. High-speed steel is usually only

Parallel ground in chip breaker

Cemented carbide

Shank

Soldered joint Brazed carbide cutters

Fig. 4.64 Turning and milling tools with brazed carbide tipped cutters

188


soldered onto tools that have lower cutting edge requirements (e.g. in the case of saw blades for woodworking). 4.6.2.2 Tools with Detachable Inserts As opposed to tool systems with soldered inserts, tools with clamped or screwed inserts have the advantage, among others, that several cutting edges can be used on a single insert. If a cutting edge has reached the end of its service life, the insert is turned or twisted after releasing the attachment, thereby bringing a new cutting edge into action. The term “indexable insert” comes from this process. The designation system for customary indexable inserts is standardized in DIN ISO 1832 and shown in Figs. 4.65 and 4.66. On these tools, the inserts are attached with clamping devices on the tool holder. Of particular advantage is the safe and quick clamping of the inserts. In accordance with the selected tolerance class, insert replacement can be accomplished quickly during manufacture, since the cutting edge does not need to be positioned with respect to the workpiece. The manufacturing tolerances of indexable inserts influence the production accuracy of the workpieces to a very large extent when there is a change of inserts (e.g. on NC machines). For inserts, a distinction is made between normal and precision design. In the case of a normal design, tolerances are in the range of ± 0.13 mm and for precision designs ± 0.025 mm, whereby workpiece tolerances of about ± 0.1 mm can be observed after an insert change without additional positioning. According to the forms described in DIN ISO 1832, there is a variety of special designs that are non-standard inserts for machining processes with defined cutting edges. Since clamping holders make it possible to clamp inserts made of diverse cutting tool materials, they can be adjusted well and quickly to the machining task at hand. Storage costs are relatively low, since these are limited essentially to storing the inserts and replacement parts for the tool holders [Corn74]. One must bear in mind that the components of different products are not interchangeable with regard to storage. For this reason, it is uneconomical to use different holder designs in one plant for the same purpose, since this makes the storage of replacement parts considerably more expensive. There are now intelligent tool dispensing systems provided by different tool manufacturers that make it possible to manage the tools and to provide orders automatically adjusted to the purpose of the operator [Pitt04]. As shown in DIN ISO 1832, there are inserts with and without holes. Inserts without holes generally have a superimposed chip former and ones with holes have moulded chip formers. Inserts made of ceramic materials and PCBN are usually manufactured without a hole and can thus only be fixed in clamping holders and clamping fingers. Clamping holders for turning are standardized in DIN 4983. Figure 4.67 shows the standard description for clamping holders by means of an example.

4.6

Tool Designs

189

Example of insert notation: C

N

M

G

12

04

08

...

...

...

1

2

3

4

5

6

7

8

9

10

2 – Tool orthogonal clearance (α 0)

1 – Insert form (εr) 85°

A

82°

B

86°

M

C

80°

P

D

55°

R

E

75°

S

55°

K L

F 3°

B

25° G

5° C

30° N

7° D

T

H

A

O

0° P

15°

V

35°

W

80°

11°

E 20°

3 –Tolerance class Acceptable variation in mm for m

T

A C d E m F C G S H d J1 K1 D L1 V M N d m U

d / mm

m / mm

s / mm

± 0.025 ± 0.025 ± 0.025 ± 0.013 ± 0.025 ± 0.013 ± 0.05–0.152 ± 0.05–0.152 ± 0.05–0.152 ± 0.05–0.152 ± 0.05–0.152 ± 0.05–0.152

± 0.005 ± 0.013 ± 0.025 ± 0.005 ± 0.025 ± 0.013 ± 0.005 ± 0.013 ± 0.013 ± 0.08–0.202 ± 0.08–0.202 ± 0.13–0.382

± 0.025 ± 0.025 ± 0.025 ± 0.025 ± 0.05–0.13 ± 0.025 ± 0.025 ± 0.025 ± 0.025 ± 0.013 ± 0.025 ± 0.05–0.13

1 Inserts

with ground edges on the insert size (See ISO-norm 1832)

2 Depending

S

Fig. 4.65 Standard description for inserts, acc. to DIN ISO 1832

Tool holders for inserts must fulfill the following basic tasks: • Indexable inserts must always be clamped in the identical position after replacement. • Position change due to cutting forces must be avoided. • The support area must guarantee that the insert does not bend. • The mounting must guarantee that heat arising in the process is diverted well into the tool holder.

190

4 Cutting Tool Materials and Tools 5 – Size of insert

4 – Fixation and/or chip breaker A

M

R

B

N

T

C

Q

W

H

X

β = 40–60° Drawing or accurate description required

l

l

β = 70–90°

β = 40–60°

l

l

l

l l

6 – Thickness of insert

d

8 – Formation of cutting edge E

s

s

F sharp

S chamfered + rounded

9 – Cutting direction

s

s

T chamfered

rounded

right 01 s / mm

T1

02

T2

03

T3 R

1.59 1.98 2.38 2.78 3.18 3.97 left 04

05

06

07

09

s / mm 4.76 5.56 6.35 7.94 9.52

L neutral

7 – Design of cutting edge rε / mm

rε

02

0.2

04

0.4

08 12 16 24

0.8 1.2 1.6 2.4

R

00 for diameter in inches converted in mm.

M0 diameter in metric system.

N 10 – internal notations Condition of the edge, special chip former geometry, for example: P = polished W = with (expanded) finishing cutting edge

Fig. 4.66 Standard description for inserts, acc. to DIN ISO 1832 (continued)

• Cutting forces should be transferred to the holder such that the centering of the insert is supported. • The use of a chip former is required depending on the machining task at hand. Figure 4.68 shows different fixture designs for indexable inserts with (below) and without (above) a hole. The advantage of mountings for hole inserts is that all clamping elements in the holder are protected from chips. The indexable inserts

4.6

Tool Designs

191

Notation example for tool holders: Cartridge DIN 4985 –

C

S

G

N

R

32

25

M

16

Q

Kind of fixation of the insert (clamped from above)

C

Basic form of the insert

S

Form of the tool holder

G

Tool orthogonal clearance of the insert

N

Design of the tool holder

R

Height of the cutting edge (here: h = 32 mm)

32

Shank width of the tool holder (here: b = 25 mm)

25

Length of the tool holder (here: l1 = 150 mm) Size of the insert

M 16

Code letter for special tolerances

Q

Fig. 4.67 Standard description for tool holders, acc. to DIN 4983

are unalterably centred and fixed in the insert seat, for example, by toggle levers or pins. The simplest clamping is achieved with a clamping bolt. Another possibility is fixing with a clamping claw, which clamps the indexable insert and a continuously

Chip former Clamping device

Insert

Clamping screw Excentric nut Clamping claw Tool shank

Underlay

Toggle

Clamping screw Clamping screw

Fig. 4.68 Types of insert fixation (Source: Widia, Hertel)

192

4 Cutting Tool Materials and Tools Positive insert

Negative insert

–6° 5°

6°

6°

Wedge angle β0 < 90°

Wedge angle β0 = 90°

11°

11°

Number of useable edges depending on tool type

Fig. 4.69 Characteristic properties of positive and negative inserts

adjustable chip former. In this design, the chip former can be adjusted especially well to changing cutting conditions in order to guarantee a favourable chip form. Tool holders with clamping fingers are used when the chip former is gradually adjustable and working conditions are mostly constant. If damage to the tool holder occurs, the individual parts can be replaced with replacement parts offered by the tool holder manufacturer. In order to shorten setting-up and auxiliary process times in large-batch production, often special designs are used such as short tool holders or cassettes. Primarily for linked machines or tracer lathes, tools can be employed that can change used indexable inserts automatically during workpiece loading without loss of time. There are negative and positive inserts. The criterion for this distinction is the size of the tool orthogonal rake angle when fixed, i.e. in the machining position. If there is a positive tool orthogonal rake angle, the insert is referred to as a positive insert and vice versa (Fig. 4.69). Positive inserts have usable cutting edges only on the upper side. Indexable inserts for tool holders with incorporated positive tool orthogonal rake angles are provided with tool orthogonal clearances. If, as in Fig. 4.69, the tool orthogonal clearance of the insert is 11◦ (wedge angle β0 = 79◦ ) and the tool orthogonal rake angle of the holder is + 5◦ , then the tool orthogonal clearance during tool engagement is + 6◦ . Negative indexable inserts have a wedge angle of 90◦ , making cutting edges available on both the upper and lower side of the insert. Inserts with moulded or ground chip formers have the same basic shape as negative indexable inserts, but in effect they cut with a positive tool orthogonal rake angle

4.7

Tool Preparation

+ 18°

193

+ 10°

+ 12°

+ 12°

Fig. 4.70 Inserts with chip former (Source: Sandvik)

due to the geometry of the chip former (Fig. 4.70). This type of indexable insert thus combines the advantage of the positive tool orthogonal rake angle, a lower cutting force, with the higher number of cutting edges of the negative insert. The main function of the chip former is to produce short-breaking chips. On automated manufacturing devices, in which case process interference due to uncontrolled chip breakage is particularly disruptive, indexable inserts are therefore almost always provided with chip formers [Muel93]. One must bear in mind that moulded/ground chip formers should be adjusted to the respective cutting conditions [Sche78]. In order to expand the possible range of applications, often several chip breakers are arranged in a row. Small chip formers arranged near the corner radius guarantee good chip breakage in finishing operations. Figure 4.71 provides examples of tools with detachable indexable inserts for drilling, milling, sawing and turning. The large tool included angle of square inserts gives them a high level of cutting edge stability. As opposed to triangular inserts, they can only be used in form turning to a limited extent. Triangular indexable inserts have less cutting edge stability as a result of their small tool included angle. Very high surface qualities can be obtained by using round indexable inserts. The disadvantage of these however is that the smallest workpiece radius to be produced is prescribed by the cutting insert’s geometry. Rhomboid inserts have been manufactured especially for copy turning operations. With these, deep and round contours can be reproduced.

4.7 Tool Preparation Expensive tools, e.g. broaching tools, are as a rule prepared again, provided the cost to do so is lower than that of a new tool. Regrinding is almost exclusively done in

194

4 Cutting Tool Materials and Tools Drilling tool

End milling cutter

Deep hole drilling tool

Side milling cutter

Drill out tool

Grooving tool

Side and face milling cutter

Saw blade

Fig. 4.71 Typical application examples

the case of special tools and tool designs with soldered inserts. Other tools that are typically prepared are solid carbide tools like drills and shaft millers, serration tools made of HSS and indexable inserts made of PKD or PCBN. Preparing indexable inserts made of cemented carbide or HSS is not economical because of the low material value. If a tool is to be prepared, regrinding must take place in a timely fashion in order to save on grinding and tool costs. This means that the cutting edge should not be used beyond a permissible level of wear. Figure 4.72 gives an overview of standard wear data used as tool life criteria for tools. Grinding or regrinding tools is carried out as a rule on special 5-axis tool grinding machines. Only in this way we can always grind on the same angle and chip former geometries. Manually ground faces are imprecise and thus have a lot of influence on tool life and the chip forms produced. When grinding high speed steel, the hardness and grain size of the grinding wheel must be adjusted to the high speed steel, depending on which grinding type is used, such as coarse or fine grinding. cBN or precious corundum discs are used as abrasives. Extensive information on the topic of abrasives and method variants of grinding can be found in [Kloc05a]. Special manufacturing processes in the area of finishing require sharply ground tool cutting edges. Before coating the tools, their surfaces are pretreated by irradiation. This produces a pronounced cutting edge radius on the tool. In order to be able to use sharply ground tools for finishing purposes nonetheless, different grinding strategies have been developed. Figure 4.73 shows different regrinding strategies corresponding to the type of wear arising. If crater wear is predominant, the flank face is reground so that the tools maintain the more wear resistant coating on the rake face. On the other hand, if flank face wear is predominant, the rake face is reground.

4.7

Tool Preparation

195


Measurand

Width of flank High speed steel

Cemented carbide

Assumed wear data

VB

0.2 to 1.0 mm

VBmax

0.35 to 1.0 mm

Crater depth

KT

0.1 to 0.3 mm

Width of flank

VB

0.3 to 0.5 mm

VBmax

0.5 to 0.7 mm

KT

0.1 to 0.2 mm

VB

0.15 to 0.3 mm

KT

0.1 mm

wear land

wear land Crater depth Width of flank

Cutting ceramics

wear land Crater depth

Fig. 4.72 Guidelines for the tool life criterion

Dominating flank wear... Flank wear

Substrate

Dominant rake face wear...

Crater wear (rake face) Substrate

Coating

... regrind from the rake face:

Coating

...regrind from the flank:

Common processes:

Common processes:

Drill out Reaming Broaching

Grooving Parting off

Fig. 4.73 Machined surfaces in regrinding

Regr

ind

Regrind

196


In regrinding, the grinding strategy and thus the surface quality has a large influence on the realizable tool life travel paths of the tool. An error-free regrind must therefore be guaranteed by a correct choice of abrasive and adjusted machining conditions. Moreover, one must take care during the grinding process that the tool does not become heated beyond annealing temperature in order to avoid rim zone damage. The avoidance of cracks caused by heat and force is of especial importance when preparing tools with soldered cutting inserts [Weir64]. If cemented carbide is used, regrinding, fine grinding and grinding chamfers and curves is done with diamond discs. Electrolytic grinding is also a very suitable method for pregrinding and finishing tools [Rein69].

Chapter 5

Finite Element Method (FEM)

Machining processes have been modelled numerically with the finite element method (FEM) for some years, leading already to highly promising results in the modelling of cutting processes. The use of numerical models for simulating cutting processes makes it possible to illustrate complex tools while simultaneously taking plasto-mechanical and thermal processes into consideration. Besides FEM, analytic and empirical process models can also be consulted that have the advantage of providing a quick representation of the process. Empirical models are of limited use: they are generally only calibrated to be valid for a limited process range. Due to the simplification, analytical models are only partially suitable for describing complex processes such as can be described by FEM. FEM is a numerical method for finding approximate solutions to continuous field problems. Originally, it was developed to solve stress problems in structural mechanics, but its use was soon expanded to the large field of continuum mechanics [Bett03].

5.1 Basic Concepts of FEM It is helpful for the user to be acquainted with the basic concepts of FEM in order to avoid errors and to be able to evaluate the results of calculations. The user should be aware of the assumptions made by the software and their possible effects on the calculation result. The consideration of a process with FEM is also called finite element analysis (FEA). The following steps are taken in every FEA [Hueb82, Zien00, Redd93, Roll93]: 1. 2. 3. 4. 5.

discretization of the continuum, selection of interpolation functions, determination of the element properties, assembly of the element equations and solution of the equation system.

As the continuum is discretized, solution domains (e.g. the workpiece) are subdivided into a finite number of subdomains – the finite elements. The type, number, size and distribution of the elements is also determined. Continuum elements make F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_5,

197

198

5 Finite Element Method (FEM) 1D

Bar element

2D

Triangle

Quadrangle

3D

Tetrahedron

Pentahedron

Pyramid

Hexahedron

Fig. 5.1 Types of elements for the discretization of continuum problems, acc. to STEINBRUCH [Stei98]

it possible to grasp all normal and shear stresses. The complete definition of an element type includes the element form, the number of nodes, the type of node variables and the interpolation functions. Figure 5.1 shows exemplary some element types that are used for discretization. The selection of interpolation functions, which are often designated as shape, form or basic functions, is made in practice simultaneously with the selection of the element type. Interpolation functions serve to approximate the profile of state variables within an element. Element nodes function as support points for the interpolation. In the case of linear elements, these are the corner points of the element. Higher order elements have a set number of additional nodes on the element edges or inside the elements; due to their higher number of nodes, they provide a more exact solution. Because of their differentiability or integratability, polynomials are frequently used as interpolation functions. The order of the polynomial depends on the number of element nodes, the number of unknowns of each node and the continuity conditions at the nodes. Since the interpolation functions represent the properties of the state variables within the element, the values of the state variables at the nodes represent the unknowns of the discretized problem. After choosing element types and interpolation functions, the element equations (element matrices) are determined. These equations describe the relations between the primary unknowns (e.g. speed, displacement, temperature) and the secondary unknowns (e.g. stresses). To determine the unknowns, several approaches can be considered. One approach for example is using the principle of virtual work (energy). One of the fundamental differences between FEM and other numerical methods of approximation is that the solution is first formulated for each individual element.

5.3

Explicit and Implicit Methods of Solution

199

In order to approximate the properties of the total system comprising the sum of all the elements, the element matrices are combined (assembled) into the global matrix of the problem. The boundary conditions (clamps, external forces, etc.) are also defined. The assembly of element equations leads to system equations that can be solved with the help of the right methods. The numerical integration methods used to solve the element matrices require the evaluation of the integrals at certain points within an element, called integration points. The number of required integration points can be reduced while maintaining the same accuracy by careful selection of their positions. The Gauss quadrature is a very common method for numerical integration. The positions of the integration points within an element are exactly set and represent the positions at which stresses and strains are calculated [Koba89, Roll93, Zien00].

5.2 Lagrangian and Eulerian Considerations of the Continuum The continuum can be discretized from different standpoints, where the LAGRANGian and EULERian approach are the most common in FEM [Bath96]. In the case of the LAGRANG ian approach, the nodes of an element move with the material. An observer travelling on a node would see state variable changes of a particular particle throughout the entire forming process. One disadvantage of the Lagrangian method is the distortion of the mesh brought about by large plastic deformations, which sometimes requires remeshing. The now necessary interpolation of the state variables from the distorted to the newly generated mesh leads, depending on the number of remeshing cycles, to an undesirable, more or less distinct smoothing of the state variables. The Eulerian approach considers the motion of the continuum through a fixed mesh. An observer on a node of such a mesh would see the states of all particles that pass his fixed observation point. This method is especially suited to the investigation of stationary processes and is frequently employed in flow simulations. The “arbitrary LAGRANGian EULERian” method (ALE) is becoming more and more accepted, which is a combination of the above approaches and permits the mesh a motion independent of the material as long as the form of the domains under consideration remains the same [Koba89, Wu03].

5.3 Explicit and Implicit Methods of Solution Many FE programs utilized to calculate large plastic deformations make use of “implicit” methods. For highly dynamic applications on the other hand, such as crash simulation, explicit time integration is prevalent in FE programs. Explicit methods consider the process under investigation as a dynamic problem subdivided into time steps. The desired quantities at time t + t are determined solely from the values available at time t. This is done usually with the help of difference formulae. However, this method is only stable if the time step t is smaller

200

5 Finite Element Method (FEM)

than the time it takes for an elastic wave to travel a route corresponding to the shortest element edge. In this way, the possible length of the time step is a function of the sonic velocity c existing in the material. For solids: c=

E ρ

(5.1)

The maximum possible length of the time step thus depends on the density ρ and the elastic modulus E of the material. Since the length of the time step can be in the range of microseconds, a very large amount of computing steps is sometimes necessary. “Mass scaling”, i.e. artificial increase of the material’s density or artificial reduction of the process time, represents an attempt to increase the possible length of the time step. Mass effects caused by such interventions have to be compensated by appropriate countermeasures [Roll93, Chun98]. When using implicit methods, there is no such limitation. Implicit solvers look for the solution for every time t +t under consideration of the values of the desired quantities both at time t as well as at time t + t [Hueb82]. The solution of such a non-linear system of equations demands special iteration methods (e.g. NEWTONRAPHSON) [Roll93, Zien00]. The advantage of the length of the time step being up to 1000 times larger compared to the explicit method is therefore accompanied by the computing time required for the iterative equation solution.

5.4 Combined Thermal and Mechanical FEA In cutting processes, heat is generated both from inelastic deformation as well as from the work of friction on the rake and flank faces. In order to take thermal processes into account, mechanical and thermal calculations must be combined. In the case of simultaneous combination, this is accomplished by establishing a completely combined equation system. Non-simultaneous combination proceeds from a purely mechanical formulation in which temperature merely serves to help ascertain temperature-dependent material characteristic values (e.g. the flow curve). The mechanical calculation determines the heat of friction, heat from plastic deformation and heat exchange with other objects or the environment. This data serve as input quantities for the thermal calculation. The combination can take place at every iteration (iterative combination), or at every time increment (incremental combination) [Kopp99].

5.5 Nonlinearities Linear analysis is the analysis of a problem that exhibits a linear relationship between the applied load and the response of the system. Linear analysis is a simplification because every real physical system is nonlinear. These nonlinearities can

5.6

Material Laws

201

however be neglected in many cases. If we speak of nonlinear problems in FEM, these nonlinearities must be taken into consideration. Basically, there are three different types of nonlinearities: • material nonlinearities, • geometric nonlinearities and • nonlinearities in the boundary conditions. Material nonlinearities result, for example, from a nonlinear relation between stress and displacement, as occurs in the case of metallic materials after leaving the Hookian range. Further causes of material nonlinearities include material behaviours contingent on forming speed and/or temperature as well as material failure. Geometric nonlinearities are caused by change in geometry during the calculation. As soon as the displacements are large enough to influence the behaviour of the system, there is a nonlinearity. Nonlinearities in the boundary conditions arise, for example, when there is a change in external loads or new contact, or in the case of loss of contact between two objects (e.g. tool and workpiece). In the context of a typical machining simulation, usually all three types of nonlinearity appear.

5.6 Material Laws The material laws applied in FEM can basically be classified in two main groups: those concerning elastic material properties by considering the material elastically and in case of further deformation plastically, and those concerning the material as rigid until the plastic limit is reached (Fig. 5.2). The use of a rigid/plastic material law speeds up the calculation and provides satisfactory simulation results for many applications, in which the amount of plastic deformation is significantly larger than the amount of elastic deformation [Roll93]. Elastic/plastic material models are of importance when elastic effects are not negligible (e.g. the calculation of residual stresses remaining in the component). Material laws for the simulation of chip-removing machining tasks must satisfy the particular requirements of cutting. To this belongs the description of material flow stress as a function of deformation ε, the rate of deformation dε/dt and temperature T. One can see from Table 5.1 how extreme the conditions are in machining in comparison to other manufacturing processes. Large deformations (up to ε = 5) and very high deformation rates (dε/dt < 106 1/s) combined with high temperatures (T < 1500◦ C) are common in cutting processes. To determine the flow curves that are valid in these extreme conditions, special test methods are used. Among these is, for example, the use of the Split-Hopkinson pressure bar. This testing apparatus specially developed for high speed deformation can reach deformation speeds of up to dε/dt = 104 1/s. The arrangement of the Split-Hopkinson pressure bar consists of two cylindrical bars of equal diameter arranged in a line, the input and output bar. Between them is found a sample of a smaller cross-section than that of the bars. An accelerated mass strikes the input bar. This sudden stress induces an elastic compression wave

202


Material models for strong deformation

Material models without elastic behaviour

Rigidplastic

σ

Materials model with elastic behaviour

σ

ε1 , ε 2

ε1 ≠ ε 2

Elastoviscoplastic

Elastoplastic

Viscoplastic

ε3 ε2 ε1

σ

σ

ε1, ε 2

ε2 ε1

ε1 ≠ ε 2

ε1 < ε2 < ε3

ε1 < ε 2 < ε3

ε

ε3

ε

ε

ε

Fig. 5.2 Classification of the material models for high plastic deformation, acc. to ROLL [Roll93]

Table 5.1 Comparison of deformation, rate of deformation and temperature for various manufacuring processes, acc. to JASPERS [Jasp99] Manufacturing process

Deformation

Deformation rate/s

Thomologous a

Extrusion Forging/rolling Sheet forming Machining

2–5 0.1–0.5 0.1–0.5 1–5

10–1 –10–2 10–103 10–102 103 –106

0.16–0.7 0.16–0.7 0.16–0.7 0.16–0.9

aT homologous

= T/Tmelting

that passes through the input bar and is measured at the first strain gauge. When it arrives at the contact area between the input bar and the sample, the wave is divided due to the change in cross-section in the following way: part of the elastic wave is reflected and the rest goes through the sample and plastically deforms it (because of the smaller cross-section of the sample). The remaining part of the wave which proceeds to the output bar is measured by the second strain gauge. With the timestrain curve determined by the strain gauges, both the compression speed and the flow curve can be determined [Abou05]. In order to reduce the number of experiments to a minimum and to make flow curve extrapolation possible, a constitutive material law is required that can be implemented in a FE program. This law must be capable of describing mechanical material properties under tensile, compressive or torsion load for a broad range

5.6

Material Laws

203 Tempered chamber

Projectile Input bar

Output bar

Strike bar Bearing Compressed air tank

Mass of projectile: Projectile speed: Compression speed:

Specimen

Buffer

m = 0.64 / 2.01 kg v = 8–50 m/s dε/dt = 103–104 s–1

Fig. 5.3 Split-Hopkinson pressure bar, acc. to ABOURIDOUANE [Abou05]

of deformation, stress velocity and temperature. Several models have been developed for simulation, which take into consideration the influence of deformation (strain hardening), deformation speed (strain rate hardening) and temperature (thermal softening). Several of these are built on microstructure-mechanical foundations, yet most of the models used in cutting simulation are based on empirical methods. For empirical material laws, usually constitutive equations are applied to describe flow curves. These equations make associations between the current values of stress, deformation, formation speed and temperature (σ , ε, dε/dt, T as variables). The constants contained in these equations can be viewed as material parameters and adjusted to the experimental results with nonlinear regression or the method of least squares. SHIRAKASHI and USUI have suggested the following empirical relation, σ = A · εn · ε˙ m · [−λ (T − T0 )]

(5.2)

Which was successfully applied to describe dynamic viscoplastic material behaviour [Shir70]. Here, A, n, m and λ are material parameters. MOLINARI and CLIFTON introduced a similar description, σ = K(B + ε)n · ε˙ m · T −υ

(5.3)

which describes the strain hardening of the material (acc. to SWIFT) and the dependence of stress on temperature in another way [Moli83]. The material model frequently implemented in FE programs

T − Tr m σ = A + Bεn · (1 + C · ln (˙ε /˙ε0 )) · 1 − Tm − Tr

(5.4)

originates from JOHNSON and COOK [John83]. In it, A, B, n, C, m are material constants, (dε/dt)0 a reference velocity and Tr , Tm room and absolute melting temperature. Material strain hardening is described in this model according to LUDWIK, velocity dependence logarithmically and the influence of temperature by means of a power function. However, an analytically closed formulation for the adiabatic flow curve is not possible with this temperature function.

204


JOHNSON and COOK also suggested a modified, simple equation with an exponential temperature dependence,

σ = A + Bεn · (1 + C · ln (˙ε /˙ε0 )) · e−λ(T−Tr )

(5.5)

with which the temperature T by ε and dε/dt can be explicitly expressed. Here, λ is also a material constant, which takes into consideration the influence of temperature. Based on physical foundations, ZERILLI and ARMSTRONG developed a semiempirical model to describe material behaviour, which contains two equations [Zeri87]; the first type is meant for materials with fcc lattices: σ = σG + C2 ε1/2 · exp [−C3 T + C4 T ln (˙ε )] + k · l−1/2

(5.6)

and the second for materials with bcc lattice: σ = σG + C1 exp [−C3 T + C4 T ln (˙ε)] + C5 εn + k · l−1/2

(5.7)

The ZERILLI-ARMSTRONG equation consists of additive components that contain an athermal component (σG : influence of dissolved materials and the initial dislocations density of inclusions), a thermal and velocity-affected component, the LUDWIK expression and a HALL -PETCH relation. Besides σ , ε, dε/dt, and T, the grain size (l: average grain diameter) of the material to be modelled is also applied as an additional parameter. In 1995, a generalized, combined formulation was proposed by ZERILLI and ARMSTRONG [Zeri95]: σ = C0 + C1 · exp [−C2 T + C3 T ln (˙ε )] + C4 ε1/2 · exp [−C5 T + C6 T ln (˙ε )] (5.8)

5.7 Software A number of FE programs have become commercially available that are tailored to simulating the cutting process. These programs are adapted to the requirements of machining technology, thus making it easier for the user to build and carry out simulations. The simplified operation of such specialized program systems usually entails a limitation how much the model can be influenced. On the other hand, “general purpose” systems are indeed highly flexible and can be used for a wide variety of applications, but they demand a large amount of experience for setting up the model as well as a larger amount of time. Depending on the spectrum of applications, the use of different FEM programs within one company is therefore quite common. The following programs are used frequently in cutting simulations: SFTC/ TM TM DEFORM , T HIRD W AVE/ADVANT EDGE and “general purpose” systems like ABAQUS or MSC/Marc [Denk04a]. TM The FE program DEFORM was originally developed for simulating forming TM processes. A DVANTE DGE is a program designed specially for machining, which

5.9

Phases of a Finite Element Analysis (FEA)

205

facilitates setting up and carrying out cutting simulations but is limited with respect to the influence of the user due to its specialization. Both programs are based on the implicit Lagrangian formulation and include an automatic remeshing routine followed by data interpolation from the old to the new mesh in order to avoid TM highly distorted mesh topologies. The DEFORM software also includes the Euler approach for calculating the quasi-stationary process condition. Commercial FE programs make it possible to import object models from CAD volume software over standardized interfaces. Generally, the STL standard interface (Standard Triangulation Language) is used as the exchange format for threedimensional data models; often the interface generated for exchanging product data STEP (STandard for the Exchange of Product model data) is used as well.

5.8 Hardware With the rapid development of the computer industry, computational performance is being steadily improved. While a few years ago costly computer equipment was still required to carry out FEM calculations, nowadays already the majority of personal computers (PC) used at workstations fulfil the necessary requirements to carry out simple calculations with regard to memory and computing power. As a result of the low cost of hardware in the PC sector, the large majority of all commercial FE packages are also available for free or proprietary operating systems used in such computers. In order to handle very complicated problems with a large amount of elements, such hardware can also be operated in “clusters”. Here, several computers are networked in such a way that an expansive problem can be calculated on several computers simultaneously. To do this, this function must be supported by the FE software. The computing time for a single problem does not however become reduced in direct proportion with the number of nodes (here: computers) in a cluster. The effective increase in speed depends to a larger extent on the hardware and software being used. The behaviour of a software in relation to the available computing power is called its scalability. A software scales well if the computing time required for a task is approximately halved when the available computing power is doubled.

5.9 Phases of a Finite Element Analysis (FEA) A typical finite element analysis takes place in three phases from the standpoint of the user: • data preparation with the preprocessor, • calculation and • evaluation of the results with the postprocessor. With the help of the preprocessor, the user makes all the information necessary for modelling the problem available to the software. To do this, the problem must be abstract enough that the software can depict it. The simplifications made and the

206


quality of the data that the user inputs into the system is of decisive influence on the quality of the simulation results. As a rule, preprocessing has the following steps: • • • •

defining the geometry, meshing, inputting the material data and defining the boundary conditions

Modern programs include import filters for common CAD formats. The geometries of the objects involved in the simulation can thus be taken from 2D or 3D CAD data sets of the design or tool construction in electronic form. After defining the geometries of all the objects, they are meshed (discretized). For generating the mesh, most software manufacturers make tools available that make it possible to mesh geometries quickly. If the limitations of these tools are reached, separately offered commercial meshers can be used. If several element types are available, it is up to the user to make a selection between them. However, specialized software packages often specify one element type. The element type and the element density of the mesh have a significant effect on the quality of the simulation result. In principle, the use of a larger number of elements leads to more precise results. To save computing time, adaptive meshes with locally varying densities are frequently used in order to improve the resolution of local gradients in the state variables. Some programs offer the option of an automatic adaptive remeshing in which the element density is automatically adjusted to existing gradients. After discretizing, we input the mechanical and thermo-physical material data. Following this, the boundary conditions (external loads, contact conditions, friction between different objects, speeds) are entered. After calculation, the results are evaluated in the “postprocessor”. Depending on the software, a variety of graphic representations and processing options are possible. Result evaluation also includes critical assessment by the user. It is essential to compare the results with those from experience, rough calculations or experimental results. Potential sources of error in FE analyses include: • discretization errors from geometry interpolation when meshing and interpolation of the state variables, • faulty input data (e.g. material data, process data, friction conditions), • numerical errors (e.g. in numerical integration) and • rounding errors due to the limited precision of the floating point representation in the computer.

5.10 The Use of FEM in Cutting Technology For the most part, two-dimensional models are used to simulate the chip formation process. The FE model treats the free, orthogonal cross section because it is considerable less complex than, for example, external cylindrical turning (diagonal, bound cross section). The corner radius of the insert does not engage with the workpiece

5.10

The Use of FEM in Cutting Technology

207

and can therefore be ignored in the FE model. As long as the tool lead angle κ r is equal to 90◦ , the tool inclination angle λs is equal to 0◦ and the depth of cut is many times larger than the undeformed chip thickness, it is permissible to assume a plane strain state of deformation, and the simulation of the cutting process is possible with a two-dimensional FE model. This cuts down on the time required for the simulation considerably.

5.10.1 Continuous Chip Simulation Most cutting simulations use the Lagrangian method. In the Lagrangian formulation, the FE mesh follows the material, i.e. the nodes of an element move with the material, so that in the case of material deformation the elements are prestressed/distorted. Especially in cutting processes, particularly large deformations are seen in front of the cutting edge, the result of which is that the elements become highly distorted. In addition, considerable deformation and stress gradients appear in the area of the primary and secondary shear zones. For this reason, a finely structured mesh is necessary for a sufficiently precise representation in the FE model. Besides the difficulty of representing large deformations and deformation gradients in the model, there is also the problem of separating the chip from the workpiece. Definite crack formation in front of the cutting edge is still contentious and has not yet been demonstrated particularly in the case of cutting materials with ductile behaviour. In fact, in the case of ductile material, the material is extremely deformed in front of the cutting edge without forming an observable crack such as we see in forming processes. Crack developments have only been observed in the case of cutting brittle materials [Reul00, Schw36]. Basically, three different simulation methods exist for large deformations, deformations gradients and chip separation: The separation can be realized • on the basis of a geometric separation criterion, e.g. the criterion of the distance at which the separation begins as soon as the tool cutting edge has fallen short of a critical distance to the workpiece nodes lying ahead, • on the basis of a physical separation criterion, e.g. exceeding a defined maximum effective strain or a previously set maximum stress, or • by dispensing with a separation criterion (Fig. 5.4). Besides providing a separation criterion in the first two approaches, it is necessary to define a separation line, along which the nodes are separated when the separation criterion is reached. This method lends itself to the use of a geometric separation criterion, i.e. as soon as the distance between the nodes and the cutting edge is below a critical distance – the length of one element edge as a rule – separation of the mesh occurs. The separation of the chip from the workpiece can also be realized by means of erasing the elements before the cutting edge of the tool. Element erasure is undertaken as a function of equivalent plastic strain or

208

5 Finite Element Method (FEM) Geometric separation criterion

Physical separation criterion I

As Bs

Chip

Cs Ds

vc

Hs,w Gs,w Fs,w Ew

Ds

Tool X X Dw

dcr

Bs

Cw

Fs

Es

Hs,w Gs,w Fw

Bw

Cutting plane

vc

Cs

IKR

Es

X

As

Separation criterion

tool X

Ew

Dw

Cw

Bw

Cutting plane

d

Fig. 5.4 Chip formation along a previously defined split line based on a geometrical separation criterion (on the left) and a physical separation criterion (on the right), acc. to VAZ [Vaz00]

a)

Chip

Tool

b)

Fig. 5.5 (a) Distorted mesh topology before remeshing (b) New, undistorted mesh topology after remeshing

previous material damage [Sche06, Oute06]. One disadvantage of the element erasure method is that material is removed from the model – it must be guaranteed that this has no effect on the result of the simulation. When an automatic remeshing routine with subsequent data interpolation from the old to the new mesh is utilized and a purely ductile material is assumed, it is possible to simulate the cutting process without a separation criterion. The remeshing routine is invoked as soon as the elements are critically distorted (Fig. 5.5). A criterion which can be used for automatically remeshing the model and interpolating the data during a calculation should fulfill the following conditions according to HABRAKEN and CESCOTTO [Habr90]: • The criterion represents the quality of the mesh as faithfully as possible. • The value of the criterion increases with increasing mesh deformity. • If there is a remeshing, the value of the criterion is reduced. In the actual remeshing, usually the outer edge of the old mesh is used as the starting point for the new mesh. The new mesh must now approximate the given edge as exactly as possible and balance any tool penetrations [West00]. After the successful remeshing, the data must finally be transferred from the old to the new discretization. The goal of a good data interpolation algorithm is to transfer the solutions of an FE

5.10


209

calculation at the integration points (e.g. temperature, stresses etc.) and node points (e.g. velocities, displacements etc.) of the old discretization as faithfully as possible to the new discretization. In the execution of the data transfer by means of extrapolation and interpolation, faults can arise that lead to a reduction of the size of calculated gradients. To avoid this effect, a fine discretization in the primary and secondary shear zone is necessary. As a rule, the mesher of a FE program offers the possibility of graduating the fineness of the mesh, whereby one should take heed that the transition of sections of varying discretization is defined as continuously as possible. The information for the density distribution of elements is given with the so-called weight factors in the model. Besides the Lagrangian solution method, the Euler model is used to simulate static cutting processes. In contrast to the Lagrangian formulation, the material moves through a fixed mesh. The advantage of this is that separation criteria are unnecessary. By removing the material from the mesh, large deformations in front of the cutting edge no longer lead to highly distorted meshes, and no time-consuming remeshing is necessary. In order to represent a chip geometry that is as realistic as possible, an iterative adjustment of the free chip edges and chip surfaces is made in the calculation process based on an initial meshing of the chip root. Here, the stationary state is modelled: shear bands, entering and leaving processes as well as non-stationary chip formation processes, such as exist in milling, cannot be described by means of the Euler method [Leop01]. The fundamental advantage of the Euler method in comparison to the Lagrangian method is, besides dispensing with chip separation criteria, the shorter computing time for calculating a quasi-stationary process state.

5.10.2 Segmented Chip Simulation Segmented chips are either produced by cracks and pores, adiabatic shear band formation or by a combination of both mechanisms. Segmented chips can be simulated in two ways: • simulation of the segmented chip by deformation localization based on modified material characteristic values, • simulation of the segmented chip by crack initiation based on fracture and crack hypotheses and • a combination of both approaches Deformation localization can be simulated either by a corresponding modification of the flow curve, which causes a softening of the material starting from a plastic limit strain, or by an artificial reduction of the specific thermal capacity and thermal conductivity. Generally, shear localization begins when thermal softening exceeds mechanical strain hardening. This softening leads to a concentration of plastic deformation, which results in a further increase in temperature and with it further concentration of plastic deformation. The process accelerates itself, leading to the formation of an adiabatic shear band [Abou05].

210

5 Finite Element Method (FEM) 1600

Flow stress / MPa

1400 T = 0°C

1200 1000

800°C

800 600

600°C

400

900°C

200 3.50

1.75

Plastic strain

0

0

0

1000°C 0.20 0.40 0.60 0.80 1.00 1.20

Plastic strain

Fig. 5.6 Simulation of the segmented chip formation with vc = 25 m/ min and f = 0.20 mm with appropriate flow curves which contain an artificial material hardening for high deformation

Figure 5.6 shows the simulation of a segmented chip while cutting titanium alloy Ti6-4 with uncoated cemented carbide cutting edges [Mess07]. The flow curves in Fig 5.6 were measured up to a strain of ε = 0.25 with help of high speed deformation experiments at a deformation velocity of 3000 1/s [Klim00]. For larger deformations, a material softening was assumed such as has been determined in the case of Ti6-4 for lower deformation speeds [Doeg86]. Between the chip segments, a localization of the deformations can be recognized, which leads to a softening of the material in accordance with the provided flow curve; the formation of shear bands can result from this. If the segmented chip is simulated by crack initiation, such as occurs when cutting hardened steels, suitable fracture or crack hypotheses must be integrated into the FE model. In this case a distinction is drawn between macromechanical and micromechanical fracture hypotheses. Macromechanical fracture hypotheses describe the part of the form-changing energy introduced that has dissipated up to a crack, which serves as an indicator for the probability of failure of a material. Micromechanical failure hypotheses come from the consideration that a ductile fracture arises as a result of the formation, growth and consolidation of micropores. Considered microscopically, an inhomogeneous plastic deformation leads to the formation of microcracks, which form cavities under external loads [Brod01]. Macromechanical hypotheses are classified as ones that depend on the forming path or ones that are independent of it [Zitz95]. Since the capability to change form depends on the deformation history, hypotheses that only consider a momentary state of local process quantities (i.e. forming path- independent) are of only limited use for predicting the time of damage. On the other hand, hypotheses that take deformation history into account provide a damage value C, which is dependent on stress σ , expansion ε and material-specific parameters a (Eq. (5.9)). This damage value is added up over the forming path until a critical strain ε c is reached. The critical damage value Ccrit of the crack formation is a characteristic parameter of the


i Initial contact

211

ii

iii

iv

Shear initiation

Crack initiation

Sliding

90 75 62,5 50 37,5 25 12,5 0

v End of sliding

vi

vii

New segmentation

Shear initiation

Material speed/(m/min)

5.10

viii Cracking

Fig. 5.7 Simulation of the segmented chip formation with vc = 70 m/ min, f = 0.25 mm and γ = −6◦ (ADI-900)

material [Kloc07]. As soon as an element in the FE simulation exceeds the critical damage value, it is erased. εk (5.9) Ccrit = f (σ , ε, a)dε 0

Figure 5.7 provides an example of a simulation of segmented chip formation based on the integration of a fracture hypothesis when cutting ADI-900. BROZZO’s criterion was used as a fracture hypothesis, which relates the tensile stress to the largest shear stress [Broz72]. Because experimental investigations have shown that the formation of segmented chips when machining ADI-900 is based on two different mechanisms, CHUZHOY’s material model additionally assumes material damage as a function of the state values deformation and deformation velocity in order to take into account the accelerated sliding of the chip segments due to high shear stresses and the heterogeneous graphite sphere structure. In the case of the segmented chip formation under consideration, after first contact and shear initiation a crack develops on the free surface which runs along the shear zone. The stagnation area disappears and the chip segment begins to slide jerkily from the cutting edge. Due to the crack which has developed, the supporting cross-section between the chip segment and the base material becomes smaller. The shear angle gets larger so that the chip segment is further pushed out of the chip root area. Starting from a critical shear angle, the original shear zone collapses and splits

212


up: one part wanders with the chip segment across the rake face; the other part reforms after repeated compression in from of the cutting edge, and the segmentation process repeats itself [Kloc07].

5.10.3 Simulation of the Cutting Process 5.10.3.1 Simulation of the Turning Process Computing power is becoming increasingly cheaper and faster with the rapid development of the computer industry. This makes it possible to simulate cutting processes that do not allow for the assumption of a plane strain state threedimensionally. Figure 5.8 gives an example of chip formation simulation for the external cylindrical turning of a normally annealed heat-treated steel C45E+N. In order to reduce computing time, only one section of the workpiece was treated in the model. The values for friction and heat transfer were taken from the model for machining C45E+N verified by HOPPE [Hopp03]. The mechanical material properties of the normally annealed heat-treated steel C45E+N were determined by EL-MAGD with the help of tensile and compression tests [Elma06]. The tool used was a Al2 O3 /TiN-coated insert of the form CNMG120408. When wear models are integrated into the simulation, not only stresses, temperatures and chip forms can be predicted, but also possible tool wear. USUI discovered in experimental investigations that wear velocity dW/dt can be calculated with the vc = 300 m/min ap = 1 mm f = 0.3 mm

vc

Z X

Y

tc = 0 ms

tc = 0.34 ms

tc = 1.03 ms

tc = 2.46 ms

Fig. 5.8 Simulated first cut in external turning of C45E+N

5.10


213

Tool

PHASE 1

PHASE 4

Thermo-mechanical FE-analysis of the turning process

VB

Adaptation of the tool geometry dependant on wear

tool chip

tool VBK+1 VBK

B

A

Δt = tK+1 - tK tK

tK+1

Cutting time tc

Integration of the wear model

PHASE 2 Calculation of the thermo-mechanical load collective in steady state condition

dW = σ n ⋅ v ch ⋅ c 1 ⋅ e dt

PHASE 3

−

c2 ϑ

Calculation of the nodal wear rate and the resulting wear

Fig. 5.9 Simulation of tool wear

following equation when the normal stress σ n , sliding velocity vch and surface temperature of the tool ϑ are known [Usui78]: C2 dW = σn · vch · C1 · exp − dt ϑ

(5.10)

The material constants C1 and C2 are dependent on the material/cutting tool material combination. To determine them, ALTAN has suggested a numerically combined calibration method in which the specific material constants can be determined based on experimental results for wear velocity and numerically calculated values for σ n , vch and ϑ [Alta02]. Figure 5.9 shows the basic sequence of a wear simulation. After the collected thermo-mechanical load for the stationary state has been calculated in the first and second phases, the current wear rate dW/dt is calculated in the third phase based on the wear model provided in the user subroutine. This is followed by the modelling of the worn tool in the fourth phase, whereby the nodes of the tool contour are shifted as a function of the calculated wear volume W. This 4-phase wear calculation cycle is continued until a user-defined tool life criterion is reached, and the simulation is stopped. Figure 5.10 provides an example of wear prediction of uncoated cemented carbide inserts. Until the wear criterion of VBT = 200 μm is reached, a total of seven

214

5 Finite Element Method (FEM) Ti6-4, depth of cut ap = 1 mm, HW-K20 cutting speed vc = 100 m/min and feed f = 0.2 mm

Flank wear width VB/µm

300

Simulation Experiment

250 200 150 100 50 0 0

50

100 150 Cutting time tc /s

200

250

Vergrößerung 50x-magnification 50-fach

200 µm

Work piece

Vergrößerung 200-fach 200x-magnification

600 µm

tc = 250 s

Tool

200 µm

Fig. 5.10 Simulation of the flank face wear and comparison with the experimental results

4-phase cycles are calculated in the wear simulation. As the comparison of the wear curves of the simulation and the experiment shows, the simulation deviates only in the last calculation cycle. The SEM image of the cutting edge, which was taken at cutting time t = 250 s in the experiment, makes it clear that width of flank wear width VB increases erratically in the experiment as a result of the displacement of the cutting edge. As this point, the crater broke through the cutting edge and thus led to an acceleration of flank face wear. Since wear in the area of the rake face was ignored because of lacking material constants for crater wear in the FE simulation, simulating cutting edge displacement in the direction of the flank face was impossible. 5.10.3.2 Simulation of the Milling Process The following example shows how cutting simulation has been applied in the industrial sector (Fig. 5.11): Based on a numerical chip form simulation, the tool geometry of an insertedtooth cutter was optimized with respect to chip flow in the product development

5.10


215 Z X

Y

Chip shape simulation for the calculation of the thermo-mechanical load spectrum

Ff,c,p

CAD-Model of the tool holder: miller Ø 100 mm equipped with 10 PCD die heads

Ultimate design of tool holder and cutting insert

Fig. 5.11 Chip form simulation for the development of the tool body and the tool holder (Source: Kennametal)

phase. In the typical product development process for indexable inserts for metal machining, prototypes are manufactured and tested in several iteration stages in order to identify the ideal design. The development process is time-consuming, extremely costly and usually lasts up to 8 weeks. By integrating FE simulations of chip formation into the development process, the number of iteration cycles in the design process can be significantly reduced. Furthermore, the design of the tool holder can be improved with the help of FE simulation; however, this type of simulation requires very exact input data regarding force and directions of force, which in the past could only be acquired with costly cutting experiments. Chip form simulation can be of assistance here. By simulating the chip flow, which essentially is determined by the lead angle of the cutting edge to the cutting direction as well as the chip former geometry, which is fixed in the tool body, predictions can be made not only about chip form but also

216


about forces and temperatures. These can then be utilized to help design the tool holder. By incorporating chip formation simulation into the tool prototype development process, in the above example of an inserted-tooth cutter, we can find an optimal combination of adjusted, positive axial and radial lead angles as well as an optimized cutting edge bevel. This in turn leads to a considerably better surface quality after cutting. The chip is removed from the surface being produced, resulting in a very smooth surface. The two-tier cutting edge bevel results in lower cutting forces, which can minimize burr formation. 5.10.3.3 Simulation of the Drilling Process Drilling is the most common means used in tool production. From the standpoint of cost and productivity, modelling and optimizing drilling processes is thus an of Temperature T/°C

Simulation

400

Drill

200

Experiment d 20

Workpiece 6 ratio (dQ / d) %

Specific feed force kf,max /(kN/mm² )

32

5 4

28 24 20 1

3 2 1

0

3

5

7

9

Diameter d / mm

Material Cutting speed Feed Cutting tool material Cutting edge radius Cooling

1

2

3

: C45E+N : vc = 35 m /min : f = 0.012 * d : HW-K20 : rn = 4 µm : dry

Experiment Simulation kf,max = 2 * Fz,max / ( d * f )

4 5 6 7 Drill diameter d/mm

8

9

Fig. 5.12 FE-model of the drilling process compared with the experiment

10

5.10


217

enormous importance in the manufacturing industry. The drilling process is particularly challenging for 3D-FE simulation, demanding a lot of computing power from the hardware and efficient simulation tools due to the numerous influencing parameters involved (e.g. the complex geometry of the drill, cutting edge rounding, different contact friction processes, difficulty definable heat transfer and thermo-mechanical material properties) [Kloc06]. Due to its complexity, the process can not be represented by 2D simulation as can the orthogonal cross section in turning. A plane strain deformation state cannot be assumed in the case of drilling, since there are different cutting speeds along the drill radius and usually very complex, curved cutting edge geometries, causing a transverse material flow. For this reason, only a three-dimensional approach is purposeful. However, this increases the costs of both implementation and computation exponentially, whereby the computing time with today’s computer technology is increased by the third power of the model size. Nevertheless, the tool and workpiece must be discretized in detail as volume bodies into finite elements. In order to reach a satisfactory level of accuracy in the simulation results, the most crucial area of the major cutting edge should be meshed especially finely [Kloc06]. Figure 5.12 gives an example of an FE model of the drilling process. In order to shorten computing time, the drilling process is considered from the point at which the entire major cutting edge is first engaged. The cylindrical workpiece model is adjusted to the material of the first cut process by notching a conic section on the cylinder. To verify the drilling model, both the chip form and the calculated feed force are compared with experimental results. The high level of agreement between them shows that even for such complex processes as drilling, numerical simulation of diverse target figures is possible.

Chapter 6

Cutting Fluids

6.1 The Functions of Cutting Fluids By fulfilling its main functions, cooling and lubricating the machining site as well as carrying away the chips, modern cutting fluid systems make a substantial contribution to the high performance level of many manufacturing processes. This is achieved by removing process heat from the tool/workpiece contact area by cooling and slowing the progress of heat via lubrication. Not only must excessive heating of the workpieces, which leads to expansion, be avoided, but also temperature load on the cutting tool material be reduced. The fulfilment of these functions may at first sound simple, but often requires cutting fluid properties which are not readily combined with each other.

6.2 Types of Cutting Fluids According to DIN 51385, cutting fluids are classified as non water-miscible, watermiscible or water-based. Water-based cutting fluids are fabricated simply by adding the water-miscible concentrate to water (Fig. 6.1). Cutting fluid

Non water-miscible cutting fluid

Emulsifiable cutting fluid

Water-miscible cutting fluid

Water-soluble cutting fluid

Water based cutting fluid

Coolantemulsion

Coolantsolution

Fig. 6.1 Division of the most important metalworking cutting fluids, acc. to DIN 51385 F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_6,

219

220

6 Cutting Fluids

6.2.1 Non Water-Miscible Cutting Fluids Non water-miscible cutting fluids are mostly mineral oils that contain additional active agents to improve lubrication, wear protection, corrosion protection, durability and foaming properties. Additives that improve lubrication (“antiwear additives” or AW-additives) help to reduce friction at the cutting location. To this end, natural fat oils (palm oil, rapeseed oil) or synthetic fatty substances (ester) are added. The polar structure of these additives gives the additives good adhesive properties on the metal surface, on which they form a half-solid lubricating film called “metal soap”. However, the effectiveness of this lubricating film is reduced in temperatures higher than its melting point (120–180◦ C). EP-additives (EP = extreme pressure) are also added to the mineral oils. Compounds containing phosphorous and sulphur as well as free sulphur are used. The chlorine compounds also shown in Fig. 6.2 are now of secondary importance in Germany. Burning used cutting fluids containing chlorine is now only allowed in special incineration sites, since toxic dioxins can possibly be generated in case of uncontrolled burning. This makes disposal much more expensive, so additives containing chlorine are largely avoided. On the metal surface, the additives form metal salts at different temperatures. These salts can absorb high pressures and exhibit only a low level of shear strength. In this way, not only are forces lowered but also heat arising at the cutting location is reduced. The temperature spheres of action of the individual additives can be seen in Fig. 6.2. In addition to mineral oils, low-viscosity ester oils are also used as cutting fluids in machining. Presently, mostly mineral oils (about 90%) are utilized. The reason for this is the lower cost required to procure it in comparison to synthetic ester oils. This

Kind of additive

Lubrication improving additives

EP-additives

Fatty oils (animal, vegetable) Synthetic Fats (ester)

Chloric compounds

Phosphoric compounds Sulfidic compounds

Free sulfur

Temperature sphere of action

until ca. 120 °C

until ca. 180 °C until ca. 400 °C

until ca. 600 °C until ca. 800 °C until ca. 1000 °C

Fig. 6.2 Temperature range of coolant additives (Source: Mobil Oil)

6.2

Types of Cutting Fluids

221

advantage of mineral oils compared with synthetic ester oils is however becoming increasingly smaller due to the steady increase in raw oil prices in the last several years. Compared with mineral oils, ester oils are characterized by a lower evaporation tendency, a higher flash point, more favourable lubrication properties, they are skin-friendlier and they are biologically more degradable [Frei00]. Although ester oils have much better lubrication properties in comparison to mineral oils because of their polar chemical structure, as a rule tribologically active, surface-active additives in the form of phosphorous or sulphur compounds are also added to them to improve these properties further. These additives reduce friction and wear, but they reduce the biological degradability of the oils considerably. With this in mind, the goal of current research [Murr07] is to develop a quickly biodegradable family of fluids based on renewable raw materials that fulfils all of the cooling and lubricating functions required in a machine tool. That means, nonadditive ester oils should not only be used as a cutting fluid in machining, but as lubricating and pressure transmission media in all other tribosystems of a machine tool. The loss of tribological functions of the lubricating medium caused by dispensing with additives should be compensated by appropriate PVD wear protection coatings, i.e. the functions of the cutting fluid should be relocated to the tool surface [Krie01, Kloc06a].

6.2.2 Water-Based Cutting Fluids Emulsifiable cutting fluids are supplied as a concentrate and thinned with water to become emulsions prior to use. The high amount of water used (up to 99%) is the cause of the good cooling effect of emulsions, but they further corrosion for that reason as well. The fluid’s lightly alkaline character (pH-value of 8–9) helps protect against rust. Higher alloyed emulsions contain the additives shown in Fig. 6.2 to improve lubrication and compressive strength. Infestation by micro-organisms such as bacteria, yeasts and fungi are a major problem in emulsions. The result is a decreased pH-value and thus also a reduction of the fluid’s ability to protect against corrosion. There is also an odour nuisance and hygienic conditions for the operating staff are worsened. In addition, the emulsion becomes unstable, i.e. oil is deposited on the surface, and deposits are formed that clog the filter and thus lead to malfunctions. Biocides, which are added to the emulsion at a percentage of about 0.15%, provide a remedy to this. If this value is significantly exceeded, it can lead to dermatological problems in the personnel, while a percentage that is too low has no effect. The most important coolant additives are emulsifiers. They have the function of dispersing the oil in the water so that a stable oil-in-water emulsion is produced after mixing with water. A distinction is made between ionogenic and non-ionogenic emulsifiers. They form a relatively stable film on the boundary surface between the oil droplets and the water, which prevents the oil droplets from coalescing. Examples of emulsifiers include, for example, alkali soaps of fatty acids or naphthenic acids (ionogenic) and reaction products of alkylphenols and athylene oxide.

222

6 Cutting Fluids

The amount of emulsifier determines the size of the oil droplets. In the case of the coarsely dispersed emulsions used in metalworking, this amount is between 1 and 10 μm [Mang01]. Cutting fluid solutions are produced by mixing a water-soluble concentrate with water. They are not generally used for cutting with geometrically defined cutting edges. Their main area of application is in grinding.

6.3 Guidelines on the Use of Coolant Emulsions When coolant emulsions are used, there are some guidelines that should be considered in order not to impair their stability and performance. As the main component of the emulsion, the quality of the water is of decisive importance. The hardness of the water, its most important property, is the result of the content of water-soluble calcium and magnesium salts. It is given in ◦ dH (degree of German hardness) or mmol/l (1◦ dH = 0.179 mmol/l). The water’s hardness should be between 5 and 20◦ dH. If the water is too hard, the emulsifiers react with the calcium and magnesium salts, which leads to the formation of water-insoluble soaps (creaming on the emulsion surface) and reduces the emulsifier content. The useful life of the emulsion is shortened drastically by this. Soft water on the other hand promotes unwanted foaming. There are further water requirements concerning nitrate and chloride content. In general, drinking water quality is sufficient in order to keep the initial load on the emulsion by bacteria, yeasts and fungi at a low level. The nitrate content should not exceed 50 mg/l. When applying coolant emulsions, the water must first be added to the container, then the concentrate. The mixing process should always be carried out with vigorous stirring so that a good oil-in-water emulsion without flocculation is formed. One test point of the finished emulsion is concentration. This can either be done with a hand-held refractometer or an emulsion test flask. With the hand-held refractometer, the concentration can be quickly determined on the spot. It utilizes and displays the relation between the refraction index and concentration. A more exact method consists in separating oil and water in a test flask by adding hydrochloric acid. The disadvantage of this method is that it is more time-consuming. The pH-value of emulsions should in their freshly prepared state be between 8 and 9. This is tested in practice usually with indicator paper, the coloration of which when dipped into the fluid serves a measure of pH. Potentiometrical determination of pH is only recommendable when higher precision is required. During operation, impurities such as chips or dust constantly get into the cutting fluids. They need to be removed from the emulsion, as they have a negative effect on tool life and the manufacturing product in addition to clogging the pumps. In order to get rid of impurities, one can choose between gravitational purification in sedimentation tanks, centrifugal force purifiers (e.g. centrifuges, separators), magnet filters or various designs of belt filters.

6.4

Effects of the Cutting Fluid on the Machining Process

223

When changing the coolant, one should always carefully clean the container, since nests of bacteria immediately infect the new filling, significantly reducing the service life of the emulsion. System purifiers or hot water jet devices have proved effective in this task. Disposal of the coolant is a major cost factor. The separation of the oil from the water necessary to do this can be accomplished chemically (salt or acid separation), by ultrafiltration (membrane technology) or by vaporization/incineration.

6.4 Effects of the Cutting Fluid on the Machining Process As explained in more detail in Chap. 3, the tool is exposed to very high mechanical and thermal loads during the cutting process, whereby the mechanical energy created to form the chips is almost completely converted into heat in the shear and friction zones. The result of these loads are wear phenomena such as mechanical wear and the shearing off of adhered materials, both of which occur within the entire effective cutting speed range, as well as diffusion processes and scaling – effects which manifest themselves only above certain temperatures (Sect. 3.7). By virtue of their lubricating effects, cooling fluids primarily influences adhesive wear, which occurs due to the periodical migration of built-up edges within certain speed ranges [Opit70a, Köni66, Prim69]. Especially wear phenomena associated with adhered material particles in the lower speed range (when using HSS) can be effectively countered with lubrication. To counteract surface pressure, there should be solid layers with high compressive strength and low shear strength on the metal surface, which prevent a direct sliding of materials into each other. This can stop or at least reduce welding. If necessary, this can be accomplished with high pressure additives in the cutting fluid. However, one must bear in mind that sulphur or phosphorous additives only become effective at certain temperatures, and for this reason the composition of the lubricant must be calibrated to the respective operation. The most important prerequisite is that the lubricant has access to the contact zone. In the range of increased built-up edge formation, this condition is met by the fluctuation of the built-up edges. As the cutting speed increases – in the range of reduced built-up edge formation – the conditions for the formation of high pressure lubricating films becomes progressively more unfavourable because the increased chip flow shortens the time for potential reactions between the additives and the metallic surface. At the same time, the increase in temperature leads to diffusion phenomena between the friction partners or, in the extreme case, to plastic deformation of the cutting edge, making it necessary to cool down the cutting area. Accordingly, at these speeds begins the range in which tool life is improved not so much by the lubricating effect of a fluid but by its ability to remove heat, i.e. by cooling. On the other hand, it is thoroughly possible that wear on the tool is considerably increased by the cooling and the tool life is accordingly reduced. This is clarified in Fig. 6.3. There is much more crater wear in wet cut than in dry cut operations. By cooling, the temperature of the flowing chip is lower and thus its strength is

224

6 Cutting Fluids

Fig. 6.3 Wear-cutting time diagram for dry cut and with application of cutting fluid

300

Process:

Cylindrical turning Material: 42CrMo4+QT Cutting tool material: Coated carbide 100 Cutting speed.: vc = 240 m/min f = 0.2 mm 80 Feed: ap = 2 mm 60 Depth of cut:

Crater depth KT / µm

200

40

20 Coolantemulsion 10 8 6 4 Dry cut

2

1

1

2

4

6 8 10

20

40

Cutting time tc / min

higher, which manifests itself in increased forces. Since the cutting fluid mainly cools the top of the chip, and the bottom of the chip remains largely unwetted due to its intense contact with the rake face, a larger temperature gradient is formed in the chip as in the uncooled process. This results in a larger chip curvature, so that the contact surface between the chip and the tool is reduced in size. On the whole therefore, there is an increase in the specific stress on the rake face and consequently more crater wear. One must also consider machining operations with low cutting speeds, which generally are designed so that the built-up edge range is avoided. But when cooling lubricants are used, the temperatures prevailing during the formation of built-up edges shift to higher cutting speeds, so that a process optimized for uncooled cutting may be incorrectly designed [Opit64]. On the other hand, we can expect clear improvements in service life when the cutting temperature is near the softening point of the cutting tool material without the use of a cutting fluid. Figure 6.4 shows how effective the cooling effect of an emulsion is in this case using the example of drilling copper alloys with HSS drills. The use of emulsions permits higher cutting speeds and larger feeds and a significantly augmented tool life [Bömc85]. A further aspect to be taken into consideration is the transport of chips from the cutting area. For example, it can be advantageous when groove milling with

6.4

Effects of the Cutting Fluid on the Machining Process

225

100 80 Dry cut Emulsion

60 50

Tool lifeT(VB = 0.3 mm) / min

40 30

20

G-CuSn12Ni 10 8 G-CuAl10Ni 6 5 4 3 20

30

40 50 60

80 100

200

300

Cutting speed vc / (m/min) Tool: Cutting tool material: Material:

Twist drill (d = 11 mm) HS6-5-2-5 G-CuSn12Ni

G-CuAl10Ni dry

wet

dry

wet

f = 0.1 mm

f = 0.2 mm

f = 0.2 mm

f = 0.4 mm

Drilling depth: l = 30 mm

l = 30 mm

l = 45 mm

l = 45 mm

Feed:

Fig. 6.4 Influence of cooling on tool life during the drilling of copper

cemented carbide-fitted shaft millers to remove chips from the cutting edge with compressed air or cutting fluid so that no chips are pulled into the cut, increasing wear. The disadvantage that cooling increases temperature change stress on the indexable inserts is more than compensated by this (Fig. 6.5) [Köll86]. When drilling, the cutting fluid also fulfils the main function of leading chips out of the drill hole and thereby helping to avoid any clogging of the flutes. Some machines are designed in such as way that specially arranged nozzles clean the working space with cutting fluid so that chips do not hinder subsequent operations or impede the clamping of new workpieces.

Fig. 6.5 Time course of wear during groove milling of titanium in dry cut and wet cut with external compressed air

6 Cutting Fluids Maximum width of flank wear land VBmax / µm

226 600 Dry

400

Compressed air Cutting fluid 200

100 80 60 40

20

10

1

2

4

6 8 10

20

40 60

Milling time t f / min Cutting speed: vc = 50 m/min Feed per tooth: fz = 0.17 mm Tool diameter: d = 32 mm

Width of cut: Depth of cut:

ae = 32 mm ap = 30 mm

6.5 Selection of Cutting Fluids As a summary, those aspects of cutting fluid selection will be mentioned that are of significance from a technological and economical point of view. Emulsions are characterized by a number of properties that predestine them as cutting fluids for use in machining, above all their favourable cooling and cleansing effect as well as the low purification cost for workpieces, tools and chips (Fig. 6.6). Their high cost of maintenance and the working area and environmental problems associated with their use have led to emulsions being replaced with non-water soluble cutting fluids [Kloc96a]. Oils are already being used as cutting fluids in many areas, for example in processes for manufacturing tooth systems, broaching and high-speed grinding with PCBN. The advantage of non-water soluble cutting fluids is on the one hand their technical properties – particularly their good lubricating effect – but especially their much lower maintenance and disposal costs compared with emulsions. Their service life is practically unlimited. Only the discharge need be replaced. From the standpoint of occupational health, the superior skin-compatibility of oils is another essential advantage. The use of non-water soluble cutting fluids is however also associated with a series of disadvantages, such as higher maintenance and refilling costs. The higher viscosity of oil requires an adjustment of the machine with respect to the cutting fluid and filters. The strong tendency to form vapour and mist requires a complete enclosure of the machining area with corresponding suction. The good adhesion

6.6

Reducing or Avoiding the Use of Cutting Fluids

227

Oil

Emulsion

Consumption

10%

90%

Cooling effect

low

very good

Lubricating effect

very good

low

Skin compatibility

good

problematical

Bacteria tolerance

good

less good

WGK 1–2

WGK 3–4

Preparation costs

high

low

Maintanance costs

low

high

unlimited

2 –24 Months

low

high

Resultant wastewater

very low

very high

Machine compatibility

good

problematical

Corrosion protection

good

low

Fire protection requirements

high

not necessary

Effort for workpiece and chip cleaning

high

low

possible

not possible

Water endangering class

Service life Disposal costs

Reuse of cutting fluid

Fig. 6.6 Advantages and disadvantages of using oil or emulsion as cutting fluid

of the oil to the workpieces leads to large drag-out losses, demands long periods of draining in basins as well as extensive turning, spinning and cleansing of the workpieces. The chips must also be spun and, if necessary, washed. Moreover, oil vapour and mist can form a combustible mixture with air, leading to deflagration. In conclusion, a general statement regarding the economical use of this or that cutting fluid must be made on a case-by-case basis.

6.6 Reducing or Avoiding the Use of Cutting Fluids Besides the technological benefits provided by cutting fluids, they also represent a considerable hazard to the environment and humanity. Cutting fluid components such as bactericides and fungicides, reaction products originating in the cutting fluid and included foreign substances can all become the root cause of illnesses. Leakage

228

6 Cutting Fluids

and drag-out losses, emissions, wash water and the disposal of used cutting fluids are a burden on earth, water and air. The handling of cutting fluids and their disposal is therefore being regulated by lawmakers and professional associations with strict requirements. Guidelines regarding cutting fluid requirements and associated devices, regarding maintenance and disposal of the fluids and regarding the laws, regulations and the requirements and rules of professional associations are given, among other places, in BG rule 143, VDI guideline 3397 and the VDMA brochure “Cutting Fluids – Fresh Air at the Workplace” [VDMA 2002]. For companies, these regulations and requirements mean not only great responsibility towards their workers but in particular growing financial burdens. For manufacture, the task of satisfying requirements concerning environmental protection without putting the profitability of the production in danger is becoming increasingly crucial. An approach for this purpose represents the reduction or the avoidance of the cutting fluids use.

6.6.1 Reducing Cutting Fluids A reduction of the use of cutting fluids aims to make avail of only the technologically necessary amount of cutting fluid. Secondary tasks of the cutting fluid, such as chip transport within the machine or bringing those chips to the right temperature must be executed by other measures which concern not only the tool but also the machine tool (Fig. 6.7) [Köni93a]. There are different possibilities for the machine tool. In order to transport chips, inclined machine bed concepts, conveying systems, special chip guiding systems Dicontinuous CF-use

CF-supply change

Optimize filtration

New tool materials

Tool coatings

Alternative machine tool cooling

Control feed pump (NC)

alternative tool cooling

Change chip transport

alternative workpiece cooling

Fig. 6.7 Tool and machine related means of reducing the demand of cutting fluid

6.6


229

or other alternative media (e.g. compressed air) can be used. The machine tool’s temperature consistency, important for form and dimensional tolerance, can be achieved, for example, by means of closed cooling cycles. With respect to tools, there are a number of different approaches for reducing the amount of cutting fluid used. One example is the internally cooled tool. Besides drills with cooling ducts, which have been state-of-the-art for a long time, indexable inserts with “internal” cutting fluid supply are also being used. The cutting fluid is supplied directly to the cutting area by a duct in the supporting tool and the insert. This more effective cutting edge cooling requires not only a significantly lower volume of cutting fluid, but also has, in grooving and parting-off operations, a positive effect on tool wear, tool life, chip removal and the surface quality of the workpiece [Köni93a].

6.6.2 Minimum Quantity Cooling Lubrication (MQCL) A further measure to reduce the amount of the cutting fluid represents the so-called Minimum Quantity Cooling Lubrication (MQCL) [Kloc96a, Köni93a, Wein04]. In this cutting fluid technology, the tools are supplied with the smallest amounts of a coolant and/or lubricant. Usually, oils are used, but also emulsions, water or air. They are supplied to the tool and/or cutting area in the smallest possible quantities. This is achieved either with or without a transport medium. In the case of “airless” systems, the tool is supplied by means of a pump with a medium in the form of individual, rapid, successive finely dosed droplets, usually of oil. In the second case, the medium is atomized into ultrafine droplets with the help of compressed air in a nozzle and supplied as an aerosol to the machining location. In the sphere of dry machining, minimum quantity lubrication is generally understood as the supply of a cooling lubricant medium in the form of an aerosol. Depending on the type and primary task of the added medium, we can draw a distinction between a minimum quantity lubrication (MQL) and a minimum quantity cooling (MQC) (Fig. 6.8). When oils are used, it is their good lubricating effect that stands in the foreground. Their task is to reduce friction and adhesion processes between the workpiece, chip and tool. By reducing friction, there is less frictional heat. The result is less heating of the tool and component in comparison to pure dry machining (Fig. 6.9) [Eise00]. Due to the low thermal capacity of oil (cp,oil = 1.92 KJ/kgK) and air (cp,air = 1.04 KJ/kgK) and the small quantity applied, the direct cooling effect of the oil/air mixture is of only secondary importance. Because of the very small cooling effect of the oil/air mixture, the use of oil as a medium is referred to as a minimum quantity lubrication (MQL). In comparison to oils, emulsions or water are used as media for minimum quantity cooling lubrication much less often. They are generally only used when the tool or component must be cooled more intensively than is possible with oils. Due to the much lower lubricating effect of emulsions and water’s complete lack thereof, the use of these media are also referred to as minimum quantity cooling (MQC).

230

6 Cutting Fluids

Minimum quantity cooling lubrication (MQCL)

Minimum quantity cooling (MQC)

Minimum quantity lubrication (MQL)

Emulsion Water (cp, Water) = 4.18 kJ/kgK Air (cp, Air) = 1.04 kJ/kgK

Oil (cp, Oil) = 1.92 kJ/kgK

Medium

Cooling

Lubrication

Chip transport

Emulsion

very good

good

good

Oil

very good

very good

good

limited

nothing

limited

Compressed air

Fig. 6.8 Definition of “minimum quantity cooling lubrication”

Tool temperature Tmax / °C

Material: C45E

Drill with thermocouples

400

Dry

200

MQL

300 250 200

Thermocouples near the main cutting edge

CF

150 0

40

60

80


100

Process: Drilling Drill: Solid carbide TiN coated d = 11.8 mm f = 0.2 mm Feed:

Fig. 6.9 Influence of minimum quantity lubrication on the tool temperature

Minimum quantity cooling (MQC) with emulsions, water (with a corrosion inhibitor), cold air or fluid gases is still a relatively seldom used component of minimum quantity cooling lubrication technology (MQCL) and is thus little known to users. MQC technology can however most certainly make a contribution to solving thermal problems in the tool or component in dry machining operations. Due to their great importance in dry machining, the following remarks are concerned exclusively with minimum quantity lubrication technology (MQL). The

6.6


External MQL supply 1-channel

231

Internal MQL supply

2-channels

1-channel 1-channel rotary feedthrough

2-channels 2- channel rotary feedthrough

Injector orifice

Supply-systems Compressed air supply

Oil supply

Compressed air supply

Compressed air supply Unpressurized tank

Mixing chamber

Oil

Oil Pressure tank

Pump

Fig. 6.10 MQL supply systems

MQL medium can be supplied to the cutting area from the outside by means of nozzles affixed separately in the machine space or by means of the tool spindle and internal cooling ducts inside the tool (Fig. 6.10). Both systems have their fields of application. In optimally adjusted MQL systems, less than 50 ml/h of the lubricating medium is used. Measured by the fact that up to 6 m3 cutting fluid can be discharged daily in the case of a transfer line with a cutting fluid volume of 60 m3 , this means an enormous reduction of the quantities used in MQL technology. The essential feature of MQL is that, when correctly used, the tools, workpieces and chips all remain dry. In the case of external supply, the aerosol is sprayed onto the tool from the outside with one or several nozzles. The number and orientation of the nozzles as well as the spray pattern, which is dependent on the nozzle design, all have a significant effect on the result. This method is used, for example, in sawing, shaft and knife head face milling and also in turning. In the case of internal machining operations, such as drilling, reaming or tapping, external supply of the medium is only practical up to length/diameter ratios of l/D < 3. For larger l/D ratios, the tool has to be retracted for a rewetting several times if necessary, which can lead to a considerable elongation of the machining process. The use of external supply is also problematic in machining tasks in which tools are used that vary significantly with respect

232

6 Cutting Fluids

to length and diameter. The supply nozzles need to be manually aligned or with the help of positioning systems, which, linked with the machine control, move the nozzles in an axial or radial direction or pivot them by a certain range of angles depending on the tool length and diameter. External MQL supply can however also be indispensable if the tools used do not have internal cooling ducts. In the case of drilling, reaming and tapping with larger l/D ratios, internal supply of the medium by means of a spindle and the tool is advantageous, since the medium is continuously available near the cutting area independently of the drilling depth, and chip removal from the drill hole is supported. This is also similar for tools with highly varied dimensions. In the case of deep-hole drilling, internal MQL supply is indispensable because of the large l/D ratios. Further advantages of the internal supply of a MQL medium is that positioning fields such as are seen when nozzles are used, are avoided, and the integration of the MQL into the machine tool does not require the working space to be restricted by feed lines. With internal MQL supply, we make a further distinction between 1-channel and 2-channel systems (Fig. 6.10). In the case of the 1-channel variant, the aerosol is produced outside the spindle and supplied by the latter to the tool. In 2-channel systems, oil and air are conveyed separately by the spindle. The air-oil mixture is produced directly in front of the tool. The essential requirement of both system variants is that the medium be available at the cutting location at the moment the cut begins in sufficient quantity. Internal MQL supply requires tools with cooling ducts. Currently, drilling tools of < 1 mm diameter with internal cooling ducts are already available. In the case of tools without internal cooling ducts – be it drills, tappers or end milling cutters – external MQL supply is absolutely essential. But even in this case, tool manufacturers are offering special solutions that make it possible to conduct the MQL medium flowing through the spindle within the tool holder outside and then on the tool circumference lengthwise to the machining location. The media used in MQL are primarily fatty alcohols and ester oils (chemically modified vegetable oils). Medium selection depends on the type of supply, the material, the machining method and the aftertreatment of the component (annealing, coating, varnishing). In many material/method combinations, the use of a minimum quantity lubrication is vital to the realization of a dry machining operation (Fig. 6.11). From the standpoint of the material, this is especially true for dry machining aluminium wrought alloys, and from the methodological standpoint, largely independently of the material, for drill hole manufacture and drill hole aftertreatment. The classic application area for MQL technology is sawing. Due to the high hot wear resistance of the coated cemented carbide tools available today, the turning and milling of steel and cast iron materials are done to a large extent completely dry. Figure 6.12 shows the effect of MQL on the tool’s condition and tool life quantity when drilling into an aluminium wrought alloy. In the case of dry machining without MQL, the tool was already unusable after 16 holes due to material adhesion in the flute. When MQL was used, neither wear nor adhered material could be detected after 128 drill holes.

6.6


233

Steel

Aluminium

Process

Wrought alloy

High alloyed steel, bearing steel

Maching steel, heat-treatable steel

GG20 – GGG70

Casting

Cast alloy

Material

Drilling

MQL

MQL

MQL

MQL / dry

MQL / dry

Reaming

MQL

MQL

MQL

MQL

MQL

Thread cutting

MQL

MQL

MQL

MQL

MQL

Thread moulding

MQL

MQL

MQL

MQL

MQL

Deep hole drilling

MQL

MQL

MQL

MQL

Milling

MQL / dry

MQL

dry

dry

dry

Turning

MQL / dry

MQL / dry

dry

dry

dry

dry

dry

dry

MQL

MQL

MQL

MQL

MQL / dry

dry

Hobbing Sawing

MQL

MQL

Broaching

Fig. 6.11 Application of MQL

6.6.3 Avoiding Cutting Fluids The most decisive step in the avoidance of problems associated with the use of cutting fluids is dry machining. Many machining tasks, that still use large quantities of cutting fluids today do not technologically require them. For every current or future machining task therefore, the basic question should be posed of whether one can dispense with cutting fluids or not. This cannot however simply be realized by stopping the cutting fluid supply, even if this brings about a positive effect in isolated cases. A sensible and economically viable dry machining operation demands a very thorough analysis of given restraints as well as an understanding of the complex relations that interconnect the process, cutting tool material, component and machine tool (Fig. 6.13) [Kloc98, Tham98, Adam02, Wein04].

234

6 Cutting Fluids

300 128

Number of boreholes

250 200 150

100 50 16 0

Dry

MQL

Dry

Material: AlSi9Cu3 Tool: Solid carbide drill, d = 8.5 mm Coating: (Ti,Al)N + MoS2

Bling hole: Cutting Speed: Feed:

MQL l = 30 mm vc = 300 m/min f = 0.5 mm

Fig. 6.12 Drilling of aluminium with and without minimum quantity cooling lubrication

Process


Qsp Qws - Heat - Friction - Adhesion - Chips

QW

Dry-

- Hot hardness - Hot wear resistance - Resistance against Adhesion

- thermal - Warming processing - Material behaviour - Surface, Rimzone - Chip transport - Machine accuracy - Dimensional accuracy - Machine concept Surface

Fit Tolerance Workpiece

Fig. 6.13 Dry processing – requirements and constraints

Machine tool

6.6


235

In dry machining, the primary cutting fluid functions of lubrication, cooling and cleaning and rinsing, are omitted. This means for the cutting process on the one hand that stronger frictional and adhesive processes can take place between the tool and the material. It also means that part of the heat produced in energy conversion locations and dissipated by the chip, tool and workpiece is no longer absorbed by the cutting fluid and hot chips are no longer rinsed out of the cutting area or the machine tool. The consequence of this is higher thermal loading of the tool, component and machine tool, which in turn has a negative effect on tool life and component/machine precision. When planning and designing single processes or manufacturing sequences without cutting fluid, the goal must therefore be not only to lay the technological foundations, but to create the prerequisites necessary for dry machining from the standpoint of the component and the machine tool. At present, cutting materials provide the best basis for dry machining. Cemented carbides, cermets, cutting tool ceramics and polycrystalline boron nitride have sufficient hot wear resistance to be used without cutting fluids. Tool coating is particularly important in this regard. This reduces the thermal load on the substrate and reduces frictional and adhesive phenomena between the material and the cutting tool material. Dry machining also leads however to an alteration of the heat flows between the tool and the chip. Since there is no cutting fluid to absorb the heat, more heat must be dissipated by the chip with a comparable heat conversion. This requires in turn that the hot chips are removed as quickly as possible from the working space by a suitable machine tool concept [Kloc98]. While we can do without the use of cutting fluids in many cases in turning and milling cast iron materials, steels, aluminium alloys and non-ferrous metals, conditions are generally more difficult in the case of processes like drilling, reaming and tapping (Fig. 6.11). Problems in dry machining include higher thermal loading of the tool, component and chip as well as the poor chip removal. Chips caught in or welded to the flute reduce the quality of the drill hole and can lead to tool damage. In tapping, compression, friction and adhesion phenomena lead to higher amounts of mechanical tool load. There are a number of drills, taps, fine boring and reaming tools available with special substrates, coating systems and tool geometries adjusted to the particular requirements of dry machining. Dry machining tools exhibit much better wear and performance properties than conventional wet machining tools. Despite promising attempts to expand the field of application of dry machining and to make it more economical by finding appropriate tool geometries, coatings and cutting parameters, it is incontestable that a complete relinquishment of cutting fluids will not be possible for all machining tasks. Restrictions may derive from the method, material or required component precision. Process substitution is a possible alternative. One example of this is the manufacture of internal threads by thread milling and combination drill taps. Interrupted cuts and the use of coated milling tools made of cemented carbide are favourable prerequisites for manufacturing threads by dry cutting, improving surface quality and even reducing manufacturing times [Kloc98]. Effects on the rim zone and form/dimension faults in the component represent further potential restrictions on dry machining processes. Since the component’s

236

6 Cutting Fluids

quality is affected by the amount of heat that flows into the component, the process must be planned such that as little as possible heat enters the workpiece. Because they shorten the contact time between the tool and the workpiece, higher cutting speeds and larger feeds contribute just as much here as larger positive tool orthogonal rake angles, which reduce cutting work. This same is true for the reduction of friction, adhesion and wear by the use of coated tools. The distribution and number of cuts is very important. The operation should take place in one cut if possible. This demands components with volumes to be machined that are as small as possible as well as equal machining allowances. “Near-net-shape” parts, the final contours of which are frequently created with one cut, provide optimal conditions for this. Manufacturing processes will in future no longer only be assessed from the standpoint of the improvement of performance but also with respect to ecological safety. In light of the problems associated with the use of cutting fluids, dry machining is surely the most effective approach in cutting technology to combining ecological objectives with economical advantages. Dry machining is feasible in numerous manufacturing methods. Many companies have recognized this and converted at least some components of their production to dry machining or minimum quantity lubrication. But it is also indisputable that cutting fluids will remain necessary for some machining tasks. For these processes therefore, we must look for alternative, environmentally friendly media that not only fulfil the functions of traditional cutting fluids but are also not potentially ecotoxic [Krie01, Kloc05].

Chapter 7

Tool Life Behaviour

The term tool life behaviour was introduced to describe material and cutting tool material behaviour during the machining process. The following definition applies: Tool life behaviour is the ability of a working pair (tool and workpiece) to withstand a certain cutting process [DIN6583]. This is influenced by the cutting edge durability of the tool, by the machinability of the workpiece and by tool life conditions (Fig. 7.1). Machinability is the property of a workpiece or material which allows chip removal under specified conditions. Cutting edge durability is the 4x ability of a tool to retain its cutting ability during machining. Cutting ability is the ability of a tool to machine a workpiece or a material under specified conditions [DIN6583]. Tool life behaviour is evaluated by means of the tool life conditions, criteria and parameters. Machinability and cutting edge durability are both functions of the state variables force and temperature. These state variables are influenced in turn by the tool life conditions. The tool life conditions are all the conditions present during the cutting process or operational test. They comprise multiple components [DIN6583]: • • • • •

of the tool, e.g. its form, cutting edge geometry and cutting tool material of the workpiece, e.g. its shape and material of the machine tool, e.g. its static and dynamic stiffness of the cutting process, e.g. its kinematics and cutting edge engagement of the environment, e.g. the type of cutting fluid and thermal marginal conditions

In order to judge the tool life behaviour of the system encompassing the workpiece, tool, clamping, machine tool and coolant, tool life criteria are used which represent limiting values for undesired changes to the tool, workpiece or cutting process caused by machining. Examples of tool life criteria are: • all measurable tool wear values, e.g. width of flank wear land, • all measurable workpiece data, e.g. changes to roughness, • all measurable values of the cutting process, e.g. changes to cutting power, chip temperature and form. To describe the tool life of the system encompassing the workpiece, tool, clamping, machine tool and coolant, i.e. from the beginning of use until the achievement of F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_7,

237

238

7 Tool Life Behaviour

Tool life conditions Tool

- Form - Cutting edge geometry - Cutting tool material

Work piece

- Material - Geometry

Cutting edge durability

(Tool) Tool life criterions Tool wear Resultant force, cutting power Surface roughness Chip form and temperature

Machine

- Dynamic a. static stiffness

Tool life behaviour

Cutting process - Kinematics - Cutting conditions

Environment

- Cutting fluid - Thermal boundaries

Machinability

(Material)

Tool life parameter Tool life Tool life travel path Tool life quantity Tool life volume

Fig. 7.1 Tool life behaviour

the tool life criterion under the influence of tool life conditions, tool life parameters are used. Tool life parameters are times, quantities or paths achieved in chipping under specified conditions until a tool life criterion is reached. These parameters include [DIN6583]: • • • •

the tool life, the tool life travel path, the tool life volume and the tool life quantity.

In order to describe the tool life behaviour of the system of the workpiece, the tool, the clamping, the machine tool and the coolant in a clear way, the tool life conditions, criteria and parameters must always be specified. For example, when describing the tool life behaviour via tool life parameter, the tool life criterion and tool life condition are indicated in the index. If the tool life condition for describing machinability is selected, the tool life parameter and tool life criterion must be taken into account in the index. If the tool life criterion is used for description, the tool life parameter and condition must both be entered in the index. To conclude this section, the two following examples illustrate the relations explained above.

7.2

Machinability

239

(a) description by means of the tool life condition: vcT15; VB 0.2 = 200

m min

(7.1)

Specified parameters: T = 15 min (tool life parameter) and VB = 0.2 mm (tool life criterion) (b) description by means of the tool life criterion: Fcap 2; N 500 = 4000 N

(7.2)

Specified parameters: ap = 2 mm (tool life condition) and N = 500 (tool life parameter) Due to the interaction between machinability and cutting edge durability, both parameters can be used to evaluate the machining system, assuming that a parameter is held constant with respect to the tool life behaviour (constant machinability for evaluating the cutting edge durability of different cutting tool materials given constant tool life conditions or constant cutting edge durability for evaluating the machinability of different materials given constant tool life conditions). Tool wear is highly significant. In contrast to the time-varying state variables of the machining system, such as mechanical stress or temperature, tool wear can be defined relatively easily. The following sections will consider this more closely.

7.1 Determining Tool Life Parameters Tests to determine tool life parameters can basically be executed using either slow testing methods or quick testing methods. Long-term cutting tests are executed for detailed descriptions of the stress of tools on machine tools. Such tests are costly in terms of both time and materials. As an alternative, different quick testing methods were developed in order to evaluate and compare the cutting edge durability and machinability of different materials while minimizing the required time and materials as much as possible.

7.2 Machinability Parameters subjected to state changes during machining can be used as evaluation parameters for judging machinability. One must strictly define, however, whether the object of evaluation is the material (the workpiece) or the cutting tool material. This section will focus on the material, while the cutting tool material will be assumed to be constant. The following parameters can be used to evaluate machinability: • cutting force, • tool life (or tool life travel path, quantity, etc.),

240


• the surface value of the workpiece and • the chip form, etc. It is often sufficient to use a single dominant parameter to evaluate machinability.

7.2.1 Tool Life The tool life Tc of the tool is the most significant parameter for characterizing the machinability of a material. The tool life Tc is the time in min in which a tool performs from its first cut to its becoming unusable due to a specified tool life criterion under specified machining conditions. 7.2.1.1 Temperature Tool Life Rotation Test The temperature tool life rotation test was developed for cutting tool materials with low temperature resistance (tool steels and high speed steels). The test is executed whenever not wear, but rather the influence of cutting temperature is the predominant factor for terminating tool life. The test is executed under constant tool life conditions until the cutting edge becomes unusable due to thermal conditions. This process of succumbing to thermal influences is also referred to as bright braking. Bright braking can be recognized by the formation of bright lines or lines of temper colour on the cut surface or on the machined workpiece surface or by the appearance of surface alterations. Altered chip forms and noises are also indicative of an advanced stage of damage to the cutting edge. 7.2.1.2 Wear Tool Life Rotation Test The wear tool life rotation test is executed for cutting edge materials with a great temperature resistance (cemented carbide, cermet, ceramics, CBN). The test is executed whenever wear instead of cutting temperature is the predominant influence on tool life leading to the unusability of the tool. It is held using a longitudinal round cut with constant tool life conditions. After different cutting times, wear is measured on the flank and rake faces until the previously determined tool life criterion has been reached. It is generally sufficient to determine the width of flank wear land VB, the crater depth CD and the crater mean CM. The measurement results can be represented in a diagram (Fig. 7.2). Using the wear curves in Fig. 7.2, respective value pairs can be formed for the tool life criterion from the cutting speed and cutting time which together form the tool life curve (Fig. 7.3). The curve in Fig. 7.3 can be described approximately by means of a general exponential function. In a double logarithmic system, this function assumes the approximate form of a straight line (Fig. 7.4).

Machinability Width of flank wear land VB/µm

7.2

241

800 600

Tool life criterion

vc3

vc2 >

vc1 >

400 200 0 0

2

4

tc1 tc2

6

8

10

12

14

16

18

tc3 Cutting time t c /min

Fig. 7.2 Wear of uncoated carbide during machining heat treatable steel 60

Fig. 7.3 Tool life curve (heat treatable steel/carbide)

Tool life Tc / min

45

30 tc3 /vc3

15 tc2 /vc2

tc1 /vc1

0 0

200

400

600

800


A straight line can generally be described by the linear equation: y=m · x + b

(7.3)

log Tc = k · log vc + log Cv

(7.4)

The following applies:

And after taking the antilog, the following equation applies: Tc = Cv · vkc

(7.5)

242

7 Tool Life Behaviour 100 80

Material: C53E Feed: f = 0.8 mm Depth of cut: ap = 2.5 mm Cutting tool material: HW-P15 Insert geometry: SNGN120412 Tool life criterion: VB = 0.5 mm

60 40

Tool life Tc / min

20

α 10 8 6 4 Tool life graph

Taylor equation

2

1 50

100

200

300 400 500


Fig. 7.4 Tool life curve in a logarithmical system (heat treatable steel/carbide)

This is referred to as the TAYLOR equation. The parameter Cv (ordinate intercept) indicates the tool life at a cutting speed of vc = 1 m/min (standard tool life) and the parameter CT (abscissa intercept) the cutting speed at Tc = 1 min (standard cutting speed). Thus the increase value k can also be expressed as follows: tan α = k = −

log Cv log CT

(7.6)

The tool life function in Eq. (7.4) is designated as a simple tool life function, since it only takes into account the influence of cutting speed on wear. This simple tool life function was developed by TAYLOR [Tayl07]. It is also known in an expanded form which takes the influence not only of cutting speed, but also of feed and cutting depth. Tc = C · vkc · fzkfz · akpa

(7.7)

Research in to these basic relations essentially goes back to the American TAYLOR [Wall08]. Further research became necessary in his wake to make the tool life equations practically useful in a broad field of application. KRONENBERG

7.2

Machinability

243

was to exert the most influence on this research [Kron27]. Further research by KRONENBERG also showed that the application of similarity mechanics to chipping leads to the relations named in Eqs. (7.5) and (7.7). It must be generally borne in mind that the parameters in these functions are not constants. They can only be assumed to be approximately constant in certain areas.

7.2.2 Resultant Force Knowledge of the magnitude and direction of the resultant force F or its components, the cutting force Fc , the feed force Ff and the passive force Fp , is a basis for • constructing machine tools, i.e. designing frames, drives, tool systems, guideways etc. in line with requirements, • determining cutting conditions in the work preparation phase, • estimating the workpiece accuracy achievable under certain conditions (deformation of workpiece and machine), • determining processes which occur at the locus of chip formation and explaining wear mechanisms. Furthermore, the magnitude of the resultant force represents an evaluative standard for the machinability of a material, since greater forces tend to arise during the machining of materials which do not easily chip. Resultant force is described by its amount and direction. In addition to amount, the force’s effective direction can also have a significant effect on mechanically related changes to the tool or workpiece. Also, one can determine the stress state in the material in front of the workpiece cutting edge from the resultant force (Chap. 3). The following will focus primarily on the influence of materials on the resultant force; geometrical and kinematic influences of the machining process will remain largely excluded from consideration. Exceptions to this will be noted where appropriate. In practical applications, cutting force is often used instead of resultant force as an evaluation parameter. The cutting force is the component of the resultant force in the direction of primary motion. This procedure is reliable when the other components of the resultant force remain negligibly small. The specific resultant force or the specific cutting force may also be used as evaluation parameters (Chap. 3). With respect to machinability tests in which the resultant force is used as an evaluation parameter, the distinction must be made between two different types of cut which respectively cause fundamentally different stress states in the material: • Those that cause a biaxial stress state or one which lies at least in direct proximity to a biaxial stress state. • Those that are far removed from a biaxial stress state.

244


7.2.2.1 Resultant Force Measurements Measurements of resultant force are carried out by means of dynameters, which measure the average mechanical strain on the cutting tool in three directions which are orthogonal to each other [Schl29], preferably in the direction of the axes of the machine tool (Chaps. 3 and 8). SALOMON discovered that there is an approximate exponential relationship between specific force and chip thickness [Salo26, Salo28]. Since this discovery, force measurements have been represented in relation to chip thickness values (Fig. 7.5). It must be borne in mind that extrapolations are not permissible, especially in the region of small chip thicknesses, since in this case at least the same exponential function is no longer valid. By means of a representation in a double logarithmical diagram in which the exponential function follows a straight line, the specification parameters of the straight line may simply be determined by means of the axial sections and the gradient (see Fig. 7.6). The following equations apply: log kz = log kz1.1 + mz · log h

(7.8)

kz = kz1.1 · hmz

(7.9)

mz = tan α

(7.10)

5000

Specific resultant force kz / (N/mm2)

4500 4000 3500 3000 2500 2000

vc1 vc2 vc3

1500 1000

vc1 < vc2 < vc3 500 0 0

200 400 600 800 1000 1200 1400 1600 Undeformed chip thickness h / µm

Fig. 7.5 Gradient of specific resultant force depending on the chip thickness (schematic diagram)

7.2

Machinability

245

Specific resultant force kz / (N/mm²)

10000

α

kz1.1 1000

100 100

1000 Undeformed chip thickness h

10000

Fig. 7.6 Specific resultant force depending on the chip thickness in a double logarithmical diagram

7.2.3 Surface Quality Surface quality can also be used to estimate machinability. The most important factors for this are the elastic and plastic deformations of the material in the area of the minor cutting edge [Djat52]. Low cutting speeds and certain material-tool combinations may lead to the adhesion of material particles on the rake face. This is referred to as the growth of built-up edges (Fig. 7.7). Due to mechanical and thermal stresses, the material which builds up on the rake face is sporadically stripped off und transferred to the workpiece surface. Built-up edges are undesirable. They increase tool wear and lead to a poor surface quality (Fig. 7.7). With increased cutting speeds, this influence becomes increasingly insignificant. The kinematic roughness is yielded by the relative motion between workpiece and tool and by the edge radius. During turning, it is primarily influenced by the form of the cutting edge and the feed. Figure 7.8 compares calculated and measured roughness values given a constant cutting speed without process disturbances caused by built-up edges. BRAMMERTZ’s theory accounts for plastic deformations and also the elastic spring-back of the material effected after the cutting edge reaches a certain area

246


25

Material: C45N Cutting tool material: HW-P10 ap = 3 mm Depth of cut: f = 0.25 mm Feed:

15

Cutting section geometry: γo = 6° αo = 5° λ s = 0° κ r = 70° ε r = 90° rε = 1.2 mm

Built-up edge

Average roughness Rz / µm

20

10

5

0 0

50 100 150 200 Cutting speed vc / (m/min)

250

Fig. 7.7 Influence of cutting speed on surface quality

of the material [Bram61]. A minimum cutting depth must be achieved in order to guarantee chip formation. Otherwise, the material is only deformed elastically by the cutting edge. The yield point can be used to judge the elastic behaviour of the material to be machined. 28

Roughness R /µm

24

rε = 0.25 mm 0.5

1

20 16

2

Rt = 8rf ε

12 8

Measured roughness

4

Fig. 7.8 Comparison of calculated and measured roughness

2

Theoretical roughness (Rt)

0 0

0.1

0.2

0.3

Feed f /mm

0.4

0.5

0.6

7.2

Machinability

247 Trace of the plane of the rake face Aγ

Tool

Trace of the plane of the flank face A α

fmin ≡ hmin rβ

Work piece

assumed working plane

≡ Tool-cutting edge normal plane

Fig. 7.9 Schematic sketch of minimum feed and minimum chip thickness

Since the material is not only stressed on the finished surface, but also on the cut surface, elastic deformation must also be taken into account in this case. This means that a minimum feed fmin must also be achieved as to allow a material separation to take place. This material separation depends on the cutting edge radius rβ and the yield point Re of the material. In theoretical cutting theory, minimum feed is frequently projected to the tool-cutting edge normal plane Pn and referred to as minimum chip thickness hmin (Fig. 7.9). There is no material separation below the minimum chip thickness value and thus no chip formation. As soon as the cutting edge radius assumes higher values than the chip thickness, the influence of the nominal tool orthogonal rake angle becomes insignificant, since the effective tool orthogonal rake angle becomes increasingly negative with an increasing cutting edge radius. The effective tool orthogonal rake angle represents a major influence on minimum chip thickness. If the minimum chip thickness is not reached, material accumulates in front of the cutting edge radius, as a result of which the workpiece material becomes pressed, squeezed and conveyed to the flank face (ploughing effect) [Albr60]. This compromises the achievable surface quality. The literature varies with respect to the minimum chip thickness hmin . As reference data for hmin , König and Klocke cite a factor of two to three of the cutting edge radius or the bevel width. In reference to the turning of steel, Sokolowski cites a dependence of minimum chip thickness on cutting speed, with hmin decreasing with increasing cutting speed [Soko55]. He investigated cutting speeds between 8 and 210 m/min, leading to values for hmin /rn between 0.25 and 1.125. Further notable influences on surface quality are material inhomogeneities and hardening.

248


7.2.4 Chip Form In the machining of different materials, different chip forms are formed under the same tool life conditions. Examples of typical chip forms are shown in Fig. 7.10. Long chip forms make the evacuation of accumulating chips difficult. Flat helical chips tend to migrate outside the engagement length via the flank face, thus causing damage to the tool holder and the cutting edge. Ribbon, snarled and discontinuous chips represent an increased hazard to machine operators. The formation of the different chip forms depends greatly on the friction conditions in the contact area between the chip and the rake face, the tool orthogonal rake angle, the cutting parameters and the material properties. Chip forms can be altered through alloying different chemical elements, such as phosphorous, sulphur and lead, or by means of a targeted heat treatment of the material. Chip breakage is generally favoured by the material’s increasing strength and decreasing toughness. Especially advantageous are chip forms that do not inhibit the machining process. These may not damage the tool system, the machine tool and the surface of the processed component. The spatial requirement for the chips and chip evacuation are also important parameters (chip volume ratio). Unfavourable

1

2

Favourable

3

4

1 Ribbon chip 2 Snarled chip 3 Flat helical chip 4 Angular helical chip 5 Helical chip

5

Favourable

Good

6

7

8

6 7 8 9 10

9

10

Helical chip segment Cylindrical helical chip Spiral chip Spiral chip segment Discontinous chip

Fig. 7.10 Conventional chip forms

7.2.5 Cutting Speed The following will evaluate the influence of cutting speed on machinability and discuss this influence via and analysis of specific cutting force. Figure 7.11 shows a typical curve for the specific cutting force when machining non-heat-treated steels. Adhesion and the growth of built-up edges occur with lower cutting speeds. The adhesion tendency decreases with increasing cutting speed. A temperature range is

7.3

The Machinability of Steel Materials

249

2000

1000

500

Cutting speed range of blue brittleness

1500

Propensity to adhesion

Specific resultant force kz/( N/mm²)

2500

Thermal softening

0 0

25

50

75 100 125 150 175 200 225 250

Cutting speed vc/ (m/min)

Fig. 7.11 Material properties depending on the specific resultant force

reached in which the specific resultant force again rises until a maximum value is achieved. This is the area in which blue brittleness occurs. Blue brittleness refers to a material condition in which the dislocation mobility is strongly limited by the interaction with nitrogen present in the microstructure. As a result, deformability decreases, thus increasing the specific resultant force. If the cutting speed is further increased, the temperatures in the shear zone increase sharply and the material is weakened, which causes the resultant force to be lowered again. The following will describe materials frequently processed by chipping methods and their machinability.

7.3 The Machinability of Steel Materials Steels can be classified according to their alloying elements, their metallographic constituents and their mechanical properties. Such a classification aids both in selecting a material with respect to the properties required for their future function and in determining machining conditions. Depending on their alloy content, steel materials are divided into the following groups: • unalloyed steels, • low-alloyed steels (alloy content < 5%) and • high-alloyed steels (alloy content ≥ 5%). In the case of unalloyed steels, we must furthermore differentiate between those steel materials that are not to be heat-treated (common construction steel) and those that are (grade and special steel). Steels referred to as common construction steels

250


(e.g. S235JR, S355J2G3) those whose mechanical properties exhibit minimum values. Construction steels are used in practice when no special requirements are made with respect to their structure.

7.3.1 Metallographic Constituents The machinability of steels essentially depends on their respective crystalline structure. The crystalline structure of steel is mainly composed of the following components: • • • • • •

ferrite, cementite, perlite, austenite, bainite and martensite

Depending on carbon content, alloying element content and heat treatment, one or more of these metallographic constituents predominate, whose mechanical properties (Table 7.1) affect the machinability of the given steel. Table 7.1 Mechanical properties of metallographic constituents, acc. to VIEREGGE [Vier70] Hardness/HV 10 Rm /(N/mm2 ) Rp0,2 /(N/mm2 ) Z/% Ferrite Cementite Perlite Austenite Bainite Martensite

80–90 >1100 210 180 300–600 900

200–300 – 700 530–750 800–1100 1380–3000

90–170 – 300–500 300–400 – –

70–80 – 48 50 – –

Ferrite (a-ferrite, cubic body-centred, maximal solubility for carbon: 0.02%) is characterized by relatively low strength and hardness and by a high deformability. Ferrite complicates the chipping process through the following: • its strong tendency to adhere, which favours material smearing on the tool and the formation of built-up edges and built-up edge fragments. • its tendency to form undesirable ribbon and snarled chips due to its high deformability • poor surface qualities and increased burr formation on the workpieces Cementite (iron carbide, Fe3 C) is hard and brittle and thus cannot be machined. Depending on carbon content and cooling speed, cementite can occur either by itself or as a structural constituent of perlite or bainite. Perlite is a eutectoid phase mixture composed of ferrite and cementite. The eutectic point lies at 723◦ C and 0.83% carbon. Perlite appears in iron materials given a carbon content between 0.02 and 6.67%. Given a carbon content of 2.06%, it is the

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251

only metallographic constituent; above 2.06%, it is a constituent of ledeburite. For the most part, lamellar, linear cementite appears in perlite. After to a corresponding heat treatment (soft annealing) however, globular (spheroidal) cementite can also be formed. Heat treatable steels with a carbon content < 0.8% are referred to as sub-eutectoid. During machining, the low deformability and greater hardness of perlite causes, on the one hand, • significant abrasive wear and • high resultant forces. On the other hand, • • • •

Perlite reduces the adhesion tendency and the formation of built-up edges, promotes the formation of favourable chip forms, causes less burr formation on the workpiece and improves the surface quality.

Austenite refers to the γ-mixed crystal of iron. It has a face-centred cubic structure. The maximum solubility for carbon is 2.06%. The structure exhibits only minimal hardness, although its strength can be increased by means of cold-forming. Austenite is the main constituent of many non-rusting steels and is not ferromagnetic. In unalloyed and low-alloy steels below approximately 723◦ C, austenite is transformed into perlite and, depending on the carbon content, into ferrite or cementite. Thus austenite is only present at room temperature in alloys. Examples of austenite builders are nickel (Ni), manganese (Mn) and nitrogen (N). Essential properties of austenitic steel materials relevant to machining are: • their high deformability and toughness: The high ductility of this material is based on the good plastic deformability of its face-centred cubic crystal lattice, which has four slip planes each with three slip directions and thus 12 slip systems in total. The result is a strong tendency of the material to form built-up edges, gluing-points and built-up edge fragments and to form unfavourable ribbon and snarled chips. • their strong tendency to adhere to the cutting edge material: face-centred cubic materials tend to adhere much more than hexagonal or space-centred cubic metals, which means that built-up edges, gluing-points and built-up edge fragments occur more frequently when machining these steel materials. • their tendency towards strain hardening: in the area of the workpiece surface being created, the deformation of the material caused in chip formation causes a hardening of the material forming the surface. This results in additional stress to the tool cutting edge, especially in subsequent cuts. • their relatively low heat conductivity, which is 1/3 lower than in unalloyed steels. It impairs heat removal through the chip and increases the thermal stress of the cutting edge.

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Bainite forms in the temperature range between those of perlite and martensite: iron diffusion is no longer possible, and carbon diffusion is already considerably hindered. There are two main forms of bainite: • acicular bainite (with continuous cooling and isothermal transformation) • granular bainite (only with continuous cooling) Independently of the form, bainite consists of carbon-supersaturated ferrite, with the carbon partially precipitated as carbides (e.g. Fe3 C) whose size (from rough to very fine) is determined by the conversion temperature. Among the acicular bainite forms, one distinguishes according to the transformation temperature between lower bainite (strong similarity to martensite) and upper bainite (strong similarity to perlite). Martensite forms when a steel material with a carbon content of > 0.2% is rapidly cooled from the austenite temperature range to a temperature below the martensite starting temperature. Due to the rapid cooling, the carbon dissolved in austenite is forced to remain dissolved in the mixed crystal. By means of a diffusionless transformation, a tetragonally deformed, body-centred martensite lattice forms from the face-centred cubic austenite lattice. Martensite has a fine-acicular, very hard and brittle microstructure which is difficult to machine. The cutting tools used are subject to • an increased abrasive wear and • high mechanical and thermal stresses

7.3.2 Carbon Content The metallographic constituents of steels can be seen in the iron-carbon phase diagram (Fig. 7.12). Steels with C-content < 0.8% precipitate ferrite when cooling from the austenite zone. The remaining austenite disintegrates into perlite below 723◦ C. With a carbon content of 0.8%, only perlite is formed as a eutectoid mixture of ferrite and cementite. In steels with a C-content > 0.8%, perlite and secondary cementite are formed. The secondary cementite is precipitated from the austenite predominately at the grain boundaries. The machinability of steels with a carbon content < 0.25% is essentially characterized by the properties of the free ferrite. Because of the high deformability of the material, the resulting cut surface roughness is high. Since ferrite has an intrinsically large adhesion tendency, built-up edges form with low cutting speeds. Tool wear and cutting temperature only increase slowly with a rising cutting speed. Because of ferrite’s low strength, only tools with the greatest possible positive tool orthogonal rake angle (e.g. with turning: γ o > 6◦ ) should be used. In order to decrease the adhesion tendency and to improve the surface quality, oils are generally used as cutting fluids, with their lubrication properties being more significant than their cooling effect. Steels with a carbon content of < 0.25% pose particular problems for grooving and parting off, as well as for boring, reaming and thread die cutting.

The Machinability of Steel Materials Steel A

H

B

Melt

δ-Mc + γ -Mc

1147

Melt + Austenite

Austenite (γ -Mc)

911

G

769 723

M P

α -Mc

α-Mc + γ -Mc O S

Ferrite + Perlite Q

500 0

Perlite

Temperature T/°C

Melt + δ-Mc I

δ -Mc

Cast iron

0.8

E

Austenite + Secondary cementite

Melt + Primary cementite C

Austenite + Secondary cementite + Ledeburite

Perlite + Secondary cementite + Ledeburite

Perlite + Secondary Cementite

F

Ledeburite + Primary cementite

Primary cementite + Ledeburite

Cementite (Fe 3 C)

1600 1536

253

Ledeburite

7.3

K

L

2.06

4.3

6.67

Carbon concentration / mass-%

Fig. 7.12 Section of the iron-carbon phase diagram (metastable system)

Due to high deformability and relatively low cutting speeds, the achievable surface qualities are markedly inferior. Moreover, burr formation occurs to a greater extent. The amount of perlite in the structure increases with increasing carbon content (0.25–0.4%). In this way, the special property of this metallographic constituent is granted a stronger influence on the machinability of the material. The strength of the structure increases, its deformability decreases. This results in the following: • a lower adhesion tendency and thus a shift of built-up edges to lower cutting speeds, • increasing tool wear and higher temperatures in the tool contact zones due to the higher material strength, • an increase in abrasive wear, • an improvement of surface quality and chip form. Machinability can be improved by means of coarse grain annealing given a low C-content and by means of normalizing given a C-content above 0.35%. Tools with a positive orthogonal tool rake angle should be used. Cold-forming has a positive effect on machinability, especially with respect to chip formation. This is important to the extent that these materials are frequently deformed through cold extrusion and then finished through cutting.

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A further increase of carbon content (0.4–0.8%) causes a further decrease of the amount of ferrite with a corresponding perlite increase until only perlite remains at 0.83% carbon. Due to increased strengths, high rake face temperatures already arise at low cutting speeds. At the same time, the increasing mechanical stress of the rake face causes increased wear, also in the form of crater wear. Problems related to chip form are more rare. The chip form improves with increasing carbon content, with wear also increasing. Steels with a carbon content of 0.4–0.8% are generally regarded as having good machinability only with respect to surface quality and chip form. After slow air cooling, the structure of supereutectoid heat-treated steels (C > 0.8%) consists of secondary cementite and perlite. Perlite formation initiates directly from the austenite grain boundaries. Secondary cementite precipitates at the grain boundaries given a carbon content significantly exceeding 0.8%. The free cementite forms shells around the austenite/perlite grains [Schu04]. Such shells cause very strong wear in cutting processes. In addition to the strongly abrasive effect of the hard and brittle metallographic constituents, the high pressures and temperatures which arise cause an extra stress for the cutting edge. Strong crater and flank face wear is to be observed even at relatively low cutting speeds. The cutting edge parts must have a stable design. For turning, tools with positive orthogonal tool angles γ o up to 6◦ and mildly negative tilt angles λs up to –4◦ should be used.

7.3.3 Alloying Elements Alloying and trace elements can influence the machinability of steel by changing the composition or by forming lubricating or abrasive inclusions. The following will discuss the influence of some important alloying elements on the machinability of steel. 7.3.3.1 Manganese Manganese improves temperability and increases the strength of steel (approx. 100 N/mm2 per 1% alloying element). Because of its high affinity to sulphur, manganese forms sulphides with sulphur. Manganese contents up to 1.5% facilitate machinability in steels with low amounts of carbon due to good chip formation. In the case of steels with larger amounts of carbon, however, machinability is negatively affected by increased tool wear. 7.3.3.2 Chrome, Molybdenum, Tungsten Chrome and molybdenum improve temperability, thereby influencing the machinability of case-hardened and heat-treated steels in terms of structure and strength. In the case of steels with a greater carbon or alloy content, these elements and also tungsten form hard special and mixed carbides which impair machinability.

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255

7.3.3.3 Nickel Nickel is among the group of elements which expand the γ-phase zone in iron alloys. By adding nickel, the strength of steel materials increases. Nickel increases toughness, especially at low temperatures. This generally leads, especially in the case of austenitic nickel steels (with larger amounts of nickel), to an unfavourable machinability. 7.3.3.4 Silicon Silicon increases the strength of ferrite constituents in steel. With oxygen, it forms, in the absence of stronger deoxidation agents like aluminium, hard Si-oxide (silicate) inclusions. This results in increased tool wear. 7.3.3.5 Phosphorous Alloying phosphorus, which is carried out only in some free-cutting steels, leads to segregations in the steel which cannot even be removed with subsequent heat treatment and heat deformations and to an embrittlement of the ferrite. This allows for the formation of short-breaking chips. Up to an amount of 0.1%, phosphorous has a positive effect on machinability. While higher amounts of phosphorous lead to an improvement of the surface quality, they also cause increased tool wear. 7.3.3.6 Titanium, Vanadium Titanium and vanadium, even in small amounts, can increase strength considerably due to the extremely dispersed carbide and carbon nitride precipitations. They lead to a high grain refinement, which has a negative effect on machinability with respect to mechanical stress and chip form. 7.3.3.7 Sulphur Sulphur is only slightly soluble in iron, but it forms, depending on the alloying components of the steel, various stable sulphides. Iron sulphides (FeS) are undesirable, as they exhibit a low melting point and deposit primarily at the grain boundaries. This leads to the unwanted “red brittleness” of steel. What are desirable on the other hand are manganese sulphides (MnS), which have a much higher melting point. The positive effects of MnS on machinability are short-breaking chips, improved surface quality and a smaller tendency towards forming built-up edges. With an increased inclusion length, MnS exerts a negative influence on mechanical properties like strength, strain, area reduction and the impact value, especially when it is included transversely to the strain direction. In practice, however, special alloying additives (e.g. tellurium, selenium) can effectively help to counteract the deformability-related extension of MnS.

256


7.3.3.8 Lead Lead is not soluble in iron; it is present in the form of submicroscopic inclusions. Because of the low melting point, a protective lead film forms between the tool and the material. This film reduces tool wear. The mechanical stress of the tool can be lowered by up to 50%, with the chips becoming short breaking. This effect is exploited especially in machining steels (see Sect. 7.4.1). 7.3.3.9 Non-metallic Inclusions The elements added to the steels for deoxidation, i.e. aluminium, silicon, manganese or calcium, bind the oxygen released by steel solidification. The hard, non-deformable inclusions then found in the steel, e.g. as aluminium oxide and silicon oxide, diminish machinability, especially when the oxides exist in the steel in larger amounts or in linear form [Wink83]. However, by choosing a suitable deoxidatizing agent, the machinability of steel can also be positively influenced. For example, under certain machining conditions, wear-inhibiting oxidic and sulphidic layers may form after deoxidation with calcium-silicon or ferro-silicon [Köni65, Opit67]. One measure for improving the machinability of steels deoxidized with aluminium is calcium treatment. In this process, calcium is added to the steel melt by means of secondary metallurgy. This converts the sharp-edged aluminium oxides present in steel in conventional melting which act abrasively during cutting into globular and, in machining conditions, plastifyable calcium aluminates. Analogously to manganese sulphides among machining steels, these calcium aluminates are able under certain cutting conditions to form friction-reducing and/or wear-inhibiting coatings in the contact zones of the tool cutting edge [Köni65, Töns89, Kloc98a, Zink99].

7.3.4 Types of Heat Treatment When executed in a targeted manner, heat treatments can influence microstructures with respect to quantity, form and the configuration of their constituents. In this way, mechanical properties and thus machinability can be customized to the given requirements. According to DIN EN 10052, heat treatment is defined as follows: Heat treatment is a sequence of heat treatment steps during which a workpiece is in whole or in part subjected to time/temperature sequences in order to cause a change of its properties or its structure. In certain circumstances, the chemical composition of the material may be changed during the treatment (thermochemical treatment). Three basic categories of heat treatment may be distinguished: • establishing an even structure in the entire cross-section, which to a great extent is in a state of thermodynamic equilibrium (e.g. soft annealing structures, structures consisting of ferrite or austenite) or in thermodynamic disequilibrium (e.g. perlite, bainite, martensite),

7.3


257

1300 Austenite (γ –Mc) 1200 (E) Homogenisation

1100

ide

t ne

arb

Temperature T / °C

1000

sit

G

Coarse grain annealing

Har

900

den

α−Mc + γ−Mc P

800 700

ion

c of

ing

Ac3 723 °C

co

De +N

orm

o mp

alisi

O Ac1

Ac m Austenite + Secondary cementite

ng S

Soft annealing

(K)

Recrystallisation annealing

α−Mc 600

Low-stress annealing Recrystallisation annealing

500 (Q)

400 0

0.4

0.8

1.2

1.6

2.0

Carbon concentration / mass-%

Fig. 7.13 Partial iron-carbon diagram with heat treatment range data

• establishing a hardening structure that is restricted to smaller areas of the cross-section at unaltered chemical composition (in particular: surface layer tempering), • establishing structure types that are starkly differentiated over the cross-section, especially in the rim zone, as a result of a change of the chemical composition (nitriding, carbonitriding, carburizing, carburizing hardening). There are different heat treatment methods with a broad field of application which, depending on the chemical composition of the steel, can affect machinability in a targeted way with respect to, for example, chip form and tool wear. The temperature ranges for the individual types of heat treatment can be found in Fig. 7.13.

7.3.4.1 Homogenization Homogenization is defined according to DIN EN 10052 as follows: annealing at a high temperature with a hold sufficiently long enough to minimize local differences in chemical composition due to diffusion-related segregations.

258


7.3.4.2 Coarse Grain Annealing Coarse grain annealing is defined according to DIN EN 10052 as follows: annealing at a temperature usually considerably above Ac3 with a hold sufficiently long enough to achieve coarse grain. Coarse grain annealing followed by isothermal transformation is utilized for subeutectoid steels with a C-content of 0.3–0.4% (ferritic-perlitic steel) in order to produce a coarse-grain structure with a ferrite network which is as closed as possible in which either perlite or bainite is enclosed [Hors85, Schu04]. Tool wear when machining such a structure is relatively low, and chip formation generally good. High surface qualities remain achievable. The application of coarse grain annealing to improve machinability is, however, limited by interference by strength properties and due to financial considerations. 7.3.4.3 Normalizing Normalizing is defined according to DIN EN 10052 as follows: a heat treatment comprising austenitizing and subsequent cooling in still air. Through normalizing (+N), a nearly even, fine-grain structure is achieved whose machinability is, depending on the carbon-content, determined by the predominant structural constituent, i.e. either by ferrite (low wear, poor chip formation) or perlite (increased wear, improved chip formation) [Hors85]. A α/γ transformation takes place in the process. Subeutectoid steels are heated to temperatures above Ac3 . Supereutectoid steels are heated to temperatures 30–50◦ C above Ac1 or, if the carbide network is to be dissolved in structures with higher carbon content, to temperatures exceeding Ac3 . In general, this heat treatment should be suited to achieving an even, fine-grain perlitic-cementitic structure. Stronger carbide bands cannot be fully dissolved, which means that normalized supereutectoid steels can still cause relatively high tool wear. However, they allow a high surface quality of the finished surfaces. 7.3.4.4 Soft Annealing Soft annealing is defined according to DIN EN 10052 as follows: a heat treatment for reducing the hardness of a material to a specified value. Soft annealing (+A) is applied in order to take away the high hardness and low deformability from structures with lamellar perlite or lamellar perlite and cementite. By means of annealing at temperatures just below the PSK line (Fig. 7.13) – or, if necessary, oscillating at the PSK line – and subsequent slow cooling, a soft and low-stress state can be created as the lamellar perlite and strands of cementite are caused to disintegrate. The goal is to achieve a perlite structure which is as granular as possible consisting of ferrite with globular cementite. Such a structure is soft and easily deformable. The machinability of such a structure becomes more favourable with respect to wear effects on the tool, while chip formation worsens to the extent that ferrite predominates in the structure.

7.3


259

7.3.4.5 Annealing to Specific Properties In practice, other types of treatment are required in addition to classic soft annealing which target specific properties or structures. A special type of soft annealing is annealing to spherical carbides (+AC) or spherical cementite (also referred to as GKZ annealing). The goal of this type of annealing is to mould the cementite completely as a sphere. Temperatures are held in the region of the PSK line for a longer period of time, if necessary oscillating at this temperature. Since, when the cementite is fully moulded, its machinability approximates that of pure ferrite, it may worsen [Opit64]. Given steels with a carbon content of 0.10–0.35%, materials with Widmannstätten structure may be created by means of high austenitizing temperatures, long holding times and rapid cooling. The result is an acicular ferrite with extraordinarily finely distributed lamellar cementite. Such a structure is characterized by good chip formation and chip form, though it exhibits poor usage properties [Schu04]. Case-hardened steels are heat treated to a ferrite/perlite microstructure (+FP). In this state, they can achieve a similarly good machinability as with machining steels with low carbon content, and this with respect both to low tool wear and to good chip formation. A further heat treatment for improving machinability used primarily with case-hardened and heat-treatable steels is annealing to a specific tensile strength (+TH). For energy-saving reasons, targeted heat treatments are executed directly using the heat of the forge, e.g. controlled cooling from the heat of the forge (BY annealing) [Fasc80]. Machinability investigations [Wink83] have shown that heat-treatable steels cooled from the heat of the forge (e.g. C45E+BY) may exhibit a more favourable wear behaviour than the same materials in a heat-treated or normalized state. Differences with respect to chip formation could not be ascertained. The reason lies in the relatively coarse-grain structure of the steels used and in the fact that the ferrite network also encloses the perlite grains during the shear process. 7.3.4.6 Recrystallization Annealing Recrystallization annealing is defined according to DIN EN 10052 as follows: a heat treatment with the goal of achieving the formation of new grain in a cold-formed workpiece by means of nucleation and growth without a phase change. Recrystallization annealing refers to annealing following cold forming at a temperature below Ac1 , which for steel is usually between 500 and 700◦ C, without causing a α-γ-transformation of the crystal lattice. This is usually applied between the individual shaping stages, e.g. when cold rolling or cold drawing metal sheets and wires [Schu04, Meta06]. Strain hardening becomes noticeable in the structure through the occurrence of displacements, glide lines and the splitting of brittle crystal types (cementite). Recrystallization annealing prevents changes to properties associated with this, such as increased hardness and strength and decreased strain and toughness. If the material reaches its forming limit in the process of cold forming, a recrystallization

260


must be executed to form new grains. The composition of the microstructure is not newly formed during recrystallization; only the grains are newly formed. The greater the strain is, the greater the tendency towards the formation of new grains. At high strain levels, the size of the newly formed grains can lie below that of the original grains [Schu04]. 7.3.4.7 Low-Stress Annealing Low-stress annealing is defined according to DIN EN 10052 as follows: a heat treatment comprising heating and holding the material at a sufficiently high temperature and a subsequent cooling appropriate for eliminating internal stresses as much as possible without essentially changing the structure. The usual temperatures for low-stress annealing steel workpieces lie between 550 and 650◦ C. A microstructural transformation does not take place. Low-stress annealing is mainly applied to workpieces which exhibit high internal stresses either as a result of irregular cooling following casting, welding, forging or another thermal process or after strong mechanical processing through milling, turning, planing, deep drawing, etc. Low-stress annealing serves to reduce these stresses. This prevents the release of existing internal stresses during the further processing of such workpieces and the formation of geometrical deviations resulting from warpage. The strain hardening in the deformation zones induced by cold-forming, however, is reduced by means of recrystallization annealing. 7.3.4.8 Hardening Hardening is defined according to DIN EN 10052 as follows: a heat treatment comprising austenitizing and cooling under conditions conducive to an increase in hardness caused by the more or less complete transformation of austenite into martensite and, if applicable, bainite. When hardening steel (+Q), the precipitation of carbon from the γ-mixed crystal which happens at a normal cooling speed counteracted through a high cooling speed. At supercritical cooling speeds, martensite forms after falling short of the Ms-temperature (Ms = martensite starting) [Schu04]. At cooling speeds lower than the critical cooling speed, the conversion processes proceed in the intermediate stage and in the perlite stage [MPI61, MPI72, MPI73]. Transformation in the intermediate stage is basically characterized by that fact that only the carbon can diffuse. Processes of continuous and isothermal transformation for the heat-treatable steel C45E are shown in Figs. 7.14 and 7.15. These structure types are not as easy to machine because of their greater strength. Chip formation can be considered good. The cutting edge must have a stable design. 7.3.4.9 Heat Treatment Heat treatment is defined according to DIN EN 10052 as follows: Hardening and tempering at higher temperatures in order to achieve the desired combination of mechanical properties, especially high toughness and ductility.

7.3


261

1000 900

Temperature T / °C

800 700

50

30 20

40

50

70

80

F

600

P

179

500 400

B

A Ms

2 5

300 200

M 210

100

548 224

274 0 0.1

Ac3 Ac1

60

1

102

10 1

A F P B M

Austenite Ferrite Perlite Bainite Martensite Hardness/HV 20 Part of the structure/ % 104

103 10

10

2

s

105

min 103

Time Austenitisation temperature 880 °C (holding time 3 min) heated up in 2 min

Specimen dimension / mm: Ø 3 x 15 Material: C45E

Fig. 7.14 Time-temperature-transformation diagram for continuous cooling of C45E steel, acc. to MPI

The strength values of steel can also be increased through quenching and tempering (+QT). When the material is tempered, the martensite formed during hardening is broken down again in a targeted way through re-heating. At low tempering temperatures, carbon precipitates in a finely distributed form, while at higher temperatures coarser cementite grains develop [Schu04]. The machinability of the tempered structure increases with increasing martensite decay. Several possibilities for a targeted influence on structure by various heat treatments are shown in Fig. 7.15 using the heat-treatable steel C45E as an example. The allocation of the partial images to the heat treatments is as follows (Fig. 7.16): 1. Coarse grain annealing. Metallographic constituents: coarse-grain perlite with lamellar cementite, a ferrite network between the grains (white). 2. Normalizing. Metallographic constituents: perlite with lamellar cementite, ferrite. These are the same constituents as in coarse-grain annealing, but the structure is finer-grain and more homogeneous. 3. Quenching and tempering. Structure: tempered martensite. 4. Soft-annealing. Metallographic constituents: ferrite (white) with globularly shaped cementite.

262


Start of ferrite formation

Start of perlite formation

800

Ac3 Ac1

Temperature T / °C

700 A 600 500

25

F

28 29 30 33

75

15 P 85

A F P B M

Austenite Ferrite Perlite End of Bainite transformation Martensite Hardness/HRC 20 Part of the structure/ %

33 42

B 400 Ms M 300

50 50 % 90 %

200

Specimen dimension Outside diameter 4 mm Insidediameter 3.2 mm Length 30 mm

100

1

102

10 1

103 10

104 2

10 Time

105 10

3

s min 10

106 4

Material: C45E Austenitisation temperature 880 °C, (Holding time 5 min) heated up in 1 min

Fig. 7.15 Time-temperature-transformation diagram for isotherm conversion of C45E steel, acc. to MPI

Fig. 7.16 Crystalline structure for different heat treatments (C45E)

7.4

Machinability of Various Steel Materials

263

7.4 Machinability of Various Steel Materials In addition to the categorization of steels according to their alloy content, they are also categorized in practice with a view to their uses and applications. The categories are: • • • • • •

free cutting steels, case-hardened steels, heat-treatable steels, nitrided steels, tool steels and rust- and acid-proof, heat-proof and highly heat resisting steels.

7.4.1 Machining Steels The free cutting steels are materials which are distinguished by especially good machinability. Characteristics in this respect are a favourable breakage, clean workpiece surfaces and low tool wear. They can be machined without difficulty on automatic lathes and multi-spindle lathes and are thus highly suited to serial and mass production. The high cutting speeds can be used, these being often limited by the workpiece diameter and the respective machine tools (multi-spindle lathes). Depending on their chemical composition and field of application, free cutting steels are subjected to heat treatments like carburizing hardening or quenching and tempering. These favourable properties with respect to machinability are given to machining steels primarily by adding different alloying elements. The main alloying elements, sulphur, lead and phosphorous, as well as tellurium, bismuth and antimony reduce the strength of the material in the shear zone. The reduced strength results in shortbreaking chips with a low level of compression. Machining steels are predominately cut with coated cemented carbide tools and moulding tools consisting of high speed steel. The development of wear on these tools is relatively slow. The low-carbon machining steels in widespread use (e.g. 11SMn30, 11SMnPb30, 9S20) are characterized by high ferrite and low perlite content. The result of this is a low level of abrasive tool wear. Even at low cutting speeds, adhesion in the form of built-up edges plays an important role with respect to tool wear. When machining low-carbon machining steels at low cutting speeds (vc < 100 m/min), friction-reducing layers (MnS, Pb) form which contribute to reducing tool wear [Zink99]. The machinability of free cutting steels is relatively low for machining purposes. Reasons for this are the large amount of ferrite in the microstructure as well as the strength-reducing properties of the main alloying elements. The achievable surface qualities are considerably influenced by the adhesion tendency of the material caused by the formation of built-up edges. MnS layers, Pb and P reduce this adhesion tendency, positively influencing the surface quality of the workpieces.

264


As a function of the steel production process, phosphorous causes a formation of segregations (demixing) in the steel because of its relatively small diffusion coefficient. This undesirable property can only be partially removed through subsequent heat treatment. High-temperature diffusion annealing can only compensate microsegregations (crystallization segregations, concentration differences). In contrast, segregations at the macro-scale (macrosegregations) cannot be removed extensively by means of a subsequent heat treatment. This is because of the large distance from the middle to the edge of the steel ingot. Sufficiently eliminating macrosegregations would require annealing times of such a length as would lead to a significant coarsening of the grains. Smaller macrosegregations can be achieved by deoxidizing the steel prior to casting. Nitrogen and carbon precipitations are promoted by phosphorous and induce an embrittlement of the α-mixed crystal (ferrite embrittlement). These precipitations increase with increasing temperature, which means that tempering brittleness occurs to a greater degree and the impact value is already reduced from 100◦ C. The degradation of strength properties caused by phosphorous are responsible for the embrittlement of machining steels at the average shear level temperature of approximately 200–400◦ C. The result is a more favourable chip formation with short-breaking chips when machining. The adhesion tendency in the contact zone continues to be reduced and the surface quality is positively influenced. Machining steels contain up to 0.1% phosphorous. The positive effect of MnS on machinability is explained by the fact that both the internal friction of the material in the shear zone and the friction in the contact zone are reduced. Recent investigations prove the inclusion of MnS as the basis for glide line formation in the shear zone. Since sulphur reduces tool wear but hardly influences tempering brittleness, it is favoured over phosphorous as alloying element. MnS causes the chips to be short, improves the surface quality of the workpiece and reduces the tendency toward the development of built-up edges. The form of sulfidic inclusions in machining steel is determined by the quantity of oxygen released upon solidification. Three different types of sulphides are found in steel: • Type 1: This type of sulphide forms as a liquid phase at an oxygen content of > 0.02% corresponding to the quasiternary system Fe-MnO-MnS in iron-rich melts. After solidification, this sulphide appears in the form of regularly distributed, globular particles or irregularly rounded particles. These brittle type 1 sulphides are clearly separated from each other in the steel, enclosed in cells. • Type 2: In iron melts with an oxygen content of < 0.01%, this form of sulphide precipitates from melts rich in manganese sulphide at the primary grain boundaries as a MnS phase similar to an eutectic point. Type 2 sulphides are not present in steel in enclosed cells, but rather grow as if radiating from individual centres. In the process, furcations form or, in some places, the sulphides growing out of the funnel-shaped cells fuse together. • Type 3: This type of sulphide crystallizes from Fe melts with a reduced melting point. An essential requirement for the formation of these sulphides is a carbon

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and silicon content between 0.1 and 0.4% and an aluminium content between 0.05 and 0.3%. Type 3 sulphides precipitate in interdendritic regions. However, like those of type 1, they are evenly distributed. Their composition and form correspond to that of the angular, face-centred cubic α-MnS. With respect to machinability, type 1 sulphides are considered to be the most favourable. As a result, different metallurgical measures are taken to create this sulphide type for free cutting steels. The formation of round manganese sulphides is favoured, for example, by alloying tellurium. The machinability of machining steel improves in an essential way with an increasing quantity of larger manganese sulphides. Such sulphides prevent the pressure welding of the ferrite grains by forming scale-like layers and create a protective zone between the chip and the tool [Köni66]. Machinability is improved by adding up to 0.35% lead. Lead does not dissolve in α-Fe and penetrates the steel microstructure in the form of sub-microscopic inclusions. The strength properties and toughness of steels are negatively influenced, especially in the region of 250–400◦ C. Lead liquefies at relatively low temperatures (TS = 326◦ C). When machining steels containing lead, a thin lead film may wet the contact surfaces between the tool and the workpiece. This reduces the tendency towards pressure welding and facilitates the shearing-off process. At cutting speeds exceeding 100 m/min or with large feeds, the lead film becomes ineffective and wear accelerates. The specific cutting forces fall by up to 50% and the chips become short-breaking. Adding lead (about 0.25%) to machining steels can increase tool life by approximately 50–70%. The effect of lead on tool wear depends on the cutting speed. Given a lead-content increase of 0–0.29% and cutting speeds below 100 m/min, there is a reduction of flank face wear on HSS tools. In the cutting speed region above 100 m/min, an increased lead content has a negative effect on the development of flank face wear. With respect to crater wear, the effect of lead is not affected by the cutting speed. As opposed to sulphur, lead can be alloyed with almost all steels, but these steels may only be stressed in a temperature range below 200◦ C. Typically, about 0.15–0.30% Pb is added to machining steels. Substituting lead as an alloying element in machining steels gains in importance with a view to its harmfulness to health [Wink83]. Similarly to lead, tellurium, bismuth and antimony also help to achieve improved chip breakage and bring about a lubricating effect in the contact zone which reduces tool wear. Investigations have shown that adding bismuth considerably improves the machinability of lead-free machining steels (e.g. 16SMn30) with respect to their chip formation, surface roughness and forces arising during machining. In comparison with alloys containing lead, however, the onset of abrasive wear on uncoated cemented carbide tools is considerably more aggressive when turning. This effect is not to be observed with coated cemented carbide tools. Machining steels in which lead has been replaced by tin exhibit a significantly reduced tendency towards chip breakage and are thus only recommendable for methods with interrupted cutting, such as milling [Esse06]. Free cutting steels in common use include 9SMn28, 9SMnPb28, 35S20, 45S20.

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7.4.2 Case-Hardened Steels The case-hardened steels include unalloyed machining steels, grade and special steels as well as alloyed special steels. Common to all of these is a relatively low carbon content (C < 0.2%). Case-hardened steels are predominately utilized in the manufacture of wear-stressed and variably stressed parts like cogwheels, gear shafts, joints, connectors etc. Case-hardened steels are almost exclusively processed by chip removal prior to case-hardening. Since the microstructure of these steels contains mostly ferrite and only a small amount of perlite, the onset of wear on the tool is low. Beyond this, cutting speeds should be kept above 200 m/min in order to avoid the growth of built-up edges. The machining of case-hardened steels is mostly achieved with coated cemented carbide tools belonging to application group P (e.g. HC-P10) or with cermets in order to withstand the thermal stress caused when machining at high cutting speeds. Because of the high ferrite content and low perlite content in their microstructure, the machinability of case-hardened steels in a soft state is low. The adhesion tendency of case-hardened steels leads to the formation of builtup edges and to bad surface qualities when procedures with low cutting speeds are used, such as tapping, boring, broaching and shaping. In order to improve machinability with respect to surface roughness, case-hardened steels are heat treated to a certain ferrite/perlite structure (BG) or to a certain strength (BF), depending on their alloying elements. Coarse-grain annealing is often applied to alloyed case-hardened steels in order to reduce difficulties caused by their adhesion tendency. At the same time, this reduces the strong tendency of the structure towards linearity, which is highly disadvantageous for machining purposes, especially for reaming and broaching, since individual lines may be cut out. This linearity can be partially suppressed by means of a rapid cooling during heat treatment, but it appears again when reheating over the transformation point. The best possible surface quality is achieved furthermore by applying suitable cutting fluids, by changing the tool geometry (positive tool orthogonal rake angle) and through reducing the feed [Opit64, Peke74]. The applicable cutting conditions are little influenced by the heat treatment used – provided that machining is executed with cemented carbide tools and the tensile strength of the material is under 650 N/mm2 [Well72]. Tools made of HSS, on the other hand, react more sensitively to differences in strength, so that different cutting speeds are required depending on the heat-treatment condition of the material. Because of their very high toughness and low carbon content, case-hardened steels tend towards forming long chips. This lends especial importance to the selection of a suitable chip breaker when turning carbide indexable inserts. Chip breakage can be improved by alloying, for example, sulphur and lead (e.g. 16MnCrS5). Alloying sulphur reduces the strength of steel. This is because of the formation of manganese sulphides (see also Sect. 7.3.3). A reduction of strength by alloying lead is only observable from a temperature of over 300◦ C. After machining, the case-hardening process follows in three steps: carburizing, hardening and tempering. The workpiece rim zones are carburized to obtain

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0.6–0.9% carbon. After case-hardening (direct hardening, single hardening, double hardening), the hardness values in the rim zone increase up to 62 HRC. As a result of the warpage of the components caused by the case-hardening, a cutting postmachining process must be executed in some cases. Finest-grain carbides, mixed ceramics and PCBN cutting tool materials are especially suited to finishing highly heat-treated or hardened steels (> 45 HRC). Very high resultant forces arise in the process. Because the chip is annealed in the clearly high temperatures, chip breakage does not become a problem. As a rule, very high surface qualities are achieved in this way (see also Sect. 7.4.6). Frequently used case-hardened steels include C15E, 16MnCr5, 20MoCr4, 18CrNi8.

7.4.3 Heat-Treatable Steels Heat-treated steels have carbon contents between 0.2 and 0.6% and are therefore stronger than case-hardened steels. The main alloying components are silicon, manganese, chrome, molybdenum, nickel and vanadium. The machinability of heat-treatable steels depends primarily on their crystalline structure, which is a result of the respective heat treatment applied, and can thus vary to a great extent. The influence of the material structure on machinability is generally stronger than that of the alloying elements. The development of wear is essentially determined by the ferrite- and perlitecontent in the microstructure. In the case of unalloyed heat-treatable steels with a carbon content up to about 0.5%, regularly formed ferritic/perlitic microstructures have a positive effect on machinability. Increasing the perlite content accelerates the development of tool wear and increases the resultant force. This limits the applicable range of cutting speeds. At higher cutting speeds, the end of tool life is brought about through crater lip breakage, especially when the steel contains larger amounts of chrome, manganese and vanadium (alloyed heat-treatable steels). The length of the chips when machining heat-treatable steels depends considerably, as do other machinability criteria, on the respective type of heat treatment and the crystalline structure of the steel. Similarly to case-hardened steels, chip breakage can be improved by means of tool geometry and by alloying lead and sulphur. The intended use of the heat-treatable steels is of particular importance when selecting alloying elements which promote chip breakage, since this can lead to a reduction of strength. Cementite spheroidization (soft annealing) is advantageous with respect to the wear behaviour of steels with larger ratios of carbon. This allows higher cutting speed to be used. However, the adhesion tendency also increases, which degrades the surface quality of the workpiece. Soft-annealed structures with a mixture of lamellar and granular cementite also lend themselves to machining at higher cutting speeds. Abrasion and thermal wear become insignificant when machining heat-treatable structures (predominantly tempered martensite). The cutting speed should be

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reduced correspondingly. The most appropriate cutting edge materials for machining heat-treatable steels are coated cemented carbides for rough-machining and, for planing, cermets. Tools made of high speed steel are used in many cases for boring and thread die cutting. In view of abrasion and thermal wear, the selection of a suitable coating is recommendable here. Material hardness values > 45 HRC require highly wear-resistant cemented carbides, cutting ceramics or CBN as cutting materials. For turning, milling and drilling, sulphur additives at a ratio of approximately 0.06-0.1% effect a clear improvement of the steel’s machinability. At higher values, this improvement is reduced and the steel’s strength is lowered. Heat-treatable steels are heat treated in order to regulate their mechanical properties with respect to their intended purpose. They can only be adjusted for good machinability in few cases. Moreover, heat treatment procedures for achieving good machinability vary. For example, the machinability of steel C60E is improved through soft-annealing, that of steel C22E through coarse-grain annealing (or cold-forming). In some cases, hardening and tempering takes place between rough-machining and planing or precision machining. Rough-machining, for which the most important factor is a high chip-removal rate, is performed on materials in a normalized state whose machinability is characterized, because of their ferritic/perlitic microstructure, by relatively low wear. Most components made of heat-treatable steel are machined in a heat-treated state. The associated strength values exceed those of the annealed state. For this reason, high cutting speeds induce a strongly increase in tool wear. Among the heat-treatable steels frequently used in practice for machining are C45E, 42CrMo4, 30CrMoV9 and 36CrNiMo4. These materials are utilized for components of medium and high strain, especially in automobile and aircraft construction (connecting rods, axles, axle-pivots, rotor and crank shafts, springs, cogwheels).

7.4.4 Nitrided Steels The carbon content of nitrided steels lies between 0.2 and 0.45%. They are heattreatable and are alloyed with Cr and Mo for improved hardenability, as well as with aluminium or vanadium (nitride formers). Nitriding is carried out at temperatures between 500 and 600◦ C, i.e. below the α-γ-transformation temperature of the material [MDH80]. As opposed to case-hardened steel, for which high levels of hardness are achieved by means of a γ-α-phase transformation and the production of the metastable phase martensite, nitrided steel has a very hard surface traced back to the brittle metal nitrides. The nitrogen diffusing into the surface layer during the nitriding process forms with the alloying elements Cr, Mo and Al special nitrides. These mostly precipitate in submicroscopic form and cause high latticework tensions, i.e. high surface hardness.

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The machining of these steels is done prior to nitriding, and usually in a heattreated state. This structural state (i.e. fine, regularly distributed carbids, tempered martensite), which is favourable for subsequent nitriding, exhibits unfavourable machinability properties. The heat-treated structure usually found when machining and the carbides distributed in the structure lead to high mechanical and thermal stresses on the tools. A short tool life is to be expected, especially at high cutting speeds. Due to the high strength values of heat-treated steels, the resultant force is relatively high. If the nitrided steels are machined in a soft state, burr formation may cause a degradation of the surface quality and an impairment of the quality of the component. When machining nitrided steels in a heat-treated state, one can expect predominately acceptable chip forms. In a non-heat-treated state, however, problems related to chip breakage will arise during machining. Greater precipitations of ferrite in nitrided steel lead to an embrittlement of the rim zones and to an irregular transition in the core zone. Coarse-grain annealing for achieving good machinability is not to be recommended with respect to the steel’s later use, since the ferrite would become even more coarse-grain and the strength would further decrease. Nitrated steels with increased nickel content, such as 34CrAlNi7 with approx. 1% Ni, are difficult to machine. Nitrided steels containing aluminium are fundamentally more difficult to machine than aluminium-free ones, such as 31CrMo12, which exhibits a lower adhesion tendency. The addition of sulphur (34CrAlS5) has a positive effect on machinability. Nitrided steels are used in a similar range of applications as case-hardened steels (cogwheels, guide strips, etc.).

7.4.5 Tool Steels One generally distinguishes between non-alloy and alloyed tool steels. Tool steels are required for different stresses. On this basis, the following categories are used: • cold-working steels, • hot-working steels and • high-speed steels. These categories are also useful for describing the machinability of tool steels. Non-alloy tool steels in a forged or rolled state with a carbon content up to 0.9% contain lamellar perlite and ferrite, while those with higher carbon contents have lamellar perlite and a cementite network. Irrespective of carbon content, softannealed steels should have more or less regularly distributed cementite grains in a ferritic matrix. With an even higher carbon content, the cementite network cannot be removed by means of conventional soft annealing. In a hardened state, the structure consists primarily of martensite in the rim layers. The martensite gradually transforms into intermediate structures as well as

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fine-lamellar perlite in the direction of the workpiece interior. In the case of supereutectoid steels, cementite grains are embedded in the matrix as well, if the steel was soft-annealed prior to hardening. Should this treatment be left out, then remnants of the brittle cementite network take the place of the cementite grains. Non-alloy tool steels with a carbon content between 0.5 and 1.5% are machined in a soft-annealed state. Subeutectoid, non-alloy tool steels can also be machined in a normalized state or in the condition of delivery after hot working. In both cases, a relatively inferior machinability is to be expected because of the increased adhesion tendency and the growth of built-up edges. The resultant force when machining tool steels is determined to a great extent by the special alloy composition and the type of heat treatment used. When machining alloyed tool steels, the dissolution of carbide formers and the increase in strength associated with this leads to an increase in the resultant force. When machining tool steels in a normalized or soft-annealed state, the increased adhesion tendency and associated growth of built-up edges have a negative effect on surface quality. This can be partially remedied by means of a quenching and tempering to a higher strength. Because of the high deformability of ferrite, long chips with bad breakage form when machining tool steels in a soft-annealed state. An increasing carbide moulding degrades the chip breakage. If machining is executed in a heat-treated state, chip breakage is not to be considered a problem. The amount of carbide formers is bears little importance for the machinability of alloyed tool steels. Carbide formers only increase the wear effect on the steel in an obvious way when they have dissolved during austenitizing and have not formed any carbides during subsequent annealing. The alloyed tool steels, especially high-alloyed high speed steels, are poorly machinable in an annealed state. This is due, as with unalloyed tool steels, to the marked formation of gluing-points and built-up edges. Disruptions may form at the outlet points of the tool. The adhesion tendency can be reduced by quenching and tempering to greater strengths (1200–1400 N/mm2 ). This increases abrasive wear and the thermal stress on the cutting edge, however. The cutting speeds which are applicable in machining tool steels are, as a rule, relatively low and increase with the level of carbide moulding. However, the adhesion tendency of these steels, with finely distributed granular carbides, increases to an equal extent. Cutting materials most often used for machining tool steels are cemented carbides containing titanium carbides and tantalum carbides with medium toughness (e.g. from application group P20) as well as cermets. Following machining in an annealed state, tool steels can also be machined in a heat-treated state (Rm < 2000 N/mm2 ) using cutting edge materials made of CBN. Selecting alloying additives for tool steels is based first and foremost on their influence on surface hardness, hardness penetration depth, tempering consistency, toughness and wear resistance, whereby a suitable coordination with the carbon content is necessary, especially for higher-alloyed steels. The carbon content of the steel determines the ratio of carbides, which significantly promote abrasive wear. Carbon also influences hardenability and contributes decisively to tempering consistency and toughness via carbide reactions during hardening and tempering.

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7.4.6 Hardened Steels For a long time, grinding used to be the only way to machine hardened steels. The development of ultrahard cutting edge materials in the 1980’s led to the machining of these materials with geometrically defined cutting edges. Since then, the machining of hardened materials exceeding a hardness of 50 HRC has come to be known as hard machining. While hard turning had been implemented almost exclusively in single and smallbatch production up to a few years ago, it is now used increasingly for large-batch and mass production. As a rule, hard turning is advantageous if a machining multiple sides can be achieved in a single clamping. When machining in one clamping, a very high accuracy of position of the functional surfaces can be realized. Hard turning procedures for machining interiors and contours are still very important. For turning hardened steel, almost the only options are ultrahard non-metallic cutting edge materials, such as crystalline cubic boron nitride (PCBN) and – with some limitations – mixed ceramics. This is because of the extremely high requirements on hot hardness and the cutting tool material’s consistency against diffusion resulting from the high temperatures and from cutting pressures in the machining zone. Because of their low hot hardness, however, finest-grain cemented carbides and ultrafine-grain cemented carbides are only suited to hard machining with low cutting times, such as turning with interrupted cutting, hard milling, broaching, gear skiving and -shaping. Because of their toughness, which clearly exceed both that of the very heat-resistant, though relatively brittle PCBN and, even more so, that of cutting edges made from mixed ceramics, they exhibit very good working behaviour in these applications [Koch96, Wina96]. The spectrum of chemical and physical properties of commercially available PCBN cutting materials is relatively large. In addition to clear differences in hardness and resistance to bending, properties that vary especially are heat conductivity, resistance to thermal shock and thermo-chemical inertia with respect to the material to be machined. This variance can be attributed to the cBN grain size and to the type and concentration of the binder phase [Kloc05b, Joch01]. When machining hardened steels with PCBN, the surface quality can be improved by modifying the cutting edge geometry. By selecting a large tool orthogonal clearance, one can principally increase the cutting time without changing the amount of free flank wear and simultaneously increase the displacement of the cutting edge. For the cutting edge fillet, an optimum radius of approximately 20 μm is recommended. Small bevels and radii lead to disruptions and wavering tool life, while larger bevels and radii cause larger forces and may induce oscillations [Kloc05c]. Mixed ceramics represent a further application group which can be used for hard machining with continuous cutting. In comparison to PCBN cutting edge materials, the use of mixed ceramics as cutting edge materials places less stress on the cutting edge with the result of clearly smaller chip thicknesses of a maximum 100 μm. The use of round indexable inserts or the selection of the greatest possible corner radius is recommended for pre-machining with smooth cutting [Abel93]. Because

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of the low toughness, anything but smooth cuts cannot be recommended [Köni89a, Momp87]. In consideration of the fact that tool costs are lower by a factor of 10–20, mixed ceramics for smooth-cutting represent a widespread alternative to cutting edge materials with a small ratio of CBN. High resultant forces and high rake face temperatures arise during hard machining with rough cutting. Specific cutting forces kc between 4000 and 4700 N/mm2 and rake face temperatures which may even, in exceptional cases, exceed the local melting temperature of the material to be machined have been recorded [Berk92]. As opposed to soft machining, the passive force takes on very high values in this case. These values may even be higher than those of the cutting force [Köni84b, Töns93, Wobk93]. The normal compression stresses in the region of the wear mark on the side of the minor flank lead to a high mechanical and thermal stress on the workpiece rim zone. The mechanical stress is the result of flank load similar to H ERTZian pressure. The stress state in the workpiece caused by this pressure induces a residual austenite transformation and cold-forming in the rim layer of the workpiece [Gold91]. Through this process, compressive stresses are induced which increase with increasing flank wear of the minor cutting edge. For this reason, when hard-machining, the maximum amount of compressive residual stresses are shifted to the greater depths of the workpiece rim zone with an increase in flank wear. The thermal stress is a result of the friction between the flank face and the workpiece. Both high normal stresses and high shear stresses lead, in combination with the relative movement between the tool and the workpiece, to high friction forces. High contact temperatures are the result, and these cause, via rapid cooling, the formation of martensite when the α-γ-transformation temperature is exceeded. The martensite is recognizable in the microsection as a “white”, unetched layer (Fig. 7.17, right). A darker zone with a tempered structure can be seen below the white rim layer. Phenomena which occur at high temperatures because of transformations and high cooling gradients cause tensile residual stresses which are superposed on mechanically induced compressive residual stresses [Jonk87]. The magnitude of the flank face wear on the minor cutting edge has a significant influence on the distribution of stress in the workpiece rim zone. Metallographic studies

20 µm

Fig. 7.17 Formation of the rim zone during hard turning

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Machinability of Various Steel Materials h > 20 µm

273 h < 20 µm

Fig. 7.18 Longitudinal section of a saw tooth chip (left) and a ribbon chip (right)

of hard-machined workpiece surfaces confirm the influence of tool wear on the formation of a new hardening zone. How the distribution of stress and the new hardening zones ultimately affect the behaviour of the component has not been sufficiently explained yet. Of particular interest are the effects on the fatigue strength of rolling contact stressed functional surfaces. The formation of new hardening zones during hard machining cannot be inhibited, not even by using coolants. Since the use of coolants effects no improvement of the tool life behaviour of the cutting edge material, hard-machining is generally performed in a dry state. When machining with interrupted cutting, the thermo-shock sensitivity of superhard cutting tool materials even forbids the use of coolants [Acke89, Töns91]. Highly heat-treated steels which receive their hardness mainly from their martensitic crystalline structure are practically non-deformable in room temperature and environmental pressure. This deformability behaviour results in a special chip formation, one different from the classic chip formation of unhardened steels. When machining hardened steels, two different chip forms may develop depending on the chip thickness. With high chip thicknesses (h > 20 μm), a “saw-tooth” chip develops [Acke89, Berk92, Naru79] (Fig. 7.18), while smaller chip thicknesses (h < 20 μm) tend to cause the formation of a ribbon chip [Köni93b]. The development of “saw-tooth” chips initiates material separation via a crack formation on the surface which transmits itself into the workpiece under an angle of the greatest shear stresses [Acke89]. The released chip segment slides off between the rake face and the newly developed separating surface. If the chip segment has slid so far that the increasing compressive stresses in turn cause the critical shear stress to be exceeded, the next crack forms. One theory ascribes the connection of chip segments to the fact that the crack is caught up in front of the cutting edge and a plastic deformation takes place due to the compressive stresses and high temperatures. Chip segments on a plasticized material thus slide off [Berk92]. In the case of high chip thicknesses, the transition from a ribbon chip to a saw-tooth chip is a function of the hardness of the material. With the predominately martensitically hardened, unalloyed tool steel C105W2 (1.1645), for example, this transition takes place at approximately 50 HRC [Naru79]. At low chip thicknesses (h < 20 μm), the chip thickness is in the region of the cutting edge radius rn with a tool orthogonal

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rake angle which acts in a correspondingly strongly negative way. A triaxial stress state with a high ratio of hydrostatic compressive stress is thereby caused in the area of the material directly in front of the cutting edge with no material separation ensuing. The shear stress hypothesis put forward by MOHR offers an explanation of these phenomena. VON KÁRMÁN already indicates the possibility of a plastic deformation of brittle materials under pressure on all sides [Karm11]. If such a state of compressive stress exists, the shear yield stress becomes a decisive criterion. If, however, the strain on the material is characterized by a monoaxial compressive stress state as it occurs on the workpiece surface, the failure of the material is caused by a cleavage fracture. Precision hard turning (vc ≈ 100–200 m/min, f ≈ 0.05–0.15 mm, ap ≈ 0.1–0.5 mm) is used for manufacturing ready-to-use components. Surface roughness values in the region of Rz = 2.5–4 μm can be achieved in series manufacturing in a procedurally reliable way. A further development of precision hard turning tending toward higher component quality is high-precision hard turning (vc ≈ 150–220 m/min, f ≈ 0.01–0.1 mm, ap ≈ 0.02–0.3 mm), which has the potential of achieving surface roughness values which lie clearly below those of conventional precision hard turning processes. The surface roughness Rz achievable when using special machine tools lie below 1 μm. However, one must also expect that increasing tool wear will bring about a degradation of surface quality. The procedure shows potential for creating surfaces with Rz ≤ 3 μm with long tool life.

7.4.7 Non-rusting Steels Non-rusting steels contain at least 10.5% chromium and a maximum of 1.2% carbon. Their main alloying elements are chrome and nickel. Amounts of chromium exceeding 12% render the steel material corrosion-resistant. Nickel expands the γ-region and, in the case of steels with a high chromium content, leads to the stabilization of the austenitic structure, which is generally unstable in the case of low-alloyed carbon steels and which disintegrates into ferrite and cementite below the Ac1-temperature. Nickel starkly decreases the heat conductivity of the steel material. “Chrome steel” is the usual term for ferritic steel types and “Cr-Ni steel” for austenitic steel types. According to their essential usage properties, DIN EN 10088-1 categorizes non-rusting steels into: • corrosion-resistant steels, • heat-resistant steels and • high-temperature steels. Non-rusting steels are distinguished by good persistence against chemically aggressive substances. In general, they have a chromium content of at least 12%. Protection is achieved with a chromium content of more than 10.5% by means of the spontaneous formation of a protective chromium oxide layer [DINEN10088]. With respect to their metallographic constituents, the corrosion-resistant steels are subdivided into ferritic, martensitic, austenitic and ferritic-austenitic steels. Figure 7.19


275 40

60

7.4

l ke Nic

Iro nc on ten t/ 80 %

70

30

n nte co 20

Austenite

t/ %

Martensite + 90

1

10

Austenite

Austenite +

Martensite

Ferrite Ferrite

0 10

0

10

1 : X5CrNi18–10;

2

0

3

20 Chromium content / % 2 : X6Cr17;

30

40

3 : X39Cr13

Fig. 7.19 Phase diagram of Cr-Ni-steels with metallographic constituent data after quenching, acc. to IGNATOWITZ [Igna97]

shows the phase diagram of Cr-Ni steels. One can see from this diagram which structure is to be expected, depending on chromium and nickel content, with a nonrusting chromium-nickel steel after quenching. Point 1 corresponds to the austenitic steel X5CrNi18-10 with 18% chromium, 10% nickel and 72% iron, point 2 corresponds to the ferritic chrome steel X6Cr17 and point 3 to the martensitic steel material X39Cr13 [Igna97]. The ferritic steels are primarily chrome steels with a chromium content of 12.5–18% and a C-content below 0.1% (e.g. X6Cr13, material no.: 1.4000, X6Cr17, material no.: 1.4016). They are magnetic and not hardenable. Martensitic steels are primarily chrome steels with a chromium content of 12–18% and a C-content of 0.1–1.2% (e.g. X12Cr13, material no.: 1.4006; X39Cr13, material no.: 1.4031). Depending on the quality, these steels also contain additional Ni and Mo. They are magnetic and can be heat-treated and tempered by means of a corresponding heat-treatment. With increasing C-content, the hardness of the steel increases when in a hardened state. A hardness of 40 HRC is achievable with a C-content of 0.1% and a hardness of 59 HRC with a C-content of 0.9%. The steels the most used by far are austenitic steel materials. They contain approximately 17–26% Cr, 7–26% Ni, less than 0.12% C and, in some cases, small

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amounts of Si, Mo, V, Nb, Ti, Al or Co (e.g. X5CrNi18-10, material no.: 1.4301; X6CrNiMoTi17-12-2, material no.: 1.4571). They are not magnetic and cannot be hardened by means of heat treatment. In an annealed state, they are characterized by very good toughness properties which are retained even at extremely low temperatures. Especially given a high carbon content, they tend when cold-formed towards considerable strain hardening. Through cold-forming, therefore, their strength values can be increased dramatically. When strain-induced martensite forms, however, the strain is reduced to a very considerable degree. Among the products made from the abovementioned corrosion-resistant steels are household devices, razor blades, knives, surgical instruments, parts for automobile construction, agricultural and materials handling technology, mechanical and plant engineering and devices and instruments for the food industry, the chemical industry, the textile industry and for shipbuilding [Stah01]. A further group of corrosion-resistant steels are the ferritic-austenitic steel materials, also either referred to as duplex steels (e.g. X2CrNiMoN22-5-3, material no.: 1.4462) or super duplex steels (e.g. X2CrNiMoCuWN25-7-4, material no.: 1.4501). The name “duplex” refers to their two-phase crystalline structure consisting of ferrite and austenite. These steels have optimal properties at a balanced ferrite/austenite ratio of approximately 50/50%. In comparison to normal austenitic steels, the duplex steels contain less nickel (about 4–8%), though usually a significantly higher amount of chromium (about 18–25%). In order to increase their resistance to intercrystalline corrosion, a certain amount of nickel is exchanged for additional nitrogen as an austenite former and molybdenum. The optimal microstructure is created by means of a heat-treatment at 1000–1100◦ C. These steels find the most use in the gas and oil industry, in the petrochemical industry, in chemical tankers and in sewage treatment plant construction. In comparison to austenitic rustproof steels, the duplex materials have a strain limit which is almost double as high (Rp0,2 about 400–550 MPa) given similar or markedly higher strength values. Heat-resistant steels are mainly ferritic and austenitic steels with a high resistance against oxidation as well as against the influence of hot gases and combustion products above 550◦ C. As a rule, the heat-resistant ferritic steels contain at least 12% chromium as well as aluminium and silicon (e.g. X10CrAlSi13, material no.: 1.4724; X10CrAlSi25, material no.: 1.4762). Heat-resistant austenitic steels are additionally alloyed with at least 9% nickel (e.g. X8CrNiTi18-10, material no.: 1.4878; X15CrNiSi25-21, material no.: 1.4841). These steel types are used, for example, in ovens and apparatus construction for annealing bells, baskets and tubes [Stah01]. The high-temperature steels are mainly martensitic and austenitic steel types with a high long-time rupture strength with a long-term mechanical strain above 500◦ C. Among the products made from high-temperature martensitic steels (e.g. X20CrMoV11-1, material no.: 1.4922; X20CrMoWV12-1, material no.: 1.4935) are components for thermal power plants, steam boilers and turbines, as well as for the chemical industry and nuclear technology. High-temperature austenitic steels (e.g. X6CrNi18-10, material no.: 1.4948; X5NiCrTi26-15, material no. 1.4980) are used in the construction of pressure vessels and armatures, pressure tanks and steam

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boilers, but also in the manufacture of blades, discs, axles and bolts for steam and gas turbines as well as in shipbuilding [Stah01]. Non-rusting steels with a ferritic structure lend themselves to machining relatively well. The onset of wear through abrasion and adhesion is comparatively low. The machinability of martensitic steels is a function of hardness, which is in turn a result of the heat treatment used. Depending on the heat treatment, the structure consists either of martensite (hardened) or tempered martensite with chromium carbides and ferrite (quenched and tempered). Duplex steels are considered to be extremely hard to machine. Characteristic reasons for this are a marked adhesion tendency, high strain hardening of the workpiece rim zone and an unfavourable chip formation. Because of the eminent importance of corrosion-resistant austenitic steel materials, the following will go into more detail on their machinability properties. Austenitic steel materials are machined in either a quenched or a solution-annealed state. In comparison to ferritic-perlitic or quenched-and-tempered steels, they are significantly harder to machine. This is due to their high deformability and toughness, their tendency towards strain hardening and towards adhesion with the cutting material, as well as their lower heat conductivity, which is about 1/3 lower than with non-alloy steels. The latter impairs heat removal through the chip and increases the thermal stress of the cutting edge. In spite of the comparatively low tensile strength of austenitic Cr-Ni steels, the results of these specific material properties are a high thermal stress of the tool cutting edge, pronounced flank face and/or rake face wear, material gluing, notch wear, pitting of the cutting tool material, cutting edge breakages and unfavourable chip forms, which allow for machining tasks to be performed only at relatively low cutting speeds, for short tool lives and with insufficient surface qualities. In the case of coated tools, one can observe in many cases that the adhesion between the material and the hard material coating is stronger than their adhesion on the substrate and that this leads to the delamination of the coating. Austenitic steel materials are machined preferably with uncoated or coated WC-Co cemented carbides of the main application group M. Because of the high thermal stress of the cutting edge, cutting speeds are used which, in comparison to those used in turning ferritic-perlitic steels, are lower by a factor of 2–5. Applicable cutting speeds range from vc = 50 m/min (X5NiCrTi26-15) to vc = 160 m/min (X6CrNiMoTi17-12-2), depending on the alloy. The tool life is between 5 and 15 min. A type of wear which limits tool life when turning austenitic steel materials with uncoated cemented carbides is the formation of a pronounced crater on the rake face (Fig. 3.41, left). Due to this strong crater wear, uncoated cemented carbides can only be used at relatively low cutting speeds (vc < 100 m/min) [Gers04]. The use of high cutting speed is only made possible when turning austenitic steels through the application of coated tools (Fig. 3.41, right). Significantly improved performance in turning austenitic steels requires cutting tool materials with properties customized with respect to the respective machining task (Fig. 7.20). The more the cutting tool material system reaches the physical limits of the substrate and the hard-material coating in the chipping process, the

278


A

20

Tool life Tc / min

B

10 9 8 7 6

R

R

5 4

*

*

3 X5CrNi18-10 2 100

150

X6CrNiMoTi17-12-2 250 300 100 200 Cutting speed vc / (m/min)

150

200

250

300

Cutting tool materials: A: WC-7.5Co-6.8MK-0.75Cr3C2 / TiN-TiCN-Al2O3-HfCN B: WC-7.5Co-6.8MK-0.75Cr3C2 / TiAlN R: Reference cutting tool material Tool life criterion: flank wear VB = 0.3 mm, *: plastic deformation of cutting edge Process: external cylindrical turning Cutting parameters: feed: f = 0.32 mm, depth of cut: ap = 2 mm Cutting fluid: emulsion 6%

Fig. 7.20 Examples for the performance of carbides with tailored substrates and coating systems compared to a common cutting tool material during roughing of austenitic steel, acc. to GERSCHWILER [Gers04]

more carefully the substrate, the hard-material coating and the tool geometry must be customized to the individual machining task [Gers04]. In addition to a high resistance to wear, machining tools used to turn austenitic steel materials are also required to exhibit “performance” in terms of chip formation and chip breakage. Because of their high deformability, the austenitic steels tend more than any others towards the formation of long ribbon and snarled chips. The latter represent a high hazard potential not only for the component and the machine, but also for the operating personnel. It is thus essential, and not only in automated manufacturing plants, that this type of chip formation is avoided through all available means to secure a higher level of machining safety.

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279

A safe chip breakage is guaranteed in two ways. Firstly, on the material side, through alloying measures. In many ways, however, there are narrow limits on these measures placed by the usage properties of the material and the component. The more promising way in most cases is thus the optimization of the cutting part geometry. In case of varying cutting depths and/or small feeds, however, this places an extremely high demand on the geometric design of the cutting edge. Especially with small chip cross-sectional areas, the tools must possess chip shaping components which lie close enough to the cutting edge so that they come into any contact at all with the fine chips [Köni90a]. The high ductility of austenitic materials requires tools with the sharpest possible cutting edges. Sharp tool cutting edges facilitate the separation of the material during chip formation and, by reducing the resultant forces, contribute in a highly essential way to lowering the plastic deformation of the workpiece rim zone, to the creation of high-quality workpiece surfaces and to the reduction of burr formation. On the other hand, however, one can also observe that, with increasing mechanical stress, very sharp cutting edges lead also contribute to an increased pitting of the cutting tool material, which in turn leads to accelerated wear. As experiments have shown, a defined cutting edge fillet can stabilize the cutting edge and significantly improve tool life. When finish turning with small chip cross-sectional areas, however, a cutting edge fillet no greater than 30 μm should be selected. When rough turning, on the other hand, a highly stable cutting edge can be designed with a 40–60 μm fillet in order to prevent edge breakages and pittings of the cutting tool material. Burr formations become a further serious problem when machining ductile austenitic steel materials. In the case of components which are to be burr-free, up to 20% of manufacturing costs may result from burr removal [Köni93]. In addition to the costs for burr removal, increasing demands on component quality and workplace attractiveness entail combating burr formation through suitable measures. As a rule, austenitic steels are machined by wet cutting. However, dry machining is thoroughly possible with corresponding process dimensioning. One must consider, however, that due to the lack of the cooling effect of the cutting fluid, both chip and tool exhibit higher temperatures than is the case with wet machining. The results of this are an increased in smeared material on the tool, clearly inferior chip formation, more aggressive notch wear on the major and minor cutting edges and, as a result, increased burr formation and inferior surface qualities. In comparison to wet machining, dry machining austenitic steels can lead to a drastic reduction of tool life, the latter being limited as a rule not by flank face wear, but rather by the degradation of the surface quality associated with notch formation on the minor cutting edge. Machining austenitic steels in an economical way thus requires special measures. In addition to a tool geometry customized to the machining task, suitable coating systems and cutting parameters, the most important of these is the use of a minimum quantity lubrication. This lubrication in particular can reduce rewelding of the material and significantly improve the surface quality. Mediums used for this are, as a rule, synthetic ester oils [Gers04] (also Chap. 6).

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7.5 Machinability of Cast Iron Cast iron refers to iron-carbon alloys which have a carbon content above 2.06%. In the iron-carbon two-material system, cast iron becomes a steel (up to 2.06% C). The material is usually shaped by means of casting and a final machining operation. Cast-iron alloy material groups are classified on the basis of the appearance of their fracture faces into white and grey cast iron (Fig. 7.21). Among the white cast iron materials are chilled cast iron and malleable cast iron, both of which solidify according to the metastable iron-iron carbide system and in which carbon is thus present in the form of cementite (Fe3 C). As opposed to this, the solidification of grey cast iron takes place according to the stable iron-graphite system, with the carbon thus being present as graphite. Depending on the form of the graphite, grey cast iron materials are divided into cast irons with lamellar graphite (GJL), cast irons with vermicular graphite (GJV) and cast irons with spheroidal graphite (GJS). Cast iron

White cast iron

White malleable cast iron (GJMW)

Black malleable cast iron (GJMB)

Grey cast iron

Chilled cast iron (GJH)

Lamellar graphite (GJL)

Vermicular Spheroidal cast iron graphite (GJV) (GJS)

Fig. 7.21 Cast iron classification

In addition to the metallic matrix, the machining properties of iron-cast materials are also heavily influenced by the amount and formation of the embedded graphite. Graphite inclusions, in the first place, reduce friction between the tool and the workpiece and, on the other, disrupt the metallic matrix. This leads to improved machinability in comparison with graphite-free cast iron or steel materials. The results are short-breaking chips, small grinding forces and higher tool lives. The metallic matrix of cast iron is, to a high degree, crucial to its machinability. The matrix is influenced by the respective chemical composition (alloying elements) and heat treatment and consists, in the case of materials with low strength, predominately of ferrite. The material becomes stronger as the amount of perlite rises, which also leads to a more significant abrasive tool wear. Cast irons of high strength and hardness often have a bainitic, ledeburitic or martensitic structure and are therefore very hard to machine.

7.5.1 White Cast Iron As mentioned above, the solidification of white cast iron takes place according to the metastable Fe-Fe3 C system (Fig 7.22). If the cooling speed is high, a eutectic

Machinability of Cast Iron

281

Steel A I

769 723 α-Mc 500

Melt

δ-Mc + γ-Mc

1147

911

B

Austenite (γ-Mc)

G M P

α-Mc + γ-Mc O

S

Ferrite + Perlite Q 0

D

Melt + Austenite

E

Austenite + Secondary cementite Perlite + Secondary cementite

0.8

Austenite + Secondary cementite + Ledeburite

Perlite + Secondary cementit + Ledeburite

Melt + Primary cementite C

F Ledeburite + Primary cementite

Primary cementite + Ledeburite

4.3 2.06 Carbon concentration / mass-%

Cementite (Fe3C)

Temperature T / °C

H δ-Mc

White cast iron

Melt+ δ-Mc

Perlite

1600 1536

Ledeburite

7.5

K L 6.67

Fig. 7.22 Section of the iron-carbon phase diagram (metastable system, fast cooling)

(1147◦ C, C-content = 4.3%) forms which is referred to as ledeburite (austenite + cementite). Depending on the carbon content, additional perlite and secondary cementite is found in hypoeutectic compositions and additional primary cementite in hypereutectic alloys. The high ratio of iron carbide in white cast iron leads to great hardness with a pronounced level of brittleness. If the white cast iron is not subjected to further heat treatment, it is referred to as chilled cast iron. If a further heat treatment is executed (tempering), malleable cast iron forms. 7.5.1.1 Malleable Cast Iron The usage properties of malleable cast iron are adjusted by means of a metastable solidification and subsequent heat treatment (tempering). On the basis of the fracture appearance, malleable cast irons are subdivided into white malleable cast irons (GJMW) and black malleable cast irons (GJMB). In order to produce white malleable cast iron, the white cast iron is annealed in an oxidizing atmosphere (approx. 1000◦ C, 60–120 h). The aim is to produce a partially or completely decarburized material by removing the carbon from the casting [DINEN1562]. The degree of decarburization of the material is strongly dependent on the duration of the tempering process and on the section thickness of the casting. A complete, uniform decarburization is only achievable with low section thicknesses; with thicker castings, a rim decarburization and a disintegration

282


of the cementite in the workpiece interior take place. The remaining graphite is then present as temper carbon. In the case of partial carburization, therefore, there is a purely ferritic rim zone structure and a perlitic-ferritic core structure. As opposed to white malleable cast iron, black malleable cast iron is heat-treated in an oxygen-free atmosphere. The cast iron is thus not decarburized. The cementite disintegrates entirely into temper carbon [DINEN1562]. A characteristic property of black malleable cast iron is that, because of the non-decarburizing annealing, there is a uniform structure across the workpiece cross section, irrespective of the section thickness. As a function of cooling speed, a ferritic matrix forms given slow cooling, with the temper carbon evenly distributed in the form of nodular formations. Given a more rapid cooling, a perlitic or even martensitic matrix forms. Perlitic-ferritic mixed matrices are also possible. In contrast to steels, malleable cast iron is excellent for machining purposes. In spite of the good plastic deformability of the different types of malleable cast iron, the manganese sulphides and temper carbons embedded in the steel-like ferrite or perlite matrix cause short-breaking chips which can be removed without difficulty [Wern83]. Moreover, the temper carbon reduces the friction between the workpiece and the tool and interrupts the base metallic material, which leads to low resultant forces and higher tool life. Given identical workpiece hardness, black malleable cast iron is more easily machined that white. This circumstance is attributable to white malleable cast iron’s purely ferritic rim zone structure (without temper carbon). Problems arising when machining ferrite are the growth of built-up edges (adhesion tendency of ferrite), the formation of ribbon and snarled chips (high deformability of ferrite) and the reduction of the surface quality with increased burr formation. Cutting tool materials usually employed for machining malleable cast iron are uncoated and coated cemented carbides, cermets, oxide ceramics and PCBN of the main application groups P and K. As a result of heat treatment, a remarkable consistency is achieved in the production of malleable cast irons with respect to their mechanical/technological material properties, which allows optimal cutting conditions for economic manufacturing. Furthermore, malleable cast irons are characterized by high strength and toughness properties as well as excellent casting processability. Workpieces can be achieve with accuracy of form with high surface quality, even thin-walled and intricate workpieces. Because of their high ductility, malleable cast irons are especially suitable for components which are subjected to dynamic stresses – oscillating or impulsive – and which must resist mechanical forces of great magnitude [Wern00]. White malleable cast iron is both weldable and hardenable. When selecting a hardening procedure, one must consider that white malleable cast iron exhibits a structure which is dependent on section thickness and a carbon content which increases from the outside to the inside. Thus only a thermochemical procedure can be applied for the low-carbon rim layer, e.g. case-hardening or nitriding [Schü06]. 7.5.1.2 Chilled Cast Iron White malleable cast iron without heat treatment is referred as a result of its high zementite-content as chilled cast iron (GJH). It is both very hard and brittle and

7.5


283

thus suitable for components which are to exhibit a high resistance to wear with low dynamic stresses, e.g. for rollers and grinding tools. Because the cementite is very hard, chilled cast iron is very difficult to machine. Tool cutting edges are subjected during machining to high mechanical and thermal stresses. Therefore, high demands must be made on the cutting tool material with respect to wear resistance and pressure resistance. Materials used for machining chilled cast iron are, almost exclusively, cemented carbides, and, especially with high hardness values, cutting ceramics (oxidic mixed ceramics) and PCBN cutting tool materials of the main application group H. In comparison to cemented carbides, the use of cutting ceramics allows for an increase of the cutting speed by a factor of 3–4. The higher material removal rate, however, comes with the disadvantage of an increased susceptibility to breaking. In order to keep the mechanical stress of the tool cutting edge as low as possible when machining chilled cast iron and thus to improve the tool life parameters, the cutting speed and the feed should be reduced with increasing material hardness. Commonly used in chilled cast iron machining are a lead angle of 10–20◦ and a tool orthogonal rake angle of –5 to 5◦ .

7.5.2 Grey Cast Iron In grey cast irons, the carbon is ideally present in the structure exclusively in the form of graphite. The melt solidifies according to the stable iron-graphite system (Fig. 7.23). In reality, however, because of the final cooling speed, small amounts Grey cast iron 1600 1536

D′

Temperature / °C

Melt Melt + Austenite 1153

E′

Melt + Graphite

C‘

F′

Austenite + Graphite (+ Fe3C)

911 769 738

K′ Ferrite + Graphite (+ Fe3C) L′

500 0

0.8

2.06

4.25

Carbon concentration / Mass-% Fig. 7.23 Section of the iron-carbon phase diagram (stable system, slow cooling)

7

284


Fig. 7.24 Cast iron with lamellar graphite (left) and spheroidal graphite (right)

of Fe3 C are found in the structure in addition to graphite. In order to maximize the amount of graphite, so-called graphitizing elements such as silicon can be added to the melt. As mentioned above, both the mechanical properties and the machinability properties of cast iron materials containing graphite depend in a very clear way on the crystalline structure of the matrix and the form of the graphite inclusions. The type of metallic matrix can vary broadly according to the selection of chemical composition and heat treatment. Cast iron types with a ferrite matrix have the lowest strength and the highest plasticity and toughness; the types with a perlitic matrix have the highest strength and the lowest plasticity and toughness. Types with a mixed matrix (with different ferrit-perlite percentages) fall somewhere between these limits [Herf07]. In practice, the matrix is usually either ferritic-perlitic or purely perlitic. Only after an especially long annealing process does a purely ferritic matrix form. The graphite inclusions take on varying forms depending on the conditions of their origination, which are also controllable through metallurgical measures (Mgcontent). These forms vary from disc-like forms (lamellar) to fissured forms (wormlike, i.e. vermicular) to globular particles (Fig. 7.24) [Hors85]. One can say in sum that the machinability of grey cast irons is not determined by the hardness and strength of the alloy alone; the form, quantity and distribution of the graphite inclusions in combination with the given type of matrix play an important role [Stäh65, Töns91]. The following will treat the different types of grey cast iron. 7.5.2.1 Cast Iron with Lamellar Graphite Cast iron with lamellar graphite (GJL) is the most frequently produced type of cast iron. The graphite is present in the form of thin, unevenly formed discs referred to as “graphite lamellae”. These graphite lamellae interrupt the base metallic mass to a very great extent. Also, when stress caused by external forces is present, stress peaks form at the margins of these graphite lamellae. The lamellae then act as internal

7.5


285

notches (predetermined breaking points), for which reason the tensile strength of this type of cast iron is relatively small (100–350 N/mm2 ), the material being designated as brittle [Herf07a]. On the other hand, lamellar graphite gives the material a high heat conductivity, favourable damping properties and, due to the embrittlement, high stiffness. Cast irons with lamellar graphite are especially suitable for machine beds and stands and bearing, gear and machine housings. When machined, GJL is distinguished by excellent self-lubricating properties. These result from the circumstance that the graphite lamellae are cut during the machining process, with the graphite consequently forming a lubrication layer on the tool. This leads to lower wear and higher tool life. In addition, the graphite lamellae act favourable on the chip form, since they stop any incipient shear early on and induce cracking leading to the formation of segmented or discontinuous chips. Short-breaking chips develop, usually spiral chip segments or discontinuous chips. The mechanical strain on the tool cutting edge when machining GJL tends to be low, leading to low resultant forces. A major advantage when machining GJL is derived from the formation of wear-reducing manganese sulphide layers on the flank and rake faces of the cutting insert. With increasing cutting speeds, the manganese sulphides form a layer on the insert which becomes thicker and thicker. This layer reduces the friction coefficient on the one hand and acts as a diffusion barrier and a protection from wear on the other. The formation of wear-reducing coatings only sets in with higher cutting speeds (from approximately 200 m/min onwards), i.e. due to the resultant increase in the machining temperature. This effect ensures long tool lives when machining GJL. The surface quality of machined workpieces made of GJL is a function of the finishing method, the cutting conditions and the composition, fineness and evenness of the cast structure [Opit70]. There is generally no burr formation at the workpiece edges during machining; instead, because of the material brittleness, there are edge disruptions. The hardness of the material is a reference value in the first approximation for the applicable cutting speed. Cast iron with lamellar graphite only having a small amount of perlite (about 10%) after annealing treatment, for example, can be machined, with the same tool life, at a cutting speed three times faster than with a cast iron with a large amount of perlite (about 90%). Other hard metallographic constituents, e.g. the phosphide eutectic steadite, increase tool wear and reduce tool life in the same way cementite does in perlite. They significantly reduce the applicable cutting speeds (Fig. 7.25). The rim zone structure of cast iron workpiece generally exhibits lower machinability than the core zone. This can be attributed, on the one hand, to non-metallic inclusions and, on the other, to the altered graphite structure and microstructure directly beneath the outermost cast layer, as well as to high temperature oxidations. This results in stronger abrasive wear and the formation of a wear notch on the tool cutting edge. In practice, this is often compensated by means of a reduction of the cutting parameters.

286

7 Tool Life Behaviour Tool life criterion: VB = 0.75 mm Cutting tool material: HW-K20 Cross-section of undeformed chip: ap·f = 2.5 · 0.25 mm² 300

Cutting section geometry γo αo λ s χ r ε r rε 6°

6° 0° 90° 84° 1.2 mm

vc = 20 m/min vc = 60 m/min

Ferrite

Cuttingspeed vc/(m/min)

250

200 Ferrite + Perlite Perlite (coarse grained)

150

Perlite (fine grained) 100

50 Quenched structure 0 0

125

150 175 200 225 250 Brinell hardness HB 30/5

275

300

Fig. 7.25 Influence of the crystalline structure and hardness on the cutting speed during turning of cast iron with lamellar graphite

A rich array of cutting tool materials lends itself to machining cast iron with lamellar graphite. The use of high speed steels is generally limited to tools with very fine cutting edges. Typical procedures used are boring, reaming and thread die cutting. The classic types of cemented carbide for machining cast irons with lamellar graphite are those in the main application group K. Types K01-K05 and especially cermets are especially suited to precision and ultra-precision machining. The material removal rate can be increased considerably by coated cemented carbide, cutting ceramics and especially PCBN cutting tool materials of the main application group K. Coated cemented carbide indexable inserts are used for turning, boring and milling grey casts. The longest tool lives and highest cutting speeds can be achieved in the process by using cemented carbide with ceramic multi-layer coatings. Silicon nitride ceramics are suitable for rough machining with large interrupted cuts or extreme irregularities in the contour of the workpiece as well as for machining with cutting fluids. 7.5.2.2 Cast Iron with Vermicular Graphite In cast irons with vermicular graphite (GJV), the graphite doesn’t form in lamellar or globular form, but rather in coral-like form, worm-shaped in microsection

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287

(lat. Vermicular: worm-shaped). The rounded edges of the graphite inclusions are reduced by internal stress peaks, as they appear on the tapered ends of graphite lamellae in GJL. With respect to its mechanical and thermal properties and its machinability, vermicular cast iron lies between cast iron with lamellar (GJL) and with globular (GJS) graphite inclusions. In comparison to lamellar cast iron, GJV has higher strength (up to 70% higher), higher toughness, higher stiffness, higher hardness, higher endurance strength, higher oxidation-resistance and higher thermal shock resistance. In comparison to cast iron with spheroidal graphite, GJV has better pouring properties, a better machinability, a better damping ability, a lower tendency to warp, a lower thermal expansion and a better deformation resistance with temperature change. Because of these properties, cast irons with vermicular graphite can be cast more with thinner walls than GJL and GJS, which allows for lighter components. These combinations of properties render GJV suitable as a construction material for combustion engines. These highly resistant types of cast iron are particularly suited to the demands of modern direct injection diesel engines. Particularly when faced with the conflicting interests of high component strength vs. low component weight, GJV offers a good alternative to traditional construction materials such as aluminium and cast iron with lamellar graphite. In comparison to traditional motors made with GJL, the use of GJV allows for a weight reduction of up to 20%. In addition to vehicle-related areas of application, such as engine blocks, crankcases, cylinder heads, cylinder liners, exhaust manifolds, brake and clutch discs, there is a growing general interest in using GJV in other applications than those in automobile construction [Kloc01, Leng06, Röhr06]. The industrial implementation of GJV in the automobile industry is slow because of the uneconomical and difficult machinability of this material in comparison with cast iron with lamellar graphite. The main difference between the machinability of GJL on the one hand and that of GJV and GJS on the other is that, when machining GJL, wear-reducing manganese sulphide coatings form on the flank and rake faces of the cutting insert. When machining cast iron with vermicular and spheroidal graphite, no sulphur and no manganese sulphide is present and thus no wearreducing layer. This is because of the magnesium treatment necessary for the formation of graphite. With lower cutting speeds (vc < 200 m/min), at which, because of the low temperatures, no wear-protective coating develops when machining GJL, tool life travel path differences when machining different cast iron materials are essentially determined through differences in the mechanical characteristic values (hardness, tensile strength) as well as in the graphite morphology. An essential impact on tool life is made by alloying elements, e.g. titanium. The alloying element titanium forms titanium carbides which are harder than WC-Co cemented carbides and thus cause an increase in the abrasive wear of the tool. Studies have shown that doubling the titanium content reduces the tool life travel path by about half. In GJV types used most frequently at present, the titanium content is less than 0.015% [Kopp04]. When machining GJV, chip formation is discontinuous within a broad range of cutting speed, with the result of short-breaking chip forms. The inhomogeneity of

288


the material resulting from the different mechanical properties of the phases ferrite, perlite and precipitated graphite encourage this form of chip development [Röhr06]. Among the cutting tool materials for machining cast iron with vermicular graphite, coated cemented carbides (HC) and, in particular, aluminium oxide ceramics (CA) show great potential for machining GJV. When comparing the tool life travel paths of HC and CA at different cutting speeds, however, one can see that each of the two cutting tool materials is suited to a respective cutting speed range. Coated cemented carbide is suitable for machining at conventional cutting speeds, while aluminium oxide ceramics is the preferable cutting tool material for highspeed machining. Because of the dynamic excitation caused by the discontinuous chip formation, tougher WC-Co-based cutting tool materials can better compensate for the high stress reversals in lower cutting speed ranges than ceramic cutting tool materials. At higher temperatures, however, the hot hardness and the chemical resistance of the cutting tool material become decisive factors, giving CA tools good usage behaviour at high cutting speeds and temperatures at which HC tools become weak and thus fall victim to increased chemical and abrasive wear [Röhr06]. Generally, because of their high hardness and wear-resistance, PCBN cutting tool materials allow for significantly cutting speeds and a correspondingly higher productivity compared to cemented carbides. Even when using PCBN as cutting tool material, because of the absence of an MnS protective layer when machining GJV, the optimal cutting speeds for machining GJV lie at an approximate value of 300 m/min, clearly lower than those used for GJL (up to 1500 m/min). If high cutting speeds are used (500–1000 m/min), significantly reduced tool live travel paths can be expected compared to GJL [Leng06, Reut02]. Initial studies have shown that PKD appears to be an interesting cutting tool material, one with which the tool lives achieved when machining GJV are comparable to those for GJL. In comparison to cemented carbides, PKD has the capacity to machine workpieces at twice the speed and achieve a 10- to 20-fold increase in tool life in the process. The machining temperatures, however, must be kept under a certain limit value in order to prevent a sudden change from mechanical to thermal wear. It is recommendable to use compressed air or minimum quantity lubrication as a coolant/lubricant. PKD tools have a great potential for milling applications with GJV, with the cutting temperatures generally much lower than with continuous cutting. Since PKD cutting tool materials have a higher fracture resistance and strength than ceramics, it is recommendable to apply neutral to positive geometries to PKD tools. This can reduce the resultant force and contact zone temperature, which in turn allows for the application of higher cutting speeds [Pret06].

7.5.2.3 Cast Iron with Spheroidal Graphite In cast iron with spheroidal graphite, also referred to as spheroidal cast iron, graphite is present in the form of spheroidal inclusions. In comparison to other cast irons with graphite inclusions, the base mass is less often interrupted by the graphite spheres and the internal notching is almost nonexistent. This increases the strength, with spheroidal cast iron reaching a high ductility [Herf07b].

7.5


289

The mechanical properties of this material, such as tensile strength and toughness, are determined by the ratio of ferrite to perlite in the matrix. Types with low strength and good toughness qualities (e.g. GJS 400-15) consist predominately of ferrite. High strength in combination with low toughness is typical for types with a predominately perlitic matrix (e.g. GJS 600-3). This ratio of ferrite to perlite is the result of the amount of bonded carbon in the matrix, which can be altered by means of a heat treatment. The machinability is also influenced by the perlite-ferrite ratio in the material matrix. With an increasing perlite content, the strength of the cast iron rises along with the abrasive tool wear. The result of this is low tool life [Stau84]. Particularly suited to machining cast iron with spheroidal graphite are uncoated cemented carbides (HW), coated cemented carbides (HC) and oxide ceramics (CA) of the main application group K. The resultant forces and the resultant effective mechanical stress of the tool cutting edge are relatively low when machining the GJS cast iron. Interrupted cuts can be used without difficulty because of the good damping property of the material. Increasing the cutting speed lightly reduces the values of the individual resultant force components. Surface roughnesses of just under Ra = 1 μm are achieved in the finishing process. On a microscopic level, the surface is characterized by disembedded and in part, lubricated graphite inclusions. This effect only marginally degrades the measured roughness values. Chips formed when machining cast iron with spheroidal graphite are, first and foremost, segmented chips. Only with very sharp cutting edges does a continuous chip develop which turns into a segmented chip given even a small cutting edge fillet. Helical chips may form which are slightly brittle because of the reduction in chip strength due to graphite inclusions. Because of the high temperatures in the contact zones between the workpiece and the tool, the material plasticises and is discharged between the workpiece cut surface and the flank face or between the lower side of the chip and the rake face. This phenomenon, referred to as the growth of built-up edge fragments, can occur when dry cutting at high cutting speeds. In comparison to steels of the same hardness and similar strength, cast iron with spheroidal graphite must be machined at somewhat lower cutting speeds, especially when fine-machining with cemented carbides. Higher feeds, optimized tools and fewer steps as a result of a smaller material allowance can compensate for this disadvantage [Schm07]. A special heat treatment in the bainite stage can produce an austenitic/ferritic (also called ausferritic) matrix in cast iron with globular graphite. The term ADI (austempered ductile iron) is common for this. Figure 7.26 shows the structure of ADI. It consists of acicular ferrite and a high carbon-containing stabilized austenite. The material was erroneously referred to as bainitic cast iron in older publications due to the visual similarity of its structure with the ferritic carbide structure of steel. In addition to its increased strength and hardness, the matrix is characterized by increased toughness (e.g. GJS 900-7). The special combination of abrasive and adhesive wear behaviour when machining ADI leads to extensive crater wear near the cutting edge, which leads in turn

290

7 Tool Life Behaviour 50 µm

Spheroidal graphite Austenitic-ferritic matrix

10 µm

Acicular ferrite

Initial grain size

Rest austenite

Stabilized austenite

Fig. 7.26 Crystalline structure of austenitic-ferritic cast iron with speroidal graphite (ADI)

to a destabilization of the cutting edge. Therefore, the weak edge breaks before the critical flank wear will be reached. The use of a cutting fluid is recommendable for continuous cut operations. Mostly, coated cemented carbide tools of the main user group K, cBN and cutting ceramics are used as cutting tool materials [Kloc03]. The average forces arising when machining ADI are comparable to those of other cast irons with globular graphite. Due to its particular finely-striped grain structure, the average cutting forces are superimposed with a high dynamic. Surface quality is comparable to that of conventional cast iron with globular graphite. In the rim zone, the austenite converts into martensite because of mechanical stress, which is accompanied by an approximately 200–300 HB increase in hardness. Segmented chip formation is similar to that of conventional cast iron with globular graphite. ADI’s high fracture strain has a negative effect on this [Klöp07]. ADI is basically not more difficult to cut than higher-strength GJS types. However, cutting strategies specially adjusted to the material are required. Lower cutting speeds due to high amounts of heat release can be compensated by increased feeds. Tool geometries specially tuned to ADI machining, such as drills with a radius or facette shape have a longer tool life. The use of cutting fluids is beneficial in continuous cut operations. For discontinuous processes such as milling, cutting fluids are not recommendable due to thermoshock stress on the tool [Schm07].

7.6 Machinability of Non-ferrous Metals 7.6.1 Aluminium Alloys Mentionable areas of application of aluminium alloys as constructional materials include the following industries:

7.6

• • • • • • •

Machinability of Non-ferrous Metals

291

automotive engineering aerospace engineering installation and apparatus engineering electrical engineering food technology chemical industry optical industry

From the realm of aeronautic engineering, the example of a modern commercial aircraft, which consists of up to 65–81 mass percent aluminium, makes it clear how many aluminium components are used [Star96, Töns01]. Aeronautic engineering traditionally uses a large amount of aluminium components, which are often characterized by a large percentage of machined semifinished products. In the chipremoving manufacture of structural components, this percentage can be in the order of 90–95% [Schu96, Berk01]. Aluminium alloys are distinguished by a low density of ρ Al = 2.7 kg/dm3 but also by low strength. In order to expand the use of aluminium alloys, obviously their strength must be increased. Alloy technology makes this increase possible with the main alloying elements manganese, magnesium, silicon, zinc and copper. A distinction is made between wrought alloys and casting alloys. Alloys of both groups can exist in both a self-hardening or hardened state. In the case of wrought alloys, the semifinished products are manufactured by forming. Here, the properties of the formed structure are decisive for machining. In the case of hardening (hot age hardening or room temperature precipitation hardening), strength is increased by depositing hard structural components, which generally are located on the grain boundaries. The intermetallic phases should also be mentioned as another important group, the properties of which are very different from those of its components. Materials whose structure primarily consists of intermetallic phases are often abbreviated simply to “intermetallic phases” in common usage. NiAl, Ni3 Al and TiAl are key examples. The machinability of aluminium alloys depends on their composition and structural state. Compared with cutting steels, boundary surface temperatures are much lower, around 350◦ C. Due to the low melting point of aluminium alloys, one must take care that the boundary surface temperature does not get too close to the melting point. The boundary surface temperature is very much contingent on the tribological conditions. Because of its low strength compared with steel, lower mechanical and thermal stresses on the cutting edge are to be expected for the same cutting parameters. The low boundary surface temperature makes it possible to use higher cutting parameters than in steel machining, most importantly higher cutting speeds, which are delimited from below by adhesive wear and from above by temperature resistance of the material. The tool life parameters of machining tools used on aluminium alloys vary depending on the alloy and the structural state. Generally, more favourable tool life parameters can be expected than in steel machining. Adhesion is mostly dominant in

292


the lower cutting speed range, although, as mentioned already in Chap. 3, abrasion is always active as well, increasing especially with a larger distribution density of hard particle inclusions. Such particles can be intermetallic compounds and nonmetallic inclusions such as impurities of the molten bath. Hardened wrought alloys and casting alloys with a silicon content of up to 12% cause increased tool wear with increasing amounts. Due to the higher boundary surface temperatures, plasticized material can escape between the cut surface and the flank face and/or between the chip and the rake face. This phenomenon contributes to the deterioration of surface quality. The cutting speed must be reduced in these conditions. Cutting tools made of uncoated cemented carbide (HW) and diamond (DP) are the most commonly used cutting materials for machining aluminium alloys. Cemented carbides are used as cutting tool materials for machining wrought alloys and sub-eutectic casting alloys due to their wear resistance and hardness. In contrast to diamond, here the main stress is on toughness and the ability to manufacture complex tool geometries such as highly twisted end mill cutters with sharp edges. Diamonds, in polycrystalline form (DP) and in monocrystalline form (DM), are the first choice for machining strongly abrasive super-eutectic casting alloys. For all aluminium alloys, the use of cutting tool materials of the material group DP is not to be recommended for drilling into solid blocks due to compression processes in the area of the chisel edge. In the case of boring, especially aluminium alloys with high amounts of silicon, cutting tool materials of the material group DP are superior tot hose of the HW group with respect to tool life parameters and material removal rate. All cutting tool materials for main application group K can be used sensibly. Selection is based on general criteria such as cutting speed, cross-section of undeformed chip, and continuous or interrupted cut [Bech63, Opit64b, Zoll69, Bömc87]. Cutting trials have shown that CVD diamond coatings on cemented carbide substrates have the potential to combine the advantages of cemented carbides with those of polycrystalline diamond. When machining the wrought aluminium alloy AlCu4Mg1 (2024), the CVD diamond coating makes it possible, for example, to improve the tool life in comparison to uncoated tools. In interrupted cut with low cutting speeds and high material removal rates, high-speed steel (HSS) can be used advantageously to machine alloys with small amounts of silicon. The specific resultant force of the types AlMg5 (5019), AC-AlSi6Cu4 (AC-42000) and AC-AlSi10Mg (AC-470000) is about 25% below those of the heat-treated steel C35. Diamond cutting tool materials are often used to create highly reflective surface properties. Generally, these processes use high cutting speeds and small crosssections of undeformed chip. The surface quality depends to a great extent on the wear mechanism of adhesion, which manifests itself in built-up edges and can be influenced by the process kinematics. Adhesion prevents the attainment of an optimal surface quality. When there is a high material removal rate in an aluminium alloy machining operation, a large amount of chips has to be removed. For the sake of undisturbed manufacture, the chip form is a particularly important criterion for judging machinability.

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293

The chip form can be affected by the alloy components, by heat treatment and by the process kinematics. When machining non-hardenable and hardenable alloys in a soft state without any percentage of silicon, long ribbon chips arise which make the machining process difficult. These alloys should be avoided as much as possible for components requiring machining [Bech62]. In hardened wrought alloys and casting alloys with a silicon content of up to 12%, increased silicon content leads to a more advantageous chip breakage. Hard and brittle inclusions such as Al2 O3 and silicon also benefit chip breakage. In summary, it can be said that cutting aluminium alloys forms the more favourable chips the harder they are. The most difficult to cut are especially nonhardenable aluminium alloys and hardenable aluminium alloys in a soft state. For this reason, it is recommendable if possible to cut these alloys in a condition of increased strength (cold-twisted/hardened) [John84].

7.6.2 Magnesium Alloys Magnesium alloys have been used in Germany since the end of the 1930s especially in automotive engineering as light constructional materials in Volkswagen manufacture. Approximately 80% of the cast pieces produced in Germany currently consist of magnesium alloys [Wolf03]. Examples include transmission housings, steering parts, seat frames, inner linings, instrument supports and engine mounts in magnesium-aluminium composite structures [Dörn04]. Other products using magnesium include manual electronic devices, laptop housings, mobile phones and other devices in the area of consumer electronics as well as machine parts and sports articles. In comparison, sand casting is of only minor importance. The same is true of products made of magnesium wrought alloys. It has become clear that magnesium alloy machining is focused primarily on die-cast components. The low density of ρMg = 1.74 kg/dm3 in comparison with ρAl = 2.7 kg/dm3 and steel with ρsteel = 7.8 kg/dm3 has predestined magnesium alloys for applications in lightweight construction in particular. Alloy engineering aims to improve the strength properties of magnesium, especially for applications at temperatures above ca. 100◦ C. Another important goal of alloy technology is the improvement of the corrosion resistance of magnesium alloys, with respect to both inner corrosion and contact corrosion. The most important alloying elements include aluminium, zinc, manganese, silicon and rare earth elements such as cerium, lanthanum, neodymium, praseodymium and yttrium. Of the many types of magnesium alloys, the following will take a closer look at the example of one alloy group. The alloy AZ91hp is the most frequently utilized magnesium alloy for die-cast parts. The notation “hp” (high purity) means that only trace elements of iron, copper and nickel are found. These elements cause inner corrosion in magnesium alloys. AZ91hp is characterized by good mechanical properties and very good castability [Kamm00, DINEN1753]. Solidifying the alloy in a temperature range of 465–598◦ C favours inhomogeneous crystalline structures in the cast piece and the

294


formation of porosities. This can impair the density of cast pieces, especially after machining. Magnesium-aluminium-manganese alloys (AM-alloys) have very good strength properties together with higher fracture strain and notched-bar impact work. The corrosion resistance of alloys of this AM group is better than that of group AZ. The alloying element manganese bonds corrosive elements like iron, copper and nickel. Manganese and aluminium combine to form hard MnAl deposits, which accelerate tool wear. Because of their good castability, AM-alloys are used in automotive engineering, e.g. for structural parts, steering parts and rims. AS-alloys (magnesium-aluminium-silicon) have similar mechanical parameters as AM-alloys. They are inferior however with respect to toughness. Their particular advantage is improved creep strength for temperatures up to about 150◦ C. Further alloy systems that aim to improve mechanical properties and creep strength include: • • • •

Magnesium-zinc-RE-Zr (ZE-group) Magnesium-RE-Ag-Zr (QE-group) Magnesium-yttrium-RE-Zr (WE-group) Magnesium-aluminium-RE (AE-group)

The rare earth elements (RE) include cerium, lanthanum, neodymium, praseodymium and yttrium. The alloy AE42 contains RE as a misch metal with the composition 50% cerium, 25% lanthanum, 20% neodymium and 3% praseodymium. With the exception of higher fracture strain and higher creep strength, it has similar attributes as the alloy AS41. In order to classify material properties into a material spectrum, fracture strain at room temperature is a helpful mechanical parameter (A = 7−14% for magnesium alloys). Beyond a certain temperature, about 225◦ C in the case of magnesium alloys, material behaviour shifts from brittleness to toughness. This phenomenon can be explained by the material structure. Magnesium has a hexagonal lattice structure at room temperature. Only the base planes act as slide planes at room temperature. Above ϑ = 225◦ C, it changes into a cubic lattice structure. This increases deformability. When cutting magnesium, lamellar chips are formed (Fig. 7.27). The distance between the lamellae or frequency of chip formation depends essentially on the tribology of the boundary surface, which can in turn be affected by feed and cutting speed. Characteristic of magnesium alloy cutting is the marked stick-slip friction, which is indeed contingent in a certain way on the material-cutting tool material combination, but can also be influenced by process kinematics. This means that a smart choice of process kinematics can adjust the amplitude and frequency of dynamic stress on the cutting edge to its loadability. Near the tool-chip boundary surface, friction causes temperatures to increase, plasticizing a thin material layer which holds the individual chip lamellae together. Sharp tool cutting edges and smooth rake faces lead to good machining results. The tool geometry is based on those of tools for machining aluminium (Sect. 7.6.1). In drilling and milling with small tool diameters, the maximum spindle rotation speed puts a limit on the cutting speed. When rotating tools with large diameters are

7.6

Machinability of Non-ferrous Metals Chip bottom side

Material: Cutting tool material: Tool geometry:

AM60 HC-N10 (CVD-Diamant) DCGX11T308-AL

295 Chip upper side

Cutting speed: Feed: Depth of cut:

vc = 1000 m/min f = 0.5 mm ap = 1.0 mm

Fig. 7.27 Chip of the magnesium alloy AM60

used, the maximum allowable tool rotation speeds are the limit criteria. Workpiece stiffness can limit the chip cross-section and material removal rate, especially in finishing. With respect to machinability, magnesium alloys are characterized by the fact that they contain few abrasive components. This is also true for the rim zones of the pieces to be machined, since these are predominately manufactured by die casting. When cutting magnesium alloys have only a slight tendency to adhesion, so no builtup edge formation is to be expected. The melting point of this alloy is in the range of 420–435◦ C. This makes it clear that the thermal load on the tool is relatively low in machining magnesium alloys. Fitting cutting tool materials for machining include high speed steel (HSS), uncoated and coated ultrafine-grain cemented carbide (HF, HC), polycrystalline diamond (DP) and diamond-coated ultrafine-grain cemented carbides. In practice, ultrafine-grain cemented carbides of the application group N10/20 and polycrystalline diamond (DP) are the usual cutting tool materials. Tools made of these materials allow for high cutting speeds and feeds. They are characterized by especially high wear resistance and contribute to a great extent to process safety. The resultant force is relatively low in magnesium machining, below those of sub-eutectic aluminium alloy machining. Figure 7.28 compares experimental results from dry face milling aluminium alloy AlSi9Cu3 and magnesium alloy AZ91hp [Kloc00]. Both alloys are used, for example, for transmission housings. When drilling both of these alloys with tools made of cemented carbide, the resultant force of the magnesium alloy is also below that of the aluminium alloy. Differences in cutting momentum are marginal [Wein00]. The possibility of machining magnesium alloys with high and extremely high cutting speeds is a good foundation for creating high surface qualities on machined components. Milling produces satisfactory surface qualities both in wet cutting when oil or emulsions are used and in dry cutting. In the case of the cutting

296

7 Tool Life Behaviour 200 AlSi9Cu3

Resultant force Fz / N

AZ91hp 150

Hardness of materials: AlSi9Cu3 102HBW 5/250 AZ91hp 67HBW 5/250 vc = 2000 m/min ae = 30 mm ap = 3 mm

100

Tool: Face mill, D = 80 mm, z = 5 Down milling in dry cut Cutting tool material: DP

50

0

0

0.1 0.2 Feed fz / mm

0.3

0.4

Fig. 7.28 Comparison of force components during face milling (dry cut)

conditions given in Fig. 7.28, an arithmetical average roughness of Ra = 0.6 μm was measured after wet face milling. For seal faces of motor vehicle components, Ra -values of about 1.5 μm are often required. These values can be safely adhered to with feeds of fz = 0.3 mm [Kloc02]. In drill hole finishing, top surface qualities in the range of Ra = 0.2 − 0.4 μm can be safely obtained under production conditions, even when the highest cutting parameters are used [Wein00]. To achieve this, the process must be stable. Due to the formation of lamellar chips, longer chips are almost completely absent, mostly spiral chip segments and discontinuous chips occur. Longer chip pieces break into smaller pieces when they make contact with the tool, the workpiece and the cutting medium because of their low stability. In general, the chip form poses no problem in magnesium machining.

7.6.3 Titanium Alloys Titanium alloys are still a relatively recent group of construction materials. The first alloys were developed at the end of the 1940s in the USA, including the classic alloy TiAl6V4. With a density of ρ Ti = 4.51 kg/dm3 , titanium is the heaviest element in the light metal group. The great technical importance of titanium alloys is based above all on their high strength, but in particular on the ratio of yield strength to density, which is not even closely approached by any other metallic material (Fig. 7.29). Even highstrength steels with yield strengths of ca. 1000 Mpa are still little more than half that of a TiAl6V4 titanium alloy with respect to this ratio. Another important property of titanium alloys is their good corrosion resistance. They are resistant to temperatures of about 550◦ C. The maximum operating temperature is limited by the increase

7.6


297

Comparison of thermophysical and mechanic characteristics 451 442

Ti99.7G TiAl6V4G C45+QT

Specific heat [J/(100g·K)]

100−145

50 52 56.6 46


180

87 80

Ti99.7G TiAl6V4G

Heat dimension [µm/(10m·K)] 120 150

785


17.2 5.8

Thermal conductivity [W/(m·K)]

100

0

Young s modulus [GPa]

110−140 205

800−1100 370−490 40 55 0

200

400

Yield strength Rp0,2 [MPa]

Ratio of yield strength and density

215

C45+QT 50

Density [g/100 cm³]

600

800

Fig. 7.29 Mechanical and physical properties of titanium materials in comparison to the heat treatable steel C45+QT (reference data)

in oxidation beyond 550◦ C [Essl91]. Also worthy of mention is the excellent biocompatibility of titanium alloys. Titanium alloys are used above all in these areas: • • • • •

aerospace industry chemical industry medical technology power engineering sports equipment technology

Titanium appears in various lattice modifications, of which each is only stable within a certain temperature range. At lower temperatures, pure titanium and most titanium alloys exist in the modification of a hexogonally close sphere packing, which is called α-titanium. The high-temperature phase on the other hand has cubic body-centred crystal lattice and is called β-titanium. The conversion of the α-phase into the β-phase takes place above 822◦ C, the “transus temperature”. The existence of these two phases and the associated structural transformation is the foundation for the diverse properties of titanium alloys [Pete02]. In accordance to their influence on transus temperature, a distinction is drawn between neutral (Sn, Zr), α-stabilizing (AL, O, N, C) and β-stabilizing (Mo, V, Ta, Mn, Cr, Cu, Fe) alloying elements. The α-stabilizing elements – which include aluminium, the most important alloying element of titanium – extend the α-phase zone to higher temperatures and form a two-phase (α + β)-zone [Pete02]. The properties of titanium alloys are basically determined by their chemical composition and structure. The chemical composition primarily determines the parts by volume of the α und β phases. Structure is defined here in the case of almost

298


exclusively two-phase titanium alloys primarily as the size and arrangement of both of these phases. The two extreme forms of phase arrangement are the lamellar structure, which arises by simple cooling from the ß-zone, and the globular structure, which results from the recrystallization process. Both structural types can exist in a fine or coarse form. Basically, the always varying structure is created by means of thermomechanical treatment. This involves a complex sequence of solution heat treatment, deformation, recrystallization, aging and stress-relief heat treatments [Pete02]. Titanium materials can be classified in the following groups: • • • • • •

pure titanium α-alloys (α + β)-alloys near-α-alloys metastable β-alloys γ-titanium aluminide alloys

Pure titanium (e.g. ultrapure titanium: 99.98% Ti, pure Ti: 0.2Fe-0.18O or 0.5Fe-0.40O) and titanium alloys (e.g. Ti-5Al-2.5Sn, Ti-0.2Pd), which only contain α-stabilizing and/or neutral alloying elements, are designated as αalloys. Because they are single-phase, α-alloys are relatively low in strength (Rm = 280−740 N/mm2 ). They are primarily used in the chemical industry and process engineering, since what is of chief interest here is good corrosion resistance and deformability. Pure titanium types contain up to 0.40% oxygen, which, as an interstitial alloying element, drastically increases the yield strength. Whereas oxygen is the only element intentionally added in order to meet strength requirements, there are other elements, such as iron or carbon, which represent impurities caused by the manufacturing process. The high level of corrosion resistance is achieved by alloying palladium (up to 0.2%). These titanium types are generally used in low temperatures. If higher strength is required, titanium materials alloyed with aluminium and tin (TiAl5Sn2,5) are also available [Pete02, Schu04]. The microstructure of (α + β)-alloys is characterized by a bimodal structure. It consists partially of globular (primary) α-phase in a matrix of lamellar α- and β-phase. This microstructure combines the good properties of lamellar structures, such as high resistance to creeping and fatigue crack growth, with those of globular structures, such as higher strength and fracture strain. The (α + β)-structure is fabricated by thermo-mechanical treatment, consisting of a combination of forming and heat treatment. By hardening, (α + β)-alloys reach very high strengths in the range of Rm = 900–1300 N/mm2 . Moreover, these alloys are characterized by high temperature resistance. Among the (α + β)-alloys is the by far most common titanium alloy TiAl6V4 (Ti-6Al-4V). More than half of all titanium materials are melted with this composition. It was already developed in the early 1950s in the USA and is both the most researched and tried titanium alloy. It is used above all in the aerospace industry. Titanium alloys are used to manufacture guide blades, impeller blades, plates, housings, pipes and distance rings between the rotor stages [Adam98, Pete02, Schu04].

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299

Near-α-titanium alloys (e.g. Ti-6Al-2Sn-4Zr-2Mo-0.1Si) are the classic high temperature alloys, used in operating temperatures of 500–550◦ C. This alloy class combines the good creep properties of α-alloys with the high strength of (α + β)alloys. Also, they exhibit a favourable combination of mechanical properties and oxidation resistance. These alloys have become very important in engine construction. They make it possible to increase operating temperatures up to 600◦ C. While the standard (α + β)-alloy TiAl6V4 can be utilized only in the relatively cold temperature range of the compressor, near-α-alloys can also be used in hotter areas, e.g. in high-pressure compressors. Two essential representatives of this alloy group are the titanium alloys Ti-6-2-4-2 (Ti-6Al-2Sn-4Zr-2Mo) and IMI834 (Ti-6Al-4Sn3.5Zn). Both alloys have a bimodal structure (Fig. 7.30), however with a much higher percentage of α-phase compared with the standard alloy TiAl6V4. Metastable β-alloys (e.g. Ti-4, 5Al-3V-2Mo-2Fe, Ti-15V-3Cr-3Al-3Sn) contain vanadium, molybdenum, manganese, chrome, copper and iron. Vanadium and molybdenum form a continuous series of mixed crystals with titanium, which remain stable at low temperatures. The mixed crystals decompose with the other alloying elements eutectoidically at low temperatures. Metastable β-alloys can be hardened to extremely high strengths of over 1400 MPa. Their complex microstructure allows for an optimization of the ratio of high strength to high fracture toughness. These alloys are used to produce foils, aerospace components and automobile components. Their use is limited however by their larger specific weight, moderate weldability, poor oxidation properties and the complexity of their microstructure [Pete02, Schu04]. Another material group based on titanium are the titanium alumnide alloys (α2 -Ti3 Al, γ-TiAl) first developed in the 1970s. Titanium alumnides are counted among the “intermetallic” phases. These are defined as compounds of metals with metals or non-metals, which have a different lattice structure than the initial component, from which their specific properties result. Intermetallic phases of the TiAl Crystalline structure of TiAl6V4 Ti-6Al-2Sn-4Zr-2Mo: heterogeneous duplex Material

Primary α -phase, hexagonal, brittle, bad deformable, tends to strain hardening

β -phase, bodycentered cubic, good deformable, distinctive adhesion tendency White needles: secondary α -phase Melting point: 1677 °C

100 µm

Fig. 7.30 Typical (α+β) mixed structure as in a compressor blade made of the near-α-alloy Ti-62-4-2, acc. to ADAM [Adam98]

300


type are characterized by low density, high high-temperature strength, high oxidation resistance and creep strength up to temperatures of 650◦ C. They thus are more temperature-resistant than the above-mentioned titanium alloys and are in direct competition with established nickel alloys with high density, which can be used in temperatures up to about 700◦ C [Pete02]. γ-titanium alumnide alloys are also characterized by their very low fracture strain at room temperature, in the area of A = 0.5–1%. Titanium alloys are difficult to machine because of their mechanical and physical properties. Their strength is high, and their fracture strain (A5 = 5–15%) is low. Their Young’s modulus is almost half that of steel. The hexagonal α-phase is relatively hard, brittle, poorly deformable and has a high tendency towards strainhardening. This phase affects the active tool cutting edge like the strongly wearing cementite lamellae in the perlite grains of carbon steels. The cubic body-centred ßphase is very similar in its machinability to ferrite, which also crystallizes into the krz lattice type: it is easily deformable, relatively soft and ductile and has a high adhesive tendency (Fig. 7.30). One important physical property for the machinability of titanium alloys is their low thermal conductivity, which is only about 10–20% that of steel (Fig. 7.29). As a result of this, only a small amount of the arising heat is removed with the chip. In comparison to machining the steel material C45E, about 20–30% more heat must be absorbed by the tool, depending on the thermal conductivity of the cutting tool material, when the titanium alloy TiAl6V4 is machined (Fig. 7.31). The result of this is that the cutting tool is subjected to high thermal stress, much more than is the Wear relevant characteristics of TiAl6V4

80

1300 TiAl6V4

Chip

60

C45E

40 Tool 20 0

Thermal conductivity λ / W/(m·K) 42 HW-P10 HSS Ceramic

84

126

HW-K10

Comparison of tool and chip heat distribution while turning steel and TiAl6V4

168 PKD

Tool temperature t / °C

Heat flow ratio αw / %

100

1100

TiAl6V4

900

f . ap = 0.25 . 1.5 mm²

700 500

C45E

300

f . ap = 0.21 . 2 mm²

100 0

20

40 60 80 100 120 Cutting speed vc / (m/min)

Comparison of tool temperature while turning steel and TiAl6V4

Fig. 7.31 Heat distribution coefficient of heat flow on the rake face and the thermal tool stress during turning of TiAl6V4 compared to C45E steel, acc. to KREIS [Krei73]

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301

case when cutting steel (Fig. 7.31, right). This means that cutting titanium alloys not only exposes the cutting tools to considerable mechanical strain, but also to extremely high thermal stress [Krei73]. Another characteristic of titanium alloy machining under conventional cutting conditions is the formation of lamellar chips. The cause of this is a constant shift between compression and sliding phenomena in the shear zone (Fig. 7.32). In position I, the shear zone is already fully formed. The lamella slides over the rake face while a new lamella is compressed. Since the deformation resistance of the titanium material is rapidly decreased immediately after the shear zone is formed as a result of the high shear/deformation speed, the cutting force is steadily reduced. The simultaneously introduced compression of the newly forming lamella causes the cutting force to go up again (position II), until the cutting force or shear force is so high in position II that the shear strength of the material is exceeded and a new lamella is formed [Krei73]. Such discontinuous chip formation subjects the tools to a mechanical and thermal alternate load, the frequencies and amplitudes of which depend directly on the cutting conditions. The dynamic cutting force can amount to about 20–35% of the static cutting force. Mechanical and thermal alternating stress can lead to tool fatigue and encourage tool failure due to the formation of cracks,

Cutting force Fc

I

II

Work piece

Work piece vc

III

Chip Tool

vc

Work piece Chip

vc

Tool

Cutting time tc

Fig. 7.32 Chip formation during turning of titanium, acc. to KREIS [Krei73]

Chip Tool

302


shell-shaped spalling, the fracture of cutting tool material particles or cutting edge fracture [Krei73, Yang99, Kita97]. Titanium materials are generally turned with uncoated cemented carbides of the main application group S. Due to the high thermal and mechanical load, the tools can however only be used at relatively low cutting speeds. The range of usual cutting speeds is about 50–70 m/min for rough turning and 60–90 m/min for finish turning. Cemented carbides containing titanium (P-types), cermets or coated cemented carbides are generally not suitable for turning titanium alloys. Elements of the substrate (Ti) or the finishing material (Ti, O, N) react with the workpiece material, which greatly reduces the wear resistance of the tools. In isolated cases, good results have been reported in the case of machining with TiB2 -coated tools. Cutting tool materials based on Al2 O3 and Si3 N4 do not come into considerations for machining titanium materials due to their low thermal conductivity and the great affinity of aluminium, silicon, oxygen and nitrogen to titanium. One alternative to uncoated cemented carbide tools in the case of finish turning titanium alloys are tools made of monocrystalline diamond, polycrystalline cubic diamond (PCD), CVD-diamond thick films and polycrystalline cubic boron nitride (PCBN). These cutting tool materials are characterized by high hardness and wear resistance, excellent thermal conductivity compared to other cutting tool materials (Fig. 7.31), low thermal expansion and low friction between the rake face and the chip (or between the flank face and the workpiece). Compared with cemented carbides, they make it possible to use higher cutting speeds, thus clearly reducing production time while maintaining the same or even improving cutting quality. The range of speeds applicable for tools made of these cutting tool materials is vc = 100−200 m/min. Of the PCBN cutting tool materials, above all type containing high level of cBN are good for finish turning several types of titanium alloys. When turning titanium alloys with PCD tools, the interactions that effect wear taking place between the workpiece material and the cutting tool material are extraordinarily complex. They are characterized by diffusion and graphitization, thermally caused cracking, surface damage as a result of lamellar chip formation and the potential formation of wear-reducing reaction films on the diamond grains [Bömc89, Neis94]. Due to these diverse interactions between the workpiece and cutting tool materials, the performance capacity of PCD cutting tool materials in the case of titanium machining is highly dependent on the composition of the cutting tool material. Especially mentionable in this context are above all the composition of the binder phase, its quantitative amount as well as the size of the diamond grains [Neis94]. The most dominant form of wear when turning titanium alloys with PCD tools is the formation of craters on the rake face. Flank face wear is of secondary importance, especially at high cutting speeds. In experiments where the titanium alloy TiAl6V4 was cut using external cylindrical turning, the lowest amount of crater wear was measured in the case of a PCD variety with SiC as binder. In the case of PCD types with cobalt-containing binders, it was seen that crater wear was greatly effected by the binder content and the grain size. The largest amount of crater wear

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303

was observed when using the PCD type with the largest quantity of binder and the smallest grain size [Köni93c, Neis94, Kloc07]. Due to its catalytic effect, cobalt promotes the graphitization of diamond. The result is that the cutting tool material becomes less resistant to abrasive wear. Furthermore, cobalt and diamond have different thermal expansion coefficients, which favours the formation of microcracks. This process can be observed above all in the case of fine-grain types. In conjunction with dynamic stress on the cutting tool material by lamellar chip formation, microcracks further the separation of individual diamond grains or entire grain bonds from the cutting tool material compound [Bömc89, Köni93c, Neis94] (Fig. 7.33). The lower amount of crater wear when turning with SiC-containing or coarsegrain cobalt-containing PCD is attributed in the literature to the formation of a wearinhibiting film made of titanium carbide on the diamond grains. It is believed that there is a reaction caused by diffusion between the titanium from the material and the carbon from the cutting tool material in the area of the crater. The TiC film that is formed in the area of the contact zone on the rake face adheres tightly to the diamond grains that make up the surface during the machining process. Since the speed of diffusion of carbon to titanium carbide is more than 10 times slower than that of carbon to titanium, the wear process is significantly reduced [Hart82]. Thus, to reduce crater formation when turning titanium alloys with PCD as much as possible, PCD cutting tool materials with large diamond grains, low amounts of cobalt or with a binder phase made of β-SiC should be used.

Width of flank wear land VB/µm Crater wear KT/µm

100 80

Width of flank wear land VB

PCD 1 PCD 3

60 PCD 2 40 Crater wear KT

20

PCD 3 PCD 2 PCD 1

0 200

600 800 400 1000 Path length due to primary motion lc/m

Composition

PCD 1

PCD 2

PCD 3

Binder phase: WC-Co * Volume percentage

Diamond *

92 %

92 %

80 %

Face turning, TiAl6V4 vc = 110 m/min ap = 2.0 mm f = 0.1 mm

Binder * Grain size

8%

8%

20 %

6–10 µm

2–6µm

0.5–1 µm

Fig. 7.33 Diamond as cutting tool material for turning of titanium alloys

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Weight reduction is an essential criterion for modern airplane engines. This can be realized by using materials of lower density, but also with the help of an innovative component design. One example of a modern component concept is the blade integrate disk or “blisk”. To manufacture a blisk, the blades must be carved from the solid material. This is an extremely demanding machining task, especially considering that the blades must meet the highest demands regarding surface quality, rim zone formation and formal/dimensional accuracy. For blade lengths up to 80 mm, blisks are produced using end milling or form cutting. For blisks with a diameter of 600 mm and about 70 blades, up to 55 h of pure milling time is required. Cutting becomes progressively more difficult with increasing blade lengths and thus with increasing milling cutter protruding lengths. In such cases, ECM machining is preferred, though all the blades must first be premilled to an overmeasure of 3 mm. Here too, machining also requires up to 45 h depending on the size of the blisk, as well as 7 or 8 more hours should an ECM operation be included. Blisk production has thus proven to be very time-consuming and cost-intensive. Blisk pre-milling is done with end milling cutters made of high performance high speed steel or conventional cemented carbides. Due to the large volume of material to be removed, tools with relatively large axial (D) and radial (D/2) depths of cut are used to realize large material removal rates. On the other hand, in order to finish the blades only very small overmeasures need to be removed. The main concern in this case is above all high process safety, surface quality as well as form and dimensional accuracy. For this type of machining, tools made of ultrafine-grain cemented carbide is recommendable due to its excellent wear and toughness properties. Because of their high level of wear resistance and bending strength, higher speeds and larger material removal rates can be realized with end milling cutters made of ultrafine-grain cemented carbide than with ones made of conventional cemented carbide. In the example provided (Fig. 7.34), it was possible to reduce the total machining time required for blade finishing by about 50% by using ultrafine-grain cemented carbide milling cutters in conjunction with an adapted machining strategy. Cutting blades with these tools at high speeds lead moreover to still further advantages, the most important of which include improved surface quality of the milled blades and lower resultant forces, which make it possible to realize higher levels of component precision. Trochoidal milling is one way to reduce milling time when pre-machining blisk blades (Fig. 7.47). Using this method, even curved blade surfaces can be manufactured in one cycle to nearly the final contour with end milling cutters on 5-axis machine tools by superimposing a wobbling motion. Analogously to the remarks given above on milling nickel-based alloys (Sect. 7.6.5), the formation of the cutting edge has a key role with respect to tool wear when milling titanium alloys as well. Here too, results from research and industrial praxis have shown that stabilizing the cutting edge with a defined rounding in the order of 5–20 μm leads to a significant increase in tool life not only for long-protruding tools when end milling titanium alloys. When milling materials that are difficult to machine, of which titanium alloys are an example, down milling should generally be given preference. In up milling, the

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Blisk production: Working time with conventional milling up to 55 h

Material: TiAl6V4 Tool: End mill, D = 8 mm

Finishing

Conventional Conventional milling 100 % Planing annular space

Finishing

High finish milling Performance Carbide S10/20 milling vc = 100 m/min fz = 0.045 mm ap = 0.80 mm ae = 0.30 mm

High performance finish milling

45 %

Micrograin carbide S10/20F vc = 350 m/min fz = 0.045 mm ap = 0.80 mm ae = 0.30 mm

Rough milling

Roughing

Working time

Carbide S10/20 vc = 100 m/min fz = 0.045 mm ap = 4 mm (D/2) ae = 8 mm (D)

Fig. 7.34 Reduction of working time by applying high performance milling during finishing a blisk made of the titanium alloy TiAl6V4 (Source: MTU)

tool enters with a chip thickness of h = 0 mm. Due to the high elastic deformability of titanium alloys, chip formation is preceded by a long friction phase between the tool and the workpiece, which greatly promotes wear on the flank face. The tool exits with h > 0 mm except for the fluting from the solid. Chip root [Peke78] and tensile residual stresses can develop in the cutting edge as a result of this. Tensile residual stresses can lead to crack formation and then to local cutting tool material fracture and tool failure. Chips adhering to the cutting edge are partially compressed upon re-entry of the cutting edge into the material on the up milling flank, or they wind up between the cutting edge and the material. Faulty component surfaces and/or cutting tool material fractures on the cutting edge are the result. In down milling on the other hand, the cutting edge enters with h > 0 mm and exits with h = 0 mm. Chip root formation is not possible when the tool exits the material. Potentially adhering chips are only connected with the cutting edge by a thin strip of material and are usually wiped from the tool upon its re-entry. The resultant surface quality of the down milling flank is much better than that of the up milling flank. The special characteristic of down milling is that the cutting edge exits with a chip thickness of h = 0 mm. This is of great importance when milling grooves, also during the phase in which the tool enters the workpiece. If the tool enters in a straight line, the cutting edge always exits with h > 0 mm until the groove is fully cut, such as in up milling. If the tool enters on the other hand with an “arc lead”, the

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7 Tool Life Behaviour 400

Work piece

Straight lead Circular-arc lead

350 Circular-arc lead

VBmax/µm

Straight lead

A M

M

300 250 200 150 100 50 0

8

16 24 32 40 48 56

64

Number of leads h>0

End mill

h=0

h : cross-section of undeformed chip at tool exit while lead

Material: Cutting tool material: Cutting parameters:

Process: Tool diameter:

TiAl6V4 HW-K40 ap = 5 mm vc = 50 m/min fz = 0.17 mm end milling D = 25 mm

Fig. 7.35 Influence of the entering path on tool wear during milling the titanium alloy TiAl6V4

cutting edge will exit the material with h ≈ 0 mm, analogously to down milling. Both first cut strategies have a significant effect on the wear and performance of end milling cutters. Compared to the conventional lead, the arc lead results in much lower tool wear (Fig. 7.35). In contrast to turning, coated tools can also be employed for many machining tasks when milling titanium alloys. Causes of the good wear and performance properties of coated tools in milling could include the lower tool temperature caused by the interrupted cut and the compressive residual stresses characteristic of PVD coatings. Because of the high thermal stress on the tool, titanium alloys are usually machined using wet cutting. Due to the long engagement times, an intensive cooling is necessary, especially when turning. One extremely effective method in this context is supplying a cutting fluid under high pressure (p > 80 bar). In experiments in turning the titanium alloy TiAl6V4 with uncoated cemented carbides, the cutting fluid was supplied both conventionally with a pump pressure of 6 bar and under high pressure with 140 bar (Fig. 7.36). Under the conditions of high pressure lubricoolant supply the resultant tool life was higher by a factor of 2.3 (VB = 0.3 mm) compared with conventional cutting fluid supply. As this example shows, a cutting fluid jet supplied with high pressure can be used to improve not the tool life, but above all the cutting speed and thus the process’s productivity [Gold07]. The cutting fluid jet supplied to the gap between the chip bottom side and the rake face not only cools the tool intensively, it also improves chip fracture. Instead of long ribbon and snarled chips, short-breaking chips are formed under the selected conditions, the removal of which from the chip formation location and the machine tool no longer presents any problems. Especially in grooving operations, high pressure lubricoolant supply

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VBmax /µm

500 400

pKSS = 6 bar

300 200

pKSS = 140 bar

100

Material:

TiAl6V4

Cutting tool material:

HW-K10/20

Cutting speed:

vc = 90 m/min

Feed:

f = 0.4 mm

Depth of cut:

ap = 0.5 mm

0 0

5

10

15

Cutting time tc / min pKSS = 6 bar

pKSS = 140 bar

tc = 4 min

tc = 14 min

Fig. 7.36 Increase of tool life due to high pressure lubricoolant supply

can lead to enormous improvements in chip formation and tool performance. One method variant is supplying the cutting fluid through the insert. The cutting fluid exits near the cutting edge in the area of the flank or rake face. The use of the high pressure lubricoolant supply is not limited to turning. It can also be used for drilling and milling operations. In the case of drilling, the cutting fluid is supplied via the spindle and the cooling ducts in the tool. The use of the high pressure lubricoolant supply when drilling into titanium-based or nickel-based alloys increases performance significantly. Compared with conventional supply, the resulting tool life is many times higher. While in the case of internal supply via the machine spindle, the pressure applicable through the pivoting feedthrough is currently limited to about 140 bar, in the case of external supply in turning or milling, the cutting fluid can be supplied with much higher pressure. At present, high pressure units with pressures of up to 1000 bar are available [Fili02]. Whether it is technologically necessary and economically sensible to work with such high cutting fluid pressures is currently being researched. As the first results have shown, cutting fluid pressures in the range of 100–200 bar should be sufficient for most machining tasks. In high pressure lubricoolant supply not only the pressure is important, but also the volume flow. The latter must also be optimally adjusted to the respective machining task. Analogously to turning, cutting fluids are regularly used when milling titanium alloys as well. As in the case of milling steel materials, the basic problem is that the highly heated cutting edge emerging from the material is abruptly cooled down by the cutting fluid. The thermoshock caused by this promotes the formation of comb cracks and hence tool wear. In unfavourable cases, comb cracks, in conjunction with mechanical shock load of the cutting edge when entering the material, can lead to

308

7 Tool Life Behaviour Down milling with ae < D

Width of flank wear land VBmax / µm

600

ae = 15 mm wet

400

Groove milling in TiAl6V4 600 400

* Fracture

Dry

Compressed air

ae = 7.5 mm wet

Cutting fluid

200

*

200 100 80 60

100 80 60 40

ae = 15 mm dry ae = 7.5 mm dry Cutting edge phase: 0.08 x 26°

40 30 2

4 6 8 10

20

40 60 100 200 400

20 10 1

Cutting time tc / min Material : TiAl6V4 Cutting tool material : HM-S40 Cutting parameters : vc = 50 m/min fz = 0.17 mm

2

4 6 8 10

20

40 60

Cutting time tc / min Down milling ae < D D = 25 mm ap = 3 mm Z =1

Groove milling ae = D D = 32 mm ap = 30 mm ae = 32 mm

Fig. 7.37 Influence of wet or dry cutting conditions on the tool wear development during milling TiAl6V4

tool failure due to fracture. Numerous investigations into dry milling steel materials have proven that dispensing with the cutting fluids leads to a significant increase in tool life. In order to diminish thermal alternate stress and the resulting formation of comb cracks, the question is whether it makes technological sense to dispense with cutting fluids when milling titanium alloys too. Comparative studies [Köll86] in down milling TiAl6V4 with ae < D have confirmed the general results gathered when dry milling steel materials (Fig. 7.37). In comparison to wet machining, dry cutting also leads here to a significantly lower amount of wear on the end milling cutters used. In the case of down milling with ae /D = 7.5/25, the tool’s life had ended already after 8 min with wet cutting, while in dry cutting its life was extended to 200 min. This result changes fundamentally when milling in full groove cutting, i.e. as soon as the cutting edge penetrates the material with a chip thickness of h = 0 mm. In this case, the lowest amount of wear was recorded when milling with a cutting fluid. The reason why wear is increased when dry milling is that chips adhering to the cutting edge or found in the groove end up in the material between the cutting edge and the groove flank, are compressed on the surface of the up milling flank and cause cutting tool material fracture on the cutting edge. When milling in full groove cutting, the cutting fluid has the essential task of separating the chips from the cutting edge and transporting them out of the slot [Köll86]. In principle, it is also possible under certain conditions to dry machine titanium alloys. In most application cases in praxis however, cutting fluids are used, not the least because of the safe chip removal and therefore for reasons of higher process safety. In these cases, one should take care that cutting fluid supply to the tool is carefully adjusted in order to reduce thermoshock and that the tools are intensively cooled.

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60

Width of flank wear land VB/µm

50

CF3, 10% CF1, 10%

40

Water CF2, 10%

30 CF1, 5% 20 Liquid nitrogen 10

0 0

20

40

60

80

120

90

Cutting time tc/s Material : Prozess : Tool :

TiAl6V4 End milling, down milling D = 50 mm, Z = 7

Cutting parameter : vc = 32 m/min fz = 0.1 mm

ap = 50 mm ae = 15 mm

Fig. 7.38 Influence of cutting fluid on the tool wear development during milling TiAl6V4

One alternative to conventional flood cooling when milling is, in analogy to turning (Fig. 7.36), supplying the cutting fluid under high pressure. The use of cold or liquid gases is an extreme form of tool cooling. Investigations with various cutting fluid media, concentrations, water and liquid nitrogen had the following results (Fig. 7.38). There are at points significant differences in performance between the cutting fluid media. When milling titanium alloys, unsatisfactory results can also stem from the use of an unsuitable cutting fluid. Since tool cooling is especially important in titanium machining, the amount of oil in the emulsion should not be too high. In this investigation, a 5% emulsion lead to less tool wear than a 10% emulsion. As the more pronounced wear formation when milling with pure water containing only a rust inhibitor proves, a certain amount of oil must however be present in the cutting fluid in order to reduce friction. In this comparison, cooling with liquid nitrogen delivered the best results (Fig. 7.38). The boiling point of liquid nitrogen is –195.8◦ C. The use of this cooling medium led to an extreme cooling of the tool and the chips. Similarly positive results can be obtained by the use of cold gases. One example of this is the use of CO2 snow (dry ice) when turning a duplex steel. By means of the intensive cooling, tool wear and burr formation could be reduced in addition to chip formation and surface

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quality being improved [Wein07]. As these examples show, the use of extremely cold media can contribute to the solution of difficult machining tasks. Whether the cost associated with their use is economically worthwhile must be decided on a case-by-case basis. Titanium alumnides may have the potential to raise the operating temperatures of titanium alloys to 800◦ C. They are for this reason a possible alternative to nickelbased alloys at about 50% the weight, not only for high pressure compressors but also for the low pressure turbine. Their low elongation at break (< 1% at ca. 700◦ C) and low thermal expansion (λ = 10 W/(mK)) make these alloys problematic for machining. The problems include high tool wear, extremely low applicable cutting speeds and insufficient surface quality. Components made of titanium alumnide have predominantly been machined with uncoated cemented carbides (HW-K10/20) [Ecks96, Aust99]. When turning however, tools made of PCD or with CVD thick diamond films can also be successfully employed. The main problem in machining these materials under conventional process conditions is the formation of defects in the form of microcracks and micro-fractures on the surface of the machined workpieces (Fig. 7.39). The cause is the high level of brittleness of the material and its incapacity to deform plastically. This becomes very clear in consideration of the discontinuous and segmented chips that are formed. Chip formation such as it is familiar when machining steels or titanium alloys, is not found in conventional machining. Instead, angular, needle-shaped chip lamellae

30 µm

60 µm

2 mm

10 mm

Fig. 7.39 Surface formation and chip forms during turning of γ-titanium aluminide with conventional and adapted cutting conditions

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are formed whose separating surfaces are very rough, have obvious fracture structures and extensive cracks. Plastic deformation of the material during chip formation has not been observed. The material to be removed is, so to speak, pried out of the workpiece material. Such surface structures are unusable above all for components requiring high levels of safety such as are needed in engine construction. As the basic research has shown, the machining result can be significantly improved by an adjustment of the cutting edge geometry and process parameters. With sharp-edged tools and the choice of a small ratio of depth of cut ap to the corner radius rε , it is possible to produce smooth damage-free surfaces by turning workpieces made of Ti-45Al-8Nb-0.2C. Under these conditions, coiled chips are formed such as one finds in steel machining. When cemented carbides of the S10/20 group are used, cutting speeds of up to 100 m/min can be used.

7.6.4 Copper Alloys Copper alloys are privileged because of their excellent thermal conductivity and corrosion resistance in the following areas: • • • • • •

air-conditioning hydraulic engineering recuperator technology food technology chemical equipment and apparatus technology auxiliary equipment

Copper alloys are defined as alloys in which there is at least 50% copper. They can range from the superhard two-material system copper/aluminium to types of pure copper with low strength and high fracture stress. In comparison to other metallic construction materials, most copper alloys, as will be explained below, are considered easy to machine. A distinction is drawn between wrought and cast materials. Within this classification, it is generally sensible to classify according to alloy groups. Although the properties of copper alloys are essentially determined by their chemical composition, this organizing principle is not suitable for classifying the machinability of copper alloys because of highly varying machinability of copper alloys of the same type [DKI83]. With respect to machinability, the following categorization is suggested: • Pure copper and copper alloys with zinc, tin, nickel and aluminium, the additive elements of which are adjusted to each other such that they only form a homogeneous mixed crystal. Such alloys are easy to deform when cold and have a high level of deformability. This group is considered to be moderately to poorly machinable. Alloys of this group are especially familiar in the form of the two-material system copper-zinc, known as brass.

312


• Alloys with the elements zinc, tin, nickel, aluminium and silicon, however without chip-breaking additives, which form a second mixed crystal. Such heterogeneous alloys are harder than the previously mentioned group. They have less deformability than the previous group, but have higher machinability. This group is especially characterized by the three-material system copper/tin/zinc and copper/nickel/zinc, also known as nickel silver. • Alloys of both above groups, to which are added lead, sulphur, selenium and tellurium as insoluble components in order to improve chip breakage. Strength is hardly affected, notched impact strength and deformability are reduced [Seid65]. This group is the most machinable because of its improved chip fracture. This group is made up of automatic alloys, to which the elements lead, tellurium or selenium are added so that an unproblematic chip breakage can reduce disturbances in the manufacturing process. Special alloys can contain nickel, cobalt, tin or vanadium. Lead, bismuth, antimony and cadmium are contained in free cutting alloys as chip-breaking additives [John84]. Beryllium, boron and natrium are supplemented as trace additives in order to affect crystallization. The emphasis of the main criteria of machinability changes depending on the machining task at hand, especially when machining copper alloys. For this reason, it is difficult to make a general classification of these alloys with respect to their machinability. However, there are a few essential points which influence the machinability of copper alloys that should be considered: • the manufacturing process for producing semifinished products, e.g. primary forming or shaping • heat treatment, e.g. hardening • the chemical composition of the alloy

7.6.4.1 Influence of Semifinished Product Manufacture When machining cast copper alloys, one must consider the structure of the rim zone, which is different from that of the core structure and is called casting skin in practice. This rim zone is characterized by a higher level of hardness and strength compared to the core structure, resulting in an acceleration of tool wear. The core structure of cast alloys is generally more machinable than that of wrought alloys [DKI83]. In the case of cold forming copper alloys, their hardness and strength are increased while their deformability is reduced. This leads to a positive influence on machinability due to improved chip fracture compared with materials that have not been shaped [DKI83]. 7.6.4.2 Influence of Heat Treatment Machining hardenable copper alloys is preferably executed with a cold-worked material prior to heat treatment, since machining after hardening would cause

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313

accelerated tool wear. On the other hand, the harder material state is preferable for subsequent grinding and polishing work. 7.6.4.3 Influence of Chemical Composition Brass and pure copper (Group 1) are difficult to machine because of their high levels of toughness and deformability. They are characterized by high chip compression, which has a large impact on the tribology in the chip/rake face boundary surface and causes high mechanical stress on the cutting edge. In practice, these phenomena are known as “snagging”. The effect of the alloying elements lead, sulphur, selenium and tellurium on the form of insoluble components in copper alloys is comparable to the effects of these elements in automatic steel [Isle73, Lore74]. 7.6.4.4 Wear When machining tough alloys in the case of continuous chip formation and low temperature at the boundary surface between the chip and the tool, built-up edges can be observed that lead to accelerated wear of the cutting edge [Klei66, Mess69]. Due to hardness and deformability, the tool life parameters are less favourable when machining nickel silver than when machining brass [Vict72]. Built-up edges are formed in brass machining as a result of the wear mechanism of adhesion. In the case of high speed milling, uncoated cemented carbide of the application group K10/20 is recommended in order to guarantee good cutting edge durability. In the case of materials that are difficult to machine and tend to adhesion, such as pure copper or brass containing large amounts of copper, polycrystalline diamond (DP) has proven superior as a cutting tool material not only in terms of better wear resistance but also by its favourable tribological conditions, leading to improved surface quality and lower resultant forces. Ceramic cutting tool materials are not suited to cutting copper alloys because of their adhesive tendency [grei91]. 7.6.4.5 Resultant Force As the cutting speed rises the specific cutting force kc falls in the case of copper alloy machining as well. When the cutting speed is increased from 5 to 160 m/min when performing an external cylindrical turning operation on copper, the specific cutting force is reduced, for example, by about a third [DKI83]. Further increase of the cutting speed causes the specific cutting force to converge asymptotically towards a constant value. It has become clear that, in the case of the cutting speeds commonly used today with cemented carbides as cutting tool materials (> 160 m/min), the influence of cutting speed can be neglected. Since the specific resultant force in the case of copper alloys is generally considerably lower than that in cutting of steel, difficulties due to insufficient driving power of the machine tool can hardly be expected in practice. For cast alloys, the casting method used to produce the semifinished products has a considerable effect on the amount of specific resultant force. Machining experiments have determined that with constant cutting values parameters the amount of specific resultant force is lower for centrifugal casting than for

314


sand casting, although the centrifugally-cast workpieces had higher tensile strength, Brinell hardness and strain. Due to the finer crystalline structure of centrifugally-cast parts, chip formation also proved more favourable. 7.6.4.6 Surface Quality Built-up edge formation and flank face wear of the minor flank face lead to poor surface quality [Klei66, Mess69]. In the case of thin-walled workpieces, deformations of the workpiece due to the cutting force can appear as a result of the low elastic modulus of copper alloys (e.g. CuZn30: 115.000 N/mm2 with RT). These deformation not only endanger dimensional accuracy, but also induce undesired residual stresses in the rim zone. Lowering the cutting force can lead to improved quality. As a rule, the use of a cutting fluid improves the surface quality as well [Grei91]. 7.6.4.7 Chip Form The machinability of nickel silver (Group 2) can vary to a great extent depending on the respective amounts of the alloying elements zinc, tin, nickel, aluminium and silicon. Usually, acceptable chip forms are obtained however. The chip formation of pure and homogeneous copper is relatively unfavourable. Large cross-sections of undeformed chip and an unhindered chip flow promote long ribbon chips. With additional alloying elements (Pb, Te, S, Se), chip fracture can be significantly improved, and thus the chip form as well. For example, an alloy made of telluriumcontaining copper CuTeP can even be machined on automatic machines because of the short-breaking chips. In contrast to unalloyed copper, the alloy CuTeP has only slightly less thermal conductivity [DKI83].

7.6.5 Nickel Alloys Nickel alloys are materials with nickel as their main component, which are alloyed with at least one other element. They are characterized by good corrosion resistance and/or excellent high temperature strength. Nickel has a cubic body-centred crystal lattice and, at low temperatures, also has excellent ductility and cold-formability. It is one of the ferromagnetic materials. Some nickel alloys have special physical properties. One example is the softmagnetic nickel-iron alloy with about 15% Fe, 5% Cu and 4% Mo, also familiar under the name Mumetall, which is used for shielding magnetic fields because of its high magnetic permeability [Schu04]. Nickel alloys are often applied as construction materials due to their superior properties. Examples include: • chemical industry (cauldrons, heat exchangers, valves, pumps), • environmental protection and waste management (flue gas desulphurisation plants), • power production (power plant generators), • aerospace (engines).

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The specific properties, adjusted to the respective application area, are essentially dependent on the chemical composition, possible cold-shaping and the type of heat treatment. In accordance with their most important alloying elements, nickel alloys can be classified in the following main groups (Fig. 7.40) [DIN17742, DIN17743, DIN17744, DIN17745, Ever71]: • • • • •

I nickel-copper alloys II nickel-molybdenum alloys and nickel-chrome-molybdenum alloys III nickel-iron-chrome alloys IV nickel-chrome-iron alloys V nickel-chrome-cobalt alloys

Materials of group II are not hardenable by heat treatment. Groups I, III and IV comprise both non-hardenable and hardenable alloys. Materials of main groups III and IV, which can be hardened provided they have a corresponding amount of aluminium and/or titanium in conjunction with heat treatment, are called “super alloys” as do those of group V. The respective trade name of each alloy to have been offered on the market was assigned to each main groups I–V. The classification in Fig. 7.40 is not very precise; for example, not all materials classified as nimionic contain cobalt as an alloying element, yet this classification is still good

Main group II Alloy of Type HASTELLOY non-hardenable

Main group Ib Alloy of Type K-MONEL hardenable Al Ti

Cr Mo Fe Cr

Cu

Main group Ia Alloy of Type MONEL non-hardenable

Al Ti Main group IIIb Alloy of Type INCOLOY hardenable

Ni

Cr Fe

Main group IVa Alloy of Type INCONEL non-hardenable

Main group IIIa Alloy of Type INCOLOY non-hardenable

Cr Co Al Ti Main group IVb Alloy of Type INCONEL hardenable

Al Ti Main group V Alloy of Type NIMONIC hardenable

Fig. 7.40 Classification of nickel alloys in main groups (Source: Wiggin Alloys, Huntington Alloys), acc. to EVERHART [Ever71]

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for orientation purposes. It should be noted that every manufacturer provides its products with its own trade name. The main stress of this chapter is on hardenable high temperature resistant nickelbased alloys, which are chiefly utilized in aeroplane engines and stationary turbines. The alloying elements of these materials can be grouped into three categories in accordance with their effect on the microstructure. Some elements effect it in several respects. The elements Cr, Co, Mo and W form mixed crystals with nickel. Besides an increase in strength at low temperatures, they cause an increase in the alloy’s creep strength at high temperatures because dislocation creep is limited in the γ mixed crystals [Schu04]. Chrome improves oxidation and corrosion resistance, cobalt promotes the stability of the γ phase. The most essential strength-increasing mechanism among high temperature resistant nickel-based alloys is the deposition of intermetallic phases γ (Ni3 (Al,Ti)) and γ (Ni3 (Nb,Al,Ti)). The γ phase, which is coherent to the γ matrix, causes particle-hardening of the structural matrix that remains effective up to the high temperature range. It is formed by alloying aluminium, which can also be substituted with titanium and tantalum. The γ phase can already originate from the molten bath. In the cast state however, it exhibits an uneven particle size, formation and distribution in the microstructure. For this reason, wrought alloys are subjected to a multistage heat treatment after melting in a vacuum and solidification. Solution annealing first dissolves the γ phase in the mixed crystal matrix. During cooling from solution annealing temperature, it is deposited again in a – in comparison to the cast state – more consistent form in the structural matrix. Final storage at high temperatures brings about a further improvement of the consistency of the particle size and form of the γ phase in the structure [Schu04]. In the case of materials that have an increased amount of niobium, such as Inconel 718, there is also a hardening via the γ phase in addition to the γ hardening. In contrast to the γ phase however, the γ phase tends toward a more rapid coarsening and thermal instability, which limits the long-term use of Inconel 718 to 650◦ C [Bürg06, Kenn05]. More recent nickel-based alloys, such as Allvac 718Plus, thus have a larger amount of titanium and cobalt than Iconel 718 with the same Nb-content in order to improve the formation and stabilization of the γ phase. For example, the continuous operation temperature of Allvac 718Plus was able to be raised to 704◦ C. In the case of the nickel-based alloy Udimet 720, the γ phase appears in the shape of Ni3 (Al,Ti) and (Ni,Co)3 (Al,Ti). A higher amount of cobalt (14.7%) reduces the solubility of titanium and aluminium and thus makes it possible for the γ phase to form in average temperatures. Udimet 710 is applicable for continuous operation temperatures of up to 730◦ C [Tors97, Helm00, Mark05, Bürg06]. In the case of polycrystalline alloys it is differentiated between forging alloys and casting alloys. Forging alloys, such as Inconel 718, Waspaloy, Udimet 720LI and Allvac 718Plus contain a volume percentage of about 30–40% of γ phase. With increasing amounts of γ phase, deformability and machinability become increasingly limited. In the microstructures of cast alloys however, such as Inconel

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317

713C and MAR-M-247LC, there are larger amounts (up to about 70%) of γ phase [Schu04]. Because of the larger amount of γ phase, cast alloys are no longer forgeable and are generally machined using grinding methods. Machining using methods with geometrically defined cutting edges is possible to some extent, depending on the cast alloy. Due to the poor machinability of cast alloys, usually fine casting methods are used close to the final contour that do not require extensive mechanical after treatment [Schu04]. The elements Cr, Ti, Mo, W and Ta form carbides of various chemical compositions (MC, M6 C and M23 C6 ) with carbon, which is present in small concentration in the alloys. In polycrystalline alloys, these carbides are formed chiefly at the grain boundaries. They bring about an increase of creep resistance because they inhibit sliding of the grain boundary [Schu04]. In addition to cast-metallurgical fabrication, parts made of nickel-based alloys can also be produced using powder-metallurgical methods. The goal of PM engineering is to avoid all forms of segregation during solidification as well as concentration and structural gradients. In addition, it is possible to increase the strength of alloys by pulverizing and compacting ultrahard, non-forgeable alloys (cast alloys) generally containing high amounts of γ phase. PM technology makes it possible to develop alloys with a higher high temperature strength than forging alloys. One central problem of PM technology is that no foreign components may infiltrate the powder, since even those of the same size as the powder particles can lead to cracking in the sense of fracture mechanics. The powder is encapsulated, hot-isostatically compressed and then forged in order to decrease the possibility of a larger foreign component inclusion or remaining inhomogeneity. One method variant in this context is “gatorizing”. This is a process in which the powder is isothermally forged using a creep forming process. PM alloys are used to manufacture turbine blades for both civil and military aeroplane engines [Adam98]. The varying chemical compositions and crystalline structures of nickel-based alloys have a direct effect on their machinability. With respect to their machinability, nickel-based alloys can be divided into five machinability-groups (Fig. 7.41). Here, group 1 represents easier, 3 average and 5 difficult levels of machinability. Representative materials are listed for each machinability-group, which are designated by their trade name as is customary for nickel-based alloys. In the case of alloys of machinability-groups 1 and 2, cold-shaping (strainhardening) has a positive effect on chip formation and surface quality. The machinability of these groups is comparable to that of corrosion-resistance austenitic steel. Machinability-groups 3 and 4 comprise nickel-based alloys that are hardenable by heat treatment, also called “super alloys”. With respect to tool wear and potential surface quality, roughing of the materials of both groups should take place in a solution-annealed state and finishing in a hardened state. With respect to machining, there is practically no difference between wrought and cast alloys of groups 1 and 4 with the same composition. Due to their higher strength and usually high amount of γ phase in the structure, PM alloys are assigned to machinability-group 4. Cast alloys of group 5 are very difficult to machine because of the large amount of γ

318


Machinability group 1

2

3

4

Wrought alloy

Alloy of main group I.) Ni-Cu Leg.

Non-hardenable alloys of main group I.) Ni-(Cr)-Mo Leg. III.) Ni-Fe-Cr Leg. IV.) Ni-Cr-Fe Leg.

5 Cast alloy

Hardenable alloys of main group II.) Ni-(Cr)-Mo Leg. III.) Ni-Fe-Cr Legierungen IV.) Ni-Cr-Fe Legierungen V.) Ni-Cr-Co Legierungen

High temperature cast alloys

Examples

Examples

Examples

Examples

Examples

Monel 400 Monel 401 Monel 404 Monel R 405

Hastelloy B Hastelloy X Incoloy 804 Incoloy 825 Inconel 600 Inconel 601

Incoloy 901 Inconel 718 Inconel X750 Nimonic 80 Waspaloy Allvac 718Plus

Nimonic 90 Rene 41 Udimet 720LI Astroloy LC PM René 95 PM

IN100 Inconel 713C Inconel 718C Mar-M247LC Nimocast 739

Fig. 7.41 Classification of nickel-based alloys in machinability-groups (Source: Machining Data Handbook, Huntington Alloys)

phase and of carbides, their coarse-grain microstructure and low grain boundary strength; fractured material particles and cracks in the grain boundaries frequently cause problems when manufacturing functional surfaces [Lenk79]. Due to their mechanical, thermal and chemical properties, nickel-based alloys are generally included among materials that are hard to machine. Their high high-temperature strength in comparison to steel, low thermal conductivity, their considerable tendency to form built-up edges and to strain hardening – as well as the abrasive effect of carbides and intermetallic phases – lead to extremely high levels of mechanical and thermal stress on the cutting edge during machining. Under the machining conditions shown in Fig. 7.42, 60% higher cutting forces and about 35% higher temperatures were measured on the chip bottom side when turning Inconel 718 with vc = 50 m/min compared to the heat-treated steel 42CrMo4 + QT [Kloc06b]. Due to the high thermal and mechanical stress, tools made of high speed steel and cemented carbide can only be used at low speeds when machining nickel-based alloys. Common cutting speeds when turning Inconel 718 and Waspaloy with uncoated cemented carbides of ISO-application group HWK10/20 are vc = 20−50 m/min. In the case of finish turning, cutting speeds of up to 100 m/min can be reached with coated cemented carbides. Presently, turning operations are still executed to a large extent with cemented carbides. In the case of pretreatment under average roughing conditions however, ceramic materials are being used increasingly – in the case of finishing both ceramic and PCBN cutting tool materials.

7.6


319

Temperature at the chip bottom side Tch /°C

800 X5CrNi18-10 250

700

Inconel 718 150

42CrMo4+QT

100

200 150

600

100 Δ Tch

500

50

+ 35 %

200 150 50

Δ Tch

100

20

50

+ 10 %

400

Δ Fc

– 15 %

+ 60 %

0 0

400

500

600

700

800

900

1000

1100

1200

1300

Cutting force Fc / N Depth of cut: ap = 2 mm, feed: f = 0.1 mm, insert geometry : CNMG120412, coated cemented carbide, dry cut, tc = 10 s

vc = 20 m/min vc = 50 m/min

vc = 150 m/min

vc = 75 m/min

vc = 200 m/min

vc = 100 m/min

vc = 250 m/min

Insert with chip breaker Quartz fibre

2-colour pyrometer

Fig. 7.42 Comparison of measured cutting forces and chip bottom side temperatures during turning of different types of steel and the nickel-based alloy Inconel 718

In many machining tasks today, the use of cutting ceramics is the state of the art. Of the ceramic cutting materials, composite ceramics ductilized with SiC whiskers (CW) have shown the greatest performance potential so far. They are suited both for finishing and for pretreatment of turbine parts under average roughing conditions (vc = 150−300 m/min, f = 0.12−0.3 mm, ap = 0.5−2 mm). Of the silicon nitride cutting ceramic group (CN) especially the α/β-SiAlONs have proved equal to whisker-reinforced cutting ceramics in their wear properties (vc = 150−200 m/min, f = 0.12−0.3 mm, ap ≤ 4 mm). α/β-SiAlONs consist of needle-shaped β-Si3 N4 and globular SiAlON crystals. In this way, they combine the toughness of β-Si3 N4 -ceramics with the high hardness and wear resistance of SiAlONs in one cutting material [Gers02]. CBN cutting materials are primarily used for finishing components made of nickel-based alloys (vc = 250–350 m/min, f = 0.12–0.2 mm, ap ≤ 0.5 mm). When machining nickel-based alloys with CBN, the selection of a type of cutting materials suited to the particular machining task is of primary importance. CBN cutting tool materials available on the market can differ considerably with respect to the modification and amount of boron nitride, grain size and the structure of the

320


binder phase. The resultant chemical, physical and mechanical cutting tool material properties affect the wear and performance of CBN tool to a large extent. For nickel-based alloy finishing, fine-grain CBN types with a TiC or TiN-based binder and a percentage of 50–65 vol% of cBN have proven especially suitable [Gers02]. The arc-shaped profile of its tool life graphs (Fig. 4.54) is characteristic of turning nickel-based alloys with ceramics and CBN cutting tool materials. This is the result of various wear phenomena which are predominant in dependence of the cutting speed. While flank face wear is the primary wear criterion limiting tool life in the case of turning with cemented carbide tools, tool life is limited at lower speeds by notch wear and only at higher speeds by rake and flank face wear in the case of ceramics and CBN cutting tool materials. The arc-shaped curve of the tool life graphs indicates that there is a range of optimal cutting speeds. The closer the ascending and descending branches of the tool life graphs are to each other, the more important it is to work within a range that is as narrow as possible around the tool life maximum. Ceramic and CBN inserts with a neutral tool orthogonal rake angle (γo = 0◦ ) have become standard for finishing nickel-based alloys. CBN inserts should have a small rounding of about 10 μm on the cutting edge. In contrast, ceramic cutting inserts are usually provided with a small protective chamfer to stabilize the cutting edge. The formation of wear notches on the major and minor cutting edges of insert is characteristic of turning nickel-based alloys with cutting ceramics or CBN cutting tool materials. This significantly affects tool performance and component quality. Notch formation on the major cutting edge leads to the formation of a burr on the workpiece edge; notch formation on the minor cutting edge leads to a deterioration of surface quality (Fig. 7.43). Notch wear at the major cutting edge

Burr formation at the workpiece edge

Notch wear at the minor cutting edge

distinctive feed ridges on the workpiece surface

Fig. 7.43 Notch wear at the major and minor cutting edges significantly affect the performance and surface quality

7.6


321

Wear mechanisms that lead to the formation of wear notches include: • fatigue, cracking and crack development phenomena – initiated by the high thermal and mechanical alternate stress on the cutting tool material as a result of lamellar chip formation and transverse material flow, • adhesion – caused by the tendency of the kfz-crystal lattice to form microwelding and the mechanical snagging of the chip edge, flowing laterally into the notches, with the cutting tool material particles, • abrasion – caused by the solidified, partially sawtooth-shaped chip edge, by the workpiece edge and the peaks of the feed groove pattern • tribooxidation – caused by the chemical reaction of the workpiece material and/or the cutting tool material with components of the surrounding medium. Wear notches arise as the result of the superimposition and mutual influence of these individual mechanisms (Fig. 7.44). A clear separation of these mechanisms with respect to their effect on total wear is only possible to a certain extent [Mütz67]. In order to reduce notch wear formation on the major cutting edge, the ratio of corner radius rε to depth of cut ap should be as large as possible and the effective lead angle in the area of the contact zone end on the major cutting edge as small as possible. A small lead angle reduces the chip thickness and thus the effect of abrasion of the chip edge on the cutting edge as well (Fig. 7.45). A lead angle of κ r ≤ 45◦ has proven effective for turning with tools made of cutting ceramics and CBN. The effective use of round inserts is explicable in this context. Depending on the geometry, they exhibit not only a high amount of mechanical stability and thus

Direction of chip flow Chip lamella Direction of feed motion Transverse material flow following upsetting of the chip lamella

Rake face Flank face

Direction of workpiece rotation Burr formation at the workpiece edge Abrasion Surface damage

Formation of a sawtooth profile at the edge of the chip lamella Abrasion Surface damage Cutting edge sector with : - high pressures - high temperatures - high shear stresses - high thermal and mechanical stress gradients - high thermal and mechanical alternating stresses Surface damage Adhesion Tribooxidation

Fig. 7.44 Schematic illustration of reasons which cause notch wear at the major cutting edge [Gers98]

322

7 Tool Life Behaviour rε

rε

RCGX090700T

κ r = 21°

κ r = 51° f

f

h = 0.06 mm

ap

Material: Feed:

VBGW160408F

h = 0.11 mm

ap

Udimet 720Li f = 0.15 mm

Cutting speed: Depth of cut:

vc = 250 m/min ap = 0.3 mm

Fig. 7.45 Influence of the corner radius on wear formation during turning of a nickel-based alloy with PCBN

more resistance to insert fracture, but, especially with small depths of cut, they also generally result in small lead angles. It is thus recommendable to use round indexable inserts if the machining task allows for it. However, the disadvantage of round indexable inserts are the high passive forces, which can cause a more pronounced plastic deformation of the rim zone. Furthermore, the larger contact arc in the area of the minor cutting edge promotes the formation of wear notches. This has been observed especially in the case of turning with round indexable inserts made of cutting ceramics. If very high demands are placed on surface quality, CBN inserts with corner radii of 0.8–1.2 mm should be used [Gers02, Kloc07a]. Notch wear formation of the major cutting edge can also be effectively reduced and tool life improved by the selection of an appropriate cutting strategy. One measure that has proven very effective particularly in pre-machining with round cutting ceramics is “ramping”. Here, the material is removed in pairwise pass with continuously decreasing (1st step) and increasing depths of cut (2nd step). As a result of the continuously changing depth of cut, the effect of the highly abrasive chip edge is concentrated not on one location of the cutting edge, but rather it extends along a larger area thereof. In this area, the cutting edge is indeed subject to more wear, but notch wear proceeds much more slowly, so that the insert can be used for a longer time (Fig. 7.46).

7.6


Conventional: Constant ap

ap = 1.8 mm

323

Modified: Variable ap (ramping)

a p1 : 0.8

ff Verschleißkerbe Notch wear

1.8 mm

f a p2: 0.8

0.8 mm

Inconel 718, whisker-reinforced cutting ceramic, vc = 180 m/min, f = 0.15 mm

I c = 261 m

I c = 1281 m

Fig. 7.46 Adapted cutting strategies allow a reduction of notch wear and a longer tool life (Source: Greenleaf)

Because of the extremely high stress on the cutting edge, milling nickel-based alloys requires cutting tool materials of high toughness and wear resistance. Because of the interrupted cut and high heat resistance of the material, the tools are subject to an extremely high mechanical and thermal alternate stress. As a result, both HSS and cemented carbide tools can only be used with relatively low cutting speeds and feeds, so that only low material removal rates can be realized. The milling of such alloys is thus an extraordinarily time-intensive machining process. HSS tools are used with cutting speeds of vc = 5−10 m/min when machining nickel-based alloys because of their low high temperature wear resistance. Their excellent toughness properties allow however for the use of relatively large feeds (fz = 0.10−0.16 mm). Due to the large number of cutting edges being engaged as well as the milling cutter with the fine knurled splines, milling with HSS tools is extraordinarily quiet. Coated end milling cutters made of PM-HSS are therefore suited above all for roughing nickel-based alloys. With a cutting edge geometry adjusted to the particular requirements of these materials, high material removal rates and a large amount of manufacturing safety is possible. End milling cutters made of superfine and ultrafine-grain cemented carbides have a much higher level of high temperature hardness than HSS tools. They can therefore be used at much higher cutting speeds (vc = 20−100 m/min), allowing for a significant reduction of machining times. They are commonly used for average

324


roughing and especially for finishing operations. When end milling cutters and inserted-tooth cutters that are fitted with HM indexable inserts are used, the performance capacity of the tools is determined to a great extent by cutting tool material spalling on the cutting edges of the indexable inserts when nickel-based alloys are machined. They are formed as a result of a mechanical overload of the relevant section of the cutting edge. For the sake of a dynamically stable and quieter milling process, one should strive to have as many cutting edge rows as possible in simultaneous engagement. This demand cannot always be fulfilled for the given machining task due to the design of milling cutters fitted with cutting inserts [Gers02]. If grooves must be generated by rough milling in full groove cut into components made of nickel-based alloys, the tools are subject to very high levels of thermal and mechanical stress. Due to the wrap-around angle of 180◦ between the tool and the workpiece, a very strong force component affects the tools in the feed direction which strongly subjects the end milling cutter to bending. Faults in shape and dimension on the groove flanks are the result. Due to the high static and dynamic stress, the tools can only be used in full groove cut at relatively low speeds (vc = 20−40 m/min) and axial depths of cut (ap = 0.5 · D). In the case of long-protruding tools, the cutting parameters must be further decreased in order to reduce mechanical stress and the danger of cutting edge fracture and total breakage of the tool. Economical alternatives to conventional end milling of grooves into materials that are difficult to machine include “trochoid milling” (Fig. 7.47) and “plunge milling”. In the former case, the feed motion of the milling cutter, whose diameter is smaller than the width of the groove, superimposes an approximately circular motion. Usually, down milling is used. What is special about this method is that, as a function of the ratio of milling cutter diameter to groove width and the chosen feeds in axial and radial direction, the wrap-around angle is significantly smaller than 180◦ , thus corresponding to finishing conditions. Trochoid milling leads to a considerable reduction of mechanical tool loading. As a result, not only the cutting speed but in particular the axial depth of cut can be significantly increased compared with conventional groove milling. Especially in the case of slender, long-protruding tools, the realizable depth of cut is several times more than that of conventional milling. Because of the larger potential depths of cut, the material volume to be machined is distributed across a larger cutting edge length. The tools are thereby used more effectively and tool life is much higher with respect to the amount of material machined, lowering tool costs drastically. A further advantage of this method is that the width of the groove that is to be manufactured is independent of the tool diameter and can be produced accurately in one milling cycle. Since in this method only low forces act upon the tools on both flanks when entering and exiting, errors in form and dimension on both groove flanks are extraordinarily few. By superimposing a wobbling motion, not only straight but also curved flanks can be created with this method. In the case of plunge milling, the end or plunge milling cutter is plunged axially into the workpiece with low radial feed. In this way, the tool is primarily stressed in the axial direction [Kloc04]. The radial toolbending forces are, dependent on the radial feed and the corner design of the end

7.6


325

Conventionel milling

Trochoide milling

Angle of contact ϕ 180°

Angle of contact ϕ 0°

°-6

30

**

**

Direction of **: rotation 2 Feed motion

1

3

1 4

0.5 x Tool-Ø 2 x Tool-Ø

15–40

vc/(m/min)

60–120

< 0.5 x Tool-Ø

ap/mm

3 mm

Fig. 7.56 Tools for machining aramid-fibre-reinforced plastics

variation as well, the permanent alternate stress prevents the fibres from eluding the cutting edge. The following describes difficulties involved with particles arising when machining fibre-reinforced plastics (dust, fibre particles). Figure 7.57 clarifies the high concentration of particles that remain in the working space even after the machining process and that is only gradually reduced. Especially the small particle sizes, thoracic (4.5–10 μm) and alveolar (< 4.5 μm) dusts, remain in the air long after machining is finished and sediment very slowly. As

336

7 Tool Life Behaviour 250 Concentration / (mg/m³)

Total dust Thoracic dust

200

Alveolar dust

150 100 50 0 0

100

200

300

400

500

Time t/s Milling time

Fig. 7.57 Dust particle size during milling of fibre-reinforced plastics

an experiment, a 90 s long machining process was conducted and the machine was immediately switched off. It is important to know that even after 500 s, the legally prescribed fine dust limit of 1.5 mg/m3 was not yet reached. These particle sizes are highly carcinogenic, so this limit must be strictly observed. The percentages of these sizes in the entire volume of dust can be influenced by tool selection and the cutting parameters. The experiment shows that it is nonetheless absolutely necessary to isolate the working area and to suction off and filter the resulting particles.

7.7.2.2 Machining Method: Drilling In the case of drilling, it is necessary just as in milling to differentiate in accordance with the materials to be machined, glass-fibre and carbon fibre-reinforced plastics or aramid fibre-reinforced plastics. To drill CFP and GFP, highly wear-resistant tools with standard cutting edge geometries can be used as long as the cutting edge radius is as small as possible. In many cases, solid cemented carbide tools produce good results. Tools that are equipped with DP cutting edges are far superior, but they can only be used for manual drilling operations to a limited extent. Standard tools cannot be utilized to machine aramid fibre-reinforced plastics. In this case, special cutting geometries should be used in order to guarantee separation of the fibres while prestressed (Fig. 7.58). Cemented carbide drills with very positive rake angles have become established for this. Besides a small cutting edge radius, high levels of surface quality on the flank and rake faces are required so that friction between the chip and the tool and between the workpiece and the tool reduces potential adhesion of the material on the cutting edge. Potential surface roughness depends heavily on the type and length of the fibres. When the tool is rotated, all fibre lengths are always cut, so that, depending on the orientation of the fibres, there is a highly aberrant surface formation in the drill hole.

7.7

Machinability of Non-metals

337

Low cutting edge radius High surface quality at rake and flank face High wear resistance

Conventional macroscopic cutting wedge geometry Side rake angle 20° .. 30° Side clearance angle 10° .. 12° Drill point angle 120°

CFRP und GFPR

Special, macroscopic cutting wedge geometry Side rake angle 30° .. 35° Side clearance angle 10° .. 14° Rear clearance angle 0° Tool cutting angle 125° .. 140° AFRP

Fig. 7.58 Tools for drilling fibre-reinforced plastics

It is characteristic of the material that, in the case of fibre-orientations of 105–135◦ with a simultaneous compressive load, only very poor surface qualities can be realized. As opposed to CFP and GFP, there are no fibre fractures in aramid fibre-reinforced plastics, but the fibres are pulled far out of the cut surface and bent against the cut surface. This is due to the tough material properties and poorer adhesion of the aramid fibres to the base material.

Chapter 8

Process Design and Process Monitoring

8.1 Finding Economical Cutting Parameters 8.1.1 Cutting Parameter Limits The further development of workpiece materials and cutting tool materials and the increasing automation of machine tools demand an occasional correction of machining standard values. Standard values should be determined under consideration of machinability and provide reference data concerning depth of cut, cutting speed, feed and the expected tool life parameters for the given tool life criteria. Variation in the machinability of a material of the same standard designation but different batch can be larger than three to one in extreme cases. This ratio is the more unfavourable the larger the permissible wear and the higher the cutting speed [Kluf83]. For an optimal cutting value determination, the tool life parameters and wear properties of the tool are the most important criteria for evaluating the tool life behaviour of a material/cutting tool material combination [Köni71]. Generally, widths of flank wear land of VB = 200−800 μm are permitted for flank face wear when cemented carbides are used. The tool life criterion for crater wear should be determined as a function of the feed f with the following equation: KT = 0.06 + 0.3 · f

(8.1)

In addition to tool wear, the other main evaluation parameters of machinability – the cutting force, surface quality and chip form – also serve, depending on the case at hand, as tool life criteria. When determining the cutting parameters, the limits of the system workpiece/cutting medium/tool/machine must be taken into consideration. Specification of these limit values refers to the description of the cross-section of undeformed chip using the theoretical parameters uncut chip thickness h and undeformed chip width b and must be converted in accordance with geometrical relations that are true for the respective cutting process into the technical parameters depth of cut ap and feed f (Chaps. 9 and 10). In order to obtain the largest possible volume removal rate within a given tool life, first a depth of cut ap should be selected that is as large as possible provided it is not F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_8,

339

340

8

Table 8.1 Maximum of undeformed chip width


Tool included angle ε r /◦

Chip width b

80–90 60 35–55

2/3 · cutting edge length 1/2 · cutting edge length 1/4 · cutting edge length

already defined by the machining allowance to be lifted in one cut. A large depth of cut reduces the number of required steps, so that besides the primary processing time (Sect. 8.1.2), the secondary processing time for lifting, returning and reengagement is economised. The maximum depth of cut apmax is limited by the stability of the tool’s cutting edge. For roughing, tools with a large tool included angle εr are used. The rules of thumb given in Table 8.1 serve as the standard values. Larger depths of cut or undeformed chip width can lead to cutting edge fractures. After determining the depth of cut, the feed should be set as large as possible in order to minimize the primary processing time. The smaller influence of feed on wear formation compared with cutting speed and the fact that specific cutting force is reduced with increased feed both speak for this sequence in determining the cutting parameters. In this way, the machine’s power can be exploited more fully. In addition to the maximum and minimum permissible feed of the machine, tool limits must again be considered in selecting the feed. In order to avoid unmachined areas on the workpiece, the following inequality can serve as a guide: fmax ≤ rε

(8.2)

The lower feed limit is determined using the minimum chip thickness hmin , which is a function of the stiffness of the system workpiece/cutting medium/tool/machine and is intended to guarantee chip removal especially in the case of rounded off or chamfered cutting edges. Following S OKOLOWSKI [Soko55], the following can serve as a standard value, for turning for example: 0.25
Pc = Fc · vc D Mspindel > Mc = Fc · 2

(8.37) (8.38)

The cutting force can be determined by the calculation basis given by SALOMON [Salo24], which is used here in KIENZLE’s notation [Kien52] in expanded form with Kvk as the correction factor for tool wear [Lang72, VDI3206, Degn00]. Fc = kc1.1 · b · h1−mc · Kvk

(8.39)

One can assume 5% per 100 μm width of flank wear land as a standard value for the correction factor Kvk .

350

8


If the spindle power is insufficient, the depth of cut or feed should be reduced accordingly.

8.1.3 Calculating the Machine Hour-Rate The machine hour-rate describes the costs to be calculated of a machine tool per hour. When determining the machine hour-rate, the • • • • •

imputed amortizations KA, imputed interest KZ, maintenance costs KI, room costs KR, and energy costs KE

are taken into consideration. The machine hour-rate KMH can be calculated as follows: KMH =

KA + KZ + KI + KR + KE TN

(8.40)

The yearly machine runtime TN amounts for example to 1600–1800 h/a for single-shift operation. In the case of multi-shift operation, the runtime is increased proportionately (e.g. two-shift operation ca. 3200 h/a or three-shift operation ca. 4800 h/a). The imputed amortizations KA are comprised of the acquisition value of the machine KAW , the amortization duration t and the sales revenue of the machine KVE after expiration of the amortization duration: KA =

KAW − KVE t

(8.41)

The duration of amortization t is relevant for determining the imputed amortization KA . The duration of amortization is calculated from the minimum of the expected technical or economical service life. The decreasing value of capital assets is designated in fiscal terms as allowance for amortization (AfD). The financial basis for the duration of amortization can be found in such AfD lists, since they turn out differently for different assets. Figure 8.6 clarifies the calculation of the imputed interest for the amortization term. The sales revenue KVE can be depreciated over the entire term. The acquisition value KAW does not remain constant over the entire term however, but gradually decreases. The imputed interest over the amortization term KZges can be interpreted geometrically as surface area under the cost function. It should be taken into consideration that the symbol I always indicates the rate of interest per year. The following is valid: KAW + KVE KAW − KVE ·i·t (8.42) ·i·t = KZges = KVE + 2 2

Finding Economical Cutting Parameters

Fig. 8.6 Imputed interest calculation

351

Costs

8.1

Cav

0.5 · (KAW – KVE) · i · t KVE KVE · i · t

Amortization period

t

Since the costs of Eq. (8.39) relate to 1 year, the imputed interest for 1 year KZ must be inserted in this equation, resulting in: KZ =

KAW + KVE ·i 2

(8.43)

Amortization, interest, room costs and fixed maintenance costs are all fixed quantities. On the other hand, energy costs, tool consumption and the variable maintenance costs are variable quantities. If the manufacturing labour is dependant on the machine run-time, it can be included in the rate of payment. In this case, we refer to it as a cost per man-hour.

8.1.4 Planning Methods and Tools According to the type of data determination, we distinguish between manual (external) and computer-aided cutting data determination. The boundaries here, depending on the computer use, are fluid (Fig. 8.7). In the case of manual cutting parameter determination and optimization, documents from various sources are utilized: • • • •

guidelines of different standardization committees catalogues and cutting parameter recommendations of the tool manufacturers trial results from research papers in-house data banks for cutting parameters

Machining standards, such as are published by the Verein Deutscher Ingenieure (VDI) (English: Association of German Engineers) for various machining processes as VDI guidelines, give the process planner only very rough reference values for the target cutting parameters [VDI3206, VDI3208, VDI3209]. Usually, cutting speeds are recommended for broadly associated material groups independently of chip cross-section, tool life behaviour, machine type and machine-hour rate that often

352

8


Methods for process optimization

Manual fixed value optimization

Manual Computer aided offoff-line optimization line optimization

Computer aided on-line optimization

Indicators - Standard value tables - Nomograms - Formulas

- Process optimization by end of tool life

- Process optimization with intermittent sensors

- Process optimization with continuous sensors and adaptive control adaption

Fig. 8.7 Methods of process optimization, acc. to GEBAUER [Geba80]

lead to an uneconomically high tool life. Many tool manufacturers today offer software systems (e.g. CIMSOURCE, CoroGuideTM , WinTool, etc.), with the help of which the determination of the cutting parameters can be combined directly with the selection of a suitable tool. More exact results can be obtained from test reports, which however are valid in most cases for special machining conditions and can only be applied to other machining cases with reservations. Based on these sources, internal cutting data collections are often made as part of a company’s process preparation. These collections make reference to the accumulated practical machining cases in manufacture and also take into consideration the long-term experiences of specialists and skilled workers. Information centres for cutting data, which collect cutting data from industry and research and make these available to companies in a form specifically tailored to them, are another modern phenomenon. Figure 8.8 shows the input mask of a data bank as an example of such a type of cutting parameter selection. These data bank systems contain cutting data for set tool life parameters for a comparable workpiece material/cutting tool material combination and under consideration of geometrical influencing variables. Such a data bank makes it possible to process large collections of previously determined machining data under consideration of many influencing parameters and to provide improved target values.

8.2 Process Monitoring Secure process monitoring has a crucial role in manufacturing technology. Uninterrupted, fully or partially automated manufacturing processes guarantee high productivity and optimal capacity utilization. Process disturbances are, however, unavoidable. Stochastically occurring tool failure, deviations in the composition of

8.2

Process Monitoring

353

Fig. 8.8 Example for technology data organization (Source: CIM GmbH)

the machined material or faulty planning/programming prior to machining are all potential causes for such interruptions. The goal of process monitoring in manufacturing technology is to recognize these process disturbances in a timely fashion and to introduce proper countermeasures. For example, extensive tool, workpiece or machine tool damage can result from tool fractures if the proper countermeasures are not promptly applied, be it manually or automatically (i.e. with the help of monitoring systems). This means, among other things, that the machining process must be monitored in order to minimize consequential costs of process disturbances. Process monitoring systems make it possible to monitor cutting tools. The tools are thereby monitored for sudden tool fractures or overloads and for increasing tool wear. Tool monitoring has been developed to varying extents depending on the various machining processes with defined cutting edge geometries. Monitoring of the drilling process has been widely researched and finds extensive use in industry. There are also different approaches from research and practice for realizing a reliable tool fracture and wear-recognition monitoring system for turning processes. In the case of milling, fracture and wear monitoring is much more costly than in the abovementioned processes due to the more complicated process kinematics. Accordingly, there are only a few monitoring solutions that exist for milling processes. Common to all monitoring systems is that process-relevant information is measured with the help of specialized sensors. The most common sensors and sensor principles today will be introduced in the next chapter.

354

8


8.2.1 Sensors for Process Monitoring A number of sensors are available on the market for monitoring manufacturing processes, of which most can be classified in six physical functional principles. Table 8.2 shows an allocation of typical measurement parameters and the physical functional principles that primarily underlie them. In order to standardize usage and to lend semantic clarity to the terms used, the essential fundamental concepts of metrology will first be defined [Pfei01]: • The measurement parameter is the physical parameter that is to be measured. • Measurement is the execution of activities planned for the quantitative comparison of measurement parameters with the unit. • The measurement result is the value taken from the measurements estimated as the true value of a measurement parameter. • The measurement principle is the physical basis of the measurement. • The measurement method is defined as a special kind of procedure used in the measurement independent of the measurement principle. • A measurement process is the practical application of a measurement principle and a measurement method. A further distinction is drawn between direct and indirect measurement methods. Direct measurement methods record the measurement parameter of interest by means of a comparison with a normal of the same physical parameter. Indirect measurement methods, in order to determine the measurement parameter, capture an auxiliary parameter which stands in a known and describable relation to the measurement parameter [Pfei01]. While for direct methods, it is often difficult to reach the measurement location with suitable sensors, the problem with indirect measurement methods is in the determination of an assignment rule, the validity of which should not be affected Table 8.2 Relation of measurement parameters to physical functional principles Physical functional principles for process description Mechanical (motion, displacement, stiffness)

Thermal (kinetic energy of atoms and molecules) Electric (electric field) Magnetic (magnetic field) Radiation (electromagnetic radiation) Chemical (forces between atomic nuclei and electrons, bond energies of molecules)

Associated measurement parameters Position, acceleration, velocity, force, torque, longitudinal/torsional stress, longitudinal/torsional strain, pressure etc. Temperature, heat flow, specific heat, thermal conductivity etc. Voltage, current, charge, conductivity etc. Permeability, magnetic current etc. Energy, intensity, emission, reflection, permeability etc. Chemical components, concentrations etc.

Process Monitoring

355

ra dia tio n

Fig. 8.9 Process emissions for monitoring of cutting process

He

at

Ac

ou

st

ic

em

Resultant force

is

si

on

Ac

ou

st

ic

Acceleration

8.2

em

iss

io

n

by changes to the process parameters. Yet the advantages and disadvantages of the direct and indirect measurement methods cannot be generalized and must be weighed for every particular case. The following is concentrated on the monitoring of cutting processes. Figure 8.9 shows the process emissions relevant for this area. Numerous sensors use the strain of an object as an auxiliary function for measuring force, torque and acceleration. For this reason, the following sections will examine the workings of strain gauge strips and piezoelectric sensors. 8.2.1.1 Strain Gauges A strain gauge converts a mechanical strain into an electrical resistance change. The OHM resistance R of a metallic conductor is a function of the material (ρ e ), length (L) and cross-section (A) according to the following relation: R = ρe ·

L A

(8.44)

In the simplest case, a strain gauge consists of a resistance wire, which is frequently incorporated into a supporting matrix. The strain gauge is fastened to the shifting position on the measuring body. The adhesive is of great importance, because the strain or compression of the measuring body must be transferred unaltered to the strain gauge via the adhesive connection. Assuming that the material properties of the strain gauge are not changed by the strain (ρ e = constant), the resistance change is dependent only on the length and cross-section change. In general, the relation between strain and relative resistance change can be described by Eq. (8.45). Strain sensitivity k, for example for a constantan strain gauge, amounts

356

8


to k = 2.044. The normal operating range of strain gauges extends up to a relative strain/compression of about 1%. R/R = k · ε

(8.45)

In order to reduce the dimensions in comparison to the single-wire form, usually meander-formed designs are utilized (Fig. 8.10). However, the deflection areas cause a sensitivity of up to 3% transverse to the measurement direction. Strain gauges are not limited to the use of resistance wires. Semiconductor, foil and thin-film materials are also employed. The resistance changes arising during the use of strain gauges are very small and are only a fraction of its total resistance. With the help of a bridge circuit, it is possible to measure the arising resistance changes in the form of a voltage variation UM . For a balanced bridge circuit with R1 = R2 = R3 = R4 = R and a feed voltage U0 , UM = 0. The WHEATSTONE bridge circuit is shown in Fig. 8.11. Also valid is the relation given in Eq. (8.46) UM = U0 ·

(R + R1 ) (R + R4 ) − (R + R2 ) (R + R3 ) (2R + R1 + R2 ) (2R + R3 + R4 )

(8.46)

Depending on whether two or four active strain gauges are used, this relation can be simplified in the form of a quarter, half or full bridge. For a quarter bridge with

Carrier foil

Fig. 8.10 Sketch of a strain gauge strip with measurement meander Connectors

Measurement meander

R3

R1 UM U0

R2

Fig. 8.11 Wiring diagram of a WHEATSTONE bridge

R4

8.2

Process Monitoring

357

R1 = R and R2 = R3 = R4 = 0 and R 0◦ ) and a negative (γf and γp < 0◦ ) cutting part geometry.

9.2

Milling

393

vf

vf

vc vc Down milling

Up milling

Fig. 9.15 Up and down milling

In all milling processes, the cutting edges, as opposed to other processes like turning or drilling, are not constantly being engaged. Rather, at least one cut interruption occurs per cutting edge during each tool rotation. Because of the constant cut interruptions, the contact conditions between the tool and workpiece are of particular importance for the wear properties of the cutting tool materials in addition to the cutting conditions. Different contacts can result depending on the geometric conditions determined by the milling cutter diameter, size of cut and cutting part geometry (Fig. 9.16). It is particularly inauspicious if the cutting edge point that is the most sensitive to impact is the first point to make contact with the workpiece. Such “S-contact” can be

Workpiece Milling cutter

γf > 0° γp > 0°

S Insert

γf > 0° γp < 0°

Reference plane S-Contact

T

γf > 0° γp = 0°

T-Contact

T S

ST-Contact

U

γf < 0° γp < 0°

γf < 0° γp > 0°

U-Contact

Fig. 9.16 Forms of contact in milling

V

V-Contact

γf = 0° γp > 0°

S

V

SV-Contact

394

9 Processes with Rotational Primary Movement

avoided by varying the cutting edge geometry and feed rate accordingly. The most favourable type of contact is “U-contact”, in which the cutting edge point that is furthest removed from the minor and major cutting edges is the first to make contact with the workpiece. All other types of point or line contact are regarded as intermediate stages between S-contact and U-contact with respect to impact sensitivity [Kron54, Beck69, Damm82]. “Helical chip milling” (γf < 0, γp > 0) has become established as a method for improving chip removal, which is often problematic in the case of rotating tools (Fig. 9.17). In comparison to the entry conditions, greater importance is attached to the exit conditions with respect to wear caused by fractures [Kron54, Okus63, Hosh65, Beck69, Lola75, Peke78, Peke79, Köll86]. In the case of tool exit with finite undeformed chip thickness (e.g. in up milling), tensile stresses can arise in the cutting edge in the unencumbered state because of resilience, which lead to cutting edge fracture. This phenomenon must be taken into account when determining milling strategies. The use of tougher cemented carbides may increase the length of tool life until failure, but that alone cannot prevent premature failure in the case of unfavourable exit conditions. In order to improve cutting edge stability, additionally stabilizing protective chamfers are fitted near the corner and on the cutting edge. Cutting interruptions mean thermal and dynamic alternate stresses for the cutting tool material, which can cause comb and parallel cracks and thus lead to cutting edge fracture. The cutting tool materials used must therefore be very tough, temperatureresistant and have high edge strength [Vier70]. For steel-working, high-speed steel and tough cemented carbides of machining application groups P15 to P40 are used; for machining cast iron, NE metals, plastic and hardened steels types K10 to K30 are used. The cutting tool materials used for milling were developed with an eye to increased thermal and mechanical alternate stress and are thus usually not directly comparable with the cutting tool material types used for turning.

Fig. 9.17 Helical chip milling cutter in use (Source: Walter)

9.2

Milling

395

Further developments in cemented carbides and coating technologies have made it currently possible to use coated cemented carbides when milling cast irons as well as steels. When fine milling steels (HB < 300), cermets have also found increasing use. For rough milling grey cast iron, Si3 N4 ceramics can be used successfully with a high material removal rate. Oxidic and mixed ceramics are suitable cutting tool materials for finish milling grey cast iron, chilled cast iron, case-hardened steels, heat-treated steels and hardened steels, while in the case of hardened or high-strength heat-treated steels (> 45 HRC) PCBN is also suitable. Supereutectic Al alloys, fibre-reinforced plastics and the milling of graphite electrodes for spark erosion are typical applications of PCD-coated tools when used for milling.

9.2.1 Process Variants, Specific Characteristics and Tools Milling processes are used most frequently to produce level surfaces (linear feed movement: slab milling). Figure 9.18 shows the most important slab milling processes, which differ by their kinematics and engagement conditions, and designates the slab milling tools associated with them. In practice, milling processes are usually named according to the type and form of milling tools used, e.g. plain milling, end milling, side and face milling, face milling, profile milling etc. Plain milling cutter

Face milling cutter

Shell end mill

Side and face milling cutter

End milling cutter

Fig. 9.18 Slab milling processes

9.2.1.1 Face Milling In the case of face milling, the width of cut ae is much larger than the depth of cut ap , and the workpiece surface is created by the minor cutting edge. If the lead angle is κr = 90◦ , this milling process is also referred to as edge milling. In this case, the workpiece surface is created with both the minor and major cutting edges.

396

9 Processes with Rotational Primary Movement vc

−Ff

vf

−Fcn ϕ

n

−Fc −Ffn

−Fa

ϕ = 0°

Fig. 9.19 Cutting force components at head face milling, according to KAMM [Kamm77]

According to DIN 6584, the resultant force F involved in milling can be broken down into an active force Fa on the working plane and a passive force Fp perpendicular to the working plane (see also Sect. 3.8.1). The direction of the active force Fa depends on the feed direction angle ϕ. The components of the active force can be related to the direction of the cutting speed vc (cutting force Fc and cutting normal force Fcn ) or to the direction of the feed velocity vf (feed force Ff and feed normal force Ffn ) as in Fig. 9.19. KIENZLE’s resultant force equation [Kien52] can also be applied to milling. For the components of the resultant force F – cutting force Fc , cutting normal force Fcn and passive force Fp , we have:

1−mi

i | = ki · b · |h| |F

(9.4)

with: i = c, cn and p Due to the wide range of undeformed chip thicknesses covered by milling, the K IENZLE relation is only valid for certain areas. The range of undeformed chip thickness of 0.001 < h < 1.0 mm is subdivided into three sections (Fig. 9.20). For each range, a line can be ascertained, which is defined by the specific resultant force ki and the exponent mi :

−mi0,01 for 0.001 < h < 0.01 mm |ki | = ki1.0,01 · |h|

−mi0,1 for 0.01 < h < 0.1 mm |ki | = ki1.0,1 · |h|

−mi for 0.1 < h < 1.0 mm |ki | = ki1.1 · |h|

(9.5)

with: i = c, cn and p As long as the characteristic values are known for the existing marginal conditions (material, cutting tool material and cutting conditions), the respective resultant

9.2

Milling

397

102 8 6

kc1.0.01 = 75 MPa 1– mc 0.01 = 0.150

kc / (GPa)

4 kc1.0.1 = 998 MPa 1 – mc 0.1 = 0.440

2 101 8 6

kc1.1 = 2100 MPa 1– mc = 0.764

4 2 Range 3 1 10−3

2

4

Range 2 6 8

10−2

2

4

Range 1 6 8

10−1

γp

αf

2

4

6 8 100

λs

κr

h / (mm)

Material: Ck45N Cutting tool material: HW-P25 vc = 190 m/min

γf

αp

εr

−4° −7° 6° 23° −6° 75° 90°

Fig. 9.20 Specific cutting force in face slab milling [Kamm77]

force component Fi can be calculated for milling. To estimate the resultant force however, characteristic values are often utilized that have been established in turning processes. To mill very small, level and right-angled surfaces, grooves with square crosssections and long slots, solid face milling cutters made of HSS or cemented carbide are used, beyond a tool diameter of D = 10−16 mm tools with clamped cemented carbide indexable inserts and in case of increased requirements on surface quality, dimensional accuracy and performance milling cutters with soldered cemented carbide cutting edges [Sack76]. The size and number of teeth of the milling tool are selected based on the dimensions of the workpiece surface to be machined and the drive capacity of the machine. The tooth pitch of the tool depends on the form and size of the tool, on the available machine power and on the chip formation of the material. Short-breaking chips require a small chip space and thus a small pitch. Large cutter heads for cast iron machining can thus be fitted with up to 200 inserts. In order to prevent chattering of the tool/workpiece/machine system with the frequency of the cutting edge engagement, face milling heads are partly manufactured with an uneven pitch on the periphery. Face milling heads up to a diameter of D = 250 mm are mounted in the usual fashion on the tool spindle. Larger face milling cutters are designed in two parts because of their great weight for better handling during tool change. The base body remains on the spindle during tool change, so only the ring with the clamped cutting edges is changed.

398


Fig. 9.21 Face milling head with tool cartridges (Source: Kennametal, Hertel)

Another potential way to increase efficiency and universality is the use of face milling heads with tool cartridges (Fig. 9.21). Depending on the requirements, tool cartridges can be inserted in a base body that can receive various kinds of indexable insert (three-corner, four-corner, round), sizes and geometries (e.g. positive, negative, κr = 90◦ , κr = 75◦ , with moulded chip breakers etc.). In general, steel-working employs positive cutting part geometries and, to improve chip removal, helical chip geometries. In the case of welded constructions or larger material inhomogeneities, a negative cutting part geometry is more advantageous for preventing cutting edge fracture. The same is true for machining materials with high strength and toughness. Milling with cutting ceramics is generally performed with a negative cutting part geometry. In order to avoid re-cutting the face milling cutter due to elastic form changes in the overall system, the milling axis can be tilted by 0.5–1◦ . However, this sets the feed direction. The tool cutting edge angle amounts to κr = 45−75◦ in face milling (special case: corner milling κr = 90◦ ). It affects to a large extent the size of the active and passive forces and thus the stability of the milling process, especially when machining thinwalled parts (e.g. welded gearboxes) or in milling operations on milling and boring machines with a widely projecting spindle. Generally, the face milling cutting conditions are selected lower than in turning processes. Smaller cross-sections of undeformed chip are chosen in particular in order to keep the dynamic stress on the cutting tool materials low and to prevent tool fracture. Face milling is used both for pre-machining and also increasingly for finishing purposes. Finishing with geometrically defined cutting edges is becoming increasingly important because of the potential of single-machine processing. Finish face

9.2

Milling

399

milling is especially used as a finishing process for large even surfaces with special surface quality and smoothness requirements when other finishing methods (e.g. grinding or shaving) are uneconomical or impossible. Such machining problems usually arise in heavy machine construction, e.g. when producing joining surfaces, machine tables and guideways on machine tools and when milling sealing faces in engine and turbine construction. Indexable inserts for finishing have an active minor cutting edge, i.e. in these tools the cutting edge angle is κ r = 0◦ , so the chamfer of the minor cutting edge lies parallel to the workpiece surface. The chamfer length of the minor cutting edge is generally Lsa = 2−3 mm, in the case of special wide finish milling tools Lsa = 10 −15 mm. The feed per tooth should not exceed 2/3 of the length of the active minor cutting edge nad. There are three kinds of finishing tools (Fig. 9.22): • Conventional finishing face milling cutter which work with small depths of cut and feeds per tooth and are fitted with a large number of teeth. • Wide finishing face milling cutters, which are equipped with a small amount of teeth (1–5) and work with very small depths of cut and high feeds (Fig. 9.23). In the case of such tools, the minor cutting edge nads are equipped with large radii to simplify tool pre-adjustment. This helps to obtain a very good surface quality

Finish milling

Wide finish milling

Combined process: Finish milling and wide finish milling

Fig. 9.22 Processes for fine milling (Source: Siemens)

Principle of fine milling with ceramics Insert ap = 0.05 mm

Workpiece

fz

Single-tooth milling cutter in use

Fig. 9.23 Wide finish milling (fine milling) with ceramics (Source: Feldmühle)

400


in cutting, similarly to shell turning. However, the jerk forces are larger than in conventional finishing face milling, which can result in an axial displacement of the tool. Ceramics are the predominate cutting tool materials in this case. • Face milling with finishing cutting edges and wide finishing cutting edges that combine the advantages of both methods. The tool is in this case only fitted with one or two wide finishing cutting edges that are radially set back and that protrude axially by 0.03–0.05 mm to produce a high surface quality. The width of the finishing cutting edges should correspond to about one and one half times the feed per rotation. For finishing, the pre-adjustment of the cutting edge has increased significance. If no special, finely adjustable tool holding fixtures are used, all finishing face milling tools should be ground and lapped prior to use on the machine tool and after fitting with the cutting edges in order to obtain the necessary face and run-out accuracy (< 5 μm) corresponding to the required workpiece surface quality. Faulty insert pre-adjustment increases kinematic roughness and shortens the tool life sometimes considerably. The cutting speed is selected high in finishing (e.g. up to vc = 300 m/min in steel-working with cemented carbide) in order to obtain a high surface quality. In the case of steel, surface finishes of Rt = 5−10 μm are obtained, in the case of grey cast iron Rt = 1−5 μm. 9.2.1.2 Peripheral Milling In the case of peripheral milling, the workpiece surface is created by the major cutting edge. There is peripheral down milling and peripheral up milling. In peripheral down milling, the cutting force acts upon the workpiece (Fig. 9.24), while in peripheral up milling it is directed away from the workpiece so that an unstable workpiece (e.g. a thin steel sheet) can be lifted from the clamping surface or be induced to chattering.

Down milling

Up milling

vf

vf n

n

fz

fz

fz Ff

Fc Ff

ae Fc

fz Fa

Fig. 9.24 Peripheral up- and down milling

Vorschub Fa

9.2

Milling

401

In the case of peripheral down milling, a table feed drive that is free of play is necessary in order to prevent vibrations and impacts. While in the case of peripheral down milling the lead takes place with an approximately full cross-section of undeformed chip, in up milling the cross-section of undeformed chip is slowly increased. This can lead to material compression and thus to the formation of a poor surface. Besides the usual HSS tools, cemented carbide peripheral milling cutters or plain milling cutters are finding increasing use. When the cutting edges have a coaxial adjustment, high dynamic stresses come into play because one whole cutting edge enters into or exits from the material at a time. In the case of helically toothed tools, the dynamic load can be reduced, but an axial force arises then which can lead to tool or workpiece displacement. Pitch-induced axial forces can be compensated by mutually bracing a right-inclined and a left-inclined plain milling cutter of the same design (Fig. 9.25). This disadvantage can be overcome with a double helical gearing with opposite pitch. Such tools are very expensive both to acquire and to prepare however (Fig. 9.25). Should sharp-edged profiles with good dimensional and formal accuracy be prepared, combined peripheral face milling cutters or plain milling cutters are used (Fig. 9.26), which are relief-ground on the front face of all cutting edges (formation of a tool orthogonal clearance).

Fa

n

Fa

Fig. 9.25 Combined plain milling cutter with opposite hand helix

Fig. 9.26 Modular shell end mill (Source: Sandvik Coromant)

9.2.1.3 End Milling End milling is a continuous peripheral face milling process which uses an end milling cutter. This process is advantageous when manufacturing mould surfaces

402

9 Processes with Rotational Primary Movement End mill with morse taper End milling cutter with parallel shank Right-hand direction of cut with right-hand helix End milling cutter with morse taper Right-hand direction of cut with left-hand helix Die-sinking milling cutter with parallel shank Right-hand direction of cut with right-hand helix Conical die-sinking milling cutter Right hand cutting with right hand helix

Fig. 9.27 End milling cutter

(e.g. in die construction) as well as forming grooves, pockets, slots and cavities of all kinds and sizes. End milling cutters have to be designed in many cases with a large degree of slenderness (l/D > 5–10) depending on the application (e.g. milling deep engravings in dies and moulds). This causes on the one hand, depending on the contact and engagement conditions, chatter vibrations during the process, which can lead to increased wear via fracture, especially in the case of hard, brittle cutting tool materials. Additionally, both chattering and bending of slender tools lead to dimensional and shape inaccuracies in the components. Measures taken to avoid these phenomena should be sought in an optimization of the tool and cutting part geometry, engagement conditions and milling strategy as well as of the cutting conditions [Schr74, Köni80, Hann83, Köll86]. End mills correspond to shell end mills in their construction; for clamping, they are equipped with a parallel shank (with side-clamping and/or fastening thread) or with a taper shank (Morse taper or steep-angle taper; sometimes with fastening thread). A distinction is drawn between right-cutting and left-cutting tools as well as between right-hand spiral, left-hand spiral and straight-toothed tools (Fig. 9.27). The mill form can be designed cylindrically, conically or as a custom design depending on the machining task. The front face of the tool is generally round or half-round; in the case of tools capable of drilling the face cutting edges must reach as far as the tool centre. HSS end milling cutters are classified into tool applications groups in accordance with DIN 1836 depending on the material to be machined (Fig. 9.28). Profiling of the cutting edges in the case of roughing tools leads to a division of the chips into smaller chips. The advantages of these chip dividers include improved chip removal and cutting fluid access as well as reduced stress on the cutting edges (Fig. 9.29). The design of individual milling cutter geometries and chip divider forms differ depending on the manufacturer. In principle, all the cutting tool materials are potentially applicable in end milling, depending on the selection criteria regarding workpiece materials and

9.2

Milling Group of application

403

Field of application

N

Machining of materials with normal strength and hardness

H

Machining of hard, hard tough and/ or short-chipping materials

W

Machining of soft tough and/ or long-chipping materials

Tool

Fig. 9.28 Tool application groups

Profiled cuttting edges Profile

Group N

Group H

Flat profile (F)

Rounded profile (R)

Fig. 9.29 Cutting profiles of roughing-end milling cutters (Source: Fette)

stability. High speed steel is still predominately used – often coated – as well as cemented carbides. Besides solid steel tools, tools with soldered cutting edges and clamped or bolted indexable inserts are used (Fig. 9.30). A more flexible adjustment of the cutting tool material to the machining task is possible by using tools with indexable inserts. Especially mentionable in this context is the use of different cutting tool materials in a single tool. This can be advantageous, for example, in the case of ball path milling cutters, which are subject to highly diverse stresses along the cutting edges. 9.2.1.4 Profile Milling Profile milling is milling with forming tools to produce profiled surfaces, e.g. for milling grooves, radii, gear wheels and gear racks as well as guideways.

404


Fig. 9.30 End milling cutter with inserts (Source: Sandvik Coromant)

Form cutter

Gang cutter

Fig. 9.31 Profile milling cutter

Profile milling tools are adjusted to the form of the profile to be produced. In most cases, a peripheral face milling process is the result. As shown in Fig. 9.31, the tools are designed in one part (form milling cutters) or in multiple parts (gang milling cutters). Profile milling cutters are in many cases manufactured as solid HSS tools due to the favourable machinability and inexpensive price. Increasingly however, cemented carbide indexable inserts or soldered cemented carbide cutting edges are being used (Figs. 9.32 and 9.33). 9.2.1.5 Hobbing In almost all areas of technology, gears are employed as components of an exact and effective transmission of motion. For high-precision gears, the gears are

9.2

Milling

405

Fig. 9.32 Gang cutter for machining of grey iron profiles (Source: Walter) Fig. 9.33 Gang cutter for machining of machine beds (Source: Walter)

manufactured primarily by machining. Because of its high efficiency, hobbing is the dominant machining process for producing externally toothed cylindrical gears. In hobbing, the coupling of a worm with a worm gear is simulated, whereby a worm interrupted by gashes represents the tool and the worm gear represents the workpiece to be manufactured. The kinematics of the process will be explained briefly with the help of Fig. 9.34. The rotary movements of the hob and the gear serve to remove the chips. Depending on the hobbing method, superimposed over these are translatory motions of the tool in the axial and tangential directions as well as in the radial direction in order to reach the depth of cutting. This results in the hobbing methods axial, radial-axial,

fa nw

vc

Fig. 9.34 Kinematics of hobbing

406


tangential, and diagonal hobbing. In industrial production, axial hobbing is currently the most commonly employed process. The feed in one direction is defined as the distance travelled per workpiece rotation. We distinguish between climb and conventional cutting by means of the direction of the axial feed fa . In climb cutting, the cutting speed and axial feed motion are directed opposite relative to the workpiece. For wear-related reasons, climb cutting is generally preferred, especially in dry hobbing. The traverse path of the hob can be subdivided into the tool inlet and outlet phases as well as that of full cut, where only in full cut does one obtain the theoretical maximum circular cut lengths and material removal rates (Fig. 9.35). There are two further process variants used to machine helical gearings, hobbing in the same direction and in the opposite direction, whereby the pitches of the tool and working gear are aligned in the same direction or in opposite directions. In order to distribute the tool wear arising during the process evenly along the hob for more a efficient use of the tool, the hob can be shifted along its axis in discrete steps. This displacement takes place as a rule after each machined workpiece and is also called “shifting”. The parameters of a hob can be seen in Fig. 9.35, which shows the tool in the machining sequence. The hob is a cylindrical screw interrupted by chip flutes which bring about the gash. The number of worm threads on the cylinder determines the thread number of the hob. The hob teeth are shaped so that tool orthogonal clearances and the potential of regrinding the rake face without altering the tooth profile are created. The pivoting angle or lead angle of the hob results from the direction and size of the helix angle and the pitch angle of the hob worm. Hobs are classified into three different groups with reference to their construction type (Fig. 9.36). Solid steel hobs are manufactured from solid material, whereby the entire body must be constructed from high-quality HSS or cemented carbide. Inserted blade hob and cutters with indexable inserts on the other hand consist of a base body made of a more inexpensive material. These hobs are especially suited to manufacturing gears Section A-A

β2 A da2

vc Recess

fa b

η = β2 γ0

fa

±

T Shift

Entry

A da0

Fig. 9.35 Terms of the hobbing process

Full cut

9.2

Milling 500

407 500

500

mm

Diameter da0

400 300

300

(HSS)

200 100 0

160

140 (HM) or (HSS) 20

60

Full steel hob

Inserted blade hob

Cutter with indexable inserts

Fig. 9.36 Hob types (Source: Fette, Saacke, Saazor)

with large diameters and large modules. Besides the potential of realizing larger constructive tool orthogonal clearances, the inserted blade hob also has a relatively large potential usable tooth thickness (large number of regrinds). Fastening is done with lateral clamp rings. In the case of cutters with indexable inserts, the tool cutting edge, which is made of cemented carbide, is used upto four times and not reground. This hob type is only suitable for workpieces beyond a modulus of 5 mm because of the resulting poorer cutting quality. With respect to the selection of the substrate/coating system, a sufficient amount of substrate toughness is of especial importance in hobbing because of the interrupted cut. At the same time, the tool must have as much wear resistance as possible against abrasive and thermal wear mechanisms in conjunction with a suitable hard material coating. By the establishment of (Ti,Al)N coatings is it possible to perform dry machining processes at cutting speeds up to vc = 200 m/min, even with tool systems based on HSS [Wink05]. More and more, coated gear tools are reconditioned. The coating is removed, the tool reground and finally recoated on the rake face after the tool operating life is reached [Klei03]. In this way, we can work with consistently high cutting speeds in all tool cycles, and the tool’s operating life can be held constant. For reasons of accuracy, ground cutters cannot be over-coated more than 5–10 times because of the coating application on the already coated rake face. Figure 9.37 clarifies the hobbing process by looking at the creation of a tooth gap. Due to the process kinematics and the resulting shifting between the tool and workpiece, the material of a tooth gap is machined in the successive engagements (hobbing positions) of the individual teeth of a hob thread as can be seen in the sketch in the bottom left of the illustration. The evolvents on the tooth of the workpiece are approximated by profiling cuts. Every cutting tooth makes a cut after one tool rotation in a further tooth gap determined by the hob thread number, but in the same hobbing position; i.e. it is always removed one chip with the same cross-section.

408

9 Processes with Rotational Primary Movement Section A Cross section of undeformed chip

Section B

Generated cut B Involute Generated cut deviation

Hob tooth

Fig. 9.37 Ratio of engagement in the hobbing process

Due to the differing penetrations between the hob and the workpiece in the particular hobbing positions, variously thick and variously formed chips result (Fig. 9.38). During the first tooth engagements, a large amount of the gap volume is machined, so that here, especially shortly before the middle position, exist the largest crosssections of undeformed chips. In the following hobbing positions, the tooth gap is mostly profiled and the cross-sections of undeformed chip are reduced. The penetration areas are highlighted in the image. During hobbing, as in other gears manufacturing processes, changes to the structurally given tool orthogonal clearance and tool orthogonal rake angles occur during the cutting process [Sand72, Sulz73, Sulz74, Köni79]. Figure 9.39 illuminates the cause of this. At the cutting point under consideration on a certain hobbing position, the effective cutting speed vr results from the cutting speed of the hob vc and the hobbing speed vA . On the entering tool cutting edge (Fig. 9.39), this leads to an enlargement of the structurally given tool orthogonal clearance but simultaneously to a reduction of the tool orthogonal rake angle. This means that, during the cutting process, the

Rake face Hob profile

Hob profile

Hob profile


Hob profile

vc

Fig. 9.38 Chip formation in the hobbing process

Upper chip side

3mm

9.2

Milling

409 cut C-C C

vA

vA

C vc vc

vA

vr

vc

vr

vr

workpiece vc

γeff

γeff

nw

Tool

EF

AF

αeff

αeff αkon

αkon

Fig. 9.39 Relative velocity and effective cutting geometry

effective tool orthogonal clearance is larger on the entering face than the structurally given one. The effective tool orthogonal rake angle, on the other hand, is smaller. There are different ratios on the exiting face, upon which the effective tool orthogonal clearance is small and the effective tool orthogonal rake angle is larger than the corresponding structurally determined angles. A penetration calculation can be applied to simulate the cutting process [Wink05]. In the penetration calculation, the workpiece is analyzed into a certain number of parallel planes. By recreating the axis motions of the hobbing machine, the cutting path of the hob teeth are generated so that they cut the workpiece planes. From the intersection we can calculate a penetration range that corresponds to the chip geometry of the respective hobbing position in the process [Weck02, Weck03]. Figure 9.40 shows a simulated chip geometry in a certain hobbing position. The 3D chip diagram reproduces the distribution of cross-sections of undeformed chips along the uncoiled cutting edge along the cutting arc, whereby the beginning of the cut is in the foreground. The calculated chip geometries form the basic data of a simulation-supported assessment of the stresses affecting the hob during the machining process. It is clear that the chip thickness (top chip thickness) in the top region of the hob can be much larger than in the edge region. In addition to cemented carbides, PM-HSS cutting tool materials have also become established as substrate materials for hobs in dry hobbing processes. Since both cutting tool materials are in competition with one another, Fig. 9.41 shows the advantages and disadvantages of both tool systems. The advantages are designated with bright points, the disadvantages by dark points.


Cut-b end

handli ng

410

h

Length of the incoming flank

Length of the outbound flank

Head

Cutting edge handling

Fig. 9.40 Tension geometry in hobbing

HSS

HM

30 µm Source: Samputensili

Tool original price Preparation charge Wear resistance Cutting speed Chip thickness Hardiness Process reliability

Fig. 9.41 Cutting tool materials for dry hobs

One of the main disadvantages of cemented carbide hobs in comparison to the HSS variant is the higher cost of acquiring the tools. They are generally at least three times more expensive. Because of the higher tool costs, it is only economical to use them if a significantly higher productivity or longer tool life is obtainable. Because of the much higher wear resistance of cemented carbide, cemented carbide tools can realize clearly higher cutting speeds. However, PM-HSS hobs can realize larger maximum head chip thicknesses. Given the high tool price, the use

9.2

Milling

411

of cemented carbide hobs is uneconomical in this case, provided the tool life is not significantly higher. In large batch production, process safety is particularly important. While HSS hobs, not least because of their high toughness, make a relatively high level process safety when their wear behaviour is monitored [Coop99, Kölk99], cutting edge fractures occur occasionally in the case of cemented carbide hobs [Kloc99a, Sulz00]. If this occurs only sporadically, such fractures cannot be detected with process monitoring systems and lead to increased tool wear. This problematizes an economical use of cemented carbide hobs given the multiple reconditioning of the tool that is required. Although dry hobbing used to always be performed with cemented carbide tools, PM-HSS tools have become continuously more popular in the last several years because of their superior toughness properties and lower tool costs [Coop99].

9.2.1.6 Skive Hobbing Skive hobbing is a continuous process using geometrically defined cutting edges for machining pre-milled gear teeth in a hardened state. Primarily, deformation caused by the heat treatment is removed and the surface quality improved. The process kinematics are identical to those of hobbing. The concept of skive hobbing is derived from the “peeling cut”. Hard finishing with geometrically defined cutting edges requires that small cross-sections of undeformed chip are selected. The skive hobs are designed with a negative top tool orthogonal rake angle, which acts as a negative inclination angle on the tooth flanks. This guarantees that the first contact between the cutting tooth and the workpiece surface is not made directly on the cutting edge but in the stable cutting part area behind it (see also contact conditions, Fig. 9.16). In order to execute a skive hobbing operation, the tooth gaps of the pre-milled gear teeth must be prepared to such an extent that the top of the skive hob does not engage and only the flank cutting edge is cut. Otherwise, the danger of fracture is increased [Faul86]. The tooth base can be free milled in two different ways: by pre-milling with tools corresponding to reference profile II acc. to DIN 3972 or by pre-milling with protuberance. In case 1 the pre-processing of the gaps is done with hobs of reference profile III, a sharp edge appears in the tooth base after skive hobbing, which can have a negative effect on tooth base strength. On the other hand, the tooth gaps pre-milled with protuberance have a rounded transition in the tooth base after skive hobbing (Fig. 9.42). In skive hobbing, it is very important for the sake of consistent cutter wear and the output that the hob tooth is exactly positioned or “centred” in the tooth gap. After centring, the same amount of material is removed on both workpiece flanks (right and left flanks) in the ideal machining case. Centring is made more difficult by the fact that the flank allowance fluctuates along the workpiece periphery and the tooth width due to pre-gear cutting deficiencies, faulty alignment for skive hobbing or as a result of deformation due to hardening.

412


Negative head rake angle

γK

Hobbing

1

Roughen BP III

Hobbing

2

Roughen with Protuberance

Fig. 9.42 Gap profiles of different rough cutters

Although only very thin chips are removed during the process, very high resultant forces are in play, which are inconstant – especially in the entrance and exit areas – and also change their signs. For this reason, a high static and dynamic stiffness must be required of the hobbing machine in addition to geometric and kinematic accuracy [Faul86]. In the case of skive hobbing, the obtainable length of tool life is depends greatly on the cutting edge, workpiece geometry, cutting parameters and the hardness of the workpiece to be machined. Favourable wear behaviour is exhibited by the ultrafinegrain cemented carbides of ISO application range K10 to K20 in conjunction with a hard material coating [Kais92, Köni95]. Skive hobbing is used as a finishing process or can also serve as a preparation process for a subsequent grinding operation. The grinding costs are thereby levelled by lowering the amount of deformation due to hardening prior to the subsequent hob grinding process. This production sequence is especially of interest for large-module workpieces (larger module 10 mm). The limits of quality in finishing are basically determined by the feed marks and profiling cut deviations (Fig. 9.37) characteristic of hobbing. With the help of a honing operation following the skive hobbing process, they can be removed at least in the case of small gears. Furthermore, gears that cannot be ground because of their geometry (large grinding wheel diameters) are also made by skive hobbing [Koep94].

9.2

Milling

413

9.2.1.7 Turn Milling Turn milling is a machining process in which the principles of turning and milling are combined in such a way that (generally) rotation-symmetric workpieces are machined with an inserted-tooth cutter on a rotating workpiece. The self-propelled cutter works like a turning tool with a longitudinal feed parallel to the workpiece. In principle, there are two process variants in turn milling. If the tool axis and the rotation axis of the workpiece are arranged perpendicularly, it is called orthogonal turn milling (Fig. 9.43, left). Alternately, in axis-parallel turn milling, both rotation axes are parallel to each other as shown in Fig. 9.43, right. Due to the arrangement of the milling tool, axis-parallel turn milling makes both internal and external machining possible. In the case of orthogonal turn milling, the milling cutter executes a screw-shaped motion relative to the workpiece due to the axial feed movement and the rotation of the shaft. In accordance with this characteristic, DIN8589c defines turn milling as a variant of screw milling. We also draw a distinction between centric and eccentric turn milling (Fig. 9.44). In centric turn milling, the workpiece axis and tool axis intersect at one point, while in the case of eccentric turn milling they are offset relative to each other by a certain amount, eccentricity e. High surface quality (low facet formation) can be obtained in turn milling only with extreme milling cutter and shaft speed settings. This can be realized with a centric cutter setting (e = 0) (Fig. 9.44): • by lowering the feed per workpiece rotation – and thus the cutter engagement – to a few 1/100 mm. We work simultaneously with very rapid workpiece rotation and slow cutter rotation. In this way, we obtain a process very similar to turning. • by very slow workpiece rotation and fast cutter rotation [Köni84a, Köni84b, Köni85, Köni86]. The workpiece speed can be lowered so far that the quickly rotating cutter edges successively cut towards the target radius almost continuously. For the sake of high cutting performance, the feed per workpiece rotation – and thus the engagement width of the milling cutter – should be maximized.

Workpiece nw

ap

nw

nF vfax

Fig. 9.43 Axle rotation in turn milling

Tool

nF vfax

414

9 Processes with Rotational Primary Movement f

f

r

r

r

f

r

e dF

e=0

ls´

e < dF /2 – ls´

e = dF /2 – l´s

e = dF /2

Fig. 9.44 Centric and eccentric turn milling

In order to produce cylindrical surfaces, a broad-tool finish tool geometry with a face cutting edge lead angle of κ r = 0◦ is unavoidable, since from now on the face cutting edges create the workpiece target radius with their full length. The feed per workpiece rotation fax is thereby limited to the length of the face cutting edge Ls [Köni84b]. Under these conditions, the process is comparable to conventional face milling. In the second case, the axial feed can be increased beyond the limit set by the length of the face cutting edge by shifting the milling cutter by the eccentricity e without hazarding cylindricity faults. In the case of an eccentricity of approximately e= the maximum of feed of

dF − Ls 2

fax = 2

2 D − e2 2

(9.6)

(9.7)

is reached. Further augmentation of the eccentricity again reduces the adjustable axial feed until finally no cylindrical surfaces can be produced [Köni84a, Köni84b]. A great advantage of centric turn milling is that sharp-edged workpiece shoulders can be created (Fig. 9.43). In eccentric turn milling, this is impossible for kinematic reasons. Shoulders are principally rounded in this case. In addition, the surface quality is not as good as in the centric process variant. Centric turn milling is thus of great interest for practical process design despite lower performance values. It is recommended therefore to implement roughing operations with an eccentrically and finishing operations with a centrically positioned milling cutter [Köni86]. Superimposition of the rotary motions of the workpiece and tool as well as of the feed motion results in highly complex cutting kinematics [Köni84a, Köni84b, Köni85]. Figure 9.45 shows an example of this process.

9.2

Milling

415 Added front side cutting tool

Range cutting edges

nF

Cutter

nW Workpiece Chip-chambers

Fig. 9.45 The turn milling process (Source: Wohlenberg, Seco)

As opposed to conventional face milling, the face cutting edges participate actively in the machining process and each produce one element of the workpiece surface, as the material is moved into the front face of the milling cutter by means of the workpiece rotation. The facet-like surface typical of turn milling can clearly be seen (Fig. 9.45). In the case of the special tool used here, the face cutting edges are extended with additional inserts in order to maximize the feed. To avoid cylindricity faults, these also have a lead angle of κ r = 0◦ . An effective and economical cutting process is also possible by using conventional milling heads. In steel machining in a smooth longitudinal cut, TiN-coated cemented carbides have proven to be particularly effective. The cutting conditions are generally somewhat higher than those of conventional face milling with TiN-coated cemented carbides [Wand92]. Steel materials that are difficult to machine can on the other hand be very effectively machined with TiCN-coated cemented carbides by means of turn milling [Stal94]. The particular advantages of turn milling are safe chip fracture when machining long-chipping materials as well as the high cutting performance in a smooth longitudinal cut at low workpiece rotary speeds. This is especially significant in the case of large or unbalanced parts. Turn milling can be used to create circular profile curves (e.g. on camshafts) and cylinder forms that run eccentrically to the component rotary axis (e.g. lifting pins on crankshafts). Lifting pins and crankshaft cheeks can be machined with externally or internally geared tools (Fig. 9.46). A flexible full-range processing of such components in one clamping is possible by exploiting the additional rotary axis for drilling, slab milling or thread cutting [Köni84b, Kauf92]. Beyond that, various mould parts can be created (e.g. longitudinal grooves) by adjusting the shaft rotary motions with the three possible, mutually independent feed motions (axial: fax , radial: ap , tangential: e) [Köni84a].

416


Fig. 9.46 Crankshaft machining with rotary milling (Source: Walter)

Extruder screws are almost exclusively produced by turn milling or planetary thread milling [Sten64]. Spherical actuators for pipeline valves as well as ellipse and eccentric forms can also be manufactured via turn milling.

9.3 Drilling The term drilling signifies the machining method with a rotary main motion in which the tool is allowed only one feed motion in the direction of the tool rotary axis. The most important process variants are shown in Fig. 9.47 along with the common respective directions of motion [DIN8589b]. The peculiarities of drilling include: • • • • •

a cutting speed that falls to zero towards the drill centre, difficult chip removal, unfavourable heat distribution at the action point, increased wear at the sharp-edged cutting edges and friction of the lands against the drill hole wall.

Different objectives with respect to material removal rate, drill depth, dimensional accuracy and surface quality have led to the development of a series of different drilling processes, which will be explored in more detail in the following.

9.3.1 Profile Counterboring Counterboring differs from drilling basically by the fact that one does not drill into solid material but rather a prepared hole which has, for example, been drilled or punched, is counterbored to a dimension smaller than specified or to finished dimensions. DIN 8589-2 distinguishes between to process variants: • planar countersinking • planar insertion

9.3

Drilling

417

Trepanning

Profile center drilling

Reaming

Tapping Drill out

Profile reaming

Profile center drilling

Non-circular drilling

Profile counterboring

Fig. 9.47 Process variants of drilling, according to DIN 8589

In the case of planar countersinking, an even surface protruding on the workpiece lying perpendicular to the rotary axis of the cutting motion is created. Planar insertion on the other hand is used to produce an even surface recessed in the workpiece lying perpendicular to the rotary axis of the cutting motion, whereby an internal cylindrical surface is formed simultaneously. For planar insertion of pre-cast or predrilled drill holes, usually three-lip spiral countersinks are used. In comparison to spiral drills, the three-blade, screw threaded design gives the spiral countersink much higher stiffness and thus leads to much higher working precision. For deburring, chamfering and inserting the seat of spherical screw heads, HSS countersinks are used that are manufactured according to standard with angles of taper σ of 60, 90 and 120◦ (Fig. 9.48). For manufacturing drill holes of fastening screws, screw head counterbores of HSS or cemented carbide design are employed, the shape and size of which is adjusted to the respective standardized screw type. Piloted counterbores are suitable not only for planar insertion but also for planar countersinking the front faces of eyes and hubs. Automation in manufacturing often presumes the use of tools that are adjusted to a particular machining task. Such special tools, to which the form countersink in Fig. 9.48 belongs, can shorten production times quite considerably, since several working cycles can be consolidated in one spindle stroke. For example, it is customary in mass production for a screw connection to drill one through-hole with a form tool, then to counterbore the cylinder for the head of a hexagon socket screw and finally to chamfer the drill hole edge.

418


Fig. 9.48 Counterboring tools

ο

Point countersink

Pilot countersink

Form countersink

9.3.2 Rotary Drilling 9.3.2.1 Centre Drilling The spiral drill occupies the position of greatest importance among drilling tools, as it is the most important tool for creating cylindrical drill holes from solid material or for enlarging a preset drill hole diameter in drilling out. It is estimated that it takes up 20–25% of machining operations, and it is today the machining tool that is produced in the largest numbers and is the most widespread [Häus79, Tika93]. Simply put, the spiral drill is composed of the shaft and the cutting part (Fig. 9.49). Only a more exact consideration reveals the complex geometric formation, especially of the drill bit. The about 150 grind types [Tika93] and numerous material-specific drill profiles represent the attempt to do justice to multifarious machining task with respect to quality and performance. For some time, analytical models have been developed to help calculate tool stresses. This requires an

Cutting part

Drill diameter

Neck

Point length

Taper shank

Cone length Neck length

Cutting length Flute length Total length

Fig. 9.49 Spiral drill with taper shank, acc. to DIN 1412

Straight shank

9.3

Drilling

419

Major cutting edge Hauptschneide Chisel edge Querschneide Freifläche Flank face

βff

αff

α

γff δ

Fig. 9.50 Cutting part geometry of a spiral drill

acquaintance with the cutting part geometry, the kinematics of the drilling process and the stresses arising during that process. Figure 9.50 shows the cutting part geometry of a spiral drill. Since according to the definition the major cutting edges point in the feed direction, the chisel edge is also part of the major cutting edge, although it hardly cuts due to its highly negative tool orthogonal rake angle, but rather deforms the material plastically and forces it to the major cutting edge. The shape and pitch of the chip flutes determine the size of the tool orthogonal rake angle γo , which is not constant along the major cutting edge but decreases from its highest value on the corner (γf ) towards the drill centre and becomes negative in the transition to the chisel edge (Fig. 9.51). The only differentiating factor however that is used is the side rake angle γf , which is identical to the helix angle δ with sufficient accuracy. The latter is varied because of the differing chip fracture of different materials and categorized into the main drill groups N for normal materials, H for hard materials and S for soft materials (Fig. 9.52) [DIN 1414a]. By means of the interaction of the cutting motion (rotation) and the feed motion, the tool cutting edge moves along a screw line. Taking the cutting conditions (effective speed) into consideration, the tool orthogonal clearance must be selected such that the effective rake angle is positive. An upper boundary of the clearance is also given however by the weakening of the cutting part and the rattling tendency. To machine steel materials, usually a point angle of 118◦ is chosen. Point angles of 90◦ are used for drilling hard, usually heavily wearing plastics in order to make the transition from the major cutting edge to the lands less sharp than is the case with very large point angles, thus reducing edge dulling correspondingly. Point angles of 130◦ result in improved free drilling in the case of resilient (“clamping”) materials; moreover, the chip clogging problem can be countered with a further enlargement to a point angle of 140◦ in the case of long-chipping light metals (Fig. 9.53). In summary, it can be seen that only a careful, automatically executed drill point grinding adjusted to the particular problem can lead to an economical machining

420


Fig. 9.51 Rake angle, clearance and the cutting speed against the drill diameter

40°

20

vc

Clearance angle α

30°

10

20°

Cutting speed vc

10°

γfe

0

0° 10 9

αfe

8 7 6 5 Drill diameter r

4

mm 1

0

–10°

Major cutting edge –20°

Chisel edge

–30°

–40°

–50° Rake angle γ –60°

Typ N

Normal workpiece material (Such as ) δ = 18 to 30°

Typ H

Hard and brittle workpiece (Such as gray cast iron) δ = 10 to 15°

Typ S

Ductile workpiece material (Such as aluminium) δ = 35 to 45°

δ

δ

δ

process. For the larger part of all machining cases, the conical relief point has asserted itself as the most consistent and suitable type. The rake faces are parts of a taper sleeve (Fig. 9.54). The advantages of these drills include easy manufacture and preparation as well as their low sensitivity to high mechanical stress.

δ = Helix angle

Fig. 9.52 Spiral drills for various materials

9.3

Drilling

421 γf e

γ

β

Cutting edge Cutting edge 2

α αf η e

f

f/2

ng direction

η

Feed direction

Effective cutting direction

Fig. 9.53 Motion sequence of the major cutting edges by double edged drilling tool

Form A

Form B

Taper sleeve grinding (Polished cone-shaped section Form A–D)

Picked out chisel edge

Picked out chisel edge with adjusted major cutting edge

Form C

Form D

Form E

Split point

Pickedout chisel edge with Drill point angle 180° with faced edge corners centered tip

Fig. 9.54 Combination of spiral drills with special polished section from model A till E compared to taper sleeve grinding, acc. to DIN 1412

422


The low self-centring and associated shape and position errors are disadvantageous. In addition, the chisel edge length is increased with increasing drill and core diameters, such that the resulting high feed forces have an unfavourable effect on machining accuracy. In this case and generally only when special demands are placed on the drilling tool, the drill point is equipped with a special grinding that either complements the taper sleeve grinding (e.g. core point thinning) or completely reshapes the drill point (centre point, Fig. 9.54). The following describes the most important point grindings in accordance with Fig. 9.54: • Form A: the taper sleeve grinding with a point-thinned core improves to a great extent the centrability of the drill and decreases the axial force corresponding to the chisel edge shortened by about 0.1·D (used in general for Type N beyond 14 mm diameter). • Form B: the taper sleeve grinding with point-thinned core and corrected rake angle makes it possible to adjust the rake angle to the machining task. However, it is customary to reduce the rake angle by about 10◦ , resulting in a very stable wedge without hindering chip transport because of a diminished helix angle. Grinding B is used in cases of high drill stress such as encountered when machining austenitic manganese steel or when drilling thin-walled aluminium sheets to reduce deformation. • Form C: A taper sleeve grinding with a split point in which case the chisel is completely eliminated. This is especially suitable for deep drill holes. The compressive chisel edge is converted into two small major cutting edges with much better cutting properties. This type also guarantees good centrability and reduced feed force. • Form D: The taper sleeve grinding with point-thinned core and bevelled corners was specially developed for machining grey-cast iron workpieces, the hard, abrasive casting skin of which stresses the sensitive corner to a particularly large extent. Here, a second taper sleeve grinding with a smaller drill-point angle provides a remedy in that it helps improve heat conduction and counters increased wear by increasing its surface area. • Form E: drill-point angle 180◦ with centred tip, used when centric drilling must be guaranteed or when round and burr-free drill holes are to be made in aluminium sheets. After total penetration of the centring cone, both major cutting edges simultaneously cut to their full length, and the corners can support themselves immediately on the drill hole wall with the lands. The drill exits again by the entire major cutting edge, whereby a ring-shaped disc is cut out with minimal burr formation. The four-face grinding, a taper sleeve grinding with a secondary face, is mentionable despite the fact that it is not standardized inasmuch as it is used when drilling below 1.5 mm diameters or with cemented carbide drills, since in this case the tapered sleeve grind causes difficulties.

9.3

Drilling

423

The special working conditions of a drilling tool place high demands on the cutting tool material with respect to hardness, toughness, wear resistance and insensitivity against thermal alternate stresses. Frequently, HSS is used as a cutting tool material. According to DIN 1414-1, high speed steels for drilling tools, contain 6% tungsten, 5% molybdenum, 2% vanadium (HS6-5-2) and for higher stresses 5% cobalt (HS6-5-2-5). The tools are hardened, topically treated (nitrated) and often equipped with wear-preventing coatings. Solid cemented carbide drills are also used however. The advantages of cemented carbides are their high hardness, compressive strength and high-temperature wear resistance. Cemented carbides have the same hardness at 1000 ◦ C as high speed steel at room temperature. As a rule of thumb, the cutting speed vc can only be increased by a factor of three. Beyond that, the machining process can continue with a feed f that is at least 30% higher. Besides this increase in cutting conditions, the tool life travel path Lf can be extended by a factor of three. Due to their high Young’s modules, cemented carbide drills are much more torsion-stiff than HSS tools. Due to their high hardness and low toughness compared to HSS tools, their use is technically meaningful and economical only on machine tools that fulfil the minimum requirements regarding accuracy, power, cooling and stiffness. One example for precision requirements is the concentricity of the drilling process. The total radial deviation measurable at the cutting edges of the drill is the result of the sum of each radial deviation of the machine spindle, interface, tool holder and tool. In current practice, the tool holder has the highest share. If the minimum requirements cannot be fulfilled, HSS drills are still preferred, not the least because of their lower price. Drilling as a machining process has several peculiarities which we will examine in the following. In comparison to internal turning, drilling with spiral drills produces a greater surface finish on the drill hole wall, which is the result of the comparatively low cutting speed, the low torsion and bending stiffness of the tool and chip transport [Spur60]. Moreover, spiral drills are subject not only to flank face and crater wear but also chisel edge wear, land wear and corner wear. In the case of the spiral drill, total tool wear is composed basically of • • • • •

flank face wear, crater wear, land wear, chisel edge wear and corner wear

and leads ultimately to relative or absolute disruption of the tool. Relative disruption is characterized by the fact that beyond a certain drilling length the machining output no longer correspond to the requirements. In the case of absolute disruption, the HSS tool becomes completely unusable because of thermal induced failure of the tool cutting edge or because of tool fracture. One essential tool life criterion in drilling is reaching pre-given limits in the case of dimension and shape faults (relative disruption). Wear on the corner and on the lands is frequently responsible for this. Due to the maximum cutting speed on the

424


outer diameter, the corner is especially stressed. This is where HSS tool often fail. On the other hand, the low cutting speed in the chisel edge area often causes built-up edge formation, which however does not have a dominantly negative effect on the process sequence nor makes itself perceptible in the output. The feed influences tool wear much less than the cutting speed, so it will no longer be considered here. The forces acting on the spiral drill are represented in Fig. 9.55. When drilling with HSS tools, the cutting speed for steels is in the range of 10–40 m/min. A series of studies has shown that here the influence on the resultant force is small, especially in the case of large drill diameters. In the case of extremely low or extremely high cutting speeds on the other hand, a considerable increase in the feed force and cutting moment has been noted, which has an especially large effect when smaller diameter drills are used. As in all machining processes with geometrically defined cutting edges, the resultant force components increase degressively with the feed (Fig. 9.55). Exhaustive investigations have shown that the surface quality of the drill hole wall cannot be significantly affected by the drill point grinding. On the other hand, the dimensional accuracy of the drill hole is dependent on the symmetry of the grinding, since straying of the tool can only be prevented when there is an extensive balancing of the radially acting passive force [Spur60]. Different passive forces Fp1 and Fp2 subject the drill to bending and lead to enlarged drill hole diameters in the output. Such passive forces arise primarily for tool-related reasons, such as • • • •

unequal major cutting edge lengths, unequal drill-point angles, unequal tool orthogonal clearances, asymmetrical point thinning,

d

Fc : Cutting force Ff : Feed force Fp : Passive force d : Tool diameter H : Distance of force-application to drill axis

Fc1

Fp1

Fp2

Fig. 9.55 Spiral drill forces, according to [Spur60]

Fp2

Ff2

Ff1

Fp1 2*H Fc2

9.3

Drilling

425

Table 9.1 Force components of the major cutting edge, the chisel edge and the land of a spiral drill

Major cutting edge Chisel edge Land

Torque (%)

Feed force (%)

65–75 10–14 15–20

17–25 65–75 7–8

• asymmetrical spiral flutes, • inconsistent cutting edge sharpness and • radial deviations [Spur60]. The increased cutting and feed forces compared to turning (at otherwise identical marginal conditions) are due on the one hand to friction of the chips in the bottom of the groove and on the drill hole wall and on the other hand to the length of the chisel edge, which should be referred to in consideration of the resultant force more than the drill diameter. The more narrow the chip space is for chip volume specified by the cutting conditions the larger are the required cutting moment and the necessary feed force. Among the feed forces that arise during turning, considerable amounts of force are also added due to friction on the land and compressive processes in the area of the chisel edge. The guidelines in Table 9.1 should be referred to when considering the percentage proportion of force of the major cutting edge, the chisel edge and the land of a spiral drill with respect to torque and feed force: These percentages were determined by means of step drill experiments. The torques and feed forces can be determined sequentially only for the major cutting edges, for the major and chisel edges as well as for the major edge, chisel edge and land. 9.3.2.2 Cutting Parameters in Drilling The cross-section of undeformed chip A has a major influence on the resultant force in drilling. Table 9.2 shows the relevant relations of the cutting parameters in drilling and the potential calculation of the cross-section of undeformed chip from the feed component per cutting edge fz and the depth of cut ap or from the chip thickness h and chip width b. 9.3.2.3 Calculating Forces, Torque and Power When Drilling with Spiral Drills Analogously to the ratios in drilling, we can approximately calculate the cutting force, feed force and torque for centre drilling and drilling out with spiral drills with the help of the K IENZLE equation (Table 9.3). For drilling, it is necessary to introduce a process factor fB in order to take into consideration the altered influences on the forces that occur in drilling as opposed to turning (e.g. cutting edge shape, cutting speed etc).

426

9 Processes with Rotational Primary Movement Table 9.2 Calculation of the cross-section of undeformed chip in drilling Centre drilling

Drilling out

Cutting parameters

h

h

A

fz

fz

b

ap

b ap

A

d D

D

ap f σ , h = fz · sin(κr ) f z = , κr = , b = z 2 sin(κr ) A = fz · ap = b · h d·f 4

Cross-section of undeformed chip

A=

D d ap b h z

Drill diameter [mm] Pre-hole diameter [mm] Depth of cut [mm] Chip width [mm] Chip thickness [mm] Number of cutting edges

A= f fz κr σ

(D − d) · fz 2

Feed [mm] Feed per cutting edge [mm] Lead angle [◦ ] Drill-point angle [◦ ]

9.3.2.4 Deephole Drilling Deephole drilling is a machining process used to produce or process drill holes. Deepholes are drill holes with a diameter between about 1 and 1500 mm and a drilling depth of about three times the diameter. It is not possible to make a general distinction between deephole drilling and other “conventional” drilling techniques by means of a universally valid definition or the like. With all deephole drilling methods, a very large ratio of drilling depth to diameter can be obtained. Further advantages of deephole drilling compared with customary drilling with spiral drills include above all the higher quality of the drill holes and its excellent cost-efficiency. Not only is deephole drilling distinguished from common drilling by its asymmetrical cutting edge arrangement, but also by the fact that in deephole drilling a cutting fluid is fed directly to the cutting edges under pressure and that its rinsing effect is the sole transport mechanism for the incoming chips. The cutting part is made of cemented carbide, so high cutting speeds can be reached, which in turn

9.3

Drilling

427 Table 9.3 Forces, torque, power required in drilling

Force application

Centre drilling

Drilling out

H = D/4

H = (D + d)/4

ap = D/4

ap = (D − d)/4 D

D D/2

H

Fcz Fcz

d ap

ap

fB = 1 (D − d) Cutting force per cutting Fcz = · f z · k c · fB 2 edge F

f B = 0.95 D · f z · k c · fB Fcz = 2

Process factor fB

cz

Feed force per cutting edge Ffz Torque

Ffz =

D · fz · kf · fB 2

Ffz =

(D − d) · f z · kf · f B 2

Fcz · z · D 4000 For z = 2: Fcz · D Md = 2000 9554 · Pc Md = n

Fcz · z · (D + d) 4000 For z = 2: Fcz · z · (D + d) Md = 2000

Pc η Fcz · vc Pc = 60000

Md · n 9554 Fcz · vc (1 + d/D) Pc = 60000

Md =

Md =

Power

Pa =

Pc =

fB

Process factor drilling

kf

H

Lever [mm]

Fcz Cutting force per cutting edge[N]

Specific feed force [N/mm2 ]

D

External drill diameter [mm]

Md Torque [Nm]

d

Internal drill diameter [mm]

Pa

Drive power [kW]

z

Number of cutting edges

Pc

Cutting power [kW]

fz

Feed per cutting edge [mm]

n

Speed [min–1 ]

vc

Cutting speed [m/min]

η

Efficiency

kc

Specific cutting force

[N/mm2 ]

428

9 Processes with Rotational Primary Movement Workpiece

Tool

Drill bushing

Bead

Shank Cooling lubricant supply

Cooling lubricant chip mixture

Fig. 9.56 The ELB process for diameters 0.8–40 mm, according to Sandvik

makes it possible to increase the material removal rate. The following three process variants are used for the industrial production of deep drill holes: • the single-lip drilling process (ELB process), • the BTA drilling process, • the ejector drilling process. The single-lip drilling process is used in the diameter range of approximately 0.8–40 mm. Figure 9.56 shows the essential characteristics of this method. The characteristic trait and main advantage of the single-lip drilling process is that the cutting fluid is supplied by means of one or several drill holes within the tool and the cutting fluid/chip mixture is safely led away by a longitudinal groove (bead) on the outside of the tool shank. Due to the shape of the lead and the large drilling depth/diameter ratio, the drill is guided on the top face of the workpiece by means of a drill bushing (Fig. 9.56). This stabilizes the start of drilling. Single-lip deephole drilling tools are used in manufacturing as solid drills, core drills, countersinks and step drills, whereby full drilling is the most common case of operation in practice. Basically, the single-lip drill consists of three components: the drill head, the shank and the clamping sleeve. In most cases, cemented carbide is used as the cutting tool material, whereby both solid cemented carbide drill heads as well as drill heads fitted with cemented carbide are employed. The BTS process (Boring and Trepanning Association – BTA) was invented at the end of the 1930s in order to prevent chip scratching on the drill hole wall during transport and the resulting damage to the surface quality. The attempt to cover the flute of the single-lip drill outwards resulted however in a drastic reduction of available chip space, which in turn limited the material removal rate. The solution was finally discovered by the “Boring and Trepanning Association”, who reversed the process characteristics of single-lip drilling and supplied the cutting fluid from outside by means of a ring-shaped crack between the drill pipe and the wall (Fig. 9.57). Reflow occurs together with the chips through the cutting jaw and the drill pipe, the

9.3

Drilling Workpiece

429 Tool

Cooling lubricant supply


Drill tube

Fig. 9.57 The BTA process for diameters of 6–300 mm, according to Sandvik

diameter of which should not be less than 6 mm. The upper diameter for full drilling tools is around 300 mm and for countersink tools around 1000 mm, whereby these limits depend to a large extent on the available machine power. Compared to deephole drilling with single-lip drills, the BTA process has the disadvantage that a complicated drill oil supply apparatus is required which takes over the sealing of the drill pipe. The process requires machines that are much different than standard drill machines. The ejector deephole drilling process is utilized for diameters of about 18–250 mm. According to VDI guideline 3209, it is a variant of the BTA process (Fig. 9.58). Cutting fluid supply is accomplished by means of a ring space between the drill pipe and an internal pipe (two-pipe process). The cutting fluid enters the drill head from the side, rinses it and flows back into the internal pipe with the chips. Part of the cutting fluid is introduced into the internal pipe via a ring nozzle. Reflow is made possible by the arising low pressure at the cutting jaw (ejector

Workpiece

Tool

Drill tube

Cooling lubricat supply

Internal tube

Ejector effect


Fig. 9.58 The ejector process for diameters of 18–250 mm, according to Sandvik

430


effect). As opposed to the BTA process, sealing against the exit of the cutting fluid is omitted. A further peculiarity is its cutting edge distribution for reducing the forces acting on the guide beads as well as the double-sized cutting jaw thus required. The cutting edge, otherwise continuous from the periphery to the centre, is subdivided such that two cutting parts are arranged at a time alternating left and right up to the centre. The consequence of this is that the stress on the guide beads is reduced by about 10% to a maximum of 50% of the otherwise expected forces, and friction, heat development and wear are reduced accordingly. Deephole drilling tools dominate the entire field of inner contours that can be manufactured by drilling (Fig. 9.59). Other drilling methods and tools are used only

200

Length/ Diameter

Length/Diameter

ELB

Deep hole Tiefbohren drilling

Überschneidungsbereich Short hole drilling zum Kurzlochbohren BTA Sintered carbide drill, HM-Bohrer, Kurzlochbohrer, Short hole drill, Wendelbohrer Twist drill 60 Diameter (mm)

ELB - deep hole drilling ELB-Tiefbohren

BTA - deep hole BTAdrilling

Bereich der konvenArea of conventional drilling tionellenBohrbearbeitung

Diameter (mm)

1500 1500

Fig. 9.59 Application of the deep hole tools compared to “conventional” drilling tools, acc. to VDI guideline 3210

9.3

Drilling

431

in the range of smaller drilling depths (up to a length/diameter of ca. 6 and a diameter of up to ca. 60 mm). Since these dimensions are predominant in general mechanical engineering, the dominance and versatility of deephole drilling processes is often not perceived. Increasingly, especially in the area of overlap between shorthole drilling (conventional drill technology) and deephole drilling, tools are being used that have the characteristics of deephole tools or are operated under conditions that resemble a deephole drilling operation. Especially in the field of deep drill holes and drill holes with large diameters, deephole techniques are used now almost exclusively. Due to its high productivity and the drill hole quality obtainable, deephole drilling is being used increasingly for manufacturing tasks in which the ratio between the drill hole depth and the drill hole diameter is larger than 6. Numerous examples of machining reveal the presence of deephole technology in the area of smaller tool diameters as well, in which naturally most application cases for drilling are to be found. Difficult-to-machine materials can as a rule be machined effectively with deephole methods. Deephole processes have the following typical characteristics: • the use of special cemented carbide tools with one or sometimes more cutting edges that lie asymmetrically to the tool axis, • self-commutation of the tool by means of a three-point mounting in the drill hole through guide beads and the minor cutting edge (cylindrical grinding chamfer), • drill start guidance of the tool in a drill bushing or a guiding bore, • continuous high-quantity cutting fluid supply under pressure resulting in constant chip removal without chip removal strokes. Some advantages of deephole drilling are: • • • • • • • • • •

very high machining performance, ideal cooling and lubricating conditions, short primary processing times, high drill hole quality with respect to diameter tolerance, surface quality and geometric contouring accuracy, high alignment accuracy, minimal drill hole inaccuracy, replacement of several operations – e.g. pre-drilling, boring and reaming – by one single operation, possibility of processing hard-to-machine materials, large drilling depths in relation to the diameter (up to a maximum of 250 times larger), cost-efficiency, even with short drilling depths, minimal burr-formation when drilling out and when overdrilling cross-holes.

By means of deephole drilling, metals of all kinds as well as other materials (e.g. plastics) can be processed both in the mass production of small parts and in the single-part production of large-scale machine parts. The process variants (including machining methods) of deephole drilling are characterized by the drilling task and the correspondingly adjusted drilling tools

432

9 Processes with Rotational Primary Movement Through boring

Blind hole

Drilling in full metal

Drilling in full metal

Drill out

Drill out

Core drilling

Core drilling

Form drilling

Form drilling

Fig. 9.60 Deephole drilling process variants

(Fig. 9.60, VDI guideline 3210). The most frequently utilized variant of drilling is full drilling. Moreover, all deephole drilling processes can be employed for drilling out and core drilling. Form drilling tools serve to produce a definite drill hole bottom defined by the tool cutting edge geometry or contour transitions in the case of step drilling. Deephole drilling machines can be constructed in a way that the primary motion can be carried out by either the tool or the workpiece or by both. In the case of a rotating tool and a stationary tool, the machines are applicable for a broad range of arbitrarily shaped parts. With an automatic loader and potentially as multi-spindle machines, they are most suitable for economical large-batch production. For both rotating and stationary tools, only rotation-symmetrical parts with small masses can be used, since there is a danger that even a small unbalance of the rotating

9.3

Drilling

433

Fig. 9.61 Deephole drilling machine (Source: TBT Tiefbohrtechnik)

workpieces may lead to poor drilling results. Figure 9.61 shows a deephole drilling machine by the company TBT Tiefbohrtechnik. Some of the technical data of this deephole drilling machine are: • • • •

full drilling rage (min.–max.) maximum drill depth with 1 steady rest maximum spindle speed drive power

: 0.9–15 mm, : 700 mm, : 24.000 1/min, : 2.4 kW.

The limits of use of deephole drilling processes are essentially determined by the following factors: • • • • •

the machinability of the material, the stability of the tool and the machine, the accuracy of the machine, the composition of the cutting fluid, the cutting tool material.

VDI guideline 3210 summarizes these limits in the form of standard values for full drills, core drills and countersinks and distinguishes according to the process and tool whether ISO degrees of tolerance above or below IT9 can be reached. This approximate indication is grounded in the fact that the machinability of the material is still clearly part of the obtainable tolerance (Fig. 9.62). For example, non-ferrous metals permit an ISO tolerance of IT6 under optimal conditions, which corresponds to a diameter range of 50–80 mm of a maximum deviation of 9 μm. On the other hand, the best possible tolerance when machining nitriding steels is IT8. Since the process combination spiral drilling/reaming can be substituted with deephole drilling, under normal conditions the surface qualities obtainable are in

434


Fig. 9.62 Tolerances obtainable in deephole drilling (Source: TBT Tiefbohrtechnik)

Non-ferrous metall Aluminium Tool steel Cast iron Tempered steel Carbon steel Cementation steel Nitrided steel Hole quality IT

13 12 11 10

9

8

7

5

6

RZ =100 RZ =25 RZ =6.3 RZ =1 N9

N8

N7

N6

N5

N4

N3

N2

N1

6.3

3.2

1.6

0.8

0.4

0.2

0.1

0.05

0.02 5

N12 N11 N10 Spiral drilling Reaming Broaching Honing Deep-hole drilling 12.5

25

50

Average thoughness

Under normal conditions achievable Under the best conditions achievable

Fig. 9.63 Obtainable surface quality in comparison to other machining methods (Source: TBT Tiefbohrtechnik)

the same order of magnitude (fine finishing). Figure 9.63 shows furthermore that under especially favourable conditions even Ra values of 0.1 μm are obtainable, which otherwise require a superfinishing operation such as honing. We can see from these considerations that deephole drilling is especially economical when subsequent processes can be reduced and/or high-alloyed materials are to be processed.

9.3

Drilling

435

The following overview summarizes the areas of application in which deephole drilling can be used advantageously [Grüb74]: • • • • • •

high material removal rate requirements machining materials with high alloy components that are hard to machine materials with a tensile strength of over 1200 N/mm2 high tolerance and surface quality requirements large drilling depth in relation to the diameter substitution of several single process steps (full drilling, drilling out, reaming) with one process step.

9.3.2.5 Reaming Reaming is a fine finishing process and serves to improve drill hole quality, whereby position and shape errors cannot be influenced. With respect to kinematics, reaming is equivalent to drilling out with small chip thicknesses (Fig. 9.64). According to DIN 8589-2, a distinction is drawn between reaming with singleblade and multi-blade tools. The single-blade reamer is guided by a guiding bead arranged on the periphery, whereby the functions of machining and guiding are divided among independent active elements (Fig. 9.65). The multi-blade reamer is guided by the minor cutting edge arranged on the periphery. Fig. 9.64 Reaming tools, according to [DIN 8589-2]

Tool Workpiece

Reaming

Form reaming

Fig. 9.65 Drill out tools with PKD – equipping and also with ISO – indexable inserts, according to Mapal

436


The cutting edges of multi-blade reamers can be arranged parallel to the axis or on a helical line. Drill holes with grooves are reamed with spiral tools in order to avoid cutting engagement impact of the cutting edge. Usually, reamers are manufactured with an even number of teeth, whereby two cutting edges face each other at a time, which makes it much easier to determine the diameter. In order to prevent clattering vibrations, an odd distribution of cutting edge distances is selected, which repeats after half of the circumference. Drill hole qualities of IT7 or better are obtainable. A distinction is made between manual reaming and machine reaming. In the case of manual reamers, the cutting tool material used is usually tool steel or HSS, while machine reamers use high speed steels or cemented carbides. Performance can be enhanced by employing coated tools in reaming as well. Insert changeability makes it possible to adjust the tool to different materials and machining tasks by means of an appropriate choice of substrate, coating and geometry. 9.3.2.6 Internal Thread Production Internal threads can be manufactured by means of primary shaping, forming or cutting. Besides the available technology, essential considerations when selecting the manufacturing method include the material to be machined, the thread type and the number of units required as well as the required tolerances, strengths and surface quality. Among the process variants used, cutting manufacture via tapping takes the leading position because of it is the most widespread. 9.3.2.7 Tapping Tapping is drilling out for the manufacture of an internal thread that lies coaxially to the rotation axis of the cutting motion (Fig. 9.66). Screw taps consist of a shank and a screwed portion. In the case of the screwed portion, a distinction should be drawn between the cutting part where machining occurs and the guide part, which is responsible for stabilizing the tool [Zura90].

Tool

Straight flute Workpiece Twisted flute

Fig. 9.66 Screw tap

Section

9.3

Drilling

437

Screw taps for through borings have a lead that guarantees good tool guidance. The screwed portion is subdivided into cutting studs by grooves. The grooves serve to receive the chips and convey them outside. Moreover, they guarantee cutting fluid supply to the cutting location. Large groove cross-sections facilitate this and are especially important when machining long-chipping materials. However, they lead to a weakening of the load-bearing tool cross-section. To process short-chipping materials, screw taps with even grooves, which are easier to manufacture, are sufficient. In the case of long-chipping materials, spiral chip spaces are necessary to facilitate chip removal and to reduce the danger of chip jams. For the sake of good tool centring, at least three cutting studs are required, and thus three chip flutes as well. In the case of manual screw taps, a set of two or three tools (pre-cutters and finish cutters or pre-cutters, intermediate cutters and finish cutters) are used to produce the thread in order to minimize tool load and the risk of fracture. The cutting tool materials employed in tapping are high speed steels and cemented carbides. The surfaces of the screw taps can be processed to increase their wear resistance. The most important processes used are nitriding, hard chrome plating and coating with hard materials. The potential cutting speeds are relatively low and depend considerably on the combination of cutting tool material, workpiece material and cutting fluid. 9.3.2.8 Thread Milling Under certain conditions, thread milling can be a good alternative to other thread manufacturing processes. Thread milling is a special screw milling process. It can be used, for example, to obtain very good surfaces on the screw flanks and to produce large numbers of units economically. Threads with large diameters can often only be fabricated via milling (number of units, tools costs, power input) [Fosh94]. The structure of a screw milling cutter is basically similar to that of a “screw tap”. As opposed to the screw tap, which consists as it were of a single spiral-shaped tooth, the consecutive teeth of a screw milling cutter do not form spirals but rather are arranged without offset (Fig. 9.67). The tooth shape corresponds as a rule to the form of the thread to be formed. In many cases it is necessary to correct the tooth profile. This is the case when the thread to be milled is not at least three times larger than the milling cutter’s diameter. Without a correction of the profile, the tool would cut freely, distorting the finished thread profile. A screw milling cutter can mill threads of various diameters. It is not

Fig. 9.67 Screw tap with coil (left) and milling cutter without coil (right), according to Fraisa

438

9 Processes with Rotational Primary Movement a

b

c

d

e

f

Fig. 9.68 The thread milling cycle

possible however to vary the pitch. The screw milling cutter is thus designated by the standard thread diameter that corresponds to this pitch. Figure 9.68 provides a graphical depiction of a milling cycle. The rotating milling cutter positioned in the centre of the drill hole is axially lowered to the desired thread depth into the core drill hole (Fig. 9.68a). The tool is then adjusted to the drill hole diameter (Fig. 9.68b). The screw milling cutter is finally radially advanced to the required standard diameter of the thread (Fig. 9.68c), so that there is a defined axis distance between the tool and the drill hole axis. To form the thread, there is a movement cycle of somewhat more than 360◦ on a coil (Fig. 9.68d), whereby the workpiece or the tool is axially shifted around a pitch. Finally, the screw milling cutter is backed out of the thread radially via a circular arc (Fig. 9.68e), driven back to the cutter axis and lifted axially out of the thread (Fig. 9.68f). 9.3.2.9 Thread Moulding According to DIN 8583-5, thread moulding is a non-cutting (forming) process in which the internal thread is created by impressing a tool (the thread moulder or groover) into the workpiece [Fieb95]. The thread moulder is separated axially into three sections. It has a conical lead part on the top which extends across several thread turns. The largest amount of forming work takes place on this area. The next section of largest external diameter has the function of removing the thread flanks from the mould. The following, slightly tapered calibration section serves to smooth the fabricated thread flanks and to guide the tool. The profile of a thread moulder is, in contrast to the round profile of a screw tap, a polygon with three or more flattened corner areas. The material is displaced on these forming edges, which serve to receive the lubricant. The thread moulding process does not require chip flutes, resulting in increased bending strength due to the larger cross-sectional area. The starting situation in the case of thread moulding is also a pre-drilled hole with a diameter corresponding approximately to the flank diameter of the thread. While

9.4

Sawing

439

Fig. 9.69 Thread tools and several formations of flanks

Thread grooving

Tapping

the forming edges of the tool penetrate into the workpiece material and the thread flanks form to the required dimensions, the displaced material flows into the tooth gaps of the thread moulder. This causes the formation of ears, typical of formed threads, in the area of the thread crests (Fig. 9.69 left). In comparison to “tapping”, thread moulding results in threads of higher strength. The cause is to be found in the strain-hardening that occurs during forming.

9.4 Sawing Sawing is cutting with a rotary or translatory main movement with a multi-blade tool of low cutting width, used for separating or slitting workpieces. Sawing is classified as a process with a rotary main movement, since even in the process variants hacksawing and bandsawing, in which there is a translatory cutting motion, the saw blades can be seen as a tool with an infinitely large diameter (Fig. 9.70). The following will be organized according to the type of tool used into bandsawing, hacksawing and circular sawing. These are the most commonly encountered process variants in praxis.

9.4.1 Bandsawing Bandsawing involves a revolving, unending saw band cutting with a continuous, mostly linear cutting motion [DIN8589g]. The saw band is supplied on large rolls and, after the cutting to the desired length, the ends are joined together with a butt-welding process. This process is distinguished by low cutting losses due to the small width of the bands. However this results as well in low tool stability against cut deviation. Figure 9.71 shows the cutting part geometry and terminology of a saw tooth. The number of teeth is commonly given in reference to 1 in. of band length (ZpZ). The size of the chip space depends on the dimensioning of the tooth base radius r and on the number of teeth. As a result, when sawing larger material cross-sections, the tooth number has to be reduced in order to obtain a sufficient chip space for the arising chips.

440


Tool

Workpiece Workpiece

Hack saw Bandsaw Tool

Workpiece

Circular saw

Fig. 9.70 Different sawing methods

In order to prevent jamming of the saw band in the cutting channel, the single teeth are bent or set to the left and to the right alternately from the cross section (see Chap. 3). The standard offset sequence for separating metals is right/left/straight. In the case of larger tooth numbers, shaft offset is also used (Fig. 9.72). The cutting motion is parallel and the feed motion perpendicular to the longitudinal axis of the band. In the case of unlimited teeth, the depth of cut ap corresponds to the width of undeformed chip b and in this case to the width of the saw blade. The engagement size ae , measured as the size of the engagement of one cutting edge in the cross section perpendicular to the feed direction, corresponds to the workpiece width. Saw bands are made either of tool steel or bimetal. In the case of bimetal bands, the blunt edge made of soft tool steel is joined by means of electron beam welding with a hardened band made of HSS, into which finally the saw teeth are inserted. Possible HSS cutting tool materials include HS6-5-2, HS2-10-1-8 or HS10-4-310. In practice, bands with soldered cemented carbide plates have also become established. Tool life criteria that define the end of a sawband’s usage are a certain wear condition, tooth fracture or a maximum permissible deviation of cut due to excessive

9.4

Sawing

441

Tool

γο

r

fz

H t

βο

vc

αο

ve

Workpiece

ae fz H r t

Tooth feed Tooth height Base radius Tooth pitch

vc vf ve

Cutting speed Feed rate Active speed

Fig. 9.71 Title and part geometry on the cutting saw band, after RENG [Reng76]

Standard restriction (SD)

Group restriction (GS)

Right-left-restriction (RL)

Fig. 9.72 Sawband restriction types (Source: Wikus)

Waves restriction (WS)

vf

442


passive force. In contrast to other cutting methods, the tool life is not given as the measure for the tool life parameter of a saw, but the cut surface (corresponding to the workpiece surface cut until reaching the end of tool life) [Reng76].

9.4.2 Hacksawing Hacksawing is a process variant with a repeated, usually linear cutting motion (Fig. 9.70). According to DIN 8589-6, hacksawing, gangsawing, and jigsawing are all hacksawing processes [DIN8589f]. From this definition, we can already see that hacksawing is a process with a discontinuous cutting motion, i.e. with machine hacksaws, material is removed only in the forward stroke. In the return stroke, the saw blade is mechanically or hydraulically lifted. The result is that the material removal rate is lower compared with bandsawing or circular sawing. On the other hand, the cutting loss is relative small. The saw blades are made either of solid HSS or of HSS segments that are riveted on a blade body. Blades made of tool steel are generally only used for manual hacksaws.

9.4.3 Circular Sawing Circular sawing is sawing with a continuous cutting motion using a circular rotating saw blade (Fig. 9.73). This high-performance process is employed for linear cuts on low-cost materials, as much material is lost due to the relatively wide kerf.

Fig. 9.73 Tooth segments of a saw blade (Source: Ohler)

9.4

Sawing

443

Both tool steel and HSS as well as cemented carbides as used as cutting tool materials. At the same time, there are fundamental differences among construction types of saw blades. Saw blades made of one material are made of tool steel or HSS. In the case of larger saw blades. For cost reasons, the blade body is made of construction steel upon which the individual HSS segments are riveted (Fig. 9.73). The material removal can be improved further by means of soldered cemented carbide cutting edges. In order to guarantee the cutting capacity of circular saw blades, it is absolutely necessary to break the chips in such a way that they are narrower than the kerf. If such measures are neglected, the chips jam in the chip space and can damage the tool. Figure 9.74 shows the two most prevalent possibilities [Schm80]. By subdividing the teeth into pre-cutting and post-cutting elements, the chips are fractured such that the pre-cutting teeth protrude at least by the amount of chip thickness and their major cutting edge length is smaller than the total width of undeformed chip. One alternative is grinding-in offset chip breaker flutes into the otherwise identically formed teeth. In this way, one narrow and one wide chip is produced per tooth that can escape towards the centre into the chip breaker flute without jamming in the kerf. The fact that a tooth in the section of the chip breaker flute of the previous

Chip cross section of the after cutting tooth

Front Incisor

After cutting incisor

Chip flute

Fig. 9.74 Cutting edge geometries of circular saw blades

Chip cross section of the front incisor

444


tooth must remove a larger chip cross-section has only an insignificant effect on the progression of wear. The advantage is that one tooth with a chip breaker flute can take on the same chip cross-section as the pre-cutting and post-cutting teeth of the other lead version. In order to increase the circular saw blades’ stability and running smoothness, it is necessary to introduce characteristic stresses into the blade in a specific fashion. This is done either by hammering or by rolling a concentric pressure zone into the side surfaces. A laser radiation process is also possible. For separating general construction steel with HSS saw blades, values of vc = 18−30 m/min and fz = 0.22−0.28 mm and with cemented carbide saw blades values of vc = 90 − 150 m/min und fz = 0.12 − 0.18 mm are common as cutting data.

Chapter 10

Processes with Translatory Primary Movement

Cutting processes using a geometrically defined cutting edge that carry out a translatory main motion include, among others: • • • •

broaching, shaving, planing and shaping as whose process variants.

10.1 Broaching Broaching is a machining process with a multi-toothed tool whose cutting teeth lie in a row, each being separated by the thickness of one chip. Tooth graduation perpendicular to the direction of the cutting speed replaces the feed motion. The cutting motion is translatory, in special cases also helical or circular [DIN8589e]. Broaching can realize a high material removal rate in one stroke, since usually several teeth are simultaneously in engaged. Moreover, high surface qualities and precision are obtainable and tolerances of up to IT 7 maintained. This method can only be utilized economically in serial production due to the high costs of tool production and preparation, as the tools can always only be used for one cross-section of undeformed chip [Kraz77]. DIN 8589-5 draws a distinction between: • • • • •

face broaching, circular broaching, helical broaching, profile broaching and form broaching.

We also distinguish between internal and external broaching, for which differently designed machined tools and tools are required. The broaching tool is, as shown in Fig. 10.1, pulled/pushed through a borehole (internal broaching) or pulled/pushed along the external surface of the workpiece (external broaching). The final contour is usually created in one stroke. F. Klocke, Manufacturing Processes 1, RWTH edition, C Springer-Verlag Berlin Heidelberg 2011 DOI 10.1007/978-3-642-11979-8_10,

445

446

10


Internal broaching

External broaching Workpiece

Tool

Workpiece Tool

vc

vc Workpiece holder

Undeformed chip thickness

Fig. 10.1 Contact ratios and process variants of broaching

To process unhardened steel materials, in most cases tools made of high speed steel are employed. In certain cases, such as in large-batch production of grey cast iron or broaching hardened steels, cemented carbide or CBN are also used as cutting tool materials. To increase performance and wear resistance, the tools can be coated with hard materials (e.g. TiN, TiCN). The tools are either designed as solid units or fitted with indexable inserts [Wege85, Merk80]. The standard cutting speeds in steel-cutting are between vc = 1−30 m/min. Powerful broaching machines can reach cutting speeds of up to vc = 120 m/min so that the surface quality is improved by avoiding built-up edge formation [Schü65, Opfe81]. Further cutting speed increase is only possible to a limited extent because of the process kinematics, since acceleration and deceleration must take place in one stroke. Cutting fluids are almost always used in broaching processes in order to guarantee chip transport from the chip space in addition to the lubricative effect [Falk70]. To this end, usually oils are used due to the low cutting speeds. When machining hardened steels, high cutting speeds of vc = 60−70 m/min and cemented carbide tools are used to reduce the high cutting forces and to increase tool life. Cutting fluids are not used in this case [Klin93]. The following description will be restricted to the most important of process variants. Technologically comparable methods will be treated jointly.

10.1.1 Face and Circular Broaching Figure 10.2 shows the typical structure of a broaching tool. Broaching tools have a roughing, a finishing and a calibrating section. The subsections differ in feed per tooth fz . In the roughing section, the feed per tooth fz is in the range of fz = 0.1−0.25 mm depending on the workpiece to be machined, while in the finishing section is it fz = 0.0015−0.04 mm and in the calibrating section equal to zero. The teeth in the calibrating section possess the geometry required for the workpiece.

10.1

Broaching

447 Finishing tooth

Rear pilot A

Roughing tooth A

Sizing tooth (fz = 0) B

t

B

bfa αf α

fz γ

Fig. 10.2 Internal broaching tool (schematic)

The calibrating section makes it possible for the expensive broaching tools to be reground and thus be economically employed. Regrinding the tool shifts the entire tool profile by one tooth towards the calibrating section. In this way, the previously first calibrating tooth now becomes the last finishing tooth. The tool rake angle γ and the tool orthogonal clearance α of the second flank face are adjusted to the material to be machined. The chamfer width bfα is parallel to the axis in the calibrating section and inclined by the tool orthogonal clearance of the first flank face α f in the roughing and finishing section. The cutting pattern determined by the arrangement of the cutting edges along the tool axis is called the offset. If the cut is perpendicular to the broaching surface, it is called offset in depth. If the broaching surface is machined from the side on the other hand, it is called lateral offset. Both offset types can be provided on one tool simultaneously depending on the workpiece geometry [DIN1415]. The pitch t of the broaching tool and thus the size of the chip space as well depend on the height of the workpiece, chip formation of the material and the maximum potential tool length. A large workpiece height and unbroken chip forms require large chip spaces. In the case of small machines and short tools, only a small pitch is required, since only workpieces with a low height can be machined. Figure 10.3 provides an overview of various broaching tools for internal and external machining. Possible tool lengths are in the area of L = 100−10000 mm. Internal broaching tools are manufactured up to a diameter of D = 500 mm (e.g. for the internal gearing of hollow wheels). To obtain chip fracture, the individual teeth are equipped with chip breakers in the form of tooth grooves [Weul85, Hoff76]. The total cutting force, which is comprised of the individual components of all the cutting teeth in action, is crucial for the minimal cross-section of the broaching tool shank, which in turn limits the maximum size of the chip space. The height

448

10


Fig. 10.3 Internal and external broaching tools (Source: Forst)

Total cutting force

b

Tool

lS1 lS0 m=3 lS2 Broaching stroke

t

Total cutting force

lS0 = 0°

lS1

lS2

t b t tan(λS2) ≥ b tan(λS1)

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