Cutting tool technology: industrial handbook

Cutting Tool Technology

Previous books for Springer Verlag by the author: Advanced Machining: The Handbook of Cutting Technology (1989) CNC Machining Technology series: Book 1: Design, Development and CIM strategies Book 2: Cutting, Fluids and Workholding Technologies Book 3: Part Programming Techniques (1993) CNC Machining Technology: Library Edition (1993) Industrial Metrology: Surfaces and Roundness (2002)

Graham T. Smith

Cutting Tool Technology Industrial Handbook

123

Graham T. Smith, MPhil (Brunel), PhD (Birmingham), CEng, FIMechE, FIEE Formerly Professor of Industrial Engineering Southampton Solent University Southampton U. K.

ISBN 978-1-84800-204-3

e-ISBN 978-1-84800-205-0

DOI 10.1007/978-1-84800-205-0 British Library Cataloguing in Publication Data Smith, Graham T., 1947–   Cutting tool technology: industrial handbook   1. Metal-cutting 2. Metal-cutting tools   I. Title   671.3'5 ISBN-13: 9781848002043 Library of Congress Control Number: 2008930567 © Springer-Verlag London Limited 2008 Apart from any fair dealing for the purposes of research or private study, or criticism or review, as permitted under the Copyright, Designs and Patents Act 1988, this publication may only be reproduced, stored or transmitted, in any form or by any means, with the prior permission in writing of the publishers, or in the case of reprographic reproduction in accordance with the terms of licences issued by the Copyright Licensing Agency. Enquiries concerning reproduction outside those terms should be sent to the publishers. The use of 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 laws and regulations and therefore free for general use. The publisher makes no representation, express or implied, with regard to the accuracy of the information contained in this book and cannot accept any legal responsibility or liability for any errors or omissions that may be made. Cover illustration: eStudio Calamar S.L., Girona, Spain Printed on acid-free paper 9 8 7 6 5 4 3 2 1 springer.com

Preface

Just over twenty years ago I began writing a book, the forerunner to this present volume for Springer Verlag, entitled: Advanced Machining – The Handbook of Cutting Technology. This original book covered many of the topics discussed here, but in a more general and less informative manner. Since this previous volume was published, many of the tooling-related topics are now more popular, or have recently been developed. Typical of these latter topics, are both High-speed and Hard-part machining that have now come to the fore. While Micro-machining and Artificial Intelligence (AI) coupled to neural network tool condition monitoring have become important, the latter from a research perspective. These machining and tooling topics, plus many others have been included herein, but only in a relatively concise manner. It would have been quite possible to write a book of this length just concerned with say, drilling techniques and associated tooling technologies alone. With the concerns raised on the health hazards to operational personnel exposed to cutting fluid mists in the atmosphere, the permissible exposure levels (PEL’s) have been significantly reduced recently. Further, with the advent of Near-dry and Dry-machining strategies, they have played a important role of late, particularly as their disposal and attendant costs have become of real consequence. Tool management issues previously discussed in the ‘Advanced Machining’ book have hardly changed, because when I wrote this chapter over two decades ago, most of today’s tooling issues by then had been addressed. However, the toolpresetting machines and associated software now, are far more advanced and sophisticated than was the case then, but the well-organised and run tool preparation ‘rules’ are still applicable today. One area of cutting tool development that has seen significant design novelty, is in the application of

Multi-functional tooling. Here, the chip control development is facilitated by both chip-narrowing and -vectoring, being achieved by computer-generated insert design, to position raised protrusions–‘embossed dimples’, on the top face. Further, some cutting insert toolholders are designed for controlled elastic compliance – giving the necessary clearance as the tool is vectored along and around the part’s profile, allowing a range of plunge-grooving and forming operations to be simultaneously undertaken by just this one tool. Coating technology advances have enabled significant progress to be made in both Hard-part machining and for that of either abrasive and work-hardened components. Some coating techniques today approach the hardness of natural diamond, particularly the aptlynamed ‘diamond-like coatings’ (DLC). Recently, one major cutting tool company has commercially-introduced an ‘atomically-modified coating’, such is the level of tool coating sophistication of late. Potential problems created by utilising faster cutting data often without benefit and use of flood coolant in cutting technology applications, has had an influence on the resulting machined surface integrity of the component. This sub-surface damage is often disguised, or not even recognised as a problem, until the part catastrophically fails in-service – as a result of the instability produced by the so-called ‘white-layering effect’. While another somewhat unusual factor that has become of some concern, is in either handling, or measuring miniscule components produced by Micromachining techniques. Often a whole month’s mass production of such diminutive machined parts could easily be fitted into a small shoebox! All of these previously mentioned tooling-related challenges and many others have to a certain extent, now become a reality. While other technical and machining factors are emerging that must be techni-

VI

Preface

cally-addressed, so that cutting tool activities continue to expand. It is a well acknowledged fact that if one was to list virtually all of our modern-day: domestic; medical; industrial; automotive; aerospace, etc; components and assemblies, they would to some extent rely on machining operations at a certain stage in their subsequent manufacturing process. These wide-ranging manufactured components clearly show that there is a substantive machining requirement, which will continue to grow and thus be of prime importance for the foreseeable future. This present book: ‘Cutting Tool Technology – Industrial Handbook’, has been written in a somewhat pragmatic manner and certain topics such as ‘Machining Mechanics’ have only been basically addressed, as they are well developed elsewhere, as indicated by the referenced material at the end of each chapter. Any book that attempts to cover practical subject matter such as that of cutting technology, must of necessity, heavily rely on information obtained from either one’s own machining and research experiences, or from indus-

trial specialist journals. I make no apology for liberally quoting many of these industrial and research sources within the text. However, I have attempted – wherever possible – to acknowledged their contributions when applicable, in either the references, or in the associated diagrammatical and pictorial figures herein. Further, it is hoped that the ‘machining practitioner’ can obtain additional information and some solutions and explanations from the relevant appendices, where amongst other topics, are listed a range of ‘trouble-shooting guides’. Finally, it is hoped that this latest book: ‘Cutting Tool Technology – Industrial Handbook’ will offer the ‘machining practioner’ the same degree of support as the previous book (i.e. Advanced Machining – The Handbook of Cutting Technology) achieved, from the significant feed-back obtained from practitioners and readers who have contacted me over the past decades. Graham T. Smith Fortuna, Murcia, Spain

Acknowledgements

First and foremost, I would like to express my sincere thanks to my wife Brenda for her support and for the time I have taken, whilst writing this book: Cutting Tool Technology – Industrial Handbook. I could not have achieved such an in-depth treatment and reasonably comprehensive account of the subject matter without her unstinting co-operation and help. A book that relies heavily on current industrial practices could not have been produced without the unconditional support from specific tooling manufacturers and the machine tool industries. I would like to particularly single-out one major cutting tool company, to genuinely thank everyone at Sandvik Coromant who have provided me with both relevant and significant: information; photographic; and diagrammatic support – the book would have been less relevant without their indefatigable co-operative help and discussion. Likewise, other tooling companies have been of much help and assistance in the preparation of this book, such as: Seco Tools; Kennametal Hertel and Kennametal Inc; Iscar Tools; Ingersoll; Guhring; Sumitomo Electric Hardmetal Ltd; Mitsubishi Carbide; Horn (USA); Shefcut Tool and Engineering Ltd; Rotary Technologies Corp; Diashowa Tooling; Centreline Machine Tool Co Ltd; DeBeers – element 6; Walter Cutters; Widia Valenite; TRW – Greenfield Tap and Die; Triple-T Cutting Tool, Inc; Hydra Lock Corp; Tooling Innovations; and Microbore Tooling Systems. Several machine tool companies have been invaluable in providing information, notably: Cincinnati Machines; Yamazaki Mazak;

Dorries Scharmann; DMG (UK) Ltd; Giddings and Lewis; Starrag Machine Tool Co; and E. Zoller GmbH and Co KG. While other tooling-based and associated companies have also provided considerable information, including: Renishaw plc; Kistler Instrumente AG; Taylor Hobson plc; Mahr/Feinpruf; Cimcool; Kuwait Petroleum International Lubricants; Edgar Vaughan; Pratt Burnerd International; Lion Precision; Westwind Air Bearings Ltd; Third Wave AdvantEdge; Susta Tool Handling; Tooling University. I have listed the main companies above, rather than attempting to name individuals within each company, otherwise the list would be simply vast. However, I would like to express my gratitude to each one of them, personally. I would also like to acknowledge the breadth and depth of information obtained from industrially-based journals, such as: Cutting Tool Engineering; American Machinist; Metalworking Production; Machinery and Production Engineering. The publishers of this book Springer, have been most patient with me as I have attempted to meet extended deadlines for the manuscript, for which I am indebted to and can only offer my sincerest thanks. Lastly, if any unfortunate mistakes have inadvertently slipped into the text, or misinterpretations in the draughting of any line diagrams have occurred, it is solely the author’s fault and does not represent any of the companies, or their products, nor that of the individuals mentioned. Graham T. Smith

Contents

1 Cutting Tool Materials  . . . . . . . . . . . . . . . . .   1 1.1 Cutting Technology – an Introduction  .. . .   2 1.1.1 Rationalisation  .. . . . . . . . . . . . . . . . .   2 1.1.2 Consolidation  .. . . . . . . . . . . . . . . . . .   4 1.1.3 Optimisation  .. . . . . . . . . . . . . . . . . . .   4 1.2 The Evolution of Cutting Tool Materials    7 1.2.1 Plain Carbon Steels  .. . . . . . . . . . . . .   7 1.2.2 High-Speed Steels  . . . . . . . . . . . . . . .   7 1.2.3 Cemented Carbide  . . . . . . . . . . . . . .   8 1.2.4 Classification of Cemented Carbide Tool Grades  .. . . . . . . . . . . .   12 1.2.5 Tool Coatings: Chemical Vapour Deposition (CVD)  .. . . . . .   14 1.2.6 Diamond-Like CVD Coatings   . . .   14 1.2.7 Tool Coatings: Physical Vapour Deposition (PVD)  . . . . . . .   17 1.2.8 Ceramics and Cermets  .. . . . . . . . . .   19 1.2.9 Cermets – Coated  . . . . . . . . . . . . . . .   23 1.2.10 Cubic Boron Nitride (CBN)   and Poly-crystalline Diamond (PCD)   .. . . . . . . . . . . . . . . . . . . . . . . . .   25 1.2.11 Natural Diamond  . . . . . . . . . . . . . . .   29 2 Turning and Chip-breaking Technology    2.1 Cutting Tool Technology  . . . . . . . . . . . . . . . .   2.1.1 Turning – Basic Operations   . . . . .   2.1.2 Turning – Rake and Clearance Angles on Single-point Tools  .. . . .   2.1.3 Cutting Insert Edge Preparations    2.1.4 Tool Forces – Orthogonal and Oblique  .. . . . . . . . . . . . . . . . . . . .   2.1.5 Plan Approach Angles  . . . . . . . . . . .   2.1.6 Cutting Toolholder/Insert Selection  .. . . . . . . . . . . . . . . . . . . . . . .  

33 34 34 34 36 39 41 43

2.2 History of Machine Tool Development and Some Pioneers in Metal Cutting  . . .   2.2.1 Concise Historical Perspective of the Development of Machine Tools  . . . . . . . . . . . . . . . . . . . . . . . . .   2.2.2 Pioneering Work in Metal Cutting – a Brief Resumé  . . . . . .   2.3 Chip-Development  .. . . . . . . . . . . . . . . . . . .   2.4 Tool Nose Radius  . . . . . . . . . . . . . . . . . . . . .   2.5 Chip-Breaking Technology  . . . . . . . . . . . .   2.5.1 Introduction to Chip-Breaking     2.5.2 The Principles of Chip-Breaking    2.5.3 Chip-Breakers   and Chip-Formers  .. . . . . . . . . . . .   2.5.4 Helical Chip Formation  .. . . . . . .   2.5.5 Chip Morphology  . . . . . . . . . . . . .   2.5.6 Chip-Breaker Wear  .. . . . . . . . . . .   2.6 Multi-Functional Tooling  .. . . . . . . . . . . . .  

50 50 51 54 62 66 66 68 69 71 75 79 79

3 Drilling and Associated Technologies    87 3.1 Drilling Technology   . . . . . . . . . . . . . . . . . .   88 3.1.1 Introduction to the Twist Drill’s Development  . . . . . . . . . . .   88 3.1.2 Twist Drill Fundamentals  .. . . . .   88 3.1.3 The Dynamics of Twist Drilling Holes  .. . . . . . . .   96 3.1.4 Indexable Drills  . . . . . . . . . . . . . . .   103 3.1.5 Counter-Boring/Trepanning  . . .   107 3.1.6 Special-Purpose, or Customised Drilling and Multi-Spindle Drilling  .. . . . . . . . . . . . . . . . . . . . . .   110 3.1.7 Deep-Hole Drilling/  Gun-Drilling   . . . . . . . . . . . . . . . . .   113 3.1.8 Double-Tube Ejector/  Single-Tube System Drills  .. . . . .   115



Contents

3.1.9

Deep-Hole Drilling –   Cutting Forces and Power  .. . . . .   3.2 Boring Tool Technology – Introduction    3.2.1 Single-Point Boring Tooling  .. . .   3.2.2 Boring Bar Selection of: Toolholders, Inserts   and Cutting Parameters  .. . . . . . .   3.2.3 Multiple-Boring Tools  . . . . . . . . .   3.2.4 Boring Bar Damping   .. . . . . . . . .   3.2.5 ‘Active-suppression’   of Vibrations  .. . . . . . . . . . . . . . . . .   3.2.6 Hard-part Machining, Using Boring Bars  .. . . . . . . . . . . .   3.3 Reaming Technology – Introduction  . . .   3.3.1 Reaming – Correction of Hole’s Roundness Profiles  . . .   3.3.2 Radially-Adjustable Machine Reamers  . . . . . . . . . . . . .   3.3.3 Reaming – Problems and Their Remedies  . . . . . . . . . . .   3.4 Other Hole-Modification Processes  . . . .   4

117 117 118 122 124 126 127 128 133 135

6

139

6.1 6.2

142 142

4.4 4.5

  172   177   177

5 5.1 5.2 5.3

Threading Technologies  . . . . . . . . . . . . . . Threads   .. . . . . . . . . . . . . . . . . . . . . . . . . . . . . Hand and Machine Taps  .. . . . . . . . . . . . . . Fluteless Taps  .. . . . . . . . . . . . . . . . . . . . . . . .

       

4.2 4.3

6.3 6.4

Milling Cutters and Associated Technologies  . . . . . . . . . Milling – an Introduction  . . . . . . . . . . . . . 4.1.1 Basic Milling Operations  .. . . . . . 4.1.2 Milling Cutter Geometry – Insert Axial and Radial Rake Angles  4.1.3 Milling Cutter – Approach   Angles  .. . . . . . . . . . . . . . . . . . . . . . . 4.1.4 Face-Milling Engagement – Angles and Insert Density  .. . . . . 4.1.5 Peripheral Milling Cutter Approach Angles –   Their Affect on Chip Thickness  4.1.6 Spindle Camber/Tilt – when Face-Milling  . . . . . . . . . . . . Pocketing, Closed-Angle Faces,   Thin-Walled and Thin-Based   Milling Strategies  . . . . . . . . . . . . . . . . . . . . . Rotary and Frustum-Based Milling Cutters – Design and Operation  . . . . . . . Customised Milling Cutter Tooling  .. . . . Mill/Turn Operations  . . . . . . . . . . . . . . . . .

4.1

5.4 Threading Dies  . . . . . . . . . . . . . . . . . . . . . . . 5.5 Thread Turning – Introduction  . . . . . . . . 5.5.1 Radial Infeed Techniques  . . . . . . 5.5.2 Thread Helix Angles,   for Single-/Multi-Start Threads  5.5.3 Threading Insert Inclination  . . . 5.5.4 Thread Profile Generation  . . . . . 5.5.5 Threading Turning –   Cutting Data and Other Important Factors  . . . . . . . . . . . . . 5.6 Thread Milling  .. . . . . . . . . . . . . . . . . . . . . . . 5.7 Thread Rolling – Introduction  . . . . . . . . . 5.7.1 Thread Rolling Techniques   . . . .

  149   150   151

6.5

  155   158   160   163   166   169

181 182 182 189

7

  189   191   193   195   195   198        

Modular Tooling and Tool Management  .. . . . . . . . . . . . . . .   Modular Quick-Change Tooling  . . . . . . .   Tooling Requirements for Turning Centres  .. . . . . . . . . . . . . . . . . .   Machining and Turning Centre Modular Quick-Change Tooling  . . . . . . . . . . . . . . . .   Balanced Modular Tooling –   for High Rotational Speeds  . . . . . . . . . . . .   Tool Management  .. . . . . . . . . . . . . . . . . . . .   6.5.1 The Tool Management Infrastructure  .. . . . . . . . . . . . . . . .   6.5.2 Creating a Tool Management and Document Database  .. . . . . .   6.5.3 Overall Benefits of a Tool Management System  .. . . . . . . . . .   6.5.4 Tool Presetting Equipment and Techniques for Measuring Tools  . . . . . . . . . . . . . .   6.5.5 Tool Store and its Presetting Facility – a Typical System  . . . . .   6.5.6 Computerised-Tool Management – a Practical Case for ‘Stand-alone’ Machine Tools   

Machinability and Surface Integrity  . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.1 Machinability  .. . . . . . . . . . . . . . . . . . . . . . . . 7.1.1 Design of Machinability Tests and Experimental Testing Programmes  .. . . . . . . . . . 7.2 Machined Roundness   .. . . . . . . . . . . . . . . . 7.2.1 Turned Roundness –   Harmonics and Geometrics  .. . . 7.3 Chatter in Machining Operations  . . . . . .

200 203 206 209

211 212 216 221 230 233 238 240 244 245 261 264

  269   270   270   285   291   294

Contents

8.8.1

7.3.1

7.4 7.5

7.6 7.7

7.8 7.9 7.10

Chatter and Chip Formation – Significant Factors Influencing its Generation  .. . . . . . . . . . . . . . . . 7.3.2 Chatter – Important Factors Affecting its Generation  . . . . . . . 7.3.3 Stability Lobe Diagrams  . . . . . . . Milled Roundness – Interpolated Diameters  . . . . . . . . . . . . . . . . . . . . . . . . . . . . Machined Surface Texture   . . . . . . . . . . . . 7.5.1 Parameters for Machined Surface Evaluation  .. . . . . . . . . . . . 7.5.2 Machined Surface Topography  7.5.3 Manufacturing Process   Envelopes  .. . . . . . . . . . . . . . . . . . . . 7.5.4 Ternary Manufacturing   Envelopes (TME’s)  . . . . . . . . . . . . Machining Temperatures  . . . . . . . . . . . . . . 7.6.1 Finite Element Method   (FEM)   . . . . . . . . . . . . . . . . . . . . . . . Tool Wear and Life  .. . . . . . . . . . . . . . . . . . . 7.7.1 Tool Wear  .. . . . . . . . . . . . . . . . . . . . 7.7.2 Tool Life  .. . . . . . . . . . . . . . . . . . . . . 7.7.3 Return on the Investment (ROI)  Cutting Force Dynamometry  . . . . . . . . . . Machining Modelling and Simulation  Surface Integrity of Machined Components – Introduction  . . . . . . . . . . . 7.10.1 Residual Stresses in Machined Surfaces  .. . . . . . . . .

XI

  297   297   300   301   305   308   317   324   326   326              

328 330 331 337 342 343 350

  360   360

8 Cutting Fluids  .. . . . . . . . . . . . . . . . . . . . . . .   8.1 Historical Development   of Cutting Fluids  .. . . . . . . . . . . . . . . . . . . . .   8.2 Primary Functions of a Cutting Fluid  .. .   8.3 High-Pressure Jet-Assisted Coolant Delivery  .. . . . . . . . . . . . . . . . . . . . . . . . . . . . .   8.4 Types of Cutting Fluid  .. . . . . . . . . . . . . . . .   8.4.1 Mineral Oil, Synthetic, or Semi-Synthetic Lubricant?  .. .   8.4.2 Aqueous-Based Cutting Fluids    8.4.3 Water Quality  .. . . . . . . . . . . . . . . .   8.5 Cutting Fluid Classification – According to Composition  .. . . . . . . . . . . . . . . . . . . . . .   8.6 Computer-Aided Product Development    8.6.1 Cutting Fluid – Quality Control    8.7 Selecting the Correct Cutting Fluid  .. . . .   8.7.1 Factors Affecting Choice  .. . . . . .   8.7.2 Selection Procedure  . . . . . . . . . . .   8.8 Care, Handling, Control and Usage – of Cutting Fluids  .. . . . . . . . . . . . . . . . . . . . .  

381 382 383 383 387 392 395 397 398 398 404 407 407 408 409

8.9 8.10 8.11

8.12

Product Mixing – Preparation   of a Aqueous-Based Cutting Fluids  . . . . . . . . . . . . . . . . . . . . . . . . 8.8.2 Monitoring, Maintenance and Testing of Cutting Fluid – in Use  . . . . . . . . . . . . . . . . . . . . . . . . Multi-Functional Fluids  . . . . . . . . . . . . . . . Disposal of Cutting Fluids  .. . . . . . . . . . . . Health and Safety Factors – Concerning Cutting Fluid Operation and Usage  . . . . 8.11.1 Cutting Fluid-Based Health Issues  .. . . . . . . . . . . . . . . . . Fluid Machining Strategies: Dry; Near-Dry; or Wet  . . . . . . . . . . . . . . . . . . . . . 8.12.1 Wet- and Dry-Machining – the Issues and Concerns  . . . . . . . 8.12.2 Near-Dry Machining  . . . . . . . . . .

  410   411   417   417   418   420   425   425   426

9 Machining and Monitoring Strategies    9.1 High Speed Machining (HSM)  .. . . . . . . .   9.1.1 HSM Machine Tool Design Considerations  .. . . . . . . . . . . . . . .   9.2 HSM Dynamics – Acceleration and Deceleration  .. . . . . . . . . . . . . . . . . . . . .   9.2.1 HSM Dynamics – Servo-Lag  .. .   9.2.2 Effect of Servo-lag and Gain on Corner Milling  . . .   9.2.3 Effect of Servo-Lag and Gain Whilst Generating Circular   Paths  . . . . . . . . . . . . . . . . . . . . . . . . .   9.2.4 CNC Processing Speed  . . . . . . . .   9.3 HSM – with Non-Orthogonal Machine Tools and Robots  . . . . . . . . . . . . . . . . . . . . .   9.4 HSM – Toolholders/Chucks  . . . . . . . . . . .   9.4.1 Toolshank Design and Gripping Pressures  . . . . . . . .   9.4.2 Toolholder Design and Spindle Taper   .. . . . . . . . . . . .   9.5 Dynamic Balance of Toolholding Assemblies  . . . . . . . . . . . . . . . . . . . . . . . . . . .   9.5.1 HSM – Problem of Tool Balance    9.5.2 HSM – Dynamic Balancing Machine Application  . . . . . . . . . .   9.6 HSM – Research Applications  .. . . . . . . . .   9.6.1 Ultra-High Speed: Face-Milling Design and Development  . . . . . .   9.6.2 Ultra-High Speed: Turning Operations  . . . . . . . . . . .   9.6.3 Ultra-High Speed: Trepanning Operations  .. . . . . . . . . . . . . . . . . . .  

431 432 434 445 446 448 448 449 451 458 458 465 467 469 472 474 474 480 484

XII

Contents

9.6.4

9.7 9.8

9.9 9.10

Artefact Stereometry: for Dynamic Machine Tool Comparative Assessments  .. . . . . HSM: Rotating Dynamometry  . . . . . . . . . Complex Machining: of Sculptured Surfaces  .. . . . . . . . . . . . . . . . 9.8.1 Utilising the Correct Tool for Profiling: Roughing and Finishing  . . . . . . . . . . . . . . . . . 9.8.2 Die-Cavity Machining – Retained Stock  . . . . . . . . . . . . . . . . 9.8.3 Sculptured Surface Machining – with NURBS  . . . . . . . . . . . . . . . . . . 9.8.4 Sculptured Surface Machining – Cutter Simulation  . . . . . . . . . . . . . Hard-Part Machining   .. . . . . . . . . . . . . . . . 9.9.1 Hard-Part Turning  . . . . . . . . . . . . 9.9.2 Hard-Part Milling  . . . . . . . . . . . . . Ultra-Precision Machining  . . . . . . . . . . . . 9.10.1 Micro-Tooling  . . . . . . . . . . . . . . . . 9.10.2 Micro-Machine Tools  .. . . . . . . . .

  486   493   496   496   498   502              

505 507 508 511 516 518 525

9.10.3 Nano-Machining and Machine Tools  . . . . . . . . . . . . 9.11 Machine Tool Monitoring Techniques  9.11.1 Cutting Tool Condition Monitoring   .. . . . . . . . . . . . . . . . . . 9.11.2 Adaptive Control and Machine Tool Optimisation  .. . . . . . . . . . . . 9.11.3 Artificial Intelligence: AI and Neural Network Integration  .. . . . . . . . . . . . . . . . . . . 9.11.4 Tool Monitoring Techniques – a ‘Case-Study’   . . . . . . . . . . . . . . . .

  526   531   531   535   538   538

Appendix  .. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .   549 About the Author  .. . . . . . . . . . . . . . . . . . . . . . . . .   587 Subject Index  .. . . . . . . . . . . . . . . . . . . . . . . . . . . . .   589



1

Cutting Tool Materials ‘What is the use of a book’ , thought Alice, ‘without pictures or conversations?’ 

LEWIS CARROLL

(1832–1898) [Alice in Wonderland, Chap. 1]



Chapter 1

1.1 Cutting Technology – an Introduction Previously, many of the unenlightened manufacturing companies, having purchased an expensive and sophisticated new machine tool, considered cutting tool technology as very much an afterthought and supplied little financial support, or technical expertise to purchase these tools. Today, tooling-related technologies are treated extremely seriously, as it is here that optimum production output, consistency of machined product and value-added activities are realised. Often companies feel that to increase productivity – to offset the high capital investment in the plant and to amortise such costs (i.e. pay-back), is the most advantageous way forward. This strategy can create ‘bottlenecks’ and disrupt the harmonious flow of production at later stages within the manufacturing environment. Another approach might be to maximise the number of components per hour, or alternatively, drive down costs at the expense of shorter tool life, which would increase the non-productive idle time for the production set-up. Here, the prime tooling factor should not be for just a marginal increase in productivity and efficiency, nor the perfection of any particular operation. If ‘bottlenecks’ in component production occur, they can readily be established by piles of machined parts sitting on the shop floor awaiting further valuedadded activities to be undertaken. These ‘line-balance’ production problems need to be addressed by achieving improved productivity across the whole operation, perhaps by the introduction of a Taguchi-type component flow analysis system within the manufacturing facility. The well-known phrase that: ‘No machine is an island’ (i.e. for part production) and that manufacturing should be thought of as ‘One big harmonious machine’ and not a lot of independent problems, will create a means by which increases in productivity can be achieved. The cutting tool problems, such as: too wide a range of tooling inventory, inappropriate tools/out-dated tooling, or not enough tools for the overall operational



Tooling refers not only to non-consumable items such as: cutting tools and inserts, tool holders, tool presetters, screws, washers and spacers, screwdrivers/Allen keys, tool handling equipment, but also consumable items, such as hand wipes, grease/oils employed in tool kitting and cutting fluids, etc.

requirements for a specific manufacturing environment, can be initially addressed by employing the following tooling-related philosophy – having recently undertaken a survey of the current status of tooling within the whole company: • Rationalisation • Consolidation • Optimisation NB  These three essential tool-related factors in establishing the optimum tooling requirements for the current production needs, will be briefly reviewed.

1.1.1 Rationalisation In order to be able to rationalise the tools within the current production facility, it is essential to conduct a thorough appraisal of all the tools and associated equipment with the company. This tooling exercise will be both time-consuming and costly, because it necessitates a considerable manpower resource and needs a means of identifying all the tools and inserts currently utilised, in some logical and tabulated manner. Such surveys are often best conducted by utilising a primitive but efficient tool-card indexing system in the first instance. Details, such as: tool type and its tooling manufacturer, quantity of tools in use and the current levels of stock in the tool store, their current location(s), feeds and speeds utilised, together with any other relevant tool-related details are indexed on such cards. Once these tooling facts have been established, then they can be loaded into either a computerized tool management system database, or recorded onto an uncomplicated tooling database for later interrogation. Having established the current status of the tooling within the manufacturing facility, this allows for a tooling rationalisation campaign to be developed. Tool rationalisation (Fig. 1) consists of looking at the results of the previous tooling survey and significantly reducing the number of tooling suppliers for particular types of tools and inserts. This initial rationalisation policy has the twin benefits of minimising tooling suppliers with their distinct varieties of tools, while enabling bulk purchase of such tools from the remaining suppliers, at preferential financial rates of purchase. Moreover, by using less tooling companies whilst purchasing bulk stock, this has the bonus of making you one of their prime customers with their undivided at-

Cutting Tool Materials

. Figure 1.  Rationalisation of cutting inserts, can have a dramatic effect on reducing the tooling and workholding inventory. [Courtesy of Sandvik Coromant]





Chapter 1

tention, should the need for later ‘tool problem-solving’ of manufacturing clichés in production occur.

1.1.2 Consolidation For any tooling that remains after the rationalisation exercise, these should be consolidated, by reducing the number of insert grades, by at least half – which often proves to have little effect on production capability. By grouping inserts by their respective sizes, shapes and say, nose radius for example, this will eliminate many of the less-utilised inserts, enabling the potential for bulk purchase from the tooling supplier, with an attendant reduction in tool costs. From this consolidation activity, it may now be possible to purchase higher-performance grade cutting inserts, that meet a wider application range, enabling the consolidation to be even more effective. Furthermore, such improved inserts, will probably have a longer tool life and can be utilised at higher speeds, which probably negates their extra cost, over the previously used inserts. If fewer grades of insert are stocked, the tooling/application engineers will be acquainted with them much more thoroughly and this will result in a added effectiveness and a consistent application, for the production of machined components – more will be said on this latter point in the next section on Optimisation.

1.1.3 Optimisation By consolidating the tooling, it allows productivity to be boosted by optimisation of the cutting insert grades. For example, in turning operations, the depth of cut can probably maximised and, as a result, the number of passes along, or across the part can be minimised. It can be argued that increasing the depth of cut leads to a reduction in the subsequent tool life (in terms of minutes of cutting per edge). However, there are fewer cuts per part, so each machined workpiece requires less overall cutting and as a result, many more parts per edge can be produced. More important, are that the cycle times for roughing operations be reduced: a reduction in the number of roughing passes from three to one, results in a 66% reduction in the cycle time. This increase in productivity may justify any potential decrease in tool life, on the basis that it could reduce, or eliminate a potential ‘bottleneck’ in latter production processes of the part’s manufacture. To extract the maximum productivity from today’s

high-performance grades, they must be worked hard and pushed to their fullest capabilities. When tool life is reduced by increasing the depth of cut, there are several ways that a such loss can be minimised. For example, it is known that the size of the insert’s nose radius has a pronounced effect on tool life, so by doubling the depth of cut this can, in the main, allow for a larger nose radius – assuming that the component feature allows access. If an increase in nose radius cannot be utilised, then increasing the insert’s size will help to offset any higher wear rates, by providing a better heat dissipation for the action of cutting. The accepted turning practice when roughingout, is that no more than half the insert’s cutting edge length should be utilised, because as the depth of cut approaches this value, a larger insert is recommended. Where large depths of cut are used in combination with high feedrates, a roughing grade insert geometry promotes longer tool life, than a general-purpose insert. Often, by using a single-sided insert rather than a double-sided one for roughing cuts, this has the twin benefits of increased productivity and longer tool life (in terms of machined parts per edge). Normally, single-sided inserts are recommended whenever the depth of cut and feedrate are so high that the surface speed must be reduced below the grade’s normal range, in order to maintain an adequate tool life. Such inserts should be considered if erratic insert breakage occurs. Later to be discussed in the chapter on Machinability and Surface Integrity, is the fact that the highest temperature region on the tool’s rake face is not at the cutting edge; but in the vicinity on the chip/tool interface where chip curling occurs – this is some distance back and where the crater is formed. The position for this highest isothermal region can vary, depending upon the feedrate. For example, if the feedrate is increased, the highest temperature zone on the insert’s face will move away from the cutting edge; conversely, if the feedrate is now reduced, this region moves toward the cutting edge. This phemomena means that if the feedrate is too low for the chosen insert geometry and edge preparation, heat will be concentrated too near the cutting edge and insert wear will be accelerated. Thus, by increasing the feedrate, it has the affect of moving the maximum heat zone away from the insert’s edge and is so doing, extends tool life – in terms of minutes of actual cut-time per edge. As a result, each machined part will be produced in less time and at higher feeds, so the tool life in terms of parts per edge will also increase.

Cutting Tool Materials

As a result of the inappropriate use of cutting data, such as incorrect feedrates employed for the chosen insert geometry, this can produce a number of undesirable symptoms. These symptomatic problems include: extremely shortened tool life, edge chipping and insert breakage are likely if feedrates are too high, whereas when feeds are too low, chip control becomes a problem. Once the insert grades have been consolidated with their associated geometries, it is relatively easy to determine the feedrates for a selected grade of workpiece materials. Tooling suppliers can recommend a potential insert grade for particular component part material, with an initial selection of insert grade, such surface speeds being indicated in the Appendix. These inserts can be optimised by ‘juggling’ the grades and geometries marginally around the specified values, this may allow feedrates to be increased and should provide a significant pay-off in terms of improved productivity, at little, or no additional capital expenditure. If the cutting speed is increased rather than the feed, a point is reached where any increase in surface speed will result in a decrease in productivity. In other words, cutting too fast will mean spending more time changing tools than making parts! Equally, by cutting too slowly, the tool will last much longer, but this is at the expense of the number of machined parts produced per shift. If these statements are correct, what is the ‘right’ surface speed? This question will now be discussed more fully. If we return to the theme previously mentioned, namely: ‘No machine is an island’ and treat the production shop as: ‘One big machine’ , it can be stated that every shop should determine its own particular manufacturing objectives – when considering both cutting speeds and tool life. Typical objectives for tool life might be the completion of a certain number of parts before indexing the insert, or adopting a ‘sister tool’ , or alternatively, insert indexing after one/part of a shift. If very expensive components are being machined, the main goal is to avoid catastrophic insert



‘Sister tooling’ is the term that refers to a duplicate tool (i.e. having the same tool offsets) held in the turret/magazine and can be automatically indexed to this tool, to minimise down-time when changing tools. Such a ‘sister tool’ , can be pre-programmed into the CNC controller of the machine tool, to either change after a certain number of parts has been produced, or if the tool life has been calculated, then when the feed function on the CNC has decremented down to this preset value, then the ‘sister tool’ is selected.



failure, which on a finishing cut, would probably result in scrapping the part. When exceedingly large parts are to be machined, the objective may simply be to complete just one part per insert, or in an even more extreme situation, just one pass over the part. When small parts are being produced, then the tool life can be controlled in order to minimise dimensional size variation with in-cut time. This strategy of tool life control, reduces the need for frequent adjustment of tool offset compensations in the CNC controller. However, one idea shared by all of these strategic production approaches, is that by optimising the surface speed, the manufacturing objectives will be realised. As a consequence of this approach to production, there is no correct surface speed for any specific combination of material and insert grade, the optimum surface speed depends upon the company’s manufacturing requirements at this time. When long production runs occur, these are ideal because it allows cutting data experimentation to discover the optimum speed for a particular production cycle. Sometimes it is not possible to find the speed to exactly meet the production demands and, a change of insert grades, to one of the higher-technology materials may be in order. If short production runs are necessary, this can often rule out any experimentation with insert grades, but by consultation with a ‘cutting tool expert’ , or reference to the published cutting literature the answer may be found to the problem of insert optimisation. However a cautionary note, care must be taken when utilising published recommendations, as they should only be employed as guidelines, to help initiate the job into production. Comparison with a known starting point within the recommended range for specific production conditions, namely for: large depths of cut, high feedrates, very long continuous cuts, significant interrupted cuts, workpiece surface scale and the absence of coolant, would all suggest that reductions in surface speed should be initially considered. Conversely, production conditions that result in: short lengths of cut, shallow depths of cut, low feedrates, smooth uninterrupted cuts, clean pre-turned, or bright-drawn wrought workpiece materials and flood coolant, having a very rigid setup, suggests that the recommended ranges for the insert could be exceeded, while still maintaining an acceptable tool life. It should be remembered that the main requirement is for an overall increase in production output and not perfection. After the analysis, when the tooling inventory has been consolidated, there will be fewer and



Chapter 1

more versatile insert grades and geometries that need to be considered. This smaller insert inventory allows a detailed appreciation of how to optimise the speeds and feeds in combination with depths of cut more effectively, for the desired production objectives. By optimisation here of the machining parameters, this allows full utilisation of the capital equipment, with the result that large improvements in overall manufacturing efficiency can be expected. It is evident from this discussion concerning optimisation, that the parameters of: tool life, feedrate and cutting speed form a complex relationship, which is illustrated in Fig. 2a. Consequently, if you change one parameter, it will affect the others, so a compromise has to be reached to obtain the optimum performance from a cutting tool. Preferably, the ideal cutting tool should have superior performance if five distinct areas (see Fig. 2b): • Hot hardness – is necessary in order to maintain sharp and consistent cutting edge at the elevated temperatures that are present when machining. NB  If the hot hardness of the tooling is not sufficient for the temperature generated at the tool’s tip, then it will degrade quickly and be useless.

• Resistance to thermal shock – this is necessary in

order to overcome the effects of the continuous cycle of heating and cooling that is typical in a milling operation, or when an intermittent cutting operation occurs on a lathe (e.g. an eccentric turning operation).

NB  If this thermal shock resistance is too low, then rapid wear rates can be expected, typified in the past, by ‘comb cracks’ on High-speed steel (HSS) milling cutters.

• Lack of affinity – this condition should be present between the tool and the workpiece, since any degree of affinity will lead to the formation of a builtup edge (BUE) – see the chapter on Machinability and Surface Integrity.

. Figure 2.  The main factors affecting cutting tool life, under ‘steady-state’ cutting conditions

NB  This BUE will modify the tool geometry, leading to poorer chip-breaking ability, with higher forces generated, leading to degraded workpiece surface finish. Ideally, the cutting edge should be inert to any reaction with the workpiece.

Cutting Tool Materials



• Resistance to oxidation – a cutting edge should mately 5 m·min–1. These early cutting tools frequently have the desirable condition of having a high resistance to oxidation.

NB  This oxidation resistance of the cutting tool is necessary, in order to reduce the debilitating wear that oxidation can produce when machining at elevated temperatures.

• Toughness – allows the cutting edge of the insert to absorb the cutting forces and shock loads produced whilst machining, particularly relevant when intermittent cutting operations occur.

NB  If an insert is not sufficiently tough, then when unwanted vibrations are induced, this can result in either premature failure, or worse, a shattered cutting edge. Cutting tool manufacturers, by careful balancing of these five factors for the ideal cutting tool, can produce grades of inserts which distinctly vary, allowing a wide range of workpiece materials to be machined through the selection of the correct insert grade for a particular material. In recent years, tooling manufacturers have produced wider ranges of workpiece-cutting ability from fewer types of inserts, across a diverse range of speeds and feeds, allowing tooling inventories to be reduced even further. This brief introduction showing how and in what manner correct tooling can be used to increase production output, needs to be considered against the current situation of advances in cutting tool materials and their selection – this will be the theme of the next section.

1.2 The Evolution of Cutting Tool Materials 1.2.1 Plain Carbon Steels Prior to 1870, all turning tooling materials were produced from plain carbon steels, with a typical composition of 1% carbon and 0.2% manganese – the remainder being iron. Such a tool steel composition meant that it had a low ‘hot-hardness’ (i.e, its ability to retain a cutting edge at elevated temperatures), as such, the cutting edge broke down at temperatures approaching 250°C, this in reality kept cutting speeds to approxi-

had quench cracks present which severely weakened the cutting edge, as a result of water hardening at quenching rates greater than 1000°C/second (i.e. necessary to exceed the critical cooling velocity – to fully harden the steel), upon manufacture. By 1870, Mushet (working in England), had introduced a more complex steel composition, containing: 2% carbon, 1.6% manganese, 5.5% Tungsten and 0.4% chromium, with the remainder being iron. The advantage of this newly developed steel was that it could be air-hardened, this was a significantly less drastic quench than using a water quenchant. Mushet’s steel had greater ‘hothardness’ and could be utilised at cutting speeds up to 8 m·min–1. This turning tool material composition, was retained until around 1900, but with the level of chromium gradually superseding that of manganese.

1.2.2 High-Speed Steels Around the turn of the century in the United States, fundamental metallurgical work was being undertaken by F.W. Taylor and his associate M. White and by 1901, these researchers had greatly improved the overall tool steel and slightly modifying its composition with a material that was to be known as High-speed steel (HSS), enabling cutting speeds to approach 19 m·min–1. High-speed steel was not a new material, but basically an innovative heat treatment procedure. The typical metallurgical composition of HSS was: 1.9% carbon, 0.3% manganese, 8% tungsten, and 3.8% chromium, with iron the remainder. Taylor and White’s tool steel mainly differed from that of Mushet’s by an increased amount of tungsten and a further replacement of manganese by chromium. By 1904, the content of carbon had been reduced, allowing for more ease in forging this HSS. Further rapid development of the HSS occurred over the next ten years, with tungsten content increased to improve its ‘hot-hardness’. Around this time, Dr J.A. Matthews found that vanadium additions had improved the material’s abrasion resistance. By 1910, the content of tungsten had increased to 18%, with 4% chromium and 1% vanadium, hence the wellknown 18:4:1 HSS had arrived, its metallurgical composition continued with only marginal modifications over the next 40 years. Of the modifications to HSS during this time, of note was that in 1923 the so-called ‘super’ HSS was developed, although this variant did not become commercially viable until 1939, when Gill reduced the tungsten content to enable the tool steel



Chapter 1

to be successfully hot-worked. Around 1950 in the United States, M2 HSS was introduced, having some of the tungsten content replaced by that of molybdenum. This gave the approximate M2 HSS metallurgical composition as: 0.8% C, 4% Cr, 2% V, 6% W and 5% Mo – Fe being the remainder. In this form, the M2 HSS could withstand machining temperatures of up to 650°C (ie the cutter glowing dull red) and still maintain a cutting edge. In 1970, Powder Metallurgy (P/M) by metallurgical processing via hot isostatic pressing (HIP), was introduced for the production of HSS, with careful control of elemental particle size; afterward the sintered product is forged then hot-rolled. This HSS (HIP) processing gave a uniformly distributed elemental matrix, overcoming the potential segregation and resulting non-homogenous structure that would normally occur when ingot-style HSS forging. Such P/M processing techniques enable the steel-making company to ‘tailor’ and specify the exact metallurgical composition of alloying elements, this would allow the newly-developed sintered/forged HSS tooling to approach that of the performance of cemented carbides, in terms of inherent wear resistance, hardness and toughness. In Fig. 3, a comparison of just some of the tooling materials is highlighted, here, fracture toughness is plotted against hardness to indicate the range of influence of each tool material and the comparative relative merits of one material against another, with some of their physical and mechanical properties tabulated in Fig. 3b. A typical sintered micro-grained HSS of today, might contain: 13% W, 10% Co, 6% V, 4.75% Cr and 2.15% C – Fe the remainder. One reason for the ‘keen’ cutting edge that can be retained by sintered micrograined HSS, is that during P/M processing the rapid atomisation of the particles produces extremely fine carbides of between 1 to 3 µm in diameter – which fully support the cutting edge, whereas HSS produced from an ingot, has carbides up to 40 µm in diameter. By way of illustration of the benefits of the latest micro-grained HSS – in the uncoated condition – when compared to its metallurgical competitor of cemented carbide, HSS has a bend, or universal tensile strength of between 2,500 to 6,000 MPa – this being dependent on metallurgical composition, whereas cemented carbide tooling has a bend strength of between 1,250 to 2,250 MPa. These metallurgical tool processing techniques have significantly improved sintered micrograined HSS enabling for example, high-performance drilling, reaming and tapping to be realised. Coating by either single-, or multiple-coating has been shown to significantly enhance any tooling mate-

rial, but this is a complex subject and more will be said on this subject shortly.

1.2.3 Cemented Carbide Possibly the widest utilised cutting tool materials today are the cemented carbide family of tooling, of which the group derived from tungsten carbide is most readily employed. Prior to discussing the physical metallurgy and expected mechanical/physical characteristics of cemented carbides, it is worth looking into the complex task of insert selection. In Fig. 4, just a small range of the material types, grades, shapes of inserts and coatings by a leading cutting tool company is depicted. Highlighting the complex chip-breaker geometries, necessary to both develop and break chips and evacuate them efficiently from the workpiece’s surface region. To give a simplified impression of just some of the tooling insert variations and permutations available from a typical tooling manufacturer, if 10 insert grades are listed, in 6 different shapes, with 12 differing chip breakers and five nose radii in the tooling catalogue, this equates to 10 × 6 × 12 × 5, or 3,600 inserts. In reality, there are a number of other important features that could extend this cutting insert permutation to well over five significant figures – for potential insert selection. When the permutated insert number reaches this level of complexity, selecting the optimum combination of insert characteristics becomes more a matter of luck than skill. Tungsten (synonym Wolfram, hence the chemical symbol W), is the heaviest metal in the group VIB in Mendeleev’s Period Chart (i.e. atomic number 74). It was named after the German word wolfram – from the mineral wolframite – as it was derived from the term wolf rahm, because the ore was said to interfere with tin smelting – supposedly devouring the tin. Whereas the term tungsten, was coined from the Swedish tung sten, meaning heavy stone. Hence, in 1923, the German inventor K. Schröter produced the first metal matrix composite, known today as cemented carbides. In these first cemented carbides, Schröter combined tungsten monocarbide (WC) particles embedding them in a cobalt matrix – these particles acted as a very strong binder. Cemented carbide is a hard transition metal carbide ranging from 60% to 95% bonded to cobalt, this being a more ductile metal. The carbides vary, ranging from having hexagonal structures, to a solid solution of titanium, tantalum and niobium carbides to that of a NaCl structure. Tungsten carbide does not dissolve

Cutting Tool Materials



. Figure 3.  Cutting tool materials – toughness versus hardness – and their typical material characteristics. [Courtesy of Mitsubishi Carbide]

any transition metals, but it can melt those carbides found in solid solution. Powder metallurgy processing route – liquid-phase sintering – is utilised to produce cemented carbides, as melting only occurs at very high temperatures and there is a means of reducing tung-

sten powder using hydrogen from chemically purified ore. Ore reduction can be achieved by the manipulation of the processing conditions, enabling the grain size to be controlled/modified as necessary. The uniform grain sizes of tungsten carbide today can range

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from 0.2 to 7 µm – enabling a final sintered product to be carefully controlled. Moreover, by additions of fine cobalt at a further processing stage, then wet milling the constituents, allows for precise and uniform control of the grain size – producing a fine powder. Prior to sintering, the milled powder can be spray-dried giving a free-flowing spherical powder aggregate, with the addition of lubricant to aid in its consolidation (i.e. pressing into a ‘green compact’). Sintering normally occurs at temperatures of 1500°C in a vacuum, which reduces the porosity from about 50% that is in the ‘green state’ , to less than 0.01% porosity by volume

in the final cutting insert condition. The low level of porosity in the final product is the result of ‘wetting’ by the liquid present upon sintering. The extent of this ‘wetting’ during liquid-phase sintering, being dependent upon molten binder metal dissolving to produce a pore-free cutting insert, this has excellent cohesion between the binder and the hard particles (see Fig. 5, for typical cemented carbide powders and resulting microstructures). It should be stated that most of the ‘iron-group’ of metals can be ‘wetted’ by tungsten carbide, forming sintered cemented carbide with excellent mechanical integrity.

. Figure 4.  Cutting inserts indicating the diverse range of: shapes, sizes and geometries available, with compositions varying from: cemented carbide, ceramics, cermets, to cubic boron nitride derivatives. [Courtesy of Sandvik Coromant]

Cutting Tool Materials

. Figure 5.  Cemented carbide powders and typical microstructures after sintering. [Courtesy of Sandvik Coromant]

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The desirable properties that enable tungsten carbide to be tough and readily sintered, also cause it to easily dissolve in the iron, producing the so-called ‘straight’ cemented carbide grades. These ‘straight’ grades normally contain just cobalt and have been used to predominantly machine cast iron, as the chips easily fracture and do not usually remain in contact with the insert, reducing the likelihood of dissolution wear. Conversely, machining steel components, requires alternative carbides such as tantalum, or titanium carbides, as these are less soluble in the heated steel at the cutting interface. Even these ‘mixed’ cemented carbide grades will produce a tendency to dissolution of the tool material in the chip, which can limit high speed machining operations. Today, the dissolution tool material can be overcome, by using cutting insert grades based on either titanium carbide, or nitride, together with a cobalt alloy binder. Such grades can be utilised for milling and turning operations at moderate cutting speeds, although their reduced toughness, can upon the application of high feed rates, induce greater plastic deformation of the cutting edge and induce higher tool stresses. These uncoated cutting inserts were very much the product of the past and today, virtually all such tooling inserts are multi-coated to significantly reduce the effects of dissolution wear and greatly extend the cutting edge’s life – more will be said on such coating technology later.

1.2.4 Classification of Cemented Carbide Tool Grades Most cemented carbide insert selection guides group insert grades by the materials they are designed to cut. The international standard for over 30 years used for carbide cutting of workpiece materials is: ISO 5131975E Classification of Carbides According to Use – which has a colour-coding for ease of identification of sub-groups. In its original form, this ISO 513 code utilises 3 broad letter-and-colour classifications (see Fig. 6 for the tabulated groupings of carbides and their various colours, designations and applications):





The workpiece categories are arranged according to their relative chip production characteristics and certain metallurgical characteristics, such as casting condition, hardness and tensile strength. ISO 1832–1991 has clesignations: ‘P’ (Steels, low-alloy); ‘M’ (Stainless steels); ‘K’ (Cast irons); ‘N’ (Aluminium alloys); ‘H’ (Hardened steelas)

• P (blue) – highly alloyed workpiece grades for cutting long-chipping steels and malleable irons,

• M (yellow) – lesser alloyed grades for cutting fer-

rous metals with long, or short chips, cast irons and non-ferrous metals, • K (red) – is ‘conventional’ tungsten carbide grades for short-chipping grey cast irons, non-ferrous metals and non-metallic materials. Under this previous ISO system (Fig. 6), both steels and cast irons can be found in more than one category, based upon their chip-formation characteristics. Each grade within the classification is given a number to designate its relative position in a continuum, ranging from maximum hardness to maximum toughness. This original ISO 513 Standard, has been modified over the years by many tooling manufacturers, introducing more discretion in their selection and usage. Typical of this manufacturer’s modified approach, is that found by just one American tooling company, forming a simple colour-coding matrix, such as the three designated manufacturer’s chip-breaker grades (such as: F, M and R) and three workpiece material grades (i.e. Steel, Stainless steel and Cast iron) – producing a ninecell grid. While another manufacturer in Europe, has produced a more discerning matrix, based upon adding the ‘machining difficulty’ into the matrix, producing a 3 × 3 × 3 matrix – producing a twenty seven cell grid. In this instance, the tooling manufacturer uses the workpiece material to determine the tool material needed. The insert geometry is still selected according to the type of machining operation to be undertaken, while the insert grade is determined by the application conditions – whether such factors as interrupted cuts occur, forging scale on the part are present and the desired machining speed being designated as: good, average, or difficult. NB  These manufacturer’s matrices for the tooling insert selection process will get a user to approximately 90% of optimum, with the ‘fine-tuning’ (optimisation) requiring both technical appreciation of information from the manufacturer’s tooling catalogue/recommendations from ‘trouble- shooting guides’ and any previous ‘know-how’ from past experiences – as necessary.

Cutting Tool Materials

. Figure 6.  Classification of carbides according to use. [Courtesy of Seco Tools]

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1.2.5 Tool Coatings: Chemical Vapour Deposition (CVD) Rather quaintly, the idea of introducing a very thin coating onto a cemented carbide cutting tool originated with the Swiss Watch Research Institute, using the chemical vapour deposition (CVD) technique. In the 1960’s, these first hard coatings were applied to cemented carbide tooling and were titanium carbide (TiC) by the CVD process (Fig. 7 shows a schematic view of the CVD process) at temperatures in the range 950 to 1050°C. Essentially, the coating technique consists of a commercial CVD reactor (Fig. 8a) with cutting tools, or inserts to be hard-coated placed on trays (depicted in Fig. 8b). Prior to coating the tooling situated on their respective trays, these tools should have a good surface finish and sharp corners should have small honed edges – normally approximately 0.1 mm. With the CVD technique, if these honed tool cutting edges are too large, they will not adequately support the coating, but if they are even greater, the cutting edge will be dulled and as a result will not cut efficiently. These tooling trays (Fig 8b) are accurately positioned one above another, being pre-coated with graphite and are then loaded onto a central gas distribution column (i.e tree). The ‘tree’ now loaded with tooling to be coated is placed inside a retort of the reactor (Fig. 8a). This contained tooling within the reactor, is heated in an inert atmosphere until the coating temperature is reached and the coating cycle is initiated by the introduction of titanium tetrachloride (TiCl4) together with methane (CH4) into the reactor. The TiCl4 is a cloud of volatile vapour and is transported into the reactor via a hydrogen carrier gas (H2), whereas CH4 is introduced directly. This volatile cloud reacts on the hot tooling surfaces and the chemical reaction in say, forming a TiC as a surface coating, is:

tooling, then the previously used methane is substituted by a nitrogen/hydrogen gas mixture. For example, if a simple multi-coated charge is required for the tooling, it is completed in the same cycle, by firstly depositing TiC using methane and then depositing TiN utilising a nitrogen/hydrogen gas mixture. As the TiN and TiC are deposited onto the tooling, they nucleate and grow on the carbides present in the exposed surface regions, with the whole CVD coating process taking approximately 14 hours, consisting of 3 hours for heating up, 4 hours for coating and 7 hours for cooling. The thickness of the CVD coating is a function of the reaction concentration, this being the subject of: various gaseous constituents and their respective flow rates, coating temperature and the soaking time at this temperature. The CVD process is undertaken in a vacuum together with a protective atmosphere, in order to minimise oxidation of the deposited coatings. However it should be noted that, in the case of high-speed steel (HSS) tooling such as when coating small drills and taps, the elevated coating temperatures employed, necessitate post-coating hardening heat treatment.

1.2.6 Diamond-Like CVD Coatings

TiCl4 + CH4 → + TiC + 4HCl

Crystalline diamond is only grown by the CVD process on solid carbide tools, because of the high temperatures involved in the process, typical diamond coating temperatures are in the region of 810°C. Such diamondlike tool coatings (Fig. 9), make them extremely useful when machining a range of non-ferrous/non-metallic workpiece materials such as: aluminium-silicon alloys, metal-matrix composites (MMC’s), carbon composites and fibreglass reinforced plastics. Although such workpiece materials are lightweight, they have hard, abrasive particles present to give added mechanical strength, the disadvantage of such non-metallic/metallic inclusions in the workpiece’s substrate are that

The HCl gas is a bi-product of the process and is discharged from the reactor onto a ‘scrubber’ , where it is neutralised. When titanium is to be coated onto the





Graphite shelves are most commonly employed, as it is quite inexpensive compared to either stainless steel, or nickel-based shelving, with an added benefit of good compressive strength at high temperature.

Some limitations in the CVD process are that residual tensile stresses of coatings can concentrate around sharp edges, possibly causing coatings to crack in this vicinity – if edges are not sufficiently honed – prior to coating. Additionally, the elevated temperatures cause carbon atoms to migrate (diffuse) from the substrate material and bond with the titanium. Hence, this substrate carbon deficiency – called ‘eta-phase’ is very brittle and may cause tool failure, particularly in interrupted-cut operations.

Cutting Tool Materials

. Figure 7.  A PVD-coating, with coated tooling, plus a schematic representation of the CVD and PVD coating processes. [Courtesy of Sandvik Coromant]

15

16

Chapter 1

. Figure 8.  Modern insert/tooling coating plant. [Courtesy of Walter Cutters]

Cutting Tool Materials

they become extremely difficult to machine with ‘conventional tooling’ and are a primary cause of heat generation and premature face/edge wear. Here, the high tool wear is attributable to both the abrasiveness of the hard particles present and chemical wear promoted by corrosive acids created from the extreme friction and heat generated during machining. Such diamond-coated tooling is expensive to purchase, but these coatings can greatly extend the tool life by up to 20 times, over uncoated tooling, when machining non-metallic and certain plastics, this more than compensates for the additional cost premium. Such diamond-like coated tools, combine the (almost) high hardness of natural diamond, with the strength and relative fracture toughness of carbide. The extreme hardness of diamond-like coatings enable the effective machining of non-ferrous/nonmetallic materials and, by way of an example of their respective hardness when compared to that of a PVD titanium aluminium nitride coated tool, they are three times as hard (see Fig. 3a). Although, these diamondlike coatings do not have the hardness properties of crystalline diamond, they are approximately half their micro-hardness value. Diamond-like coatings can range from 3 to 30 µm in thickness (see Fig. 9 – bottom), with the individual crystal morphology present measures between 1 to 5 µm in size (Fig. 9 – top). Recently, a diamond-coating crystal structure called ‘nanocrystalline’ has been produced by a specialised CVD process. The morphology has diamond crystals measuring between 0.01 to 0.2 µm (i.e. 10 to 200 nanometres), with a much finer grain structure and smoother surface to that of ‘conventional’ diamondlike coatings. This smoother ‘nanocystalline’ surface morphology presents less opportunity for workpiece material built-up edge (BUE) at the tool/chip interface, significantly improving both the chip-flow across the rake face of the tool and simultaneously giving a better surface finish to the machined component.

1.2.7 Tool Coatings: Physical Vapour Deposition (PVD) In 1985 the main short-comings resulting from the CVD process were overcome by the introduction of the physical vapour deposition process (Fig. 7), when the first single-layer TiN coatings were applied to cemented carbide. There are several differences between PVD and CVD coating processes and their resulting coatings. Firstly, the PVD process occurs at low-to-

17

medium temperatures (250 to 750°C), as a result of lower PVD temperatures found than by the CVD process, no eta-phase forms. Secondly, the PVD technique is a line-of-sight process, by which atoms travel from their metallic source to the substrate on a straight path. By contrast, in the CVD process, this creates an omni-directional coating process, giving a uniform thickness, but with the PVD technique the fact that a coating may be thicker on one side of a cutting insert than another, does not affect its cutting performance. Thirdly, the unwanted tensile stresses potentially present at sharp corners in the CVD coated tooling, are compressive in nature by the PVD technique. Compressive stresses retard the formation and propagation of cracks in the coating at these corner regions, allowing tooling geometry to have the pre-honing operation eliminated. Fourthly, the PVD process is a clean and pollution-free technique, unlike CVD coating methods, where waste products such as hydrochloric acid must be disposed of safely afterward. In general, there have been many differing PVD coating techniques that have been utilised in the past to coat tooling, briefly some of these are: • Reactive sputtering – being the oldest PVD coating method, it utilises a high voltage which is positioned between the tooling to be coated (anode) and say, a titanium target (cathode). This target is bombarded with an inert gas – generally argon – which frees the titanium ions, allowing them to react with the nitrogen, forming a coating of TiN on the tools. The positively-charged anode (i.e. tools) will attract the TiN to the tool’s surface – hence the coating will grow, • Reactive ion plating – relies upon say, titanium ionisation using an electron beam to meet the target, which forms a molten pool of titanium. This titanium pool then vaporises and reacts with the nitrogen and an electrical potential accelerates toward the tooling to subsequently coat it to the desired thickness. • Arc evaporation – utilises a controlled arc which vaporises say, the titanium source directly onto the inserts – from solid. As with the CVD process, all of the PVD coating production methods are undertaken in a vacuum. Further, the PVD coatings tend to have smoother and less

18

Chapter 1

. Figure 9.  A vast array of differing cutting inserts, together with diamond coated cemented carbide. [Courtesy of Sandvik Coromant]

Cutting Tool Materials

dimpled surface appearance, than are found by the ‘blocky-grained’ surface by the CVD technique. A typical tooling tungsten carbide substrate that has been PVD multi-coated is depicted in Fig. 10a. Such multiple coating technology allows for a very exotic surface metallurgy to be created, which can truly enhance tool cutting performance. In general and in the past, CVD coatings tended to be much thicker than their PVD alternatives, having a minimum coating thickness of between 6 to 9 µm, whereas PVD coatings tended to be in the range: <1 to 3 µm. Today, by employing sophisticated coating plant technology with lateral rotating arc cathodes, it is possible to have a nano-composite coating, typical of these coatings on the tooling, might be a nano-crystalline AlTiN coating embedded in an amorphous silicon nitride (Si3N4) matrix. This nanocomposite structure creates an enormously compact and resistance surface structure, not unlike that of a honeycomb. These nano-composite structures have been proven to deliver a coating hardness of between 40 to 50 gigaPascals (i.e. 1 GPa equals 100 HV) and a heat resistance of up to 1,100°C, enabling the tooling to be employed on dry, high-speed machining operations. An advantage of using a nano-composite surface structure, is that they can provide both hardness and toughness to nano-layers without the complexity and precision required to apply individual nano-layer coatings. The range and diversity of metallic and non-metallic coatings applied to tooling is simply vast and everchanging and is outside the present remit of this book. However, it is worth mentioning just one of the newlydeveloped ‘super-glide’ coatings that are currently utilised by tooling manufacturers today. These ‘superglide’ coatings have a hardness that is comparable to chalk, or talc and acts as a solid lubricant coating on the hard-coated substrate. This type of coating works really well when dry machining of: aluminium alloys, alloyed steels, nickel-based super-alloys, titanium alloys and copper. In particular, the more demanding machining operations such as small-diameter drilling and reaming, deep-hole drilling and tapping, etc, are particularly suited to such ‘soft’ coatings. A typical ‘su-



Smoother surfaces present in the PVD processes, create less thermal cracking which might lead to potential chipping and premature edge failure, while improving the resistance to repeated mechanical and thermal stresses thereby minimising interface friction, resulting in lower flank wear rates.

19

per-glide’ coating is molybdenum disulphide (MoS2) which is normally applied by the PVD modified magnetron sputtering process (see Fig. 11 for a schematic of a typical MoS2 ‘super-glide’ coating). The high-vacuum coating process is performed at a relatively low temperature (200°C). This low temperature coating process prevents the substrate from annealing, while maintaining dimensional stability. The applied MoS2 ‘super-glide’ coating has a micro-hardness of between 20 to 50 HV; it is deposited 1 µm thick, typically over a previous titanium nitride (TiN) coating, or a ‘bright’ tool. These MoS2 coatings can have over 1,200 applied molybdenum disulfide layers present, each measuring a few angströms (i.e. one angström – denoted by the symbol ‘Å’ – is equal to one 10-millionth of a mm). The atomic structure of the molybdenum disulfide coating, has a dendritic crystal structure, being similar to graphite and has weak atomic bonds between the crystal layers, allowing easy movement of the adjacent planes of the crystalline layers (Fig. 11). Such an MoS2 coating, tends to reduce the likelihood of adhesive wear and seizure, yet allowing sharp edges to the coated tooling.

1.2.8 Ceramics and Cermets The oldest cutting tool materials date back to over 100,000 BC and were ceramic (flints), as stone-aged people used these specially-prepared broken flints to cut and work into hunting tools such as arrowheads, spears and for knives when eating their hunted prey. The first modern-day industrial applications of ceramics as cutting tools occurred in the 1940’s. These early ceramic tools had the promise of retaining their hardness at elevated temperatures, while being chemically inert to the ferrous workpieces they were originally designed to machine. These advantages over the cemented carbide tools, allowed them to exploit higher cutting speeds that were now becoming available on the newly-developed machine tools of that time. These ceramic tools offered virtually negligible plastic deformation, with the cutting edge being inert to any dissolution wear. The main problem with the early ceramic tooling was that they lacked toughness and resistance to both mechanical and thermal shock (see Fig. 2b).



Dendritic derives from the Greek word for ‘tree-like’ (i.e. dendron), hence its appearance as a crystalline structure.

20

Chapter 1

. Figure 10.  Multi-coatings applied to cemented carbides and cermets, together with tool geometries of cermet cutting inserts. [Courtesy of Sandvik Coromant]

Cutting Tool Materials

21

. Figure 11.  A typical ‘super-glide’ coating of molybdenum disulfide (MoS2) applied to a hard-coating on a tool’s substrate – weak bonds between crystal layers allow easy movement of the planes. [Courtesy of Guhring]

Moreover, ceramic tools at this juncture, were only really employed for turning operations and in particular, in ‘stable machining’ , where interrupted/intermittent cutting operations did not occur. With the recent advances in powerful and very rigid CNC machine tools, this has opened-up the possibility of utilising ceramic tooling, either in a purely sintered monolithic tooling insert, or more recently as a multi-coated variant – more will be said on this topic shortly. Returning to the monolithic ceramic cutting tool materials, they have normally been available in three distinct grades, which will now be mentioned. These cutting inserts consist of: • ‘Pure’ ceramic – this is the traditional tooling insert material, consisting of aluminium oxide. The alumina is white in colour and is produced by cold pressing powder in the desired insert geometry

dies, with subsequent sintering, the fused alumina particulates are sintered together, thereby significantly decreasing porosity. These ceramics, have been known in the past as ‘pure oxide’ , or ‘coldpressed’ ceramics. The major disadvantage of such ceramics is their low thermal conductivity, making them highly susceptible to thermal shock (i.e. the hot and cold thermal cycles that can occur when interrupted cutting takes place). These thermal shock



These consolidated cutting inserts produced in compound, or ‘floating’ die sets from the admixed powders, are termed ‘green compacts’ and are friable, that is having very limited mechanical strength and must be gently handled, prior to sintering – thereafter the desired mechanical strength occurs.

22

•

•



Chapter 1

problems are exacerbated by short machining cycle times, variable depths of cut and higher machining speeds. ‘Pure’ alumina inserts can be improved by additions of zirconia (Zr) to greatly increase the toughness somewhat, but such cutting tool material, has been widely superseded today, by ‘mixed grade’ ceramics, or cermets – to shortly be discussed, Black, or mixed ceramics – tend to minimise the effects of thermal shock on the cutting insert, by having additions of titanium carbide added to the alumina, this causes the insert to turn black. A problem with these earlier ‘black ceramics’ was that they did not sinter as readily as the former ‘pure’ ceramic inserts. Therefore they usually had an additional ‘hot pressing’ operation to achieve the desired densities, which tended to limit the geometric shapes for such inserts. A later development of these cutting tool materials was termed ‘mixed ceramics’ , these had additions of titanium nitride, which improved thermal shock still further, with the sintered inserts tending to be brown, or chocolate in colour – the term ‘black’ for these later inserts, became irrelevant. These ‘mixed ceramics’ had good hot hardness, enabling them to machine harder steel components, or chilled cast irons and at greater temperatures, where the combinations of higher cutting forces and greater chip/tool interface temperatures would have induced cutting insert plastic deformation in their previous counterparts. Cermets – the original cermet was developed by Lucas under the trade name ‘Sialon’ which was a silicon nitride based material, having a very low coefficient of thermal expansion. This low expansion rate when in-cut, tends to reduce the stresses between the hotter and cooler isothermal zones of the insert, giving very high thermal shock resistance. Originally, it was difficult to sinter these inserts to full density, although by substituting some of the silicon and nitrogen with aluminium and oxygen, the new material ‘Sialon’  it had the added benefits of: ease of pressing and sintering, with equally as good thermal shock resistance. A notable later elemental addition was that of yttria (Y2O3), which aided sintering performance and during sintering. The silica (SiO2) on the surface of the silicon nitride

Sialon, this name was coined for the insert, as it represented the chemical symbols for the constituent elements: Si, Al, O and N.

particles will react with the yttria forming a liquid. This chemical reaction forms a ‘glass’ on cooling, so depending upon the relative proportions of the reactants, the resultant ‘Sialon’ formed may have either of the following atomic arrangements: beta silicon nitride, or alpha silicon nitride. It is possible to produce a very complex cutting insert material, having both ‘beta-’ and ‘alpha-Sialons’ in attendance. A typical ‘beta-Sialon’ might be composed of: Si6.ZAlZOZN8.Z Where: ‘Z’ represents the degree of substitution of silicon and nitrogen by aluminium and oxygen. Conversely, an ‘alpha-Sialon’ can consist of: Mx(Si, Al)12(O,N)16 Where: ‘M’ is the metal atom, such as yttrium. All this sounds quite confusing, but basically the ‘Sialon’ microstructure consists of a crystalline nitride phase, held in a glassy, or partially crystallised matrix. These crystalline grains can be either ‘beta-Sialon’ , or a mixture of ‘alpha’ and ‘beta’ , but generally it can be said that as the ‘alpha’ phase increases, the hardness of the ‘Sialon’ becomes greater. These chemical and mechanical changes, result in a higher ‘hot-hardness’ for the cutting insert when in-cut. An additional and probably greater benefit is gained by the significant improvement in insert toughness, which can rival that of cemented carbide of equal hardness. One limitation in the past to such cermets, was that they could not satisfactorily machine steels, owing to their poor performance in resisting solution wear. However, these earlier cermets when machining nickel-based alloys, or cast irons they performed very well, but even the ‘mixed’ ceramics based on alumina, having 25% additions of carbide (i.e. ‘whisker-reinforcement’) within the insert’s substrate are a direct competitor to such cermets.

Cutting Tool Materials

Today, with the more complex material technology cermet10 insert grades (Fig. 10b), they can easily machine ferrous-based workpieces at high cutting speeds, tool lives and excellent surface finishes. Complex powder particulates are utilised for the current turning inserts, such powders may have a large core of TiCN, surrounded by TiN – for superior hardness, adjacent particulates having a small core of niobium (Nb), surrounded by tungsten (W) and titanium (Ti) – for superior toughness. The sintered cutting insert product has a very complex substrate, which is further enhanced by subsequent multi-coating. Typical turning data for a high-performance steel product that can be rough-to-finish turned using the same insert on a 34CrMo4 grade workpiece has been shown to be: Cutting data: cutting speed (Vc) 140 m min–1, feed (f) 0.2 mm rev–1, depth of cut (DOC) 1.0 mm and with flood coolant. In interrupted cutting trials with the cutting data mentioned on this workpiece material (i.e. having 4 equally-spaced splines around its periphery), the cermet insert’s edge withstood over 7,000 impacts per edge. This can be considered as a ‘true’ testament to the hardness, shock resistance and life of the latest such cermet tooling materials.

1.2.9 Cermets – Coated To enable a wider range of machining applications while improving still further the original cermet grades available, tool coatings were introduced and with sophisticated high-technology cutting insert geometries (see Fig. 10b and c). The latest multi-coatings for indexable cutting inserts have individual ‘nanocoatings’11 and are extremely hard, approaching 4000

10 Cermet is derived from the two words ceramic and metallic and, the clear distinction between this and other cutting tool materials, such as cemented carbide and ceramic tooling has become somewhat ‘blurred’ , with one tooling manufacturer claiming it was developed in 1929, which is ‘at odds’ with the patented ‘Sialon’ product developed by the Lucas company – previously discussed. 11 One nanometre is equal to 10–9 m, or one millionth of a mm.

23

HV12 and the surfaces of such coatings tend to be very smooth and having a total thickness of less than 3 µm thick – allowing around 2,000 durable layers. One of the key factors in successfully applying these complex metallurgy multiple coatings, has been the development of ‘super-lattice technologies’ at medium temperatures, which do not compromise the thermal properties of the substrate. The unit cost of the cermet substrate tends to be lower than its equivalent cemented carbide grade, this accounts for the fact that at present, in turning operations in Japan 35% of all the inserts utilised for a range machining steel grades tend to be cermets, whereas, in Europe less than 5% of cermets are employed. Cermets are considerably more wear and heat resistant than tungsten carbide-based cutting materials. By way of illustration for the reason for edge failure of tungsten carbide inserts, is the heat generated at the tool/chip interface – at high cutting speeds. For example, if one considers the pre-sintering temperature for a typical tungsten carbide material it is in the region of 1,150°C and, if turning a: 0.48% C, 0.8% Mn medium carbon steel workpiece at 200 m min–1, this equates to the highest isothermal edge temperature of 1,000°C – creating the potential for localised thermal softening and edge failure. While an equivalent multi-coated Cermet, can readily turn alloy carbon steels at a depth of cut (DOC ) of up to 3 mm, with cutting speeds of between 200 to 300 m min–1, with feedrates ranging from 0.1 to 0.3 mm rev–1. Moreover, as less flank wear takes place, the dimensional size of subsequent components in a batch will not significantly ‘statistically-drift’ , producing much less tolerance variation (i.e. reliable size-forsize consistency) in the completed turned parts. This increased multi-coated cermet tool life, allows for an excellent surface finish and dimensional consistency, whether cut wet, or dry. In general, the multi-coated cermet cutting tool materials, can be consolidated (i.e. pressed) in compound die-sets with very complex tool geometries and have integrated chip-breakers present – as illustrated in Fig. 10c. Such inserts, have seen a slow take-up in Europe and offer considerable economical advantages when in particular, turning hardened steel parts.

12 A comparison of the hardness of different popular coatings may be applicable here, as TiCN coating has a hardness of around 2,700 HV and TiAlN coating has a hardness of approximately 3,200 HV.

24

Chapter 1

. Figure 12.  Ultra-hard cutting tool materials – cubic boron nitride (CBN). [Courtesy of DeBeers – element 6]

Cutting Tool Materials

25

. Table 1.  Cutting tool materials – with some important physical properties Black ceramic (Al2O3 + TiC)

Cemented carbide (ISO K10 grade)

CBN (DBC50)

CBN (DBC80)

Density [g cm–3]



4.28



14.7



4.28



3.52

Knoop hardness [GPa]



17



17



27.5



30

Young’s modulus [GPa]



390



593



587



649

Fracture toughness [MPam½]



2.94



10.48



3.7



5.90

Thermal expansion [10–6K–1]



7.8



5.4



4.7



4.6

Thermal conductivity [Wm–1K–1]



9.0



100



44



85

Cutting tool material: Physical properties:

1.2.10 Cubic Boron Nitride (CBN) and Poly-crystalline Diamond (PCD) Cubic Boron Nitride (CBN)/Synthetic Diamond – Ex traction and Sintering Cubic boron nitride (CBN) is one of the hardest materials available and for machining operations it can be considered as a ultra-hard cutting tool, it was first synthesised in the late 1950’s. In many ways, CBN and natural diamond are very similar materials, as they both share the same atomic cubic crystallographic structure (see Fig. 12a and b). Both materials exhibit a high thermal conductivity, although they have profoundly different properties. For example, diamond is prone to graphitisation and will readily oxidise in air, reacting to ferrous workpieces at high temperatures, conversely, CBN is stable to higher temperatures and can effortlessly machine ferrous components. CBN can therefore machine ferrous materials, such as: tool steels, hard white irons, surface hardened steels, grey cast irons, (some) austempered ductile irons and hardfacing alloys. Normally, CBN tools should be used on workpiece materials with hardnesses greater than 48 HRC, because if workpieces are less hard than this, the cutting edge will result in excessive tool wear. In graphite, the carbon atoms are arranged in a hexagonal layered structure (Fig. 12ai) and, by the appli-

cation of very high temperatures and pressures13, it can be transformed into the cubic structure of diamond (Fig. 12aii) – this transformation does not occur easily. As boron and nitrogen are two elements on either side of carbon in the Periodic Table, it is possible to form a compound of boron nitride, that exhibit’s a hexagonal boron nitride (HBN) as depicted in Fig. 12bi, having the characteristics of being both slippery and friable. HBN can be transformed in a similar fashion to that of CBN (Fig. 12 bii). In practice, to facilitate the rate of transformation in the reaction chamber, additions of solvents/catalysts are utilised for synthesis at more easily obtainable levels: pressures of approximately 60 GPa and temperatures 1,500°C. As this transformation proceeds in the reaction volume of a high pressure system, the CBN/synthetic diamond grows, being embedded in a portion of reaction mass and extracted afterward from this special-purpose press. By dissolving away the unwanted matrix, the CBN/synthetic diamond can be liberated and recovered for subsequent processing. Grain sizes vary from large dimensions of approximately 8 µm, down to sub-micron sizes, for fine-grain tooling. Once the synthesised CBN/diamond has been extracted, it is possible to sinter together these crystals of CBN, or diamond, with the aid of a ceramic binder, to produce polycrystalline masses. Commercially, in

13 To transform hexagonal graphite into the cubic diamond structure, requires exceedingly high temperatures > 2000°C and applied pressures > 60 GPa, to enable the conversion to take place.

26

Chapter 1

order to speed-up the rate of sintering, additions of a solvent/catalyst are utilised (i.e. normally metals, or metal nitrides), but during sintering the whole mass must be held in the ‘cubic region’ of the respective pressure/phase diagram – to prevent these hard crystals reverting back to their original soft hexagonal form. By sintering these hard particles together, it is possible to form a conglomerate of CBN/diamond, in which randomly orientated crystals are combined to produce a large isotropic14 mass. A very wide range of polycrystalline products can be produced, utilising either CBN, or synthetic diamond as a base. For example, by changing the: grain size (see Figs. 12 c and d), solvent/ catalyst employed, degree of sintering and particle size distribution and, the presence/absence of inert fillers, this will have a profound effect on the mechanical and physical properties of the final product – Table 1 lists

14 Isotropic materials can be considered to have the same properties in different directions.

the physical properties of various comparable cutting tool materials. In order to produce the required tool geometries, both the polycrystalline layer of CBN and polycrystalline diamond (PCD) are bonded to a thick tungsten carbide backing layer, then cutting inserts are wirecut out of this large blank – obtaining the maximum number of insert shapes per blank (see Fig. 13a). These CBN/PCD inserts are either full-size, or smaller tips that are then brazed onto suitably-shaped blanks, to fit the desired tool holder (as illustrated in Fig. 13b). Both CBN and PCD cutting tools can successfully machine: super-alloys (ie with low iron content), grey cast iron and non-ferrous metals, but show distinct differences when other workpiece materials are to be productively machined – as depicted in Fig. 14. Polycrystalline diamond cutting tools are not utilised for machining ferrous workpieces, this is because when machining under the high temperatures and sustained pressures that occur during cutting, the diamond has a tendency to revert back to graphite, after only a few seconds in-cut. This reversion, does

. Figure 13.  Cutting tool materials: Cubic Boron Nitride (CBN) and Polycrystalline Diamond (PCD) . [Courtesy of DeBeers – element 6]

Cutting Tool Materials

. Figure 14.  A diagram illustrating how Cubin Boron Nitride (CBN) and Polycrystalline Diamond (PCD) applications are grouped, by workpiece materials. Their effectiveness when ei-

27

ther machining highly abrasive components, or high temperatures in the cutting vicinity

28

Chapter 1

. Figure 15.  Turning operations with Cubic Boron Nitride (CBN) and Polycrystalline Diamond (PCD). [Courtesy of DeBeers – element 6]

Cutting Tool Materials

29

not take place when machining many non- ferrous and non-metallic workpiece materials. Although CBN is synthesised in a similar fashion to that of PCD cutting tool products, it is not as hard as PCD and is therefore less reactive with ferrous metals, as long as the cutting temperature is less than 1,000°C, it will not revert to its softer hexagonal form and oxidise in air. This means that CBN can machine many ferrous parts and cast iron grades. The complementary nature of both CBN and PCD is clearly depicted in Fig. 14, where the ‘cross-over’ between these ultra-hard cutting tool materials is shown. In both CBN and PCD machining applications, an excellent machined surface finish can be obtained (see Figs. 15a and b). In the case of many PCD operations, the cutting tool must not only machine widely differing materials that are situated adjacent to one another in many passes over such a diverse material workpiece, but produce an excellent machined surface finish, which really ‘challenges’ the tool. Tool life can be extended greatly by utilising either CBN, or PCD tooling, often tool lives can be increased by 50 to 200 times that of the previous cemented carbide alternatives. This boost in output, makes their additional purchase price irrelevant, when considered against the massive productive gains that are to be made by their adoption. Today, both CBN and PCD can often be found as either thin-coated layers on tooling (see Fig. 3 for their relative tool insert hardnesses/toughnesses), or as a ‘sandwich’ between metallic backing layers. These ‘sandwiched’ tool edges, permit brazing on both sides of the hardened product, which are then accurately positioned and held onto a tungsten carbide shank, making them an ideal alternative for many microdrilling operations. Such compound drilling edge technology, gives considerably improved edge retention and resistance to any abrasive particles present in the workpiece and its severely work-hardened swarf, typically found with the latest metal matrix composites (MMC’s). Such ultra-hard tooling, can be readily used on high-silicon aluminium alloys used in the automotive industries, while not discounting the wide range of workpiece composites employed by the aerospace industries and the resin-based components utilised in the furniture industry.

mond is used correctly in a very rigid machine-toolworkpiece setup for materials that require the best possible surface finish, then there is simply no alternative. By way of illustration of this fact, if production turning high-silicon content aluminium pistons with polycrystalline diamond (PCD) tooling, the best surface finish that can be obtained will be in the region of 0.4 µm, conversely using an SCD tool this will give a surface finish of better than 0.15 µm. If one really wants the ultimate surface finish currently obtainable by machining – in the ‘nano-range’ , then a monolithic diamond tool, mounted in a special-purpose diamond turning lathe is the only manner in achieving such superb ‘mirror-finish’ surfaces. SCD tool edges are produced as either razor sharp edges, or are made with a perfect radius being chip-free, imparting machined ‘mirror-finishes’ of just a few angströms (i.e. 10–10 m). The optical industries in particular find that the latest blemish-free ultra-sharp cutting edges of SCD, means that diamond (paste) polishing after machining has been virtually, if not completely eliminated, this fact in particular being a very big production cost for the final manufacture of large monolithic astronomical mirrors. A cautionary note, is that to use SCD tooling for anything other than as a finishing cut is totally uneconomic, as these precision components to be machined, should have been roughly configured to the desired shape, prior to diamond machining. Therefore, SCD tools should be employed for exceedingly light finish cuts of no deeper than 0.0008 m. Natural diamond is a truly remarkable material, that exhibit’s a diverse range of mechanical and physical properties. For example diamond has the highest known: bulk hardness, thermal conductivity, while having a very low coefficient of friction and will not corrode, these properties make it an ideal tool material for the highest precision and accuracy machined components. Of these properties, hardness is probably the most important characteristic in machining operations and, when measured by the Knoop indentor15. By way of comparison of ultra-hard cutting tool materials, the following two examples may prove informative:

1.2.11 Natural Diamond

15 Knoop indentors produce a wedge-shaped indentation in the form of a parallelogram, with one diagonal seven times longer than the adjacent one. The Knoop test method is generally considered the optimum technique, for crystalline solids – having crystallographic directionality (i.e. anisotropy).

Monolithic, or single-crystal diamond (SCD), is the hardest material available today. If such natural dia-

30

Chapter 1

• Natural diamond – has a hardness of 9,000 kg mm–2

(ie diamond orientation and test conditions): Diamond (111) surface, <110> direction, 500g load, • Cubic boron nitride (CBN) – has a hardness of 4,500 kg mm–2, (111) surface, <110> direction, 500g load. One of the main limitations of natural diamond is that it has distinct cleavage planes (111)16. This consistent cleavage plane makes it ideal for jewellery-makers to cleave the beautiful facets demanded of diamond jewellery, but this means that monolithic diamonds must be mounted in their respective tool holders in exactly the correct orientation/plane, so avoiding any potential cleavage in-cut. SCD tool cost is a draw-back, because these tools cost in the region of four times more than the equivalent PCD tool. However, despite this very high cost difference, SCD can reduce the overall operating costs and significantly improve productivity, when applied to the correct machining process. Expensive tooling such as SCD, must be handled with care, because although it is the hardest material known, it is also very brittle and subject to thermal shock, the problem being exacerbated with its very sharp tool edges. Therefore, it is essential that sudden impacts to the tool’s edge must be avoided, through either inappropriate cutting applications, or by rough handling.

References Journal and Conference Papers Boller, R. Crystal Clear – DLCoatings. Cutting Tool Engg., 36–40, May 2002. Craig, P. Behind the Carbide Curtain. Cutting Tool Engg., 26–41, Aug., 1997. Dzierwa, R. Slippery when Blue – Coatings. Cutting Tool Engg., 36–41, Jan., 2003. Eastman, M. Inserts Show their True Colors. Cutting Tool Engg., 30–36, April 1999. Feir, M. Post-treatment of PM Parts. Metal Powder Report, 28–30, Jan., 1981.

16 Miller indices determine the crystalline orientation for a plane in an atomic structure and for natural diamond it is normally on the (111) plane, although some cleavage has been observed on the (110) plane.

Fretty, P. Grade Wise. Cutting Tool Engg., 46–50, Feb., 2000. Gough, P. Tool Life Boosted by Titanium Nitride Coat. Machinery and Prod. Engg., 52–53, Feb. 1983. Gummeson, P.U. and Stosuy, A. Iron-carbon Behaviour during Sintering. In: Source Book on Powder Metallurgy, ASM Pub., 49–61, 1979. Hanson, K. Lowering your Grades. Cutting Tool Engg., 54–60, Jan., 2000. Heath, P.J. Ultra-hard Materials. European J. of Engg. Ed., Vol. 12 (1), 5–20, 1987. Israelsson, J. A Progress Report on Cutting Tool Materials. American Machinist, 39–40, Dec., 1992. Jindal, P.C. et al., PVD Coatings for Turning, Cutting Tool Engg., 42–52, Feb., 1999. Kennedy, B. Making the Grade – PCBN Applications. Cutting Tool Engg., 22–30, June 2002. Lewis, B. Fast Times in HSS. Cutting Tool Engg., 28–32, July 2001. Lewis, B. Confidence Game – Grades and Geometries. Cutting Tool Engg., 46–52, Dec., 2002. Mielert, W. Coating for Speed. Cutting Tool Engg., 40–44, Feb., 1996. Mirchandani, P.K. Making a Better Grade – Composite Carbide Substrates. Cutting Tool Engg., 58–61, Jan., 2005. Mitoraj, L. The Coating Edge. Cutting Tool Engg., 51–55, Feb., 2000. Novak, D. Single Minded – Single Crystal Diamond. Cutting Tool Engg., 38–41, June 2002. Raymond, M.K. Ceramics Ease Up the Machining of Highhardness Parts, American Machinist, May 1996. Raymond, M.K. Coatings Keep Cutting Tools Sharp. American Machinist, 40–42, May 1996. Richter, A. Raising Al – AlTiN Coatings. Cutting Tool Engg., 42–46, Jan., 2003. Richter, A. Top Coat. Cutting Tool Engg., 36–41, Dec., 2003. Sanders, E.H. Understanding Coated Carbides. Cutting Tool Engg. 3–7, Sept./Oct., 1977. Sprout, W. PVD Today. Cutting Tool Engg., 52–4057, Feb., 1994. Taylor, F.W. On the Art of Cutting Metals. Trans. of ASME 28, 31–350, 1907. Thalmann, R. Cracking the Code – Carbide Classifications. Cutting Tool Engg., 34–43, June 1995. Vasilash, G.S. The Superfard Coatings: More than Meets the Eye. 52–54, Production, Dec., 1995. Weiner, M. Coatings Move Forward. Cutting Tool Engg., 22–29, Feb., 1999. Woods, S. Coat, Please. Cutting Tool Engg., 50–56, Oct., 2004.

Cutting Tool Materials

Books, Booklets and Guides Balshin, M.Y. and Kiparisov, S.S. General Principles of Powder Metallurgy. MIR (Moscow) Pub., 1980. Chin, G.Y. Advances in Powder Metallurgy. ASM Pub., 1982. Dieter, G.E. Mechanical Metallurgy. McGraw-Hill Kogakusha 2nd Ed., 1976. Dowson, G. Powder Metallurgy – The Processes and its Products. Adam Hilger (Bristol) Pub., 1990. Kalpakjian, S. Manufacturing Processes for Engineering Materials (3rd Ed.). Addison Wesley, 1997.

31

Lenel, F.V. Powder Metallurgy, Principles and Applications. Metal Powder Inds. Princetown (NJ), April 1980. Metals Handbook – Powder Metallurgy. ASM Pub. 8th Ed., Vol. 3, 1967. Modern Metal Cutting – A Practical Handbook. AB Sandvik Coromant Pub., 1994. Reed-Hill, R.E. Physical Metallurgy Principles. Van Nostrand Reinhold (NY) Pub., 1973. Trent, E.M. Metal Cutting. Oxford: Butterworth Heinemann (3rd Ed.), 1991.



2

Turning and Chip-breaking Technology ‘Machines are the produce of the mind of Man; and their existence distinguishes the civilized man from the savage.’

WILLIAM COBBETT

(1762–1835) [Letter to the Luddites of Nottingham]

34

Chapter 2

2.1 Cutting Tool Technology In the following sections a review of a range of Turning-related technologies and the importance of chipbreaking technology will be discussed.

2.1.1 Turning – Basic Operations Turning can be broken-down into a number of basic cutting operations and in effect, there are basically four such operations, these are: 1. Longitudinal turning (Fig. 16a), 2. Facing (Fig. 16b), 3. Taper turning – not shown, 4. Profiling – not shown. NB  These turning operations will now be very briefly reviewed. In its most simple form, turning generates cylindrical forms using a single-point tool (Fig. 1.16a). Here, a tool is fed along the Z-axis slideway of the lathe (CNC), or a turning centre, while the headstock rotates the workpiece (i.e. the part is held in either: a chuck, on a mandrel, face-plate, or between centres – when overhang is too long), machining the component and thereby generating a circular and cylindrical form of consistent diameter to the turned part. Facing is another basic machining operation that is undertaken (Fig. 16b) and in this case, the tool is fed across the X-axis slideway while the part rotates, again, generating a flat face to the part, or a sharp corner at a shoulder, alternatively it can be cutting the partial, or finished part to length (i.e facing-off). Taper turning can be utilised to produce short, or long tapers having either a fast taper (i.e. with a large included angle), or slow taper (i.e. having a small included angle – often a ‘self-holding taper’ , such as a Morse taper). There are many different operations that can be achieved on a CNC lathe/turning centre,



The range of turning operations is vast, feedrates can be varied, as can rotational speeds.



Facing operations can also be used to produce either curved convex, or concave surface features to the machined part – here the surface is both generated and formed, requiring simultaneous programmed feeding motions to the Z- and Xaxes.

including: forming, while others such as drilling, boring, screw-cutting, of internal features, and forming and screw-cutting of external features, to name just a few of the traditional operations undertaken. With the advent of mill/turn centres, by having CNC control of the headstock and rotational, or ‘driven-/live-tooling’ to the machine’s turret, this allows prismatic features to be produced (i.e. flats, slots, splines, keyways, etc.), as well as drilled and tapped holes across and at angles to the major axis of the workpiece, or off-axis. Even this explanation of mill/turn centres is far from complete, with regard to today’s sophisticated machine tools. As machine tool builders today, can offer a vast array of machine configurations, including: co-axial spindles (ie twin synchronised inline headstocks), fitted with twin turrets with X- and Y-axes simultaneous, but separate control, having programmable steadies (i.e. for supporting long slender workpieces), plus part-catchers , or overhead gantries for either component load/unload capacity, to multiaxes robots feeding the machine tool. This type of machine tool exists and has multi-axes CNC controllers to enable the machine’s down-time to be drastically reduced and in this manner achieving high productive output virtually continuously.

2.1.2 Turning – Rake and Clearance Angles on Single-point Tools In order for a turning tool to effectively cut and produce satisfactory chips, it must have both a rake and clearance angle to the tool point (Fig. 17). Today’s single-point cutting tools and inserts are based upon decades of: past experience, research and development, looking into all aspects of the tool’s micro-geometry at the cutting edge. Other important aspects are an efficient chip-breaking technology, in certain instances critical control of the flexure (i.e. elastic behaviour) of the actual tool insert/toolholder combination for the latest multi-functional tooling is essential – more will be said on some of these topics later in the chapter. The rake angle is the inclination of the top face of the cutting edge and can vary according to the work-



Forming can be achieved in a number of ways, ranging from complex free-form features (externally/internally) on the machined part, to simply plunging a form tool to the required depth.

Turning and Chip-breaking Technology

35

. Figure 16.  Typical turning operations with the workpiece orientation shown in relation to the cutting insert, for either: (a) cylindrical turning, (b) facing. [Source: Boothroyd 1975]

36

Chapter 2

piece material being machined. In general, for ductile materials, the rake inclination is a positive angle, as the shearing characteristics of these materials tends to be low, so a weaker wedge angle (i.e. the angle between the top face and the clearance angle) will suffice. For less ductile, or brittle workpiece materials, the top rake inclination will tend toward neutral geometry, whereas for high-strength materials the inclination will be negative (see Fig. 17), thereby increasing the wedge angle and creating a stronger cutting edge. This stronger cutting edge has the disadvantage of requiring greater power consumption and needing a robust tool-workpiece set-up. Machining high-strength materials requires considerable power to separate the chip from the workpiece, with a direct relationship existing between the power required for the cutting operation and the cutting forces involved. Cutting forces can be calculated theoretically, or measured with a dynamometer – more will be said on this subject later in the text. Both side and front clearances are provided to the cutting edge, to ensure that it does not rub on the workpiece surface (see Fig. 17). If the tool’s clearance is too large it will weaken the wedge angle of the tool, whereas if too small, it will tend to rub on the machined surface. Most tools, or inserts have a nose radius incorporated between the major and minor cutting edges to create strength here, while reducing the height of machined cusps, with some inserts having a ‘wiper’ designed-in to improve the machined surface finish still further – more will be mentioned on these insert integrated features later.

2.1.3 Cutting Insert Edge Preparations Often, a minute edge preparation (see Figs. 17 and 18b, c and d) is created onto the sharp cutting edge of the insert, this imparts additional strength to the outermost corners of the cutting edge, where the rake and clearance faces coincide. There are four basic manners in which the honed edge preparation is fashioned, these are:



Machined cusps result from a combination of the feedrate and the nosed radius of the tool. If a large feedrate occurs with a small nose radius then the resultant cusp height will be high and well-defined, conversely, if a small feedrate is utilised in conjunction with a large nose radius, then cusp height is minimised, hence the surface texture is improved.

1. Chamfer – which simply breaks the corner – not illustrated, 2. Land – stretching back negatively from the clearance side to various lengths on the rake face (see Fig. 18b), 3. Radius – around the actual corner (see Fig. 18c), 4. Parabolic – has unequal levels of honing on two faces (see Fig. 18d). Even here, more often than not, certain combinations of these four edge preparations are utilised, so that the cutting forces are redirected onto the body of the rake’s face, rather than directed down against the more fragile cross-section of the edge. The T-lands and hones are often actually incorporated into the insert geometry of the contoured surface. Typical T-lands range in size from 0.07 to 0.50 mm, having angles varying from 5 to 25° off of the rake face (Fig. 18b). Honing which is the ‘rounding’ of the cutting edge, can be performed in one of several ways. Probably the oldest technique for honing, utilises mechanical means, which employs a vibrating tub filled with an abrasive media, such as aluminium oxide – to ‘break’ the corner on these inserts. A variation in this design, uses an identical abrasive, except here the inserts are held by centrifugal force to the inside of a rotating tank. While yet another method of honing using an abrasive media, involves spraying the inserts with fine abrasive particles – to hone the edges of the inserts. Probably the most popular method for obtaining cutting insert honed edges, uses brushes made from extruded nylon impregnated with diamond (see Fig. 18a). The inserts to be honed pass by these brushes in individual carriers and rotate as they all revolve under the brushes, thereby applying equal hones to all insert edges. Depending upon the amount of desired honing, these brushes can be either raised, or lowered, or alternatively, the inserts can make multiple passes through the machine. All of the above honing techniques produce a hone that is roughly equal on both the flank and rake faces – what is termed a ‘round hone’ (Fig. 18c). Yet another honing profile termed the parabolic hone (i.e. sometimes this honed edge is known as:



The radius is sometimes termed ‘edge rounding’ (i.e. denoted by the letters ‘ER’) – often applied to most edge preparations, enabling the cutting forces to be directed on to the stronger part of the insert.

Turning and Chip-breaking Technology

37

. Figure 17.  Typical turning ‘finishing’ insert/toolholder geometry and the insert’s edge chamfering, in relation to the workpiece

38

Chapter 2

P-hone, oval, or waterfall), is produced by a machine with a soft, diamond-charged rotating rubber wheel. Therefore, as the abrasive material rubs across the inserts, it tends to extend slightly over the inserts sides, producing a hone of uneven proportions between the two insert faces (Fig. 18d). As in the case of the T-land

cutting insert edge preparation, the P-hone directs the cutting forces into the body of the insert. Honing can be specified in a number of sizes, usually being determined by the amount of time these insert spend in the honing device. The original Standard for honing was established in the United States by

. Figure 18.  A honing machine (i.e. brush-style) and several types of honing edge preparations. [Courtesy of Ingersoll]

Turning and Chip-breaking Technology

the American National Standards Institute (ANSI) in 1981, which included dimensions and expected tolerances for these three basic hones. Today, many cutting tool manufacturers have expanded upon this Standard, or adopted their own – specifying hone manufacturing and identification methods. Hones must be applied prior to the application of coatings. Inserts that are destined to receive a CVD coating, must have a minimum hone to strengthen the edge, in order to counteract the effects of this high temperature coating process. Conversely, PVD coatings, can be equally applied either over fully-honed insert edges, or on an unhoned cutting edge. In recent years, the cutting tool manufacturers have an emphasis toward providing honed edges of greater consistency and repeatability.

2.1.4 Tool Forces – Orthogonal and Oblique The cutting forces are largely the result of chip separation, its removal and chip-breaking actions, with the immense pressure and friction in this process producing forces acting in various directions. Stresses at the rake face tend to be mainly compressive in nature, although some shear stress will be present (see Table 2, by way of illustration of the machining shear stresses for various materials), this is due to the fact that the rake is rarely ‘normal’ to the main cutting direction. This compressive stress tends to be at its greatest closest to the cutting edge, with the area of contact between the chip and rake face being directly related to the geometry here, hence the need for tooling manufacturers to optimise the geometry in this region. There are two distinct types of forces present in machining operations concerning single-point cutting tools/inserts (see Fig. 19), these are: 1. Orthogonal cutting forces – two forces (ie tangential and axial – see Fig. 19b), 2. Oblique cutting forces – three forces (i.e. tangential, axial and radial – see Fig. 19a).



As well as the tool/chip interface temperatures being up to 1,000°C, the interface pressures can reach a maximum of 3,000 MPa, these being sterile smooth surfaces makes them ‘ideal’ conditions for the occurrence of ‘pressure-welding’/seizure.

39

. Table 2.  Typical in-cut shear strengths of various materials Material:

Shear yield strength in cutting –2 (N mm )

Iron

370

0.13% C. steel

480

Ni-Cr-V steel

690

Austenitic stainless steel

630

Nickel

420

Copper (annealed)

250

Copper (cold-worked)

270

Cartridge brass (70/30)

370

Aluminium (99.9% pure) Magnesium Lead

97 125 36

[Source: Trent ( 1984)]

NB  Both of these cutting force models are heavily influenced by the: cutting tool/insert orientation to workpiece, tool’s direction of cut and its applied feedrate.

Oblique Cutting Forces Fig.1.19a, can be seen a model of the three-dimensional cutting force components in an oblique turning operation, when the principal cutting edge is at an angle to the main workpiece axis (i.e. Z-axis). These component forces can be separated into the: • Tangential force (FT) – which is greatly influenced by the contact and friction between both the workpiece and tool, as well as the contact conditions between the chip and the rake face of the cutting edge. The magnitude of the tangential cutting force is the greatest of these three component forces and contributes to the torque, which in turn, influences



Feedrates play a major role in determining the axial force in single-point cutting operations, in association with the tool’s orientation to the part being machined.

40

Chapter 2

. Figure 19.  The two- and three-force models of orthogonal and oblique cutting actions, with the component forces approximately scaled to give an indication of their respective magnitudes

Turning and Chip-breaking Technology

•

•

the power requirement for cutting. Fundamentally, the product of the tangential force and the cutting speed represent the power required for machining. The specific cutting force is a unit expression for the tangential cutting force, being closely related to the material’s undeformed chip thickness and selected feedrate, Axial force (FA) – the magnitude of this force will vary depending on the selected feedrate and the chosen tool geometry and in particular, the ‘plan approach angle’ , or ‘entering angle’ , – more will be said on this topic later. Its direction is from the feeding of the tool, along the direction of workpiece machining, Radial force (FR) – is directed at right angles to the tangential force from the cutting point. The ‘plan approach angle’ and the size of the nose radius, will influence this force. NB  These three component forces are significantly influenced by the rake angle, with positive rakes producing in general, lower cutting forces. The resultant force, its magnitude and angle, will be affected by all three component forces, in conjunction with the tool’s geometry and the workpiece material to be cut.

Orthogonal Cutting Forces In Fig. 19b the two-dimensional model for orthogonal cutting is depicted, once again, for comparison to the oblique cutting model, in a single-point turning operation. For simplicity, if one assumes that the point of the tool is infinitely sharp and that the tool is at right angles to the workpiece axis having no deflection present, then the two component forces are the tangential force and axial force (i.e. previously mentioned above). In this case, this tool geometry-workpiece configuration, allows long slender bars to be turned, as there is less likelihood of tool ‘push-off ’ (i.e. as the radial force



In reality, the specific cutting force is a better indication of the power requirement, as it is the force needed to actually deform the material prior to any chip formation. It will vary and is influenced by the: undeformed chip thickness, feedrate, and yield strength of the workpiece material. For example, if the cutting conditions are kept the same and only the material changed, then if a nickel-based alloy is machined, the initial chip forming force (i.e. specific cutting force) will be more than ten times greater than when cutting a pure aluminium workpiece.

41

has been neutralised – as indicated by the fact that the resultant force shows no X-axis offset). If any radial force was present, this would create either a ‘candlestick effect’ , or ‘barrelling’ to the overall turned length. In reality, there will always be some form of nose radius, or chamfer to the tool point, which will have some degree of ‘push-off ’ , depending upon the size of this incorporated nose feature – creating a ‘certain degree’ of radial component force affect.

2.1.5 Plan Approach Angles The manner in which the cutting edge contacts the workpiece is termed the ‘plan approach angle’ (Fig. 20a), being composed of the entering and lead angles for the selected tool geometry. In effect for singlepoint turning operations, the tool’s orientation of its plan approach, is the angle between the cutting edge and feeding direction. When selecting a tool geometry for turning specific workpiece feature – such as a 90° shoulder – it is important as it will not only affect the machined part geometry, but has an influence on consequent chip formation and the direction and magnitude of the component cutting forces, together with the length of engagement of the cutting edge (see Fig. 20b). In single-point turning (Fig. 20b), the depth of cut (DOC), or ‘cutting depth’ is the difference between an un-cut and cut surface, this being half the difference in the un-cut and cut diameter (i.e. the diameter is reduced by twice the DOC in one pass along the workpiece). This DOC is always measured at 90° to the tool’s feed direction, not the cutting edge. The manner in which the cutting edge approaches the workpiece is termed the ‘entering angle’ (i.e. plan approach angle), this being the angle between the cutting edge and feed direction (Fig. 20a – shown here in a cylindrical turning operation). Moreover, the plan approach angle not only influences the workpiece features that can be produced with this cutting geometry, it also affects the formation of chips and the magnitude of the component forces (Fig. 20b). The ‘entering angle’ affects the length of the cutting edge engaged in-cut, normally varying from 45° to 90°, as illustrated in the four cases of differing plan approach angles shown in Fig. 20b. Here, in ‘case I’ an



In single-point turning operations, the depth of cut (DOC) is sometimes referred to by the term: ‘undeformed chip thickness’.

42

Chapter 2

. Figure 20.  Insert approach angle geometry for turning operations

Turning and Chip-breaking Technology

entering angle of 45° and lead angle of 45° is utilised, giving rise to equal axial and radial component forces. In ‘case II’ , the entering angle has changed to 75° and lead angle is now 15°, these altered angles change the component forces, with an increase in the axial force while reducing the radial force. In ‘case III’ , an orthogonal cutting action occurs, with only a 90° entering angle (i.e. the lead angle reduces to zero), showing a large increase in the axial force component at the expense of the radial force component which is now zero10. In ‘case IV’ , an oblique cutting action has returned (i.e. as in ‘cases I and II’), but here the entering angle has changed to -15°, with the lead angle 75°, this produces a large axial component force, but the radial component force direction has now reversed. This last tool plan approach angle geometry (i.e. ‘case IV’), is similar to the geometry of a light turning and facing tool, allowing cylindrical and facing operations to be usefully undertaken – but the tool’s point is somewhat weaker that the others, with the tool points becoming of increased strength from right to left. Therefore, in ‘case I’ , for a given feedrate and constant DOC, the cut length/area is greater than the other ‘cases’ shown and with this geometry, it enables the tool to be employed for heavy roughing cuts. Returning to ‘case III’ , if this tool is utilised for finish turning brittle-based workpiece materials, then upon approaching the exit from a cut, if the diameter is not supported by a larger shoulder diameter, then the axial component force /pressure, will be likely to cause edge break-out (i.e. sometimes termed ‘edge frittering’), below the machined surface diameter at this corner (i.e. potentially scrapping the machined part). In mitigation for this orthogonal cutting tool geometry, if longer slender workpieces require cylindrical turning along their length, then with the radial force component equating to zero, it does not create significant ‘push-off ’ and allows the part to be successfully machined11. A single-point turning geometry is subject to very complex interactions and, as one geometric feature is modified such as changing the entering angle, or in-

10 In all of these cases, it is assumed – for simplicity – that there is no nose radius/chamfer on the tool and it is infinitely sharp. 11 In order to minimise the effects of the radial force component when cylindrically turning long slender workpieces with ‘Case I and II’ tool geometries, the use of a programmable steady, or a ‘balanced turning operation’ (i.e. utilising twin separately programmable turrets on a turning centre, with tools situated virtually opposite each other running parallel during the turning operation – see Fig. 41), will reduce this ‘push-off ’.

43

creasing the tool’s nose radius, this will influence other factors, which in turn could have a great impact on the: type of machined surface finish produced, expected tool life and the overall power consumption during the operation. In fact, the main factors that influence the application of tooling for a specific turning operation are: Workpiece material – machinability, condition I. (i.e. internal/external), mechanical and physical properties, etc., II. Workpiece design – shape, dimensions and machining allowance, III. Limitations – accuracy and precision requirements, surface texture/integrity, etc., IV. Machine tool – type, power, its condition and specifications, V. Stability – loop stiffness/rigidity (i.e. from the cutting edge to its foundations), VI. Set-up – tool accessibility, workpiece clamping and toolholding, tool changing, VII. Tool programme – the correct/specified tool and its tool offsets, etc., VIII. Performance – cutting data, anticipated tool-life and economics, IX. Quality – tool delivery system and service. In order to gain an insight into the complex and important decisions that have to be made when selecting tooling for the optimum production of either part batch sizes, or for continuous production runs, then the following section has been incorporated.

2.1.6 Cutting Toolholder/Insert Selection When deciding upon the correct selection of a toolholder/cutting insert for a given application, a range of diverse factors must be considered, as indicated in Fig. 21. As can be seen by the diagram (Fig. 21) and associated text and captions, there are many other variables that need to be considered prior to selection of the optimum toolholder/insert. Generally, the fixed conditions cannot be modified, but by ‘juggling’ with the variable conditions it is possible to accomplish the best compromise toolholder/insert geometry, to optimise these cutting conditions for the manufacture of a specific workpiece and its intended production requirements. Whenever toolholders and cutting inserts are required for a specific manufacturing process, it is important to view the tooling selection procedure as a logical progression, in order to optimise the best

44

Chapter 2

. Figure 21.  The factors that must be considered prior to commencing a turning operation, when utilising indexable inserts

Turning and Chip-breaking Technology

possible tools/inserts for the job in hand. Perhaps the following selection strategy for a ‘start point’ in choice and application of turning tools, can be undertaken according by the following step-by-step approach: Start Point  →

Edge clamping system, ↓ Toolholder size and type, ↓ Insert shape, ↓ Insert size, ↓ Nose radius, ↓ Insert type, ↓ Tool material, ↓ Cutting data  →

45

part access, as a toolholder is defined by its: effective entering and point angles12, together with the insert’s shape (see Fig. 22). Toolholders should be the largest possible size for the turning centre’s tool turret, this requirement is vital, as it reduces the ‘tool overhang ratio’ – providing rigidity and integrity to stabilise the insert’s cutting edge. NB  Appendix 1a shows the ISO ‘Code Key’ – for External Toolholders. Appendix 1b shows the ISO ‘Code Key’ – for Solid Boring Bars. Appendix 1c shows the ISO ‘Code Key’ – for Cartridges.

Insert Shape

Final Tool­ holder and Insert Selection

Edge Clamping System Initially, the tool holder clamping system should be selected to provide optimum performance in different applications over a wide range of workpiece geometries. The type of machining operation and to a lesser extent, the workpiece size determines tool holder selection. For example, roughing-out operations on big components will make considerably different demands, to that of finishing passes on small components. NB  Pin, clamp and lever are just three of the insert clamping systems available – consultation with the tool suppliers at this point might be beneficial.

The insert shape should be selected relative to the entering angle needed for the tool’s accessibility, or versatility. Here, the largest suitable point angle should be chosen for strength and economy (see Fig. 23). In Fig. 23, is illustrated a practical example of how changing only one variable – insert geometry (shape) – can influence an insert’s turning application. The shape of an insert will determine its inherent weakness, or strength, which is of particular relevance if roughturning operations are necessary. Furthermore, insert shape will influence whether it is prone to vibration, or not and its predictable tool life. Hence, if one is concerned about vibrations of either the tool, workpiece, or both, then a weaker insert such as a light turning and facing geometry with less cutting edge length exposed in-cut, might be more suitable. Variable conditions such as the selection of insert’s geometric shape can affect other machining parameters and, this is valid for other insert factors, so a compromise will always occur in any machining application.

Toolholder Size and Type Once the clamping system has been selected, the size and type of toolholder must be determined, with its selection being influenced by: feed directions (i.e. see Fig. 22 for turning insert shapes and feed directions), size of cuts, workpiece and toolholder situated in the machine for accessibility requirements. The workpiece’s shape plays a decisive role if surface contouring is necessary, this is particularly relevant for machining

12 Effective entering angles (κ1) must be carefully selected when the operation involves profiling, or copying. The maximum profiling angle (β) is recommended for each tool type – if ‘workpiece fouling’ is to be avoided. NB  κ1 = κ + β (for plunging into a surface), whereas κ1 = κ – β (for ramping-out of a surface), κ1 = κ (β = 0°) for cylindrical turning, Where: effective entering angle (κ1), entering angle (κ), maximum in-copy angle (β). Always select the smallest entering angle that the part geometry will allow.

46

Chapter 2

. Figure 22.  Tool paths in finish turning operations. [Courtesy of Sandvik Coromant]

Turning and Chip-breaking Technology

47

. Figure 23.  Selecting indexable inserts for turning operations. [Courtesy of Stellram]

NB  Appendix 1d shows the ISO ‘Code Key’ – for Indexable Inserts.

Insert Size An indexable insert size is directly related to the toolholder selected for the operation, with the entering angle and insert shape having previously been established. Only the matching-shaped insert can be fitted into the seat of a particular toolholder, as its shape and size are predetermined by the seating dimensions.

In roughing-out operations, the largest cutting depth for a given toolholder, will influence the insert size. For any insert, the effective cutting length has to be determined (see Fig. 20b), as the entering angle will influence the size of the insert selected. If the effective cutting edge length is less than the depth of cut (DOC), a larger insert should be chosen, or the DOC should be reduced. Sometimes in more demanding turning operations, a thicker insert – of the same geometric shape – gives extra reliability.

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Chapter 2

Nose Radius Of particular relevance in any turning operation is the insert’s tool nose radius (rε – see Fig. 17), as it is the key factor with regard to: • inherent strength in roughing operations, • the resulting surface texture from finishing operations. Further, the size of the nose radius affects vibrational tendencies (see Fig. 23) and in certain instances, the feedrates. The nose radius is the transition between the major and minor cutting edges, which determines the strength, or weakness of the point angle (see Figs. 16a and 17), therefore it is an imperative factor to get right. In general, roughing-out should be undertaken with the largest possible nose radius, as it is the strongest tool point (see Fig. 23). Further, a larger tool nose radius permits higher feedrates, although it is important to monitor any possible vibrational tendencies. Later in the relevant section, more will be said on the influence that the insert’s tool nose radius plays in the final machined surface texture, but it is worth mentioning here that the feedrate for roughing operations should be set to approximately half the size of the nose radius utilised. The size of the nose radius has an affect on the power consumed in turning in conjunction with the material’s yield strength and chip-forming ability, particularly in rough-turning operations. The maximum material removal rate (MMR) can be obtained by a combination of high feedrate, together with a moderate cutting speed, with other limiting factors, such as depth of cut (DOC), tool’s nose radius, under consideration. Often, the machine tool’s power (P) availability can sometimes be a limiting factor when mmR is the requirement and, in such circumstances the cutting speed is usually lowered somewhat. For a given nose radius and cutting insert geometry, the power can be derived, to ensure that the machine tool will be able to cope with this pre-selected mmR, in the following manner: Machine tool’s power requirement (P): P = tangential force (FT) x cutting speed (VC) P = FT × VC P = kC × A × VC ∴ P = kC × f × aP × VC (kW) Where: f = feed/rev aP = depth of cut

(mm/rev) (mm)

Cutting speed (VC) VC = πDN/1000

(m/min)

Where: D = workpiece diameter (mm) N = workpiece rotational speed (rpm) Specific cutting force (kC): kC = FT/A

(N/mm2)

Where: A = cutting area (mm2) For example, for finishing operations, with the nose radius in combination with the feedrate (i.e. pre-selected), this will affect the surface texture and part accuracy, in the following manner: Machined surface texture (Rt): (Rt, this parameter being: maximum profile height) (µm) Rt = f 2/8 × rε x 1000 Where: f 2 = feedrate per revolution rε = nose radius

(mm/rev) (mm)

NB  The surface texture parameter ‘Rt’ , can be converted into other surface texture parameters – as necessary. By utilising either: larger turning insert tool nose radius, ‘wiper insert’ (yet to be discussed), a more positive plan approach angle, or in certain circumstances, a higher cutting speed, the surface texture can be improved. In general, the coordination of the tool’s nose radius and the pre-selected feedrate in finishing operations, indicates that the feed should be kept below a certain level to achieve an acceptable machined surface texture value.

Insert Type The cutting insert type is for the most part determined by the previously selected geometry – see Appendix 1d for the selection of indexable inserts. In reality, various cutting conditions and workpiece materials make different demands on the insert’s cutting edge. For example, when machining hardened steel parts, this will be completely different from that to the machining of aluminium components.

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49

Once the insert shape has been established in connection with its plan approach angle together with the nose radius dimension, this just leaves the type of geometry to be found. In this instance, the type of insert geometry refers to the ‘working area’ (i.e. nominally found by its depth of cut and feedrate – more will be said concerning this topic later, when ‘chip-breaking envelopes’ will be discussed). Additional factors can influence the type of cutting geometry choice, such as: machine tool’s condition, its power, the stability of the workpiece-tool-machine set-up, other factors that could affect geometry selection include: whether continuous, or intermittent cutting occurs, any tendency toward vibration while machining. Turning operations can be separated into a number of ‘working areas’ , being based upon the removal of workpiece material and the generation of accurate machined component dimensions, in combination with specific surface texture requirements – as shown in Table 3. When establishing an insert type, the feedrate and depth of cut should be identified with one of the ‘working ranges’ (i.e. from Table 3), as the various insert types to be chosen relate to this chart. It should be borne in mind that the most suitable ‘working area’ selected, will vary, in combination with such factors as the insert’s: size, shape and nose radius.

Tool Material The penultimate evaluation to be made concerning tooling decision-making is the choice of insert mate-

. Table 3.  Typical working areas for external turning operations Type of machining operation:

Feedrate (f ):

Extreme finishing

0.05 to 0.15

0.25 to 2.0

Finishing

0.1 to 0.3

0.5 to 2.0

Light roughing

0.2 to 0.5

2.0 to 4.0

Roughing

0.4 to 1.0

4.0 to 10.0

Heavy roughing

>1.0

6.0 to 20

Extremely heavy roughing

>0.7

8 to 20

(mm)

(mm)

[Courtesy of Sandvik (UK) Ltd]

Depth of cut (DOC):

rial, or combination of materials that constitute the cutter’s tool edge. Today, manufacturers of tooling have a strategy for continuous improvement with variations in both tool matrices and coatings being considerable. Not only are cutting tool material research and development an on-going intensive activity, but their application for wider ranges of machining applications are being considerably enhanced. A brief review of just some of the current tool materials and coatings have been previously mentioned in Section 1.2, with the main range of cutting tool materials being: cemented carbides, coated cemented carbides, ceramics, cermets, cubic boron nitride, polycrystalline diamond and monolithic (i.e. natural) diamond. NB  A good ‘start-point’ for most machining operations, is to consider coated carbides initially, then if these grades prove unsatisfactory, for whatever reason, select one of the other materials – perhaps after consultation with a cutting tool manufacturer, or after a machinability testing procedure.

Cutting Data Once all of the physical, metallurgical and geometrical factors for the cutting tool have been established for the machining operation, then it is necessary to set, or calculate the cutting data – often these criteria can be found from tooling manufacturers recommendations and cutting data tables. Certain variable factors such as feedrate should have already been made, allowing the cutting speed to be calculated, from the well-known expression (below): VC = πDN/1000

(m min–1)

Where: VC = cutting speed (m min–1) D = Workpiece diameter (mm)13 N = rotational speed (rpm)

13 In the case of drilling, reaming and tapping operations, it is the diameter of the cutting tool that is used in the calculation. For any other internal machining operations – such as in boring, it is the initial hole diameter that is employed in the cutting speed calculation.

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Once again, manufacturers data tables are often useful ‘starting-points’ for estimating the initial cutting parameter information. Considerable care must be taken if the material has either a high work-hardening tendency, or intrinsic bulk (i.e. workpiece material) hardness, as this can influence the numerical data selected. Moreover, the plan approach angle also has an effect on the numerical value for the parameter, for example, oblique machining allows a higher value than for orthogonal machining.

2.2 History of Machine Tool Development and Some Pioneers in Metal Cutting 2.2.1 Concise Historical Perspective of the Development of Machine Tools Toward the end of the 1700’s, any high-quality machining at the time meant tolerances of 0.1mm being considered as ‘ultra-precision’ , with this level of tolerance having steadily improved from the beginning of the Industrial Revolution. Pioneers in machine tool development such as John Wilkinson (1774), developed the first boring machine, this being capable of generating a bored hole of 1270 mm in diameter, with a error of about 1 mm. A contemporary of Wilkinson, namely Henry Maudslay (1771–1831), invented many precision machine tools, but he was particularly noted for the design and development of the first engine lathe. Slightly later, Sir Joseph Whitworth (1803–1887), developed the first modern-day Vee-form screwthread and nut (i.e. 55° included angle – ‘Whitworth thread’), thereby enabling precision feed-motion to be achieved via suitable gear trains on such machine tools. These early fundamental advances in machine design, allowed others and in particular, Joseph R. Brown (1852) to design the ‘dividing engine’. This newly-developed equipment, allowed precision engraving of the hand dials on machine tool axes, enhancing them with much better machinist’s judgment in both rotary and linear control, in combination with consistent repeatability by the skilled operative. Shortly after these developments, Eli Whitney produced the original milling machine, which was refined still further

by the Cincinnati Screw and Tap Company in 1884. This ‘Cincinnati machine’ was a direct forerunner of today’s manual controlled knee-type milling machine tools. Of particular note was the ergonomic grouping of the controls centrally for a more efficient hand control by the skilled operator. At this time the machine tool still utilised the Vee-form screw thread, with the Acme-form (ie having the ability to take-up backlash) still someway off development. Steady development and refinement of a range of machine tools continued into the the first half of the 20th century until the next major ‘milestone’ occurred. This significant development was the ‘modern’ numerically-controlled (NC) machine. Around the late 1940’s, the ‘recirculating ballscrew’14 was designed so that it could take-up backlash in both directions of rotation for machine tool axes. These early ‘ballscrews’ were fitted to a converted Cincinnati Milling Machine Company’s ‘Hydro-Tel’ die-sinking machine tool, at MIT (Massachusetts Institute of Technology). This military research-funded project having been commissioned by the United States Air force – who required complex free-form aeronautical parts to be automatically machined for the latest aircraft. This research was undertaken by MIT, in association with ‘Cincinnati’ and the Parsons Tool Company. The binary-coded punched-paper tape, controlled the simultaneous machine tool axes using alpha-numerical characters (ie the forerunner of today’s programs using ‘G- and M-coded’ CNC controllers), through a

14 Who, when and where ‘recirculating ballscrew’ design and development took place is open to some debate. As proponents in the UK say it was Alfred Herbert and Sons, whereas in the United States, the Parsons Tool Company are often quoted as the originators. However, what is not in question, is that with its unique ‘Gothic’ arch’ (i.e. Ogival geometry), having point contacts between the screw and the adjacent recirculating balls, allows the assembly to be pre-loaded in-situ, thereby eliminating any appreciable backlash allowing accurate control of these axes. NB  The previous Acme taper thread (i.e. 29° included angle) fitted to ‘conventional’ machine tools had an efficiency of no better that 40% – with backlash present, whereas today’s hardened ‘ballscrews’ have efficiencies of ~90%, coupled to an impressive rigidity (~900 N µm–1) and minimal ‘stick-slip’ , therefore minimising the so-called ‘ballscrew wind-up’ due to the action of torque-effects in combination with the cutting forces.

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51

valve-driven hydraulically-servo controlled ‘computer’ called ‘Whirlwind’. In the late 1970’s, with the advent of microprocessor technology, these later NC machine tools were converted to Computer Numerical Control (CNC), offering a significant stride forward in operator-usability, via on-board editing – without the costly and timely re-punching of NC paper tapes each time a minor modification occurred to the NC program. Today, CNC machine tools have fast multiple-processor controls, with on-line computer graphics, enabling new programs to be written and ‘prove-out’ while the machine tool cuts other components, or the programs can be automatically down-loaded by a Direct Numerical Control (DNC) data-link from the CAD/CAM workstation, or via remote satellite-linkage from other sites either locally, or internationally. The design and development of some of today’s and the future machine tools, utilise ultra-fast CNC microprocessors, coupled to orthogonal multi-axes linear-induction motordriven slideways, that can be precisely monitored via laser-controlled positional encoders, with ultra-fast co-axial spindles. Moreover, non-orthogonal-axes controlled machine tools are under development, using simultaneous mulitple-axes slideway control, with hybrids having tool spindles that incorporate multiple angular orientation together with their linear slideways for truly sculptured free-form surface machining capabilities. Even today, operations carried out by several machine tools are now being incorporated into one hybid machine tool, with such as: turning, milling and grinding at one set-up. In the near future, the machine tools will have slideway acceleration/decelerations of faster >5g’s, with these machines having the ability to: rough-turn, mill, heat-treat, grind critical features, all remotely-controlled via satellite from the CAD/CAM designer, significantly speeding-up the product development process time-to-market.

In 1873, Hartig tabulated research into metal cutting in a book, which was the first authoritative work on the subject. A more practical metal cutting description was given by Tresca (1878), ustilising visio-plasticity models15. In 1881, a presentation at the Royal Society of London by Lord Rayleigh of Mallock’s metal cutting research findings was given. Mallock’s scientific study of carefully etched specimens of the workpiece and attached chip for both ferrous and non-ferrous metals, where he observed them using a microscope (magnification: x5). Mallock correctly surmised from his investigation of his ‘models’ that the cutting process was basically one involving shearing and, that friction occurred in forming the chip, emphasizing the importance of this friction along the cutting tool’s face – between the chip and the tool. The sharpness of the cutting edge was also mentioned and the reasons for instability of the cutting process, leading to unwanted vibrations, or ‘chatter’. Moreover, Mallock employed basic lubricants in this work, noting that the application of lubrication reduced chip/tool interface friction. These general observations by Mallock mentined above, offer a surprisingly close approximation to today’s theories on the ‘mechanics of metal cutting’ , although his equations for the work done in internal shearing and chip and tool friction were incorrect, surprisingly, he was unaware of the ‘plasticity models’ by Tresca and his theory of ‘plastic heating’. To compound the metal cutting problems still further, in 1900, an unfortunate ‘step backward’ in the understanding of the metal cutting process was taken by Reuleaux. He suggested that a crack occurred ahead of the tool’s point and likened the cutting action to that of splitting wood, regrettably having popular support for some years. In 1907, a seminal paper by the now-famous American researcher Taylor, who published his 26 years of

2.2.2 Pioneering Work in Metal Cutting – a Brief Resumé

15 Tresca’s visio-plasticity models, involved scoring a grid of accurate closely-spaced lines onto the edge of a specimen of metal to be machined, then partially cutting it at a preset depth of cut and leaving the chip attached. He then investigated the plastic deformation that had taken place as these grids were distorted and buckled by the action of machining. Both lighter and deeper cut depths were investigated in this manner, across a range of metal specimens. Tresca noted that finer depths of cut introduced greater plastic deformation than larger cut depths, stating that stiffer and more powerful machine tools were needed to benefit from these recommended greater depths of cut (i.e. undeformed chip thickness).

Basic research into metal cutting did not commence until approximately 70 years after the first machine tool was introduced. In 1851, early research by Cocquilhat was into the work required to machine a given volume of material by drilling. By 1870, the terms ‘chip’ and ‘swarf ’ were introduced by the Russian engineer Time, where he attempted to explain how chips were formed.

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. Figure 24.  The formation of a continuous chip, based upon the ‘deck of cards’ principle. [After: Piispanen, 1937]

practical experience into investigation and research findings in metal cutting. Taylor, was fascinated by the application of time-and-motion studies that could be applied within the machine shop and in particular, ‘piece-work systems’16. In order to enable the progression through optimisation of these time-and-motion studies, new cutting tool materials were employed, in particular high-speed steels (HSS). Taylor investigated the effect that tool materials and in particular, cutting conditions had, on tool life during roughing operations, in order to assist in the application of these time-and-motion studies. His principal objective was to establish empirical laws that would enable optimum

cutting conditions to be attained. By establishing optimum cutting data for metal cutting operations and employing ‘piece-work systems’ at the company, Taylor was able to increase the Bethlehem Steel Company’s output by 500%. Of particular note, was the fact that the empirical law governing the cutting tool and its anticipated tool life17 is still used today, in the study of machining economics – more will be said on this topic later in Chapter 7 (Machinability and Surface Integrity). Notable in the years prior to World War Two, were the contributions made into the generation of data on cutting forces and tool life, initially by Boston (1926)

16 Piece-work systems are where a set time allowance is given for a particular job, or a batch and, a bonus is agreed if the worker performs this task within the allotted time.

17 Taylor’s machinability work produced a fundamental discovery, namely, that the interface temperature existing at the tool’s cutting edge controlled the tool-wear rate.

Turning and Chip-breaking Technology

53

. Figure 25.  Variations in chip morphological surfaces at different cutting speeds, giving an indication of the various shearing mechanisms. [Source: Watson & Murphy, 1979]

and later, by Herbert (1928). Around this time, the cutting speeds were steadily improving with the arrival of new cutting tool materials, such as cemented carbide. In 1937, Piispanen introduced his so-called ‘Deck of Cards’ principle as an explanation of the cutting process (see Fig. 24 for Piispanen’s idealised model, with Fig. 25 depicting sheared chips at a range of cutting speeds). Here, Piispanen’s model depicts the workpiece

material being cut in a somewhat similar manner to that of a pack of cards sliding over one another, with the free surface an angle, which corresponded to the shear angle (ϕ). So, as the tool’s rake face moves relative to that of the workpiece, it ‘engages’ one card at a time, causing it to slide over its adjacent neighbour, this process then repeats itself ‘ad finitum’ – during the remainder of the cutting process. Some important

54

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limitations are present with Piispanen’s model, namely that it: • exaggerates strain in homogeneity, • shows tool face friction as elastic rather than plastic in nature, • considers shearing takes place on a completely flat plane, • assumes that BUE does not occur, • takes an subjectively assumed shear angle, • takes no account of either chip curling, or prediction of chip/tool length. NB  Piispanen’s model is easily understood and does contain the major concepts in the chip-forming process – admittedly for simple shear in the main. By way of further information concerning chip morphology: the micrographs of chip surfaces illustrated in Fig. 26 show in these cases, that the morphology indicates a semi-continuous chip form. These chip forms point towards the fact that noticeable periodic variations have occurred, perhaps as the result of the stress becoming unstable, rather than resulting from any vibrational effects produced by the machine tool. Any such instability, has the effect of causing minute oscillations (i.e. backward and forward motion) in the shear zone, while the machining takes place. The differences in segment shapes shown and their frequency occurring at differing cutting data in these micrographs, are thought to be dependent upon the frequency of the shear plane’s oscillation relative to the cutting speed. A considerable volume of fundamental work on machining research has been undertaken over the last few years, but during World War Two (i.e. from a European perspective), Ernst and Merchant (1941) produced another significant paper dealing with the mechanics of the machining process – some of these research findings will be briefly dealt with in the chapter on Machinability and Surface Integrity, along with other contributions to this subject.

picted schematically in Fig. 26. The overall machining process is well concealed behind a amalgamation of: workpiece material, high speeds and feeds, elevated temperatures and enormous pressures18. The actual cutting dynamics in contemporary machining operations, utilises just a few millimetres of physical contact between the tool and the chip of a precisely-shaped cutting edge geometry in an exotic mixture of tool material to efficiently machine the workpiece – this being an impressive occurrence worthy of note. In the early work on machining, it was thought that the chip was formed by deformation along a shear plane, elastically in the first instance, then plastically as the evolving chip passed through a stress concentration. The Piispanen model (i.e. Fig. 24) illustrates this point, where workpiece material is being cut by progressive slip relative to the tool point, an angle which corresponded to that of the shear plane. Here (i.e. Fig. 24), it shows how each chip segment forms a small, but very thin parallelogram, with slippage occurring along its shear plane. In an orthogonal cutting process19, as the workpiece material approaches this ‘shear plane’ it will not begin to deform until it reaches the ‘shear plane’. Here, it is transformed from that of simple shear, as it moves across a thin shear zone, with the minute amount of secondary shear being virtually ignored, as is the case for tertiary shear – this being the equivalent of a sliding friction but having a constant coefficient of friction. Chip deformation in reality, is produced over a zone of finite width, usually termed the ‘primary shear zone’ (see Fig. 26). As the chip evolves, the back of the chip tends to be roughened, due to the plastic strain being inhomogeneous in nature (see Fig. 25). This shearing action creates a particular chip morphology as a result of the either, stress concentrations, or by presence of points of weakness in the workpiece be-

2.3 Chip-Development

18 Interface pressures between the chip and the tool are normally exceedingly high, typically of the order of 1,000 to 2,000 N mm–1, with temperatures in certain instances at the tool’s face reaching approximately 1100°C.

Most metallic materials can be considered as relatively hard to machine and this is evident from all of the reported literature on the subject of metal cutting, indicating that shearing occurs in a concentrated region between the chip and tool, this effect being de-

19 Orthogonal machining, is when the cutting tool’s edge (i.e. rake face – see Fig. 19b) is presented ‘normal’ to the evolving chip and thus, to the workpiece, at 90° to the relative cutting motion. That is, little if any, side shearing action occurs, while the chip is being formed as it progresses up the tool’s rake face – effectively created by two distinct cutting forces: tangential and axial.

Turning and Chip-breaking Technology

55

. Figure 26.  Schematic representation of a sing-point stock removal process, during the continuous cutting of ductile metals

56

Chapter 2

ing machined20. Once the chip deformation begins, it will continue within this ‘zone’ , as though here in this vicinity, the workpiece material is exhibiting a form of negative strain-hardening. The oblique cutting process21 presents a different and much more complex analytical problem, which has been the subject of a lot of academic interest over the years. Even here, the whole cutting dynamics change, when the tool’s top rake surface is not flat, which is the normal status today, with the complex contoured chipbreaker geometries nowadays employed (typically illustrated in Figs. 4, 10 and 27a). Actual chips are normally severely work-hardened, in particular with any strain-hardening materials (for example: high-strength exotic alloys employed for heat-resistance/aerospace applications) as they evolve, by the combined action of: elevated interface temperatures, great pressures and high frictional effects. Such machined action of the combined effects of mechanical and physical work, produce a ‘compressive chip thickness’ 22, which is on average, dimensionally wider than the original undeformed chip thickness (see Fig. 26). The rake angle depicted in Fig. 26 is shown as positive, but its geometry can tend to the neutral, right through to the negative in its inclination. As the rake angle changes, so will the complete dynamic cutting behaviour also change, modifying the mechanical and

20 As the shear plane passes through a particular stress concentration point, it will deform more readily and at a lower stress value, than when one of these ‘points’ is not present. 21 Oblique machining, is when the rake face has a compound angle, that is it is inclined in two planes relative to the workpiece, having both a top and side rake to the face, creating a threeforce model (see Fig. 19a), where the cutting force mathematical dynamics are extremely complex and are often produced by either highly involved equations, or by cutting simulations. This latter simulated treatment is only briefly mentioned later and is outside the remit of this current book. However, this information on dynamic oblique cutting behaviour can be gleaned, from some of the more academic treatment given in some of the selected books and papers listed at the end of this chapter. 22 Compressive chip thickness is sometimes known as the: chip thickness ratio (r)* – being the difference between the undeformed chip thickness (h1) and the width/chip thickness of the chip (h2). *Chip thickness ratio (r) = h1/h2 ** (i.e. illustrated in Fig. 26). ** h2 = W/ρwl Where: W = weight of chip, ρ = density of (original) workpiece material – prior to machining, w = chip width (i.e DOC), l = length of chip specimen.

physical properties within the chip/tool region, as the various deformation zones are distinctly altered. In effect, due to rake angle modification (i.e. changing the rake’s inclination), this can have a profound affect on the: cutting forces, frictional effects, power requirements and machined surface texture/integrity. The chips formed during machining operations can vary enormously in their size and shape (see Fig. 35a). Chip formation involves workpiece material shearing, from the vicinity of the shear zone extending from the tool point across the ‘shear plane’ to the ‘free surface’ at the angle (ϕ) – see Fig. 26. In this region a considerable amount of strain occurs in a very short time interval, with some materials being unable to withstand this strain without fracture. For example, grey cast iron being somewhat brittle, produces machined chips that are fragmented (i.e. termed ‘discontinuous’), conversely, more ductile workpiece materials and alloys such as steels and aluminium grades, tend to produce chips that do not fracture along the ‘shear plane’ , as a result they are continuous. A continuous chip form may adopt many shapes, either: straight, tangled, or with different types of curvature (i.e. helices – see Fig. 35a). As such, continuous chips have been significantly worked, they now have considerable mechanical strength, therefore efficiently controlling and dealing with these chips is a problem that must be overcome (see the section on Chip-breaking Technology). Chip formation can be classified in a number of distinct ways23, these chip froms will now be briefly reviewed: • Continuous chips – are normally the result of high cutting speeds and/or, large rake angles (see Figs. 26 and 27b). The deformation of workpiece material occurs along a relatively narrow primary shear zone, with the probability that these chips may develop a secondary shear zone at the tool/chip interface, caused in the main, by frictional effects. This secondary zone is likely to deepen, as the tool/chip friction increases in magnitude. Deformation can also occur across a wide primary shear zone with

23 One of the major cutting tool manufacturer classifies chips in seven basic types of material-related chip formations, these are: Continuous, long-chipping – mostly steel derivatives, Lamellar chipping – typically most stainless steels, Short-chipping – such as many cast irons, Varying, high-force chipping – many super alloys, Soft, low-force chipping – such as aluminium grades, High pressure/temperture chipping – typified by hardened materials, Segmental chipping – mostly titanium and titanium-based alloys.

Turning and Chip-breaking Technology

. Figure 27.  Chip-breaking inserts and chip control whilst turning – in action. [Courtesy of Iscar Tools]

57

58

Chapter 2

. Table 4.  Strength and hardness of chips when turning mild steel. Rake angle (γ°):

45

35

27

10

10

10

0.30

0.30

0.20

0.20

0.20

0.20

50

50

168

168

168

76

None

None

None

None

75

84

91

92

93

95

272

289

302

320

314

325

HV/UTS

3.6

3.4

3.3

3.5

3.4

3.4

Shear strain (Pa-s)

1.1

1.7

2.1

2.9

3.1

4.0

Feed (f - mm rev–1) Cutting speed ( VC - m min–1) Cutting fluid: Tensile strength (UTS – kg mm–2) Vickers hardness number (Hv)

Soluble oil

Soluble oil

[Source: Nakayama and Kalpakjian 1997]

curved boundaries, with the lower boundary being below the machined surface (Fig. 26), which may distort a softer workpiece’s machined surface – particularly with small rake angles and at low speeds. Strain-hardening of this type of chip, results in it becoming harder than the bulk hardness of the original workpiece material (see Fig. 28c, where the bulk workpiece hardness is 230 HK and the workhardened chip is ≈350 HK). This increase in the chip’s strength and hardness will depend upon the shear strain (see Table 4 for details of the Rheolological24 status, related to the inclination of the tool’s rake angle). Therefore as the rake angle decreases, the shear strain will increase, causing this continuous chip to become both harder and stronger – behaving in a similar manner to that of a rigid, perfectly plastic body. In order to satisfactorily deal with long continuous work-hardened chips, that could either wrap around the machined workpiece, potentially spoiling the surface texture, or become ensnarled around tooling, or even, reduce efficient coolant delivery to the cutting edge, with integrated tool chip-breakers having been designed and developed – see Fig. 27b.

24 Rheology is a branch of science dealing with both the flow and deformation of materials, with the shear strain rate, often termed just the shear rate. (i.e usually quoted in Pascals-seconds ‘Pa-s’).

• Continuous chips with a built-up edge (BUE) – when

machining ductile workpiece materials, a built-up edge (BUE) can form on the tool’s tip. This BUE consists of gradually deposited material layers from the workpiece, hence the term ‘built-up’ (see Fig. 28). As cutting continues, the BUE becomes larger and more unstable, eventually partially breaking away, with some fragments being removed by the underside of the chip, while the remainder is randomly deposited on the workpiece’s surface (Fig. 28a). This process of BUE formation, shortly followed by its destruction, is continuously repeated during the whole cutting operation. The BUE deposited on the workpiece will adversely affect the machined surface texture. The BUE modifies the cutting geometry, creating a large cutting tip radius (Fig. 28a and b). Due to the BUE being severely work-hardened by the action of successive deposits of workpiece material, the BUE’s hardness significantly increases by around 300% over the bulk component hardness (Fig. 28c). At this severely work-hardened level, the BUE becomes in effect a modified cutting tool. Normally, an unstable BUE is undesirable, conversely, a thin stable BUE is as a rule, regarded as desirable, as it protects the top rake surface. The formation mechanism for the BUE is thought to be one of adhesion of workpiece material to the tool’s rake face, with the bond strength being a function of the affinity of the workpiece to that of the tool material. This adhesion, is followed by the successive buildup of adhered layers forming the BUE. Yet another factor that contributes to the formation of a BUE,

Turning and Chip-breaking Technology

. Figure 28.  The development of a continuous chip with Built-Up Edge (BUE), its typical hardness distribution and its affect on the machined surface

59

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is the strain-hardening tendency of the workpiece material. Therefore, the greater the strain-hardening exponent, the higher will be its BUE formation. From experiments conducted on BUE formation, it would seem that the higher the cutting speed, the less is the tendency for BUE to form. Whether this lack of BUE formation at higher cutting data is the consequence of increased strain rate, or the result of higher interface temperatures is somewhat open to debate. However, it would seem that a paradox exists, because as the speed increases, the temperature will also increase, but the BUE decreases. The propensity for BUE formation can be lessened by: I. Changing the geometry of the cutting edge – by either increasing the tool’s rake angle, or decreasing the DOC, or both, II. Utilising a smaller cutting tip radius, III. Using an effective cutting fluid, or IV. Any combination of these factors.

• Discontinuous chips – consist of adjacent work-

piece chip segments that are usually either loosely attached to each other, or totally fragment as they are cut (Fig. 29). The formation of discontinuous chips usually occur under the following machining conditions: I. Brittle workpiece materials – these materials do not have the machining capability to undergo the high shear strains, II. Hard particles and impurities – materials with these in their matrix, will act as ‘stress-raisers’ and actively encourage chip breakage, III. Very high, or low cutting speeds – chip velocity at both ends of the cutting spectrum, will result in lack of adherence/fragmentation of the chip segments, IV. Low rake angles/large DOC’s – either small top rakes and heavy DOC’s will decrease the adherence of the adjacent chip segments, V. Ineffective cutting fluid – poor lubricity, combined with a meagre wetting ability, will encourage discontinuity of chip segments, VI. Inadequate machine tool stiffness – creating vibrational tendencies and cutting instability, leading to disruption of the machining dynamics and loosening of chip segments.

As mentioned in ‘Roman II’ above, the hard particles and impurities tend to act as crack nucleation sites, therefore creating discontinuous chips. Large DOC’s

increase the probability that such defects occur in the cutting zone, thereby aiding discontinuous chip formation. While, faster cutting speeds result in higher localised temperatures, causing greater ductility in the chip, lessening the tendency for the formation of discontinuous chips. If the magnitude of the compressive stresses in the both the primary and secondary shear zones significantly increase, the applied forces aid in discontinuous chip formation, this is because of the fact that the maximum shear strain will increase, due to the presence of an increased compressive stress. NB  Due to the nature of discontinuous chip formation, if the workpiece-tool-machine loop is not sufficiently stiff, this will generate vibrational and chatter tendencies, which can result in an excessive tool wear regime, or machined component surface damage.

• Segmented chips – are sometimes termed: in-, or

non-homogeneous chips, or serrated chips. This chip form has the characteristic saw-toothed profile which is noted by zones of low and high shear strain (Fig. 30). These workpiece materials possess low thermal conductivity, as such, when machined their mechanical strength will drastically decrease with higher temperatures. This continuous thermal cycle of both fracture and rewelding in a very narrow region, creates the saw-toothed profile, being particular relevant for titanium and its alloys and certain stainless steel grades. For example, to explain what happens in realistic machining situation, the specimen Fig. 30a is displayed, for an austenitic stainless steel quick-stop micrograph. This micrograph being the result of a less than continuous machining process (1), utilising a 5° top rake-angled turning insert. Here, variations in the cutting process have created fluctuations in the cutting forces, resulting in waviness of the machined surface (2). Prior to the material yielding, then the shearing process occurring, the workpiece material has deformed against the cutting edge (3). To explain how changing the top rake angle influences the resultant chip formation for an identical stainless steel workpiece material, Fig. 30b is shown. Machining has now been undertaken with a 15° top rake, promoting a more continuous machining process than was apparent with the 5° tool (i.e. illustrated in Fig. 30a). This more efficient cutting process, results in smaller variations in the cutting forces (1 and 2). The chip is seen to flow over the rake face in a more

Turning and Chip-breaking Technology

. Figure 29.  Discontinuous chip formation. [Courtesy of Sumitomo Electric Hardmetal Ltd.]

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consistent manner (3). It was found with this workpiece material in an experimental cutting procedure, that the tangential cutting force component, was closer to the actual cutting edge than when similar machining was undertaken on unalloyed steel specimens. NB  The cutting data for machining the stainless steel specimens in Figs. 30 a and b, were: 180 m min–1 cutting speed, 0.3 mm rev–1 feedrate, 3 mm DOC.

2.4 Tool Nose Radius The insert’s nose radius has been previously mentioned in Section 2.1.6, concerning: Cutting Tool holder/Insert Selection. Moreover, the top rake geometry of the cutting insert will significantly affect the chip formation process, particularly when profile turning. In Fig. 31a, a spherically-shaped component is being ‘profile machined’ using a large nose-radiused turning insert. Here, as the component nears its true geometric curvature, the cutting insert forces will fluctuate continu-

. Figure 30.  Segmented chip formation, resulting from machining stainless steel and the work-hardening zone – which is affected by the sharpness of the insert’s edge. [Courtesy of Sandvik Coromant]

Turning and Chip-breaking Technology

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. Figure 31.  The cutting insert’s tool nose radius when either profiling, or general turning, will modify both the profile and diameter as flank wear occurs. [Courtesy of Sandvik Coromant]

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ously as the insert progresses (i.e circular interpolates with the X- and Z-axes of the machine tool) around the curved profile. If the geometry of the tool was not itself of round geometry, then the ‘point-contact’ could not be maintained, leading to significant variations in chip formation. If this lack of tool-work contact were not to occur, then the machined profile would be compromised and due to insufficient chip control, the actual cut surface profile would not have a consistent and accurate surface texture. The machined surface texture generated by the passage of the cutting insert’s geometry, is to a large extent the product of the relationship, between the nose radius and the feedrate and, to a lesser degree the cutting speed and its tool wear pattern. The size of the tool nose radius will have quite an effect on the surface texture produced, if the feedrate is set, then a small nose radius will create a different workpiece surface texture to that of a larger one (see Fig. 31b). Moreover, if a large nose radius is selected for a lighter DOC, or if the feed is equal to the nose radius, then this larger nose geometry will be superior to that of a smaller tool nose radius. This is because the ‘larger nose’ offers a smaller plan approach angle, having the pressure of the cut distributed across a longer cut length, creating an enhanced surface texture. There are several disadvantages to utilising a larger tool nose radius geometry, these are that the: • Chip formed becomes more difficult to bend and effectively break, • Radial cutting forces are greater, • Power consumption increases, • Rigidity of the set-up is necessary – leading to possible vibrational tendencies on either weaker, or unstable workpieces. Tool wear (i.e. denoted by ‘∆’ in Figs. 31ci and cii) and in particular flank wear25, can significantly influence the resulting machined component dimensional accuracy (Fig. 31cii), which on a batch of components cut with the same insert, will result in some level of ‘tool

25 Flank wear is normally denoted by specific ‘zones’ – more will be said on this topic later – but, in this example, the tool’s insert wear ‘VB’ is shown in both Figs. 31ci and cii.

drift’ which could affect the process capability26 of the overall parts produced. This flank wear ‘VB’ can be calculated and utilised to determine the anticipated tool’s life (ie, in-cut), this important factor in production machining operational procedure, will be discussed in due course. Wiper blades (Fig. 32) are not a new insert geometry concept, they have been used for face milling operations for quiet a long time, but only in recent years are they being utilised for component finish turning. The principle underlying a wiper insert for turning operations, concerns the application of a modified ‘tool nose radius’ (see Fig. 32 – bottom left and right diagrams). When a ‘standard’ tool nose geometry insert is used (i.e. Fig. 32 – bottom left), it creates a series of peaks and valleys (i.e. termed ‘cusps’27) after the passage of the ‘insert nose’ over the machined surface. Conversely, a cutting insert with wiper blade geometry (i.e. Fig. 32 – bottom right), has trailing radii that blends – beyond the tangency point – with the tool nose radius which remains in contact with the workpiece, allowing it to wipe (i.e. smooth) the peaks, leaving a superior machined surface texture. In the past, wiper insert geometries were only employed for surface improvement in finishing opera-

26 Process capability denoted by ‘CP’ , is a measure of the quality of the parts produced, which is normally found by the following simple relationship: *CP = Drawing specification tolerance/6 σ Where: σ = a statistical measure, termed the ‘standard deviation’ for the particular production process. *CP values of <1.0 denote low process capability, CP values of between 1.0 and 1.33 are moderate process capability, CP values of >1.33 are termed as high process capability. NB  Today, process capabilities of 2.0 are often demanded for high-quality machined parts for the automotive/aerospace sectors of industrial production, reducing likelihood of part scrappage. 27 Cusps are the product of the partial geometry of the tool nose radius geometry, positioned at regular intervals related to the selected feedrate. The cusp height (i.e. the difference in height between the peak and valley), will influence the machined surface texture of the component, in the following relationship: Rmax = fn2 × 250/rε (µm) Where: Rmax = maximum peak-to-valley height within the sampling length. fn = feedrate (m min–1) rε = tool nose radius (mm).

Turning and Chip-breaking Technology

. Figure 32.  The application of wiper insert geometry on the resulting surface texture when fine turning. [Courtesy of Iscar Tools]

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tions. With recent advancement in wiper28 geometry, this has allowed them to be used at double the previous feedrates for semi-finishing/roughing operations, without degrading the surface texture. The wiper geometry being in contact with the workpiece’s surface for longer than equivalent standard insert nose radius tends to wipe – hence its name, or burnish the machined surface, producing a smoother surface texture. Due to the fact that a ‘wiper’ has an extended edge, the cutting forces are distributed across a longer tool/chip contact region. The wiper portion of the insert, being somewhat protected, enables these wiper inserts to increase tool life by up to 20% more than when using conventional tool nose geometries. Wiper blades have their clearance lengths carefully designed, if they are too long, the insert generates too much heat, on the contrary, they need to be long enough to cope with relatively large feeds, while still smoothing over the surface cusps. Wipers with positive turning insert geometries, they can cope with feedrates of 0.6 mm rev–1 at DOC’s of up to 4 mm. For example, with steel component hardnesses of 65HRc, this often negates the need for any successive precision grinding operations. By designing wiper geometries with the cutting edge and nose radii to improve machined surface finish, while increasing tool life, can be considered as outstanding tool design.

2.5 Chip-Breaking Technology 2.5.1 Introduction to Chip-Breaking The technology of both chip-forming and chip-breaking has been one of the major areas of advancement in recent years. A whole host of novel toolholders and cutting inserts has been developed to enable the cutting process to be under total chip control, allowing some toolholder/inserts combinations to machine multiple component features with just one tool, removing at a ‘stroke’ the non-productive aspects of

28 Some tooling manufacturers have re-named wiper inserts as high-feed inserts, as they have demonstrated in production conditions to promote higher component output, without the recourse to expensive capital outlay.

tool-changing and setting, significantly increasing machine tool utilisation rates. Even when conventional turning inserts are employed, for heavy roughing cuts (Fig. 33a), where feedrates are high as are the large DOC’s, efficient control of the chip must be achieved. To enable excellent control of chip-breaking with roughing cuts (Fig. 33b), a similar overall insert geometry is shown to that in the previous example, but here the rake face embossed dimples/chip-breakers differ significantly. Finally, for light finishing cuts (Fig. 33c), chips are broken in a totally different manner to that of the previous examples. Hence, with all of these differing types of turning operations on workpieces, control of the chip is vital, as it can drastically impair the overall production rates and affect part quality, if not given due consideration. Chip formation is chiefly influenced by the following factors: • Workpiece material composition – its heat treatment (i.e. if any), which affects the chip’s strength, • Insert’s cutting geometry – rake and clearances, as well as any chip-formers present, the geometry being associated with the work piece material, • Plan approach angle – depending upon whether roughing, or finishing cuts are to be taken, • Nose radius – this being linked to the feedrate and here, to a lesser extent, the surface texture requirements, • Undeformed chip thickness (i.e. DOC) – this will affect the chip curling aspect of the chip’s formation – more will be said on this topic in the following section. Note:  Another important factor that can also play a significant role in chip formation, is the application of coolant and its supply velocity. The shear angle has some effect on the contact length between workpiece and the rake face and, it is in this vicinity that cutting forces and machining-induced temperatures predominantly affect the cutting insert. Moreover, the insert’s rake is significant, in that as the rake angle increases the contact length decreases, the more positive the rake, the shorter the contact length. Actual chip formation is primarily dependent upon several factors: DOC, feedrate, rake angle, together with the workpiece’s mechanical strength, noting that the chip starts forming in the primary deformation zone (see Fig. 26). Thus, the chip is subsequently formed by the bending force of the cutting action, effectively ‘pivoting’ from the chip’s roughen ‘free top surface’ ,

Turning and Chip-breaking Technology

. Figure 33.  Turning cuts and associated insert geometries for forming and shearing of a chip. [Courtesy of Sandvik Coromant]

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this being a somewhat shorter length than that of the ‘shiny’ underside at the tool/chip interface. Many theories have been given for the actual ‘cause and effect’ of preliminary chip formation which is schematically illustrated Fig. 33d – ‘A’- one such, being that any formation is related to the cutting speed. A large insert rake angle normally means that there is less tendency for chip curling through a larger radius, but it will have lower cutting forces. In Fig. 33d – ‘B’ , is depicted a somewhat ‘idealised’ view of the actual cutting process, which can be expressed via the simple relationship of ‘λ’ and ∆X/∆Y. NB:  In this schematic representation: ‘h1’ represents DOC and, ‘ϕ’ is the ‘shear plane angle’. When utilising CNC machine tools and in particular turning centres, a major problem is the variety of continuous chip forms created and the large quantity and volume of swarf29 produced. The manner to which swarf affects machining operations depends upon the operating conditions, but fundamentally there are several requirements in any form of swarf control, these are: • The swarf must flow freely away from the cutting zone, without impairing the cutting action’s efficiency, • Swarf must be of convenient size and shape to facilitate handling manually, or in swarf conveyors (i.e. if fitted), together with any future large-volume storage, then transportation and subsequent disposal, • Any swarf should drop away into the machine’s swarf tray, without snarling around, the workpiece, tool, or interfering with other functions such as: automatic tool-changing magazine/turret, in-situ touch-trigger inspection probes, component load-

29 Individual chips when in any great volume are generally termed swarf. It is important to be able to manage this swarf volume and, satisfactory chip control can be determined by ‘Lang’s chip-packing ratio’ , this being denoted by the letter ‘R’ , in the following manner: R = Chip volume (mm3)/Equivalent volume of uncut workpiece material (mm3) NB:  ‘R’ ranges from values of 3-to-10, where an R-value of 4 gives satisfactory chip-breaking control, producing neatly curled ‘6 and 9-shaped’ chips.

ing equipment, such as overhead gantries, or dedicated robotic loading devices. In terms of priority for these swarf control factors, possibly the most important one is that the swarf should flow smoothly away from the cutting area, as with the latest chip-breakers fitted to today’s cutting inserts, chips can be readily broken and controlled30, this will be theme of the following section.

2.5.2 The Principles of Chip-Breaking In machining, the cutting edge’s primary function is to remove stock from the workpiece. Whether this is achieved by forming a continuous chip, or by the flow of elemental chips will depend upon several factors, including the properties of the workpiece material, cutting data employed and coolant type and its delivery. The terms ‘long-chipping’ and ‘short-chipping’ are utilised when considering the materials to be machined. Short-chipping materials such as most brasses and cast irons, do not present a chip-breaking problem for swarf disposal, so this section will concentrate on the long-chipping workpiece materials, with particular focus on ‘steel family’ grades. Steels are produced in a wide variety of specifications and this allows their properties to be ‘tailored’ to the specific industrial applications. In addition, these steels methods of primary processing, such as: casting, forging, rolling, forming and sintering, together with the type of subsequent heat treatment, creates still further metallurgical variations that may have an even greater influence on the workpiece’s chip-breaking ability. The workpiece’s strength and hardness values describe the individual material’s character to some extent, but it should be borne in mind that it is the chip’s mechanical strength that determines whether it can be broken with ease. No absolute correlation exists between a steel com-

30 Today, many high-volume manufacturing companies have realised the benefit of the value of clean and briquetted swarf, as opposed to oily scrap swarf, which sells at just ‘fractions’ of this value. At present, briquetted and cleaned aluminium swarf can be sold for approaching £1,000/tonne, moreover, the coolant/oil can be reclaimed, further driving down the overall machining costs. For other non-ferrous ‘pure’ metals and others, such as copper alloys and brasses, the economic savings are even greater.

Turning and Chip-breaking Technology

ponent’s strength and the mechanical strength of the chip, illustrating that a complex metallurgical and cutting tool geometric relationship exists whilst machining occurs. In particular for turning operations, a conventionally-turned chip is a rather frail product of serrated appearance (see Figs 25 and 34a and b). In order to promote good chip-breaking tendencies, thus enabling short elements to be formed, it is necessary to encourage this basic character by causing these serrations to be as deep as possible and the chip sections in between to be rigid. This chip occurrence causes the chip to be inflexible, which can then subsequently be broken. There are several distinct ways in which chips can then be broken, these include: • Self-breaking – this is when the chip’s mechanical strength is not great enough to hold the chip segments together and they consequently break upon exiting the machining region (Fig. 31a), • Chip collision with the workpiece – as the chip is steered towards an obstacle such as the workpiece’s surface this provides the breaking force (see Figs. 33 and 34b), • Chip is stopped by the tool – here the chip-curling behaviour comes into play, this being a function of the: tool’s nose radius geometry, depth of cut and feedrate employed (see Fig. 34 bottom left-hand photograph), the latter two functions affecting the chip cross-section, or chip thickness31.

31 Chip thickness is influenced by the plan approach angle utilised and the DOC, in association with the selected feedrate. The chip thickness is measured across the cutting edge, perpendicular to the cut (i.e. along the main cutting edge). The chip width and thickness are the dimensions that define the theoretical cut of the edge into the workpiece material. Hence, the chip thickness will vary with the size of the plan approach angle according to the relationships involving: feedrate, DOC and the effective cutting depth. The chip thickness is related to the plan approach angle and this affects the amount of pressure bearing upon the cutting edge. Hence, the thinner the chip, the smaller the distributed pressure along the edge and the less power consumed, conversely, the thicker the chip, the greater will be the machine tool’s power consumption. A thicker chip is generally advantageous for an increased tool life, because of the improved contact between the chip and its cutting edge. Furthermore, if the plan approach angle is too small and chip thickness is thin, this will reduce tool life, however, this can be compensated for by increasing the feedrate, to produce a thicker chip.

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NB  The helical formation of this chip-curling behaviour will shortly be mentioned, but prior to this, chipbreakers/formers will be discussed.

2.5.3 Chip-Breakers and Chip-Formers Chip-breakers have been utilised on turning tools for many years, initially introduced in the 1940’s in the form of an abutment, or step, situated behind the rake face of the tool. Hence, with this type of early chipbreaker, as the continuous chip moves across the rake face it collides with this step and breaks. This original form of chip-breaker geomtery was relatively inefficient as the resultant force direction changed with the programmed tool path, this meant that the step would be approached by the chip from differing directions making chip-breaking less controlled. Such chip-breakers were superseded in the 1970’s by inbuilt ‘wavy-shaped’ chip-breakers sintered into the insert’s top face (Fig 34 bottom left-hand photograph). Recent developments in designing chip-breaker geometries by computer-generated (i.e. CAD) techniques, has shown a significant step-forward in both chipformer design enabling chip control and reduction in frictional forces across the rake face at a range of cutting data to be achieved. Such ‘automatic’ chip breaker geometry forces the chip to deflect at a narrower angle, causing it to break off, either immediately, or just after the free end of the chip has hit either the tool’s flank or, the workpiece before the first coil has formed. If such a collision does not take place, the result would be a smaller diameter spiral chip and, it can be anticipated that the chip would still break, but only when it became slightly longer – this later chip breakage is due to the increasing chip mass and the effect of gravity upon it, with, or without any further collision. Chip flow direction will depend upon several factors, such as the: chip-breaker profile, back rake and setting angles, nose radius, DOC and feedrate – these latter three factors require further discussion. The relationship between the nose radius, DOC and feedrate will often change during vectored tool paths in any machining operation. Even though the insert’s nose radius is preset, its influence on the chip direction differs for different DOC’s, depending on how much corner rounding is represented by the total engaged edge length (Fig. 34c). Further, the feedrate also affects the chip thickness: at different DOC’s and with a constant feedrate, the form of chip cross-section (i.e.

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. Figure 34.  The principles of chip-breaking and chip-breaking envelopes for ‘coma-shaped swarf’ control and insert edge preparations

Turning and Chip-breaking Technology

the ratio of chip width-to-thickness), will change and this has a deleterious effect on the insert’s chip-breaking ability.

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2.5.4 Helical Chip Formation

the insert’s flank face (see Fig. 27b)32. Only today’s very complex chip-breaker designs can reduce these outward-curling helical chips. Although such chip helices produced by combinations of the feeds and DOC’s that result in the chip width being too small in relation to its thickness must be avoided.

Conventional Turning

Grooving and Recessing

For the general turning operations, such as sliding (i.e. Z-axis tool feeding) and facing (i.e. X-axis tool path motions), the chip is rolled into a helix, simply because the chip edges are formed from different rotation radii (Fig. 34d). Here, the two edges of the chip consume different quantities of workpiece material, creating differing edge lengths, coupled to the fact that a variation in cutting speed is present, these relationships will result in a helical chip formation. The appearance of the chip’s helix depends upon the workpiece’s diameter and its metallurgical specification/condition, which means the chip helices are extremely difficult to quantify. Most common types of helical chip diameters are determined either directly by the initial curvature from its origin, or are the result of additional bending, introduced by the chip-breaker. For example, the helical chip type shown in Fig. 34c (left), has its chip segments turned inwards, this being a desirable chip form when not fully developed, that is prior to the first coil being completed. Whether, or not the chip is of this form will already be determined even before it meets the chip-breaker, this being the result of its cross-section and the natural tendency to bend according to the ‘line of least resistance’. If the chips width is no larger than its thickness, for example, the resistance to bending in the segment-stiffened thickness direction is larger than in the width direction. In this case, unless this kind of chip is broken early, by colliding with either part of the tool, or the workpiece whilst it is still stiff and short – called ‘self-breaking’ – a helical chip will be formed. In this case, the barbed, or serrated edge is turned outwards causing additional bending, this being introduced by the chip-breaker. For example, the helical chip type shown in Fig. 34c (right), becomes difficult and awkward to control. This outward-curving helical chip also has weakened sections in the serrations between the chip segments, but applied loads on it are readily absorbed by the spring action of the chip. This type of chip will break as it hits

In conventional turning operations, it is significantly easier to form a manageable chip, than for features requiring either grooving, or recessing. The chip formed during plunge grooving counter-rotates in relation to the workpiece, whereby it does not experience the same twisting force as chips produced by either Z-, or X-axes turning operations. When grooving, ideally the chip resembles a ‘watch spring’ , indicating that the chip is curling back onto itself and will ultimately break in several distinct ways: such as at the completion of the grooving cycle, or due to friction between the chip and its groove side walls – as the chip diameter becomes greater. In grooving operations, three significant factors affect chip control, these are: (i) Insert geometry – applied to the rake face, can be classified into distinct groupings: • Radial-ground top rake (not shown), producing the desired ‘watch-spring’ chip formation. This grooving insert geometry will not thin the chip, therefore surface finish passes are necessary on both groove side walls. NB  For long-chipping materials the chip-former does not provide enough resistance to produce chip curling, hence, a straight flat chip occurs, that may

32 One of the problems with this type of chip-breaking, is the potential for secondary wear on the insert’s non-cutting zone on the face, caused by the chip helix breaking locally against this face. Such an occurrence happens when the chip helix attains such a diameter and pitch that its free-end continually strikes the non-cutting portion of the insert’s edge – termed ‘chip-hammering’ – causing the edge to be locally weakened and to subsequently crumble. NB  Chip-hammering can be alleviated by slightly increasing the helix diameter (i.e. by somewhat modifying the cutting data) causing the chip to break against the tool’s flank – below the insert’s cutting edge, this being one of the previously employed and favoured chip-breaking mechanisms, as shown in Fig. 27b.

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. Figure 35.  The chip-breaking envelopes related to cutting data and chip-curling behaviour. [Courtesy of Sandvik Coromant]

Turning and Chip-breaking Technology

wrap itself around either the tool, or workpiece, but such a geometry is perfect for machining aluminium, or non-ferrous materials.

• Radial top rake (illustrated in Fig. 4 middle and

to the left – three grooving insert sizes illustrated). This radial top rake is designed to thin the chip. Such chip thinning, eliminates the need to undertake finishing passes on the groove’s side walls. Furthermore, this type of grooving insert geometry being on-centre, enables axial turning of diameters for wide shallow grooves33, or recesses. • Raised bumps on top rake (see Fig. 27a – left). This sophisticated grooving geometry is utilised for materials where chip control is difficult, as it provides an ‘aggressive barrier’ to the curling chip. The raised bumps force the chip back onto itself, either producing a tightly curled watch-spring chip, or causes the chip to break. (ii) Surface speed of the workpiece – in order to obtain full advantage of a grooving insert’s chip-forming abilities, the chip must be allowed to flow into the chip-former. This chip-flow can be achieved by either decreasing the workpiece’s surface speed, or increasing the feed – more will be said on this shortly. The former technique of decreasing the surface speed, allows the material to move slower across the top rake of the cutting edge and as a result, has greater contact time to engage the chip-former. This slower workpiece speed, has the benefit of increasing tool life, through lower

33 A groove, or recess, can normally be considered as a straightwalled recessed feature in a workpiece, as illustrated in Fig. 40. Typical applications for grooves are to provide thread relief – usually up to a shoulder – so that a mating nut and its washer can be accurately seated , or for retaining O-rings. As the groove is produced in the workpiece, the tool shears away the material in a radial manner, via X-axis tool motion. The chip formed with insert geometries having a flat top rake, will have an identical width as the tool and can be employed to ‘size’ the component’s width feature. However, this chip action – using such a tool geometry, creates high levels of pressure at the cutting edge as a result of the chip wall friction, which tends to produce a poor machined surface texture on these sidewalls. Grooving with an advanced chip-former insert geometry, reduces the chip width and provides an efficient cutting action, this results in decreasing the cutting edge pressure somewhat. Chip-formers offer longer tool life and improved sidewall finishes with better chip control, than those top-rakes that have not incorporated such insert chip-forming geometric features.

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tool/chip interface temperatures. The negative factors of such a machining strategy, are that the: • Part cycle times are increased and as a result, any batch throughput will be lessened, • As the cutting edge is in contact for a longer duration, more heat will be conducted into the tool, than into the chip, which could have a negative impact of inconsistent workpiece size control, • Due to the lower workpiece surface speed, the benefits of the insert’s coating will be reduced, as such coating technology tends to operate more effectively at higher interface temperatures. (iii) Increasing the feedrate – by increasing the feed allows it to engage the chip-former more effectively – this being the preferred technique for chip control. A heavier applied feedrate, produces a chip with a thicker cross-section. Further, a thicker chip engages the insert’s geometry with higher force, creating a greater tendency to break. Hence, by holding a constant workpiece surface speed, allows the faster feedrate to reduce cycle times.

Transversal, or Face Grooving Transversal grooving geometry has a curved tearshaped blade onto which, the insert is accurately located and positioned. The transversal insert follows the 90° plunged feed into the rotating face of a workpiece. These tools are categorised as either right-, or left-hand, with the style adopted depending upon whether the machine tool’s chuck rotates anti-clockwise (i.e. using a right-hand tool), or clockwise (i.e. left-hand). The minimum radius of curvature for such transversal grooving tooling is normally about 12mm, with no limit necessary on the maximum radial curvature that can be machined. For shallow face grooves, off-the-shelf tooling is available, but for deep angular face grooves they require specialised tools from the tooling manufacturers. If a relatively wide face groove requires machining with respect to the insert’s width, then the key to success here, is establishing where in the face to make the first plunge. This initial face plunge should be made within the range of the tool’s diameter, otherwise the tool will not have sufficient clearance and will ultimately break. Successive plunges to enlarge the face groove should be made by radially moving the insert 0.9 times the insert’s width, for each additional plunge. The rotational speed for face grooving is usually about 80% of the speed used for parting-off – soon to be

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mentioned. Feedrates are normally around 50% of parting-off values, with the proviso that for material which is subject to work-hardening, minimum feeds are necessary. In transversal grooving operations, a unique chip form occurs, because the chip is longer the further away it is from the workpiece’s centre line of rotation. This results in the chip which no longer flows in a straight line across the insert’s edge, instead it moves at an angle. Such a naturally curved chip is difficult to exhaust from the face groove, particularly if it is broken. Hence, no attempt should be made to break the chip. For deep and narrow grooves, the best solution is to retract the tool at short intervals, to check that the blade shows no signs of rubbing, this is to guard against any likely breakage that might occur when machining outside the blade’s range. Due to the fact that transversal grooving tooling is susceptible to chatter34, any excessive overhang of the tool should be minimised. The chip should never be allowed to become entangled within the transversal groove and should be ejected speedily, otherwise the tool is likely to break.

34 Chatter is a form of self-excited vibration and such vibrations are due to the interaction of the dynamics of the chip-removal process, together with the structural dynamics of the machine tool. Such chatter, tends to be at very high amplitude, which can result in either damage to the machine tool, or lead to premature tool failure. Typically, chatter is initiated by a disturbance in the cutting zone, for several reasons, such as: – Lack of homogeneity – in the workpiece material (i.e. typically a porous component, such as is found in a Powder Metallurgy compact), – Workpiece surface condition (i.e. typically a hard oxide scale on a hot-rolled steel component, utilsing a shallow DOC), – Workpiece geometry (i.e. if the component shape produces either a variation in the DOC – for example, because of uneven depth of casting material being machined, or light cuts on interrupted shapes, such as hexagon, square, or rectangular bar stock), – Frictional conditions (i.e. tool/chip interface frictional variations, whilst machining). Regenerative chatter is a type of self-excited vibration, resulting from the tool cutting a workpiece surface that has either significant roughness, or more likely the result of surface disturbances from the previous cut. These disturbances in the workpiece surface, create fluctuations in the cutting forces, with the tool being subjected to vibrations with this process continuously repeating, hence the term ‘regenerative chatter’. Self-excited vibrations can be alleviated by either increasing the dynamic stiffness of the system, or by increasing the damping. NB  Dynamic stiffness can be defined as the ratio of the amplitude of the force to the vibrational amplitude.

For any face grooving of workpiece material that is subject to a continuous chip formation, always use copious amounts of coolant and at high-pressure – if possible, to not only lubricate the cutting zone, but to aid in chip flushing from this groove.

Parting-off The parting-off process is normally considered to be a separate machining operation, but it simply consists of cutting a groove to centre of rotation of the workpiece, to release it from the bar stock, or to ‘part-off ’ to a previously formed internal diameter (shown in Fig. 40 for left-hand side operations). Essentially in a parting-off operation, two time-periods are worthy of mention, these are: (i) At separation from the bar stock – a lower spindle speed than was previously used on the workpiece, will prevent the ‘released part’ from hitting the machine and potentially damaging its surface. Moreover, it allows an operator – if present – to hear the change in the lower spindle speed tone, as it is about to separate from the bar stock, avoiding the parting-off tool from getting ‘pinched’ between the stock and the soonto-be-released component. Often, ‘Part-catchers’ are utilised to reduce any surface damage to the falling component, once it has been parted-off. NB  If the component to be parted-off is held in a coaxial/sub-spindle, at component release, the additional spindle supports the workpiece and under these conditions, the parting-off operation is virtually identical to that of found in a grooving cycle. (ii) Surface speed reduction – this effectively occurs when the machine’s spindle attains its maximum speed. For example, on a machine tool having a maximum speed of 3,000 rpm, 90 m min–1 would only be achievable until the parting diameter has reached about 8.6 mm. When parting to a smaller diameter than 8.6 mm, the surface speed would decrease at a fixed spindle speed. As the parting diameter reaches 5.8 mm the surface speed would be 55 m min–1, or 60% of the ideal, thus significantly increasing the chip loading as the tool approaches the workpiece’s centreline. In order to alleviate the increasing tool loading, lowering the feedrate by about 50% until separation is just about to occur, then finally dropping the surface speed to almost zero at this point, reduces the tendency for a ‘pip’ to be present on the workpiece. On a CNC driven spindle, it is not advisable for parting-off operations,

Turning and Chip-breaking Technology

to utilise the ‘canned cycle’ such as the ‘constant surface speed’  35 function. NB  A more serious parting-off problem has been that in order to eliminate the pip formed at the centre of the ‘released component’ , some tools have been ground with the front edge angle of between 3° to 15°. Such a front edge geometry, can introduce an axial cutting force component, leading to poor chip control, which in turn, causes the tool to deflect. This parting-off tool deflection, can lead to the component’s face ‘dishing’ , creating a convex surface on one face and a concave surface on the other – so this tool grinding strategy should be avoided. Today, parting-off inserts normally consist of two main types with top rakes that are either of, negative, or positive cutting edge chip-forming geometries. The negative-style of chip-formers are possibly the most commonly utilised. These inserts have a small negative land at the front edge which increases the insert’s strength, giving protection in adverse cutting conditions, such as when interrupted cutting is necessary during a parting-off operation. The land width – often termed a ‘T-land’ , is relative to the breadth of the parting-off tool. This width of the insert’s land has a direct correlation to the feedrate and its accompanying chip formation. The feedrate must be adequate to force the workpiece material over the land and into the chipformer36. Notwithstanding the widespread usage of negative parting-off tooling, positive-style insert geometries have some distinct advantages. The chief one being the ability to narrow the chip at light feedrates, with mini-

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mal tool pressure. If excessive tool pressure occurs, this can promote work-hardening of the ‘transient surface’ 37 of the workpiece. These abilities are important points when machining relatively low mechanical strength components, which might otherwise buckle if machined with negative-style inserts when subsequently parted-off. Positive cutting edge parting-off tooling having chip-formers, are ideal for applications on machine tools when either low fixed feedrates are utilised, or if the workpiece material necessitates lower cutting speeds. This positive-style of parting-off tooling, operates efficiently when machining softer workpiece materials, such as: aluminium-or, cooper-based alloys and many non-metallic materials, typically plastics. Feed­ rates can be very low with these positive-type parting tools, down to 0.0254 mm rev–1 with exceptional chip control and consistent tool life. One major disadvantage of using these positive tooling geometries for parting-off, is that the tool is much weaker than its equivalent negative geometry type. The concept of insert self-grip in its respective toolholder, was developed by the cutting tool manufacturer Iscar tools in the early 1970’s and has now been adopted by many other tooling manufacturers (Fig. 40 top left-hand side). These ‘self-grip’ tooling designs, rely on the rotation of the part and subsequent tool pressure to keep the ‘keyed and wedged’ insert seated in its respective toolholder pocket. Previously, doubleended inserts termed ‘dogbones’ , were often used but were limited to low DOC’s – due to the length of the secondary cutting edge, so have been somewhat overshadowed by the ‘self-grip’ varieties of parting-off tooling.

2.5.5 Chip Morphology 35 ‘Constant surface speed’ CNC capability as its name implies, allows the machine tool to maintain a constant surface speed as the diameter is reduced. The main problem with using this ‘canned cycle’ , is that as the maximum spindle speed is reached, the chip load will also increase. This is not a problem, so long as the maximum speed has not occurred, such as when parting-off a component with a large hole at its centre. 36 Parting-off operations that employ a negative-style insert (i.e. with a land and accompanying chip-former), normally have the feedrate determined in the following manner: by multiplying the width of the insert by a constant of 0.04. For example, for a 4 mm wide tool, it is necessary to multiply the insert’s width of 4 mm by 0.04 to obtain a feedrate of 0.16 mm rev–1. This will give a ‘start-point’ for any parting-off operations, although it might be prudent to check this feedrate is valid, from the tooling manufacturer’s recommendations.

The Characterisation of Chip Forms (Appendix 2) In the now withdrawn ISO 3685 Standard on Machinability Testing Assessment, of some interest was the fact that this Standard had visually characterised

37 Transient surfaces are those machined surfaces that will be removed upon the next revolution of either the: – Workpiece (i.e in rotating part operations), or – Cutter (i.e. for rotating tooling – drilling, milling, reaming, etc.).

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chip forms under eight headings, with several variations appearing in each groups (i.e. see Appendix 2 for an extract showing these chip form classifications). Although in the main, the chip forms were related to turning, some of these chip morphologies could be extrapolated to other manufacturing processes. The chip type that will be formed when any machining operation is undertaken is the product of many interrelated factors, such as: • Workpiece material characteristics – will the material that forms the chip significantly work-harden?, • Cutting tool geometry – changing, or modifying the cutting insert geometries38 and its plan approach angles will have a major influence on the type of chip formed, • Temperatures within the cutting zone – if high, or low temperatures occur as the chip is formed, this will have an impact on the type of chip formed, • Machine tool/workpiece/cutting tool set-up – if this ‘loop’ is not too rigid, then vibrations are likely to be present, which will destabilise the cutting process and affect the type and formation of chips produced, • Cutting data utilised – by modifying the cutting data: feeds and speeds and DOC’s, with the insert geometry maintained, this can play a significant role in the chip formed during machining operations. NB  Chip formation has become a technology in its own right, which has shown significant development over the last few decades of machining applications. As has been previously mentioned, chip formation should always be controlled, with the resultant chips formed being broken into suitable shape formation, such as ‘spirals and commas’ , as indicated by the resultant chip morphology shown in Fig. 35a. Uncontrolled chip-steaming (i.e. long continuous workpiece strands), must be avoided, being a significant risk-factor to both the: machine tool’s operation and its CNC setter/operator alike.

38 Chip-breaking envelopes (see Fig. 34 middle right), are the product of plotting both the feedrate and DOC on two axes, with their relative size and position within the graphical area being significantly affected by the cutting insert’s geometry – as depicted by the three cutting insert geometric versions shown by types: A, B and C (Fig. 34).

For every cutting insert geometry, there is a recommended application area – termed its ‘chip-breaking envelope’ (i.e. see footnote 38 below) – with regard to its range of feedrates and DOC’s. Within this ‘envelope’ , chips of acceptable form are produced by the cutting insert’s geometry. Conversely, any chips that are formed outside this ‘envelope’ are not acceptable, because they are either formed as unbroken strands, or are too thick and over-compressed. When component profiling operations are necessary (Fig. 31a), this normally involves several machining-related parameters: variations in DOC’s, together with path vectoring of the feeds and as a result of this latter point, changes to the resultant chip’s path on the rake face. These factors are important as they can modify the chip morphology when profiling operations include: recessed/undercut shoulders, tapers and partial arcs, facing and sliding operations with the same tool, together with many other combined profiled features. All of these operations make significant demands on the adaptability of the cutting insert’s geometry to efficiently break the chip. In general, the cutting insert’s chip formation principles are concerned with the chip-breaker’s ability to create a chip form that is neither not too tight a curl, nor too open. If chip curling is too tight for the specific machining application, the likely consequences are for a chip form creating: • ‘Chip-streaming’ – producing long chip strands that are undesirable, wrapping itself around the machined surface of the workpiece with workhardened swarf and possibly degrading this machined surface, or may become entangled around the various parts of the machine tool, which could impede its operation, • Excessive heat generation – this can decrease tool life, or be conducted into the machined part and consequently may affect specific part tolerances for the individual part, or could lead to modifications in the statistical variability39 of a batch of parts,

39 Statistical variability in component production can cause variations from one part to another, as the standard deviation and mean changes, these important factors will be mentioned later in the text.

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• Increased built-up edge (BUE) formation – which having a small angle when compared to the cutting through ‘attrition wear’ 40 may cause the risk of premature cutting edge failure.

When the chip curling is too open, this may result in the following negative tendencies: • Poor chip control – creating an inefficient chipbreaking ability by the cutting insert, • Chip hammering – breaking down the edge and causing it to crumble and as a result creating the likelihood of prematurely failing, • Vibrational tendencies – affecting both the machined surface texture and shortening tool life. Chip formation and its resultant morphology, is not only affected by the cutting data selected, but will be influenced by the plan approach (i.e. entering) angle of the insert. In most machining operations, they are usually not of the orthogonal, but oblique cutting insert orientation, so the affect is for the entering angle to modify the chip formation process. The insert’s entering angle affects the chip formation by reducing the chip thickness and having its width increased with a smaller angle. With oblique cutting geometry, the chip formation is both ‘smoother and softer’ in operation as the plan approach angle tends toward say, 10° to 60°, furthermore, the chip flow direction will also advantageously change with the spiral pitch increasing. As the nose radius is changed with different cutting inserts, this has the effect of changing both the direction and shape of the chips produced. This nose radius geometry is a fundamental aspect in the development of chips during the machining process – as depicted by Fig. 35b. Here, an identical nose radius and feedrate is utilised, but the difference being the DOC’s, with a shallow DOC in Fig. 35b (left), giving rise to a slow chip helix, whereas in Fig. 35b (right) the DOC is somewhat deeper, creating a tighter chip helix which is beneficial to enhanced chip-breaking ability. Shallow cutting depths produce ‘comma-shaped’ chip cross-sections,

40 Attrition wear is an unusual aspect of tool wear, in that it is the result of high cutting forces, sterile surfaces, together with chip/tool affinity, creating ‘ideal’ conditions for a pressure welding situation. Hence, the BUE develops, which builds-up rapidly and is the ‘swept away’ by the chip flow streaming over the top rake’s surface, taking with it minute atomic surface layers from the tool’s face. This continuous renewal and destruction of the BUE, enhances crater wear formation, eventually leading to premature cutting edge failure.

edge. Equally, a larger depth means that the nose radius has somewhat less affect from its radius and greater influence by the entering angle of the cutting edge, producing an outward directed spiral. Feedrate also affects the width of the chip’s cross-section and its ensuing chip flow41. Chip formation begins by the chip curving, this being significantly affected by combinations of the cutting data employed, most notably: feedrate, DOC, rake angle, nose radius dimensions and workpiece condition. A relatively ‘square’ cross-sectional chip normally indicates that an excessively hard chip compression has occurred, whilst a wide and thin band-like chip formation is usually indicative of long ribbon-like chips producing unmanageable swarf. If the chip curve is tight helix, coupled to a thick chip cross-section, this means that the length of the chip/tool contact has increased, creating higher pressure and deformation. It should be noted that excessive chip cross-sectional thickness, has a debilitating effect on any machining process. By careful use of CAD techniques coupled to FEA to construct the insert’s cutting edge, commashaped chips are the likely product of any machining, providing that the appropriate cutting data has been selected. In some machining operations, chip formation can be superior using a slightly negative insert rake angle, thereby introducing harder chip compression and self-breaking of the chip, particularly if utilising small feeds. Conversely, positive rakes can be give other important machining advantages, depending which chip form and cutting data would be the most advantageous to the part’s ensuing manufacture. Usu-

41 Chip-flow is the result of a compound angle between the chip’s side- and back-flow. The chip’s side-flow being a measure of the flow over the tool face (i.e. for a flat-faced tool), whilst backflow establishes the amount of chip-streaming into the chipbreaker groove. Detailed analysis of chip side-flow (i.e. via high-speed photography), has indicated that it is influenced by a combination of groove dimensions and cutting data. If the feedrate is increased, this results in a higher chip backflow angle, promoting chip-streaming into the chip-breaker groove. The ratio of feed-to-length of restricted contact has been shown to be an important parameter in the determination of chip- back-flow. Typically with low feedrates the corresponding chip back-flow is going to be somewhat lessened, resulting in poor chip-breaker utilisation. When the restricted contact between the chip and the tool is small – due to low feed – the chip-flow does not fully engage the chip-breaker and will as a result curve upward, with minimal ‘automatic’ chip-breaking effect.

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ally, for larger feedrates, a positive insert rake angle might optimise the chip-curving tendency, by not producing and excessively tight chip helix. Chip curve, its resultant chip flow direction, the chip helix and its accompanying shape are designed into each cutting edge by the tooling manufacturers. Tool companies ensure that a controlled chip formation should result if they are exploited within the recommended cutting data ranges specified. In Fig. 36a (left), effective chip-breaking decisionmaking recommendations are shown on a flow-chart,

indicating how to obtain the desired chip-breaking control. In the chart shown in Fig. 36a (right), the DOC’s indicate on the associated visual table the expected chip type showing that here types ‘C and D’ offer ‘good’ broken chips. Such chip morphology charts as these from tooling manufacturers, attempt to inform the user of the anticipated chip-breaking if their recommendations are followed. Whereas the flow-diagram illustrated in Fig. 36b, indicates that ‘good chip control’ improved productivity will result, if a manufacturing company adopts the machining

. Figure 36.  Chip-breaking control and chip morphology and its affect on productivity. [Courtesy of Mitsubishi Carbide]

Turning and Chip-breaking Technology

strategy high-lighted to the left-hand side. On the contrary, ‘poor chip control’ with an attendant decrease in productivity will occur, if the problems shown to the right-hand side transpire. Chip morphology can indicate important aspects of the overall cutting process, from the cutting edge’s geometry and its design, through to work-hardening ability of the workpiece. Many other factors concerning cutting edge’s mechanical/physical properties can be high-lighted, these being important aids in determining a material’s machinability – which will be discussed in more depth later in the text.

2.5.6 Chip-Breaker Wear Any form of tool failure will depend upon a combination of different wear criteria, usually with one, or more wear mechanisms playing a dominant role. Previously, it was found that the workpiece surface texture and the crater index act as appropriate tool failure criteria, particularly for rough turning operations. Moreover, tool life based upon these two factors, approximated the failure curve more exactly than either the flank, or crater wear criterion. In cutting tool research activities, it has been found that when machining with chip-breaker inserts, flank wear (i.e. notably VB) is not the most dominant factor in tool failure. In most cases, the ‘end-point’ of useful tool life occurs through an alteration of the chipgroove parameters, well before high values of flank wear have been reached. The two principal causes of wear failure for chip-breaker inserts are: • For recommended cutting data with a specific insert, the design and positioning of chip-breakers/ grooves may promote ‘unfavourable’ chip-flow, resulting in wear in the chip-breaker wall – causing consequent tool failure, • Alterations in the cutting data, particularly feedrate, affects chip-flow, which in turn, generates various wear patterns at the chip-breaker’s heel and edge (see Fig. 37). In the schematic diagrams shown in Fig. 37, are illustrated the concentrated wear zones on the: back wall (i.e. heel), cutting edge, or on both positions for a typical chip-breaker insert. Under the machining conditions for Fig. 37a, the chip-groove utilisation is very low, with the chip striking the heel directly. Thus, as machining continues, this results in abrasive

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wear of the heel and ultimately this heel becomes flattened and chip-breaking is severely compromised. Conversely, when the cutting data produces a wear zone concentrated at the insert’s edge (Fig. 37b), then chip side-flow occurs and poor chip-breaking results, together with low tool life. This accelerated tool wear, resulting from an extended tool/chip contact region over the primary rake face, promotes a rough surface texture to the machined part. In the case of Fig. 37c, these are ideal conditions for optimum chipbreaking action and a correspondingly excellent and predictable tool life, because the wear zones at both the heel and edge are relatively uniform in nature, illustrating virtually a perfect chip-forming/-breaking action. Higher tool/chip interface temperatures can result as the heel wears, forming a crater at the bottom of the chip-breaker groove. Combination wear – as shown in Fig. 37c – generally results in significantly improved tool wear, in conjunction with more predictable tool life. In the photographs of chip-breaker grooves shown for an uncoated and coated Cermet cutting insert material in Figs. 38a and b respectively, the relative wear patterns can clearly be discerned. In the case of Fig. 38a – the uncoated insert – the predominant wear concentration is primarily at the edge, indicating that the cutting data had not been optimised. While in the case of the coated Cermet insert of identical geometry (Fig. 38b), the wear is uniform across the: edge, groove and heel. This would seem to suggest that ideal cutting data had been utilised in its machining operation. In both of these cases some flank wear has occurred, but this would not render the chip-breaking ability when subsequent machining invalid. NB  A complex matrix occurs (i.e. Fig. 38c) with Cermets, this ‘metallurgy’ can be ‘tailored’ to meet the needs of specific workpiece and machining requirements.

2.6 Multi-Functional Tooling The concept of multi-functional tooling was developed from the mid-1980’s, when multi-directional tooling emerged. This tooling allowed a series of operations to be performed by a single tool, rather than many, typically allowing: side-turning, profiling and

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. Figure 37.  Schematic representations of differing chip-breaking insert tool wear mechanisms – due to alterations in the cutting data. [Source: Jawahir et al., 1995]

Turning and Chip-breaking Technology

. Figure 38.  Improved wear resistance obtained with an uncoated and coated cermet, when turning ovako 825B steel, having the following cutting data: Cutting speed 250 m min–1, feed 0.2 mm rev–1, DoC 1.0 mm and cut dry. [Courtesy of Sandvik Coromant]

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grooving, enabling the non-productive elements42 in the machining cycle to be minimised. In the original multi-directional tooling concept, the top rake geometry might include a three-dimensional chip-former, comprising of an elevated central rib, with negative Klands on the edges. Such a top rake profile geometry could be utilised for efficient chip-forming/-breaking of the resultant chips. This tooling when utilised for say, grooving operations, employed a chip-forming geometry – this being extended to the cutting edge, which both narrowed and curled the emerging chip to the desired shape, thereby facilitating easy swarf evacuation. A feature of this cutting insert concept, was a form of effective chip management, extending the insert’s life significantly, thus equally ensuring that adequate chip-flow and rapid swarf evacuation would have taken place. When one of these multi-directional tools was required to commence a side-turning operation, the axial force component43 acting on the insert caused it to elastically deflect at the front region of the toolholder. This tool deflection enabled an efficient feed motion along the workpiece to take place, because of the elastic behaviour of the toolholder created a positive plan approach angle in combination with a front clearance angle – see Fig. 39a and b (i.e. illustrating in this one of the latest ‘twisted geometry’ insert multi-functional tooling geometries). Any of today’s multi-functional tooling designs (Figs. 39 and 40), allow a ‘some degree’ of elastic behaviour in the toolholder, enabling satisfactory tool vectoring to occur, either to the right-, or left-hand of the part feature being machined. These multi-func-

42 Non-productive elements are any activity in the machining cycle that is not ‘adding value’ to the operation, such as: toolchanging either by the tool turret’s rotation, or by manually changing tools, adjusting tool-offsets (i.e. for either: tool wear compensation, or for inputting new tool offsets – into the machine tool’s CNC controller), for component loading/unloading operations, measuring critical dimensional features – by either touch-trigger probes, non-contact measurement, or manual inspection with metrology equipment (i.e. micrometers, vernier calipers, etc.), plus any other additional ‘idle-time’ activities. 43 An Axial force component is the result of engaging the desired feedrate, to produce features, such as: a diameter, taper, profile, wide groove, chamfer, undercut, etc. – either positioned externally/internally for the necessary production of the machined part.

tional tools are critically-designed so that for a specific feedrate, the rate of elastic deflection is both known and is relatively small, being directly related to the applied axial force, in association with the selected DOC’s. At the tool-setting stage of the overall machining cycle, compensation(s) are undertaken to allow for minute changes in the machined diameter, due to the dynamic elastic behaviour of one of these tools in-cut. For a specific multi-functional tool supplied by the tooling manufacturer, its actual tool compensation factor(s) will be available from the manufacturer’s user-manual for the product. In-action these multi-functional tools (Fig. 39b), can significantly reduce the normal tooling inventory, for example, on average such tools can replace three conventional ones, with the twin benefit of a major cycle-time reduction (i.e. for the reasons previously mentioned) of between 30 to 60% – depending upon the complexity of features on the component being machined. Some other important benefits of using a multi-functional tooling strategy are: • Surface quality and accuracy improvements – due to the profile of the insert’s geometry, any ‘machined cusps’ 44, or feedmarks are reduced, providing excellent machined surface texture and predictable dimensional control, • Turret utilisation improved – because fewer tools are need in the turret pockets, hence ‘sister tooling’ can be adopted, thereby further improving any untended operational performance, • Superior chip control – breaks the chips into manageable swarf, thus minimising ‘birds nests’ 45 and entanglements around components and lessens automatic part loading problems, • Improved insert strength – allows machining at significantly greater DOC’s to that of conventional in-

44 ‘Machined cusps’ the consequence of the insert’s nose geometry coupled to the feedrate, these being superimposed onto the machined surface, once the tool has passed over this surface. 45 ‘Birds nests’ are the rotational entanglement and pile-up of continuous chips at the bottom of both trough and blind holes, this work-hardened swarf can cause avoidable damage in the machined hole, furthermore, it can present problems in coolant delivery for additional machining operations that may be required.

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. Figure 39.  Multi-functional cutting insert geometry for efficient stock removal and increased part productivity. [Courtesy of Iscar Tools]

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. Figure 40.  Multi-functional tooling for the machining of rotational features such as: turning, grooving and profiling operations – having excellent chip control. [Courtesy of Sandvik Coromant]

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. Figure 41.  By employing twin turrets on a mill/turn centre, ‘balanced turning’ of the component can remove large volumes of stock at one pass. [Courtesy of DMG (UK) Ltd.]

serts, with improved insert security its toolholder location, • CNC programming simplified – many tooling manufacturer’s have specially-prepared software to significantly reduce CNC programming input times. NB  This latter point utilises CNC ‘canned cycles’ to reduce program lengths. In Appendix 3, a guide for ‘Trouble-shooting for turning operations’ are listed, with possible causes and remedies to potential production problems. In the following chapters, many other important chip-forming production processes will be discussed, with hole-making techniques such as drilling being around 25% of all manufacturing techniques undertaken by machining-related companies – these and associated hole-production methods will be reviewed next.

References Journal and Conference Papers Boston, O.W. A Research into the Elements of Metal Cutting Trans ASME 48, 749–848, 1926. Cocquilhat, M. Experiences sur la Résistance utile Produite dans le Forage Ann. Trav. Publ.en Belgique 10, 199, 1851. Doi, S. and Kato, S. Chatter Vibration of Lathe Tools, Trans. of ASME, Vol. 78 (5), 1127–1134, 1956. Fabry, D. The Tool Channel. Cutting Tool Eng’g, 58–64, Sept. 2003. Gadzinski, M. Parting Know-how. Cutting Tool Eng’g, 52–57, March 2001. Galloway, D.F. Some Experiemnts on the Influence of Various Factors on Drill Performance. Trans. of ASME, 191–231, Feb., 1957. Humphries, J.R. Energing Technologies and Recent Advances in Multi-functional Groove/Turn Systems. Int.

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Conf. on Industrial Tooling, Shirley Press Ltd, 65–85, Sept. 1979. Isakov, E. The Mathematics of Machining. American Machinist, 37–39, Aug. 1996. Isakov, E. Reassessing Power Factors. American Machinist, 43–45, Dec. 1996. Jawahir, I.S. The Tool Restricted Contact Effect as a Major Influencing Factor in Chip-breaking: An Experimental Analysis. Annals of the CIRP, Vol. 31 (1), 121–126, 1988. Jawahir, I.S., Ghosh, R., Fang, X.D. and Li, P.X. An Investigation of the Effects of Chip-flow on Tool Wear in Machining with Complex Grooved Tools. WEAR, Vol.184 (2), 145–154, 1995. Kasahara, N., Sato, H. and Tani, Y. Phase Characteristics of Self-excited Chatter in Cutting. J. of Engg for Ind., ASME Pub, Vol. 114, 393–399, Nov. 1992. Kennedy, B. Facing Facts. Cutting Tool Eng’g, 29–37, Feb. 2002. Kennedy, B. Take a Bigger Bite. Cutting Tool Eng’g, 25–29, Aug. 2003. King, K. Added Functionality. Cutting Tool Eng’g, 52–55, Feb. 2005. Kondo, Y., Kawano, O. and Sato, H. Behaviour of Self-excited Chatter due to Multiple Regenerative Effect. J. of Engg. For Ind., ASME Pub., Vol. 103 (3), 324–329, 1981. Lewis, B. Turn your Wipers on. Cutting Tool Eng’g, 47–51, Jan. 2003. Mallock, A. The Action of Cutting Tools. Proc. of Royal Soc. Lond. 33, 127–139, 1881–882. Paterson, H. Strictly Boring. Cutting Tool Engg., 22–30, Oct. 1995. Pekelharing, A.J. Built-up Edge (BUE): is the Mechanism Understood? Annals of the CIRP, Vol. 23 (3) 207–211, 1974. Piispanen, V. Eripanines Teknilliseslä Aikakauslehdeslä 27, 315, 1937. Reuleaux, F. Uber den Taylor Whiteschen Werkzengstahl in Verein zur Beförderung des Gewerbefleisses in Preussen. Sitzzungsberichte, 79, 179, 1900. Smith, G.T. Fundamentals of Chip-breaking for Continuous Cutting Operations. Int. Conf. on Industrial Tooling, Molyneux Press Ltd, 72–82, Sept. 1999. Teets, B. Facing up to Grooving Problems. Machinery and Prod. Eng’g., 51–52, Oct. 1988.

Time, I. Soprotivlenie Metallov I Dereva Rezaniju St. Petersburg, 1870. Tipnis, V.A. and Joseph, R.A. Testing for Machinability, in: Influence of Metallurgy on Machinability, ASM Pub. 11–30, 1975. Tresca, H. Mémoire sur le Rabotage des Métaux Bull. Soc. d’Encourgement pour I’Industrie Nationale 585–685, 1873. Venkatesh, V.C. and Satchidanandam, M. A Discussion on Tool Life Criteria and Total Failure Causes. Annals of the CIRP, Vol. 29 (1), 19–22, 1980. Watson, D.W. and Murphy, D.C. The Effect of Machining on Surface Integrity. Metallurgist and Matls, 199–204, April 1979. Webzell, S. Wiping away Cycle Times. Metalworking Prod., Oct. 2003. Books, booklets and guides Armarego, E.J.A. and Brown, R.H. The Machining of Metals. Prentice-Hall Pub., 1969. Boothroyd, G. and Knight, W.A. Fundamentals of Metal Machining and Machine Tools. Marcel Dekker (NY), 1989. Finish Turning – Application Guide. AB Sandvik Coromant Pub., 1995. Hartig, E. Versuche über Leistung und Arbeitsverbrauch der Werkzengmaschine. 1873. Kaczmarek, J. Principles of Machining by Cutting Abrasion and Erosion. Peter Pregrinus Pub. (Warsaw), 1976. Modern Metal Cutting – A Practical Handbook. AB Sandvik Coromant Pub., 1994. Shaw, M.C. Metal Cutting Principles. Clarendon Press, Oxford, 1984. Smith, G.T. Advanced Machining – The Handbook of Cutting Technology. IFS/Springer Verlag, 1989. Smith G.T. CNC Machining Technology. Springer Verlag, 1993. Smith, G.T. Industrial Metrology – Surfaces and Roundness. Springer Verlag, 2002. Stainless Steel Turning, AB Sandvik Coromant Pub., 1996. Tlusty, G. Manufacturing Processes and Equipment. Prentice Hall, 2000. Trent, E.M. Metal Cutting. Oxford: Butterworth Heinemann (3rd Ed.), 1991.



3

Drilling and Associated Technologies ‘In all things, success depends upon previous preparation and without such preparation……there is sure to be failure.’

CONFUCIUS

(c550–c487BC) [Analects]

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3.1 Drilling Technology 3.1.1 Introduction to the Twist Drill’s Development Drilling operations are perhaps the most popular machining process being undertaken today, with their origins being traced back to cutting tool developments in North America in the 19th century. In 1864 toward the latter part of the American Civil War, Steven Morse (i.e. later to design the significant ‘Morse taper’ – for accurate location of the ‘sleeved drills’ into their mating machine tool spindles) founded the Morse Twist Drill and Machine Company in the ‘North’. Morse then proceeded to develop probably the most important cutting tool advance to date, namely, the ubiquitous twist drill. In Fig. 42, several of today’s twist drills are illustrated along with just a small range of ‘solid’ contemporary designs. Morse’s originally-designed twist drill has changed very little over the last 150 years – since its conception. In comparison to the somewhat cruder-designed contemporary drills of that time, Morse stated: ‘The common drill scrapes metal to be drilled, while mine cuts the metal and discharges the chips and borings without clogging’. Morse’s statement was at best, to some extent optimistic, whereas the ‘cold reality’ tells a different story, as a drill’s performance is influenced by a considerable number of factors, most of which are listed in Fig. 43.

3.1.2 Twist Drill Fundamentals The basic construction of a conventional twist drill is depicted in Fig. 44a. From this illustration two distinct cutting regions can be established: firstly, the main cutting edge, or lips; secondly at the intersection of the clearance and main cutting edge – termed the chisel edge. In fact for a twist drill, the cutting process can be equated to that of a left-hand oblique turning tool, where the rake and clearance face geometries are identical and the correlation between these two machining processes have been validated in the experimental work by Witte in 1982. Both of these regions remove material, with the cutting lips producing efficient material removal, while the chisel edge’s contribution is both inefficient and is mainly responsible for geometric errors in drilling, coupled to high thrust loads.

The main cutting edges are accountable for a relatively conventional chip formation, as shown in the ‘quick-stop’ photomicrograph in Fig. 44b. An oblique cutting action occurs to the direction of motion, being the result of an offset of the lips that are parallel to a radial line – ahead of centre – which is approximately equal to half the drill point’s web thickness and increases toward the centre of the drill. This obliquity is responsible for inducing chip flow in a direction normal to the lips in accordance with Stabler’s Law. The increasing chip flow obliquity can be seen in Fig. 45a, by observing the flow lines emanating from the chip’s interface along the lips and up the flute face. Such an oblique cutting action serves to increase the twist drill’s effective rake angle geometry. With the advent of ‘Spherical trigonometric computer software’ for obtaining direct three-dimensional calculations – previously described by Witte (1982) in two-dimensional formulae for cutting edge performance – these calculations have been enhanced. Under the chisel point, or web, the material removal mechanism is quite complex. Near the bottom of the flutes where the radii intersect with the chisel edge, the drill’s clearance surfaces form a cutting rake surface that is highly negative in nature. As the centre of the drill is approached, the drill’s action resembles that of a ‘blunt wedge-shaped indentor’ , as illustrated in Fig. 45b. An indication of the inefficient material removal process is evident by the severe workpiece deformation occurring under the chisel point, where such deformed products must be ejected by the drill to produce the hole. These ‘products’ are extruded, then wiped into the drill flute whereupon they intermingle with the main cutting edge chips. This fact has been substantiated by force and energy analysis, based on a combination of cutting and extruding behaviour under the chisel point, where agreement has been confirmed with experimental torque and thrust measurements. The chisel edge in a conventionally ground twist drill has no ‘true’ point, which is one of the major sources for a drilled hole’s dimensional inaccuracy.



Stabler’s Law – for oblique cutting, can be formulated, as below: Chip flow (cos η) = cos I (bc/b) Where: I = inclination of cutting edge, bc = chip flow vector, b = direction of cutting vector.

Drilling and Associated Technologies

The conventional twist drill chisel point geometry can be seen in Fig. 46, together with associated nomenclature for critical features and tolerance boundaries. From the relatively complex geometry and dimensional characteristics shown in Fig. 46, the ob-

89

tainable accuracy of holes generated whilst drilling is dependent upon grinding the drill to certain limits. Any variations in geometry and dimensions, such as: dissimilar lips and angles, chisel point not centralised, and so on, have a profound effect on both the hole di-

. Figure 42.  A selection of just some of the many ‘solid’ and ‘through-spindle’ drilling varieties and ‘inserted-edge’ insert geometries currently available. [Courtesy of Seco Tools]

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Chapter 3

. Figure 43.  The principal technical drill performance criteria and factors associated with drilling operations in this case for example, on castings

Drilling and Associated Technologies

. Figure 44.  The twist drill geometry and associated chip shearing mechanism. [Source: C.J. Oxford Jr., 1955]

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Chapter 3

. Figure 45.  The twist drill shearing and extrusion mechanism at the bottom of a hole. [Source: C.J. Oxford Jr., 1955]

Drilling and Associated Technologies

. Figure 46.  Twist drill geometry

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mensional accuracy and roundness, with some ‘helical wandering’  as the drill passes through the workpiece. Hole accuracy and in particular the ‘bell-mouthing effect’ , is minimised by previously centre-drilling prior to drilling to ‘size’. The main cause of such this ‘bellmouthing’ is probably the inconsistency in the drill geometry. Such effects are exacerbated using Jobber drills, or even worse, by utilising longer-series drills, which tend to either slightly ‘unwind’ , or bend as a result of lessening rigidity promoting some drill bending/deflection. It is worth noting that the rigidity of a tool such as a drill will decrease by the ‘square of the distance’. Therefore it follows that the greater the drill penetration into the workpiece, the progressively larger the deflection and, the further from the ‘true axis of rotation’ will be the subsequent drill’s path. This deflected drilled hole slope angle ‘ϕ’ , can be defined in the following manner: Drilled hole slope angle ‘φ’ = 3/2 l × R/T (1 – I/k × tan k l) Where: l = length of deflected tool,

R = ratio of the transverse reaction at the drill point, T = thrust force, I = system’s ‘moment of inertia’ , k = √T/E I. As suggested above, this ‘axis slope error’ is initiated when the chisel edge begins to penetrate the workpiece and unless the feed is discontinued, or in some manner the error is corrected, the magnitude of deflection will increase as drill penetration continues. The drill’s magnitude of deflection can reach up to 60 µm, under exaggerated drilling conditions. The geometry of the point has been the subject of considerable research and development for many years, with some unusual departures from the ‘standard’ 118° drill point included angle. Typical of these extreme approaches were the so-called ‘Volvo point’ , having a negative 185° included angle – primarily utilised to avoid ‘frittering’  of drilled holes, or the highly positive geometries such as 80° included angle used for drilling some plastics. Not only can the point angle be modified, but the shape and profile of the chisel point, or web offers numerously-ground opportunities for detailed geometric modifications, with only some of which being shown in Fig. 47. Four of the most commonly-ground drill point geometries being: • Conventional – the ‘original’ Morse geometry, having a straight chisel edge, with poor self-centring drilling action (Fig. 46a), • Split-point – there are a range of point-splitting techniques available to alter the point profile, which has the effect of modifying the chisel point to allow a reasonable self-centring action (Fig. 47b),



‘Helical wandering’ is the result of the drill’s geometry being ‘unbalanced’ , resulting from of differing lip lengths, or an offset chisel point, causing the drill to ‘spiral-down’ through the workpiece, as it progresses through the part (see Fig. 70). ‘Bell-mouthing’ of the drilled hole is attributable to the chisel point and is produced by the line-of-contact, as the drill point initially touches the component’s surface, causing it to ‘walk’ until the feed/penetration stabilises itself at the outer corners (i.e. margins) entering the workpiece, whereupon, these margins guide the drill into the part.



‘Bell-mouthing effect’ is produced by the drill chisel point’s eccentric behaviour as it attempts to centralise its rotational motion as it enters, or exit’s the workpiece.





‘Jobber drills’ are considered to be ‘standardised drills’ that are normally utilised for most drilling general operations, unless otherwise specified.

‘Frittering’ refers to the break-out at the hole’s edge as the drill exit’s the part, on some brittle materials, such as on several Powder Metallurgy compacts.





‘Rigidity rule’: a drill, reamer, tap, or a milling cutter held in a spindle will have its rigidity decreased by the ‘square of the distance’ , namely, if a drill is twice as long it is four times less rigid.

‘Web’ refers to the internal core of the drill – which imparts mechanical strength to the drill. The web increases in thickness the further one gets from the chisel edge (i.e. shown in Fig. 47 – in lower diagrams and with cross-sections). Hence, if the drill is reground many times, the chisel point width will obviously increase, this necessitates that the chisel point must be ‘thinned’ , otherwise too high a thrust force occurs and an inefficient drilling action will result.



‘Split-point’ ground drills are sometimes referred to as ‘Multifacet drills’.

NB  A cantilevered tool such as a boring bar has its rigidity decreased by the ‘cube’ or the distance – meaning that too much tool overhang, will seriously reduce tooling rigidity.

Drilling and Associated Technologies

. Figure 47.  A range of typically ground twist drill points

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• Web-thinning – as its name implies, the chisel point edges of the drill, increasing both tool life and improvis web-thinned/notched, by regrinding to reduce the width of the chisel point, while slightly modifying the profile, giving a partial self-centring action (Fig. 47c), • Helical – the chisel point is ground to an ‘S-shape’ , which modifies both the chisel point and its profiled shape, improving the drilling performance and self-centring action (Fig.47d). NB  On some drills a sophisticated grinding action has imparted drills without a chisel point, which significantly improves their drill penetration rates into the workpiece, but requires a complex drill regrinding operation to re-sharpen them when the edge becomes ‘dulled’. Not only are drills supplied with appropriate point geometries, but for twist drills the twin spiral flutes of the drill can also be specified – from the tooling manufacturer, as this gives the drill its ‘equivalent of the rake angle’ as found on a single-point turning tool. On conventional jobber drills, the normal flute angle is 29° – giving a relatively ‘slow’ helix (Fig. 47a) and in the past, typically being utilised for drilling most plain carbon steel grades. Conversely, a drill with a ‘quick’ helix angle (Fig. 47d), might be employed to drill soft materials such as certain plastics. Brittle materials on the other hand, which might be utilised typically when drilling Cartridge Brass (i.e. 70Cu 30Zn composition), require a zero, or slightly negative helix. NB  It is possible to temporarily modify the drill’s helix angle by re-grinding, termed ‘drill dubbing’ , which refers to lightly ‘flash-grinding’ the flutes at the lips to decrease the effective flute helix angle. The main strength of a drill is via its web, or its crosssection which can be changed and as a result, will modify the flute’s geometric profile (i.e. see Fig. 48). In general, drill cross-section are classified in three groups, namely: • Axe-shaped – having well-defined margins (Fig. 48 –top), • Rounded heel – with increased web, but small margins (Fig. 48 – middle), • Rhombic – incorporating a large web, with wide margins (Fig. 48 – bottom). NB  Some twist drills feature oil/coolant holes to allow cutting fluid to reach right down toward the cutting

ing the hole’s ‘Surface Integrity’ .

3.1.3 The Dynamics of Twist Drilling Holes Introduction The term ‘drilling’ refers to all production techniques for the manufacture of cylindrical holes in workpieces using chip-making cutting tools, for short-hole10 and deep-hole drilling operations. The expression ‘solid drilling’ has been introduced in recent years – which is hole-making generation undertaken in a single operation, to differentiate it from that of the previous techniques of either: centre-drilling or, pilot-hole drilling (i.e, see Fig. 50b) prior to drilling to size. Drill technology includes a range of specialised hole-making tooling, including: twist drills, solid drills, counter-boring and trepanning tools and deep-hole drills11. In the first instance, mention will be made of twist drilling operations, then a review of these other drilling production methods will occur.

Twist Drills Twist drilling operations have been carried out for around 150 years, with a twist drill imparting ‘balanced cutting conditions’ , assuming that the drill’s geometry is symmetrical. It has been suggested that the work of drilling may be considered as two singlepoint lathe tools engaged in an internal straight turning operation. A twist drill produces both torque and thrust as it rotates and is fed into the workpiece. The main contribution to torque is through the lips, with a small amount of torque being generated by the chisel point as the drill rotates against the resistance of the



‘Surface Integrity’ has been coined to describe the ‘altered material zone’ (AMZ), for localised sub-surface layers that differ from those of the bulk material – considerably more will be said on this subject in Chapter 7.

10 ‘Short-hole drilling’ operations cover depth-to-hole-diameter-ratios of up to 6D (i.e. for diameters up to 30 mm), whilst larger drilled holes are limited to depths of 2.5D. Where: D = nominal drill diameter. 11 ‘Deep-hole drilling’ and ‘Gun-drilling’ operations are virtually the same, with the term Deep-hole drilling being the preferred term in this text.

Drilling and Associated Technologies

. Figure 48.  Symmetrical twist drill cross-sectional profiles [After: Spur and Masuha, 1981]

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workpiece (see Fig. 49). The thrust force (Fig. 49) is the result of the selected penetration rate (i.e. feed), in combination with the bulk hardness of the workpiece and its work-hardening ability and the efficiency of the coolant supply – if any – to the cutting edges (i.e lips). The resolution of the cutting resistance into their various components when twist drilling, is shown in Fig. 49 at a mid-point along the lips. The thrust force is just

one of the cutting resistances in a drilling operation, contributions to drill resistance are from the: • Lips – equal lip lengths and angles are important for a ‘balanced cutting action’ , this being considered an efficient cutting process, • Chisel edge – is highly negatively skewed and as it acts like a ‘blunt wedge-shaped indentor’ , extruding the workpiece material from this vicinity,

. Figure 49.  The balanced cutting forces resulting from drilling holes utilising twist drill geometries. [After: Kaczmarek, 1976]

Drilling and Associated Technologies

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• Land, or margin – via a rubbing, or frictional ac- A = 2Az = szd = sd/2 = 2hzb tion.

NB  The latter two are relatively inefficient processes, moreover, the resistance components of the lips and chisel edge are the product of resistance of the undeformed chip to plastic strain, in combination with resistance due to external friction. The land resistance occurs from the friction (i.e. rubbing) against the side of the drill’s hole. When symmetrical twist drilling (illustrated in Fig. 50), the undeformed chip can be characterised by its: • Cutting depth (a) – where ‘a’ = d/2, with ‘d’ being the drill’s diameter (mm), • Feedrate (s) – this being the distance the cutting edge moves in the drilling axis direction during one revolution. Normally two rigidly joined cutting edges are cutting at any instant, each one in its travel corresponds to feed ‘sz’ , which removes an undeformed chip whose size – in the direction of the drilling axis is: ‘s1’ and ‘s2’ respectively (i.e see Fig. 50a) and, as most drills are symmetrical in design, then: s = s1 = s2 = s/2

(mm),

• Undeformed chip thickness (hz) – to be removed by

each of the drill’s cutting lips, which can be determined from the following relationship:

hz = sz sinθ

(mm).

NB  With a symmetrical drill, then: Hz = h1 = h2.

• Undeformed chip thickness (b) – can be found from the following relationship:

b = a 1/sinθ = d/2sinθ

(mm).

∴ It follows from these expressions, that the transverse cross-sectional area of the undeformed chip at each of the twist drill’s cutting lips, can be shown by the following relationship: Az = szd/2 = sd/4 = hzb

(mm2).

Hence, the total transverse cross-sectional area when drilling of the undeformed chip will be:

(mm2).

Conversely, in the case of ‘Pilot’ hole drilling (Fig. 50b), the undeformed chip elements are identical to ‘Solid’ drilling, but for the exception of the DOC, which can be expressed in the following manner: a = d-do/2

(mm).

Where: d = diameter of final hole (mm), do = diameter of primary hole (mm). Thus, for example in the case of ‘Pilot’ hole drilling, the total cross-sectional area of the undeformed chip, will be: A = 2Az = sz(d – do) = s(d – do)/2 = 2hzb

(mm2).

The calculation of cutting forces in ‘Solid’ hole drilling (Fig. 50a), can be found from the general formulae for axial force (F) and torque (M), in the following manner, respectively: F = CF d bF suF KH M = CM d bM suM KH

(kg) (kg mm)

Where: CF12 and CM13 = constants (i.e. derived from Kaczmarek‘s findings), d = nominal drill diameter (mm), bF and bM = exponents characterising the influence of the drill diameter, s = feed rate (mm rev–1), uF and uM = exponents characterising influence of feedrate, = workpiece material’s correction coKH efficient (i.e. concerning mechanical properties).

12 CF is derived from experimental data, typically: Carbon steel (construction) 84.7, Grey CI 60.5, Malleable CI 52.5, Bronze (medium hardness) 31.5 – with HSS drills, ranging from φ10 to 60 mm. 13 CM is derived from experimental data, typically: Carbon steel (construction) 33.8, Grey CI 23.3, Malleable CI 20.3, Bronze (medium hardness) 12.2 – with HSS drills, ranging from φ10 to 60 mm.

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. Figure 50.  Drilling a hole with/without a ‘pilot’ hole and the cutting, rubbing and extrusion mechanism. [After: Kaczmarek, 1976]

Drilling and Associated Technologies

101

Conversely, for ‘Pilot’ hole drilling (Fig. 50b), these mathematical formulae are modified in the following manner: F = CF d bF aeF suF KH M = CM d bM aeM suM KH

(kg) (kg mm)

Where: a = DOC (mm), eF and eM = exponents indicating the DOC’s influence. These axial force and torque formulae derived in the work by Kaczmarek, are concerned with ‘so-called’ average twist drilling values. These ‘averages’ are related to drill diameters between 15 to 35 mm, having feed ranges in the vicinity of 0.2 to 0.4 mm rev–1. Therefore, the entire axial force (F) and torque (M), comprises of contributions of the lips, land and chisel point, in the following manner: • Axial force (F) – lips (50%), land (10%) and chisel point (40%), • Torque (M) – lips (80%), land (12%) and chisel point (8%). NB  These contributing factors to axial force and torque are for drill depths that do not exceed 2.5d. If the drilling force is significantly increased, then this has the effect of distorting the drill shaft. Such distortion, causes the drill’s cutting edge to move forward into the workpiece material, in this manner it jointly increases the DOC and the drilling force. Correspondingly, if the drilling force is reduced, the twist drill will recover its shape, with the cutting edge moving back from the workpiece, thus reducing both the DOC and cutting force. This stretching and compression of the drill’s shaft – somewhat like a spring14 – is unique to twist drilling, being an unstable element in the cutting process. By way of comparison, most cutting tool

14 ‘Lengthening effect’ is associated with the twist drill’s shaft being twisted by the application of torque, with elastically springs-back upon release of the drilling torque. Not only will the twist drill ‘spring’ , but it can also ‘bend’ due to the increased thrust loads produced by high penetration rates.

edges are normally deflected away from increases in the load. A common form of failure of twist drills in operation is from shattering, with such catastrophic failure being related to the dynamic nature of twist drilling. By way of illustration, a φ4.5 mm long-series twist drill is capable of withstanding a torque of approximately 6 Nm before it catastrophically fails. Normally, the torque for most drilling operations is around 1 Nm.

Temperatures in Twist Drilling The accumulation of heat in the vicinity of cutting is an important factor in the cutting process, with much of the mechanical energy necessary for machining being converted into heat, then conducted into the chip, workpiece and tool (Fig. 51). The consequential thermal phenomena are important, as they can affect the: • Mode of deformation – elastic/plastic behaviour of the chip, • Machined surface – for metals the ultimate metallurgical state of the material, • Tool wear rate – which depends upon a number of criteria, such as the tool’s coating, cutting data employed, work-hardening ability of the workpiece and coolant delivery and its efficiency. It is imperative to comprehend the factors that control both heat generation and its dissipation, together with the tool and work’s temperature distribution in and near the cutting zone. A drilling operation can be considered as a complex machining process, with specific and unique characteristics, not least of which, are the production of chips when drilling. These chips are in continuous contact with the drill flutes and the generated hole’s surface. Hence, any minute changes in the drill’s geometry, can cause enormous modifications to the either the drill’s wear rate and its predicted life. Heat generated whilst drilling will be transformed by a range of ‘states’ , including: • Conduction – through the chips, workpiece and drill, • Convection and radiation – via the ‘air-spaces’ in the hole as the drill penetrates deeper into the workpiece. The drilling temperature during a prolonged operation can approximate steady-state conditions, with the heat generated whilst cutting when employing a new drill

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. Figure 51.  The drilling process and the asociated zones of heat generation whilst hole-making. [After: Trigger and Chao, 1951]

Drilling and Associated Technologies

is associated with two distinct regions (i.e. see Fig. 51 – Section on X-X) at the: • Primary shear zone – where plastic deformation occurs, this being the major source of heat generation, • Secondary shear zone – from within the tool/chip interface, where pronounced friction takes place. NB  The drill clearance surface temperature, is significantly affected by the rake face interface temperature. The bulk rise in the drill’s temperature is multifarious, due to the necessity to consider a range of factors, including: heat flow distribution, the geometric shape of the conducting bodies, together with any variation of thermal properties of both the drill and workpiece materials with temperature changes. The generated heat distribution when drilling depends upon the thermal properties of the tool, workpiece and chip. Therefore, the thermal diffusivity (K/ρc), will determine the rate at which heat transfers through the material, while also controlling the penetration depth of the surface temperature. While the absorption coefficient (Kρc), determines the quantity of heat being absorbed by a given mass of material. Drilling temperatures vary considerably in the research work undertaken over the years, being heavily influenced by a wide range of cutting-related parameters, making it extremely difficult to obtain meaningful comparisons of local temperatures in a real-time drilling operation. For example, the scatter of ‘bulk’ temperature values for say, a φ6mm twist drill, can vary between approximately 200 to 380°C, under steady-state drilling conditions15, for comparable workpiece materials, making it very difficult to obtain meaningful drill life comparisons. Coolant delivery is imperative when drilling and to this end, through-the-nose coolant operation enables the lubrication and cooling of the drill’s point (Fig. 52c – illustrating the coolant holes behind the lips). This efficient technique of ensuring that the coolant gets to the action of drilling, gives better chip control, helping

15 Twist drill interface temperatures have been reported to be over 870°C in the workpiece’s ‘plasticity region’ , which somewhat contradicts the ‘bulk’ temperatures, although in mitigation, it should be said that these very high temperatures at the interface at somewhat localised.

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to reduce machining temperatures significantly and aid drill penetration rates, while increasing tool life. Coolant holes through-the-nose are not restricted to twist drills, as Spade-16 and Gun-drills17, together with Indexable drills also often incorporate this coolant delivery feature, to remove heat and lubricate the cutting edges.

3.1.4 Indexable Drills Indexable drills have some significant advantages over their twist drilling counterparts (i.e. a range of both indexable and twist drills are depicted in Fig.52a). These indexable drills – allowing the cutting inserts to be changed (see Fig. 52c), permit faster cutting speeds and enable a wider range of workpiece materials to be successfully drilled than when utilising conventional twist drills. Normally, indexable drills are limited to shorter hole depths of around ‘4D’ , than equivalent diam­eter twist drills. Indexable drills must be set up with care and in the correct relationship to the machine tool’s headstock/ spindle, ensuring that both the drill’s and the spindle’s centrelines are coincident, otherwise over-, or undersized holes may be produced (see Fig. 53a – top). Yet another problem that needs to be addressed when employing these indexable drills, is termed ‘radial runout’18, which affects the inserts centre height and should be limited to <0.127 mm. One advantage of being able to manipulate the indexable drill’s axis, is that it can be used to adjust the drilled hole’s diameter, by parallel adjustment of the drill’s and spindle’s respective centrelines – this being very useful for controlling

16 Spade drills are twist drills (i.e. bodies are normally manufactured from either 1018, or 1020 low-carbon steel) with a standard blade being inserted at the drill’s point, enabling holes to be generated up to 8D deep. Blades are usually coated micrograin HSS with high-cobalt content, or coated cemented carbides, ranging from stub drills to extra-long lengths, with either straight, or spiral flutes. 17 Gun-, or Deep-hole drills will be mentioned in some detail later in this chapter, but they allow considerable length-to-dia­meter ratios to be drilled in workpieces, necessitating high-pressure coolant delivery, with efficient chip-flushing capabilities. 18 Radial runout refers to misalignment in the radial direction, which should be minimised, as it alters the position of the drill’s cutting inserts.

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. Figure 52.  Short hole drilling. [Courtesy of Sandvik Coromant]

Drilling and Associated Technologies

drilled hole tolerances. On turning centres this can be readily achieved by modifying the CNC cutting program, to offset the drill with respect to the machine’s centreline. Moreover, for turning operations employing drilled features, then the indexable insert’s top surfaces must remain parallel to the X-axis of the machine tool. The arrangement of the inner and outer cutting edges of an indexable insert drill relative to each other, together with the drill’s position to the axis of rotation are vital for perfect drilling operations19 (i.e. see Fig. 53a). The cutting inserts positions by possible X-axis adjustments, are critical for the: smooth running, resultant cutting forces and, will influence the drilled hole’s alignment. Preferably, the cutting edges are arranged in such a manner that the inner- (SBI) and outer-inserts (SBA) have identical cutting widths (Fig. 53a – bottom left). When new insert cutting edges are utilised, this results in a balance of the cutting forces in the Y-axis, guaranteeing drilled holes of accurate size and surface texture without ‘retraction striae’ 20. When selecting the appropriate adjustment angles (χi) and (χA), the lines of force via feed forces (Fa and Fi) will coincide with the drill’s axis in the centre of the clamping shaft (Fig. 53a – top). Hence, the clamping shaft must only transmit torque resulting from the cutting forces and the bending moment of the resultant cutting force which will be present. Typically, the outer insert’s resultant cutting force FA (i.e. Fig. 53b) is comprised of the following forces: • Remaining cutting force (∆Fc) – generated through greater wear rates at the periphery of the outer cutting insert, • Passive force (Fp) – generated by the corner radius of the outer cutting insert. With indexable drills, the chip flute is selected so that the drill’s profile from the tip up to the chip flute, has its runout twisted by between 65° to 85°. In the vicinity of the chip flute the runout (i.e. the longest ‘lever arm’ of the force), is where the maximum resistance

19 Some tooling manufacturers recommend that an indexable drill’s inner insert is positioned slightly below the spindle’s centreline, as this allows a small core of uncut material to pass over the top surface of the insert and break off – being carried away with the rest of the chips. 20 ‘Retraction striae’ refers to the ‘trail-lines’ resulting from the outer insert’s gouging, or ploughing the previously drilled surface as it is withdrawn from the hole.

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moment to the resultant cutting force FA is found. The bending strength attained in this manner can be greatly increased by employing round profiled chip flutes. This rounded chip flute cross-section, does not significantly weaken the drill’s body and provides optimum chip-flow – even when drilling long-chipping workpieces. A taper of the tool holder behind the insert seats, prevents a ‘squeezing’ of the chips between the drill and the drill-hole wall. Due to the design of the indexable drill, the two cutting inserts are subjected to very dissimilar stresses when drilling. For example, the indexable drill (Fig. 53a – bottom), has the outer insert being subjected to greater stress than its inner counterpart, typically having both thermal and abrasive stresses, while the inner insert must have high toughness characteristics. Some cutting tool manufacturers recommend socalled ‘mixed-tipping’ 21 of inserts, where a toughened grade is used for the inner insert and a wear-resistant grade for the outer insert. However, some discretion should be used when utilising indexable drills with mixed cutting inserts, so perhaps reference back to the tooling manufacturer may be advisable if production quantities are sufficient in order to optimise this potential ‘mixed grade strategy’. Typically, by exploiting ‘mixed-tipping’ , for example when machining free-cutting steel grades cutting speeds of up to 400 m min–1 are possible, whereas when drilling low-silicon aluminium grades cutting speeds of 600 m min–1 can be achieved with tool lives of up to 45 min of cutting time per edge. Several design factors will influence an indexable drill’s performance, these include: • Sintered cutting insert chip-breakers – these will improve chip control and enable high penetration rates to be utilised, • Advanced flute design – allowing deeper chip gullets, thus minimising chip-jamming tendencies, • Faster-, slower-, or straight-fluted designs – with wider flute profiles reduce chip-binding and degradation of the drilled hole surface, whilst also improving penetration rates, • Cutting insert shape – utilising square (i.e. having 4 cutting edges), rectangular (i.e. with two edges), or triangular inserts (i.e. having three edges) – the

21 ‘Mixed-tipping’ , refers to having dissimilar grades of inserts for the outer and inner cutting edges, as they fulfil different mechanical working criteria whilst drilling.

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. Figure 53.  Indexable insert drills – insert position and flute geometry. [Courtesy of Kennametal Hertel]

latter being the most popular version for generalpurpose drilling operations.

be up to 3.8 mm, or 7.6 mm on a diameter22, which in reality, amounts to a ‘fine-boring’ operation, giv-

NB  Double-sided cutting inserts are available, but are mainly used for milling operations. One major advantage that indexable drills have over twist drills, is that they can be offset to produce different hole diameters. This offset for turning centres can

22 On machining centres this offset is somewhat less, with the maximum radial offset being approximately 1 mm, or 2 mm on dia­ meter. Of note, is that when an indexable drill is offset, then the maximum feedrate should be no greater than 0.15 mm rev–1.

Drilling and Associated Technologies

ing considerable scope in diametric sizing of holes. These drills should be selected for the minimum overhang and need to have both the drill’s and machine tool’s centrelines parallel 23 to within 0.076 mm, or better. By utilising an indexable drill on a turning operation, they can be beneficial when attempting to start drilling angled, or uneven workpiece surfaces, as illustrated by some of the surface faces depicted in Fig. 54. Indexable drills, employed for such non-flat workpiece faces, obviates the need for either a previous counter boring, or spotfacing operation. The major advantage of indexable drills over either their HSS, or carbide twist drill counterparts, is their ability to run at much higher rates. For example and, by way of simplistic comparison, if a φ18 mm hole has to be drilled in free-cutting stainless steel, then an: • HSS twist drill – can be run at 16.8 m min–1 with a 0.2 mm rev–1 feed would produce 280 rpm at 57 mm min–1, • Solid carbide twist drill – can be run at 61 m min–1 with a 0.1 mm rev–1 feed would produce 1,019 rpm at 103 mm min–1, • Indexable drill – can be run at 170 m min–1 with a 0.76 mm rev–1 feed would produce 2,852 rpm at 217 mm min–1.

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3.1.5 Counter-Boring/Trepanning

An important safety note when using these indexable drills in turning operations, is that the high penetration rates, coupled to the exit of a through hole in the part, creates a slug which is thrown clear from either the chuck, or from the rear of the bore of the machine’s headstock at tremendous force. If the machine tool is not guarded appropriately, it could prove to be hazardous to any operator in the local vicinity.

A counter-boring operation is an often utilised technique for the enlargement of previously manufactured holes, normally to provide them with accurate dimensions and/or improved surface finish (i.e. see Fig. 55a). Counter-bores are also produced to register a larger-faced shouldered shaft, or to sink a precision cap-head bolt below the clamped part’s surface. In this latter case, often the previously-drilled hole is used to align the bolt’s axis, by using a counter-boring tool with a ‘pilot-bush’ of the same drilled diameter to act as a guide to allow machining to a correct counter-bored depth. Counter-boring heads are employed to open-out existing holes (Fig. 55a). The heads were in the past, often of a single insert design, but latterly, they are produced with multiple inserts – particularly when the working clearance is such, that it cannot cope with by the single insert variety. The multi-insert counter-boring heads, can have their inserts finely diametricallyadjusted by means of precision wedge and radial adjusting screws. A trepanning operation is undertaken in one operation, but instead of machining the complete hole, only part of the hole is cut, leaving a core (Fig. 55b). For large workpiece dimension machining, a trepanning operation uses less power and axial pressure than other equivalent manufacturing processes. However, one major problem occurs with the trepanning operation, which is that the core produced as the trepanning tool penetrates into the workpiece, becomes quite difficult to handle (i.e. see Fig. 55b -lower diagram). Trepanning heads are normally utilised on: • Large workpiece diameters – greater than 120 mm, • Limited machine tool power – alternatively, if it is not prudent to switch