Metals Handbook, Volume 6: Welding, Brazing, and Soldering [9 ed.] 9780871700070, 0871700077

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ASM INTERNATIONAL

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Pu blica t ion I n for m a t ion a n d Con t r ibu t or s

Machining was published in 1989 as Volume 16 of the 9th Edition Metals Handbook. With the second printing (1995), the series title was changed to ASM Handbook. The Volume was prepared under the direction of the ASM Handbook Committee.

Au t h or s a n d Re vie w e r s • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • •

T.E. Aaron Anocut, Inc. Gary Adams Cominco Metals John Agapiou General Motors Technical Center M.S. Ahmed Transfer Technology Limited (England) G. Albares Technical Consultant Tom Andrew Harper Company James A. Aris Rockwell International William N. Ault Norton Company A. Bagchi Ohio State University J. Gary Baldoni GTE Laboratories Moshe M. Barash Purdue University Carl Bartholed Reishauer Corporation Alan M. Bayer Teledyne Vasco Abdel E. Bayoumi Washington State University Bruce N. Beauchesne Laser Services, Inc. Bruce A. Becherer Teledyne Vasco Guy Bellows Metcut Research Associates Inc. Gary F. Benedict Allied-Signal Aerospace Company Garrett Engine Division R.C. Benn Inco Alloys International, Inc. E.O. Bennett University of Houston Michael Bess Aluminum Smelting & Refining Company, Inc. Certified Alloys Company Hugh Bettis DoAll Company J. Binns, Jr. Binns Machinery Products J. Binns, Sr. Binns Machinery Products J T. Black Auburn University Mark Bobert Technical Consultant J.F. Boland Rockwell International S.P. Boppana GTE Valenite F.W. Boulger Technical Consultant K. Brach General Electric Company J. Bradley Technical Consultant José R.T. Branco Colorado School of Mines R. Bratt Technical Consultant R.W. Breitzig INCO Alloys International James Brewer Fairfield Manufacturing Company Chris Brookes The University of Hull (England) S.T. Buljan GTE Laboratories Virgil Buraczynski Besley Products Corporation Stephen J. Burden GTE Valenite Corporation John H. Burness The Timken Company A.C. Carius General Electric Company Nick Cerwin A. Finkl & Sons, Inc. Harry E. Chandler ASM International S. Chandrasekar Purdue University

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Chao-Hwa Chang University of California, Los Angeles John D. Christopher Metcut Research Associates Inc. T.J. Clark General Electric Company Hilary A. Clouser Extrude Hone Corporation Joseph W. Coniglio Gould & Eberhardt Gear Machinery Corporations John Conlon Conlon Industries, Inc. S. Cook LTV Aerospace Company David Cunningham General Electric Superabrasives Richard Dabeck Coral Chemical Company Dilip Dalal The Cross Company J. Dalton Bardons & Oliver Timothy Danielson Chem Tronics, Inc. C.V. Darragh The Timken Company D.W. Davies BNF Metals Technology Centre (England) Warren J. Demery Sossner Tap & Tool Company Amedeo deRege Domfer Metal Powders Limited (Canada) Warren R. DeVries Rensselaer Polytechnic Institute Kurt Dieme Reed Rolled Thread Die Company J. Dimitrious Pfauter-Maag Cutting Tools Phil Diskins DiCo Corporation Charles A. Divine, Jr. AL Tech Specialty Steel Corporation R. Dixon Crucible Specialty Metals Stephan Donelson Colorado School of Mines Carl J. Dorsch Crucible Materials Corporation Clifford E. Drake ENERPAC Group Applied Powers, Inc. W. Dresher International Copper Research Association D. Dykehouse Technical Consultant Robert P. Eichorst United Technologies Ahmad K. Elshennawy University of Central Florida Dana Elza Coherent General Phil Esserkaln Kempsmith Machining Company J. Richard Evans Dowty Canada Ltd. (Canada) John J. Fickers Los Alamos National Laboratory Michael Field Metcut Research Associates Inc. M.E. Finn Steltech Inc. (Canada) Thomas Fisher Surftran Division Robert Bosch Corporation Donald G. Flom Technical Consultant Thomas O. Floyd Seco-Carboloy John E. Foley S. Baird Corporation David Fordanick The Cross Company Paul Frederick Dow Chemical Company Howard Friedman Fotofabrication Corporation John E. Fuller Rockwell International Roland Galipeau ThermoBurr Canada (Canada) Douglas V. Gallagher Rockwell International Ramesh Gandhi Alliance Tool & Manufacturing Inc. Geoffrey Y. Gill Muskegon Tool Industries Inc. J. Ginsberg Photo Chemical Machining Institute M.A. Glandt Giddings & Lewis Claus G. Goetzel Technical Consultant F. Gorsler General Electric Company Leigh Gott Kearney & Trecker Corporation Dennis Grable The Cross Company Allan M. Grant Allan M. Grant & Associates

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Mikell P. Groover Lehigh University Walter W. Gruss Kyocera Feldmuehle, Inc. J. Gurland Brown University Clarence J. Hagstrom Lindmark Machine Works T.E. Hale Carboloy, Inc. James E. Hanafee Lawrence Livermore National Laboratory R. Hanson Ferranti Sciaky, Inc. R.E. Hardesty Electrofusion Corporation S.M. Harrington Thomas & Betts Corporation Derek Hartell Cleerman Machine Tool, Inc. P.J. Heath De Beers Industrial Diamond Division (FTY) Ltd. (England) Barry Heller Teledyne Firth Sterling Gene Herron Metem Corporation Thomas Hill Speedsteel of New Jersey, Inc. R.M. Hooper University of Exeter (England) L. Houman Axon EDM, Inc. David J. Howell Roll-A-Matic, Inc. Fred Huscher Rockwell International Automotive Division Richard M. Jacobs Consultant Services Institute, Inc. J. Jackson Radian Corporation E.C. Jameson Transtec, Inc. Ernest Jerome Zagar Inc. Mark Johnson Tapmatic Corporation C.E. Johnston Flow Systems, Inc. K. Jones Tooling Systems Inc. John F. Kahles Metcut Research Associates Inc. Serope Kalpakjian Illinois Institute of Technology A. Karl Garrett Turbine K. Katbi GTE Valenite L. Alden Kendall Washington State University B. Klamecki University of Minnesota J.B. Kohls Institute of Advanced Manufacturing Sciences, Inc. Ranga Komanduri National Science Foundation Yoram Koren University of Michigan Ted Kosa Carpenter Technology Corporation William P. Koster Metcut Research Associates Inc. T. Kozinski Precision Art Coordinators James E. Krejci Keystone Threaded Products Division Theodore J. Krenzer The Gleason Works Gleason Company Gerald Kusar Ajax Manufacturing Company John B. Lambert Fansteel Eugene M. Langworthy Aerochem, Inc. L.K. Lauderbaugh Rensselaer Polytechnic Institute J.A. Laverick The Timken Company Frank D. Leone Pitney Bowes, Inc. D. Levinson Taussig Associates, Inc. Terry L. Lievestro Lehr Precision, Inc. Richard P. Lindsay Norton Company Steven Lochmoeller Roton Products Inc. R. Luke DoAll Company Pel Lynah P. R. Hoffman Machine Products Gerald Makuh Weldon Tool Company Reza A. Maleki Moorhead State University Stephan Malkin University of Massachusetts at Amherst

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O. Masory Texas A&M University Larry Mayer TIMET Bob McLemore The Marquardt Company Alan McMechan McDonnell Douglas Canada (Canada) Pankaj K. Mehrotra Kennametal Inc. Fred Meyer Precitec Corporation Thomas W. McClure Balax Inc. W. Mihaichuk Eastern Alloys, Inc. James Millar Lapmaster Division of Crane Packing Company Brian Mitchell, Sr. General Broach and Engineering Company Walter R. Mohn Advanced Composite Materials Corporation Frank Moravcik The Cross Company Mary Moreland Bullen Ultrasonics Inc. Jonathon Morey Morey Machining Company R.A. Morley Reynolds Aluminum T.O. Morris Martin Marietta Energy Systems, Inc. David Moskowitz Technical Consultant Bill Murphy Rodeco Company Elliot S. Nachtman Tower Oil & Technology Company Steven J. Neter Peterson Precision Engineering Company M. Anthony Newton NItech, Inc. Ronald P. Ney Carpenter Technology Corporation Roger Nichting Colorado School of Mines P. Niessen University of Waterloo (Canada) Bernard North Kennametal Inc. Raymond J. Novotny Technical Consultant J. Padgett J.R. Padgett Associates Ralph Panfil Davenport Machine Jeffrey T. Paprocki Kearney & Trecker Corporation W. Neil Peters Corning Glass Works Robert E. Phillips Everite Machine Products Company R. Pierce Radian Corporation Kenneth E. Pinnow Crucible Metals Corporation Robert A. Powell Hoeganaes Corporation D. Powers Leybold Vacuum Systems, Inc. J. Prazniak The Timken Company Ralph E. Prescott Monarch Machine Tool Company Allen Queenen Kearney & Trecker Corporation S. Ramanath Norton Company V. Rangarajan Colorado School of Mines M.P. Ranson Inco Alloys International, Inc. James Reichman Kenworth Truck Company Lawrence J. Rhoades Extrude Hone Corporation C.E. Rodaitis The Timken Company Harvey W. Rohmiller Lodge & Shipley Division Manuflex Corporation Stuart Salmon Advanced Manufacturing Science & Technology Shyam K. Samanta National Science Foundation Ron Sanders Laserdyne A.T. Santhanam Kennametal Inc. K. Scheucher Modtech Corporation Ronald W. Schneider MG Industries Scott Schneier Regal Beloit Corporation Michael Shultz Wisconsin Drill Head Company R. Seely Corning Glass Works

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W.R. Sharpe Battelle Pacific Northwest Laboratories Chi-Hung Shen General Motors Technical Center T. Slawson Ridge Metals Inc. Ted A. Slezak Armstrong-Blum Manufacturing Company William M. Spurgeon University of Michigan--Dearborn D.R. Stashko GTE Valenite Corporation William Stasko Crucible Materials Corporation Larry E. Stockline PROMESS, Inc. Glenn E. Stork S.S. White Industrial Products Division of Pennwalt Corporation K. Subramanian Norton Company Lewis Sylvia Morse Cutting Tools D. Taylor Manufacturing Systems Extension Center R.A. Thompson General Electric Company Thomas Thompson Badger Meter Company P. Tierney Kennametal Inc. Jiri Tlusty University of Florida C. Treadwell Sonic-Mill Albuquerque J. Tulloch Wells Saw Division Charles I. Turner Kearney & Trecker Corporation William R. Tyrell Branson Ultrasonics Corporation A. Galip Ulsoy University of Michigan G.L. Van Arsdale Battelle Pacific Northwest Laboratories M.R. Van den Bergh Composites Specialties, Inc. Christopher Van De Motter The Ohio Broach & Machine Tool Company Philip A. Ventura The Cross Company Don Vick Ingersoll Milling Machine Company Craig E. Virkus Elliott Company R.J. von Gutfeld Thomas J. Watson Research Center International Business Machines Charles F. Walton Technical Consultant L. Walton Latrobe Steel Company I. Weber Technical Consultant R. Terrence Webster Teledyne Wah Chang Albany W.R. Welton Welton Rolled Thread Corporation Robert Werkema Technical Consultant Robert I. Werner R.D. Werner Company Inc. Gene White Coherent General Richard F. Williams Natco, Inc. M.L.H. Wise University of Birmingham (England) William Wonnacott Thread Grinding Service R.E. Wood Lockheed Aeronautical Systems Company Hiroshi Yaguchi Inland Steel Company Patrick Yeko ENERPAC Group Applied Powers, Inc. C. Zimmerman GTE Valenite Emory W. Zimmers, Jr. Lehigh University

For e w or d In the 22 years since the 8th Edition Metals Handbook volume on machining was published, material removal operations have undergone dynamic changes. The mechanics of the cutting process are better understood, new cutting tool materials have been developed, machine controls and computer-aided engineering have rapidly advanced, and nontraditional machining methods continue to be refined. The difficult challenges faced by industry have necessitated these developments. Requirements for high-strength materials and the introduction of difficult-to-machine structural ceramics, composites, and electronic components have placed new and greater demands on machining technology, and have also spurred continued research and development in material removal techniques.

Volume 16 of the 9th Edition describes the evolution of machining technology comprehensively, with great attention to detail and accuracy. In addition to providing valuable information on recent developments, the Handbook devotes exhaustive coverage to more standard, traditional machining methods. This new Volume is also the final step in the fulfillment of ASM's commitment to coverage of metalworking technology in the 9th Edition, taking its place alongside Volume 6 (Welding, Brazing, and Soldering), Volume 7 (Powder Metallurgy), Volume 14 (Forming and Forging), and Volume 15 (Casting). This enormous undertaking was made possible by the combined efforts of many dedicated and selfless authors and reviewers, the ASM Handbook Committee, and the ASM editorial staff. Special recognition is also due to Metcut Research Associates Inc. and its president, William P. Koster, for permission to use tabulated data published in Volumes 1 and 2 of the Machining Data Handbook (3rd edition). To all the men and women who contributed to the planning and preparation of this Volume, we extend our sincere thanks. Richard K. Pitler President, ASM International Edward L. Langer Managing Director, ASM International

Pr e fa ce Machining is one of the most important of the basic manufacturing processes. Almost every manufactured product contains components that require machining, often to great precision. Yet material removal operations are among the most expensive; in the U.S. alone, more than $100 billion will be spent this year on machining. These high costs put tremendous economic pressures on production managers and engineers as they struggle to find ways to increase productivity. Compounding their problems is the increasing use of more difficult-to-machine materials, such as nickelbase superalloys and titanium-base alloys in aerospace applications, structural ceramics, high-strength polymers, composites (both metal-matrix and resin-matrix), and electronic materials. The present Volume of Metals Handbook has been structured to provide answers to the questions and challenges associated with current machining technology. Following a general introduction to machining processes, 9 major sections containing 78 articles cover all aspects of material removal. Much of this material is new. In fact, 30 articles in this Volume were not included in its 8th Edition predecessor. Noteworthy are the articles that have been added to describe the mechanics of the cutting process and advances in new materials, new processes, new methods of machine control, and computer-aided engineering. The first Section of the Handbook reviews the fundamentals of the machining process. Included are articles describing the mechanics of chip formation, the forces, stresses, and power at the cutting tool, the principles of tool wear and tool life, and the relationship between cutting and grinding parameters and surface finish and surface integrity. In the following Section, extensive data are provided on the applications, advantages and limitations, properties, tool geometries, and typical operating parameters for seven classes of tool materials: high-speed tool steels (both conventional wrought and powder metallurgy), cast cobalt alloys, cemented carbides, cermets, ceramics, and ultrahard tool materials (polycrystalline diamond and cubic boron nitride). Recent developments in wear-resistant coatings that are applied on high-speed steel, carbides, and ceramics are also discussed. The third Section focuses on cutting and grinding fluids--their functions, selection criteria, and application. Coverage of proper maintenance procedures (storage, handling, recycling, and disposal) and the toxicology and biology associated with cutting and grinding fluids is included. The next Section contains 21 articles that summarize the process capabilities, machines, cutting parameters and variables, and applications of traditional chip removal processes, such as turning, drilling, and milling. Advanced tooling used in multiple-operation machining, proper tool fixturing, and tool condition monitoring systems are also discussed, along with computer numerical controlled machining centers, flexible manufacturing systems, and transfer machines.

Although near net shape technology, including a greater use of precision casting, powder metallurgy, and precision forging, has lessened the need for some traditional machining operations, abrasive machining is being employed to a greater extent than in the past. The fifth Section of the Handbook examines the principles, equipment, and applications of grinding, honing, and lapping as well as recent developments in super-abrasives, used for precision grinding of difficultto-machine and/or brittle materials. The sixth Section looks at a variety of nontraditional machining methods that do not produce chips or a lay pattern in the surface. Mechanical, electrical, thermal, and chemical nontraditional techniques are described. Applications of these methods are emphasized, with practical examples involving nontraditional machining of metals, ceramics, glasses, plastics, and electronic components. The next Section describes high-speed and high removal rate processes that have been developed to dramatically increase productivity. The effects of high-speed processing on chip formation and tool wear are discussed, along with materials that are being machined using these processes. The eighth Section introduces the reader to two of the most rapidly developing and important areas in machining technology: machine controls and computer applications. Although the basic configurations of many machine tools have not changed significantly, the advent of numerical control and adaptive control has substantially improved manufacturing productivity and workpiece quality. Machine controls and the integration of CAD/CAM technology into machine tools are described in articles written with the engineer, not the software expert, in mind. The last Section of the Handbook covers specific machining practices for 23 different metal systems, including all structural alloy systems, and relates the latest information on such topics as powder metals, metal-matrix composites, and honeycomb structures. Machining parameters (speeds, feeds, depth-of-cut, etc.) and the influence of microstructure on machinability are described in detail. Coverage includes difficult-to-machine aerospace alloys and high-silicon cast aluminum alloys, as well as materials such as beryllium and uranium that require special considerations during machining. Finally, an article on machinability test methods examines various types of tests used to study cutting tool and workpiece machining characteristics. Much of the credit for the content and organization of this Handbook must be given to the Steering Committee that worked with the ASM staff during the early stages of the project. This group includes Professor George E. Kane, Lehigh University; Dr. William P. Koster, Metcut Research Associates Inc.; Dr. Ranga Komanduri, National Science Foundation; Dr. Richard P. Lindsay, Norton Company; Mr. Gary F. Benedict, Allied-Signal Aerospace Company, Garrett Engine Division; and Mr. Michael E. Finn, Stelco Inc. We are also indebted to the officers of the Society of Carbide and Tool Engineers for their assistance in the planning of the Volume. Finally, we gratefully acknowledge the countless hours of time and expertise loaned to the project by the nearly 200 authors and reviewers. Without the collective efforts of all these individuals, the successful completion of this Handbook would not have been possible. The Editors

Ge n e r a l I n for m a t ion Office r s a n d Tr u st e e s of ASM I n t e r n a t ion a l ( 1 9 8 8 - 1 9 8 9 ) Office r s

• • • •

Richard K. Pitler President and Trustee Allegheny Ludlum Corporation (retired) Klaus M. Zwilsky Vice President and Trustee National Materials Advisory Board Academy of Sciences William G. Wood Immediate Past President and Trustee Kolene Corporation Robert D. Halverstadt Treasurer AIMe Associates

Tr u st e e s

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John V. Andrews Teledyne Allvac Edward R. Burrell Inco Alloys International, Inc.

National

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Stephen M. Copley University of Southern California H. Joseph Klein Haynes International, Inc. Gunvant N. Maniar Carpenter Technology Corporation Larry A. Morris Falconbridge Limited William E. Quist Boeing Commercial Airplane Company Charles Yaker Howmet Corporation Daniel S. Zamborsky Consultant Edward L. Langer Managing Director ASM International

M e m be r s of t h e ASM H a n dbook Com m it t e e ( 1 9 8 8 - 1 9 8 9 ) • • • • • • • • • • • • • • • • • • • • • • • • •

Dennis D. Huffman (Chairman 1986-; Member 1983-) The Timken Company Roger J. Austin (1984-) ABARIS Roy G. Baggerly (1987-) Kenworth Truck Company Robert J. Barnhurst (1988-) Noranda Research Centre Peter Beardmore (1986-) Ford Motor Company Hans Borstell (1988-) Grumman Aircraft Systems Gordon Bourland (1988-) LTV Aerospace and Defense Company Robert D. Caligiuri (1986-) Failure Analysis Associates Richard S. Cremisio (1986-) Rescorp International, Inc. Gerald P. Fritzke (1988-) Metallurgical Associates J. Ernesto Indacochea (1987-) University of Illinois at Chicago John B. Lambert (1988-) Fansteel Inc. James C. Leslie (1988-) Advanced Composites Products and Technology Eli Levy (1987-) The De Havilland Aircraft Company of Canada Arnold R. Marder (1987-) Lehigh University John E. Masters (1988-) American Cyanamid Company L.E. Roy Meade (1986-) Lockheed-Georgia Company Merrill L. Minges (1986-) Air Force Wright Aeronautical Laboratories David V. Neff (1986-) Metaullics Systems Dean E. Orr (1988-) Orr Metallurgical Consulting Service, Inc. Ned W. Polan (1987-) Olin Corporation Paul E. Rempes (1986-) Williams International E. Scala (1986-) Cortland Cable Company, Inc. David A. Thomas (1986-) Lehigh University Kenneth P. Young (1988-) AMAX Research & Development

Pr e viou s Ch a ir m e n of t h e ASM H a n dbook Com m it t e e • • • • • • • • • • • • • •

R.S. Archer (1940-1942) (Member, 1937-1942) L.B. Case (1931-1933) (Member, 1927-1933) T.D. Cooper (1984-1986) (Member, 1981-1986) E.O. Dixon (1952-1954) (Member, 1947-1955) R.L. Dowdell (1938-1939) (Member, 1935-1939) J.P. Gill (1937) (Member, 1934-1937) J.D. Graham (1966-1968) (Member, 1961-1970) J.F. Harper (1923-1926) (Member, 1923-1926) C.H. Herty, Jr. (1934-1936) (Member, 1930-1936) J.B. Johnson (1948-1951) (Member, 1944-1951) L.J. Korb (1983) (Member, 1978-1983) R.W.E. Leiter (1962-1963) (Member, 1955-1958, 1960-1964) G.V. Luerssen (1943-1947) (Member, 1942-1947) G.N. Maniar (1979-1980) (Member, 1974-1980)

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J.L. McCall (1982) (Member, 1977-1982) W.J. Merten (1927-1930) (Member, 1923-1933) N.E. Promisel (1955-1961) (Member, 1954-1963) G.J. Shubat (1973-1975) (Member, 1966-1975) W.A. Stadtler (1969-1972) (Member, 1962-1972) R. Ward (1976-1978) (Member, 1972-1978) M.G.H. Wells (1981) (Member, 1976-1981) D.J. Wright (1964-1965) (Member, 1959-1967)

St a ff ASM International staff who contributed to the development of the Volume included Kathleen M. Mills, Manager of Editorial Operations; Joseph R. Davis, Senior Editor; Steven R. Lampman, Technical Editor; Theodore B. Zorc, Technical Editor; Heather J. Frissell, Editorial Supervisor; George M. Crankovic, Assistant Editor; Alice W. Ronke, Assistant Editor; Karen Lynn O'Keefe, Word Processing Specialist; and Jeanne Patitsas, Word Processing Specialist. Editorial assistance was provided by Lois A. Abel, Robert T. Kiepura, Penelope Thomas, and Nikki D. Wheaton. The Volume was prepared under the direction of Robert L. Stedfeld, Director of Reference Publications. Con ve r sion t o Ele ct r on ic File s ASM Handbook, Volume 16, Machining was converted to electronic files in 1999. The conversion was based on the third printing (1997). No substantive changes were made to the content of the Volume, but some minor corrections and clarifications were made as needed. ASM International staff who contributed to the conversion of the Volume included Sally Fahrenholz-Mann, Bonnie Sanders, Marlene Seuffert, Gayle Kalman, Scott Henry, Robert Braddock, Alexandra Hoskins, and Erika Baxter. The electronic version was prepared under the direction of William W. Scott, Jr., Technical Director, and Michael J. DeHaemer, Managing Director. Copyr igh t I n for m a t ion ( for Pr in t Volu m e ) Copyright © 1989 by ASM International All rights reserved No part of this book may be reproduced, stored in a retrieval system, or transmitted, in any form or by any means, electronic, mechanical, photocopying, recording, or otherwise, without the written permission of the copyright owner. First printing, March 1989 Second printing, March 1995 Third printing, March 1997 This book is a collective effort involving hundreds of technical specialists. It brings together a wealth of information from worldwide sources to help scientists, engineers, and technicians solve current and longrange problems. Great care is taken in the production of this Reprint, but it should be made clear that NO WARRANTIES, EXPRESS OR IMPLIED, INCLUDING, WITHOUT LIMITATION, WARRANTIES OF MERCHANT-ABILITY OR FITNESS FOR A PARTICULAR PURPOSE, ARE GIVEN IN CONNECTION WITH THIS PUBLICATION. Although this information is believed to be accurate by ASM, ASM cannot guarantee that favorable results will be obtained from the use of this publication alone. This publication is intended for use by persons having technical skill, at their sole discretion and risk. Since the conditions of product or material use are outside of ASM's control, ASM assumes no liability or obligation in connection with any use of this information. No claim of any kind, whether as to products or information in this publication, and whether or not based on negligence, shall be greater in amount than the purchase price of this product or publication in respect of which damages are claimed. THE REMEDY HEREBY PROVIDED SHALL BE THE EXCLUSIVE AND SOLE REMEDY OF BUYER, AND IN NO EVENT SHALL EITHER PARTY BE LIABLE

FOR SPECIAL, INDIRECT OR CONSEQUENTIAL DAMAGES WHETHER OR NOT CAUSED BY OR RESULTING FROM THE NEGLIGENCE OF SUCH PARTY. As with any material, evaluation of the material under enduse conditions prior to specification is essential. Therefore, specific testing under actual conditions is recommended. Nothing contained in this book shall be construed as a grant of any right of manufacture, sale, use, or reproduction, in connection with any method, process, apparatus, product, composition, or system, whether or not covered by letters patent, copyright, or trademark, and nothing contained in this book shall be construed as a defense against any alleged infringement of letters patent, copyright, or trademark, or as a defense against liability for such infringement. Comments, criticisms, and suggestions are invited, and should be forwarded to ASM International. Libr a r y of Con gr e ss Ca t a login g- in - Pu blica t ion D a t a ( for Pr in t Volu m e ) Metals handbook. Vol. 16: Prepared under the direction of the ASM International Handbook Committee. Includes bibliographies and indexes. Contents: v. 1. Properties and selection--v. 2. Properties and selection--nonferrous alloys and pure metals--[etc.]--v. 16. Machining I. Metals--Handbooks, manuals, etc. I. ASM Handbook Committee. II. ASM International. Handbook Committee. TA459.M43 1978 669 78-14934 ISBN 0-87170-007-7 (v. 1) SAN 204-7586

I n t r odu ct ion t o M a ch in in g Pr oce sse s J. T. Black, Auburn Universit y

I n t r odu ct ion MACHINING is a term that covers a large collection of manufacturing processes designed to remove unwanted material, usually in the form of chips, from a workpiece. Machining is used to convert castings, forgings, or preformed blocks of metal into desired shapes, with size and finish specified to fulfill design requirements. Almost every manufactured product has components that require machining, often to great precision. Therefore, this collection of processes is one of the most important of the basic manufacturing processes because of the value added to the final product. By the same token, machining processes are often the most expensive. The majority of industrial applications of machining are in metals. Although the metal cutting process has resisted theoretical analysis because of its complexity, the application of these processes in the industrial world is widespread. Machining processes are performed on a wide variety of machine tools. Figure 1 shows an example of a machine tool--a dual-turret numerically controlled (NC) lathe. Workpieces are held in workholding devices, such as a three-jaw chuck. The tools used to cut metal are in the turrets. Other examples of basic machine tools are milling machines, drill presses, grinders, shapers, broaching machines, and saws.

Fig. 1 A dual- t urret NC t urning cent er wit h 16 t ool st at ions. Court esy of Cincinnat i Milacron

Each of the basic machine tool types has many different configurations. Lathes, for example, may be engine lathes, turret lathes, tracer lathes, or automatic-screw machines. Lathes have followed the trend of other machine tools, and NC lathes can now be routinely purchased. The primary chip formation processes are listed below, with alternative versions in parentheses. Each process is performed on one or more of the basic machine tools. For example, drilling can be performed on drill presses, milling machines, lathes, and some boring machines: • • • • • •

Turning (boring, facing, cutoff, taper turning, form cutting, chamfering, recessing, thread cutting). Shaping (planing, vertical shaping) Milling (hobbing, generating, thread milling) Drilling (reaming, tapping, spot facing, counterboring, countersinking) Sawing (filing) Abrasive machining (grinding, honing, lapping)

•

Broaching (internal and surface)

Processes can be combined into multiple-capability machines, known as machining centers. The machining center shown in Fig. 2 is capable of performing the machining processes normally performed on a milling machine, drilling machine, and a boring mill and is numerically controlled. The position and velocity of the tool with respect to the work is under feedback control. Different tools can be automatically inserted into the spindle as needed to do different machining processes. The horizontal spindle machine shown in Fig. 2 was one of the first NC machining centers to be able to change workpiece pallets.

Fig. 2 Num erically cont rolled m achining cent er t hat can change workpieces as well as cut t ing t ools. Court esy of Kearney and Trecker Corporat ion

For each of the basic machine tool types, there are many different kinds of workholders, cutting tools, and cutting tool holders, resulting in a rather formidable list of equipment and processes. In this Volume, a Section entitled "Fundamentals of the Machining Process" is presented first, with the intent of putting these processes into perspective and helping the reader to understand the problems associated with using machining processes in the manufacture of products.

Ove r vie w of M a ch in in g Pr oce ss Va r ia ble s Metal cutting processes can be viewed as consisting of independent (input) variables, dependent variables, and independent-dependent interactions or relationships. The engineer or machine tool operator has direct control over the input variables and can specify or select them when setting up the machining process. Several input variables are described below. Figure 3 summarizes the input/output relationships associated with metal cutting.

Fig. 3 I nput / out put relat ionships in m et al cut t ing ( m achining)

I n de pe n de n t I n pu t Va r ia ble s W or k pie ce M a t e r ia l. The metallurgy and chemistry of the workpiece can either be specified or is already known. Quite often, a material is selected for a particular application chiefly because it machines well. Cast iron and aluminum, for example, are known to machine easily. Other metals, such as stainless steel or titanium, are difficult to machine. They often have large cutting forces or poor surface finishes, which can result in short cutting tool life, yet these metals are selected to meet other functional design criteria. Machining practice for specific workpiece materials are reviewed in the Section "Machining of Specific Metals and Alloys" in this Volume. St a r t in g Ge om e t r y. The size and shape of the workpiece may be dictated by preceding processes (casting, forging,

forming, and so forth) or may be selected from standard machining stock (for example, bar stock for screw machines). Usually this variable directly influences the machining process or processes that are selected, as well as the depths of cut. Spe cific M a ch in in g Pr oce sse s. The selection of machining processes required to convert the raw material into a finished product must be based on the geometry of the part (size and shape, rotational or non-rotational), the required finishes and tolerances, and the quantity of the product to be made. Machining processes can be grouped into three broad categories. These include traditional chip formation processes, abrasive machining processes, and nontraditional machining processes.

Ch ip For m a t ion Pr oce sse s. As described earlier, there are seven basic chip formation processes: turning, shaping,

milling, drilling, sawing, broaching, and abrasive machining. The equipment and principles of operation associated with each of these processes (with the exception of abrasive machining, which is treated separately) are described in the Section titled "Traditional Machining Processes" in this Volume. Abr a sive m a ch in in g is the basic process by which chips are formed by very small cutting edges that are integral parts

of abrasive particles. The principles of abrasive machining, the fundamental differences between metal cutting and grinding, and the abrasives and equipment used for abrasive machining operations are described in the Section "Grinding, Honing, and Lapping" in this Volume. N on t r a dit ion a l M a ch in in g Pr oce sse s. Machining processes that involve compression/shear chip formation have a

number of inherent disadvantages. These include: • • • • •

High costs incurred with chip formation (high energy output and chip removal, disposal, and/or recycling) Heat buildup that often results in workpiece distortion High forces that create problems in holding the workpiece and which can also cause distortion Undesirable cold working and residual stresses in the workpiece that often necessitate further processing to remove the harmful effects Limitations as to the size and delicacy of the workpiece

In order to avoid these limitations, nontraditional machining processes are increasingly being used. Nontraditional methods usually do not produce chips or a lay pattern in the surface and often involve new energy modes (see the Section "Nontraditional Machining Processes" in this Volume). Volumetric material removal rates, however, are much lower than with traditional machining processes. Tool M a t e r ia ls. The three most common cutting tool materials currently in use for production machining operations are

high-speed steel (HSS), both in wrought and powder metallurgy (P/M) form; carbides; and coated tools. Cubic boron nitride (CBN), ceramics, and diamonds are also being widely employed. Generally speaking, HSS is used for generalpurpose tools, for tools of complex design or for tools used when cutting speeds are more modest. Carbide and ceramic tool materials, which can operate at faster cutting speeds, come in a wide variety of grades and geometries. Titanium nitride and titanium carbide coatings for HSS and carbides are now commonplace. Selection of a tool material that provides reliable service while fulfilling the functional requirements is still an art. The harder the tool material, the better it can resist wear at faster cutting speeds. The faster the cutting speed, the higher the cutting temperature and the shorter the tool life. Retention of hardness at elevated temperatures as well as long tool life are desirable characteristics in cutting tools. See the Section "Cutting Tool Materials" in this Volume for descriptions of the processing, properties, and applications associated with the aforementioned materials. Cu t t in g Pa r a m e t e r s. For every machining operation, it is necessary to select a cutting speed, a feed, and a depth of

cut. Many factors impinge on these decisions because all of the dependent variables are influenced by them. Proper selection of variables also depends on the other input variables that have been selected; that is, the total amount of material to be removed, the workpiece and tool materials, and the machining process or processes. These need to be selected before preliminary choices for speed, feed, and depth of cut can be made. Tool Ge om e t r y. Cutting tools are usually designed to accomplish specific operations, and thus the tool geometry (angles) is selected to accomplish specific machining functions. Generally speaking, large rake and clearance angles are preferred, but they are possible only on HSS tools. Tools made from carbides, ceramics, and other very hard materials must be given small tool angles, which keep the tool material in compression during machining and thereby avoid tensile failure and brittle fractures of the tool. The greater the precision required of the process, the better the geometry of the cutting edge itself must be. W or k h oldin g D e vice s. Workpieces are located (held in specific position with respect to the tools) and clamped in

workholding devices in or on the machine tools. For every machine tool, there are many different kinds of workholding devices, ranging from general-purpose vises to specifically designed jigs and fixtures (see the article "Proper Fixturing" in this Volume). The workholding devices are the key to precision manufacturing; thus, the selection (or design and

construction) of the correct workholding devices is every bit as important as the selection of the right cutting tool and machine tool. Cu t t in g Flu ids. The selection of the right cutting fluid for a particular combination of work material and tool material

can mean the difference between success and failure in almost every production machining process. Cutting fluids serve to cool the workpiece, tool, and chips; reduce friction by means of lubrication; carry the chips away from the cutting region; help improve the surface finish; and provide surface protection to the workpiece (a more complete discussion may be found in the article "Metal Cutting and Grinding Fluids" in this Volume). D e pe n de n t Va r ia ble s Dependent variables are determined by the process based on the prior selection of the input or independent variables. Thus, the manufacturing engineer's control over these is usually indirect. The important dependent variables are cutting force and power, size and properties of the finished product, surface finish, and tool wear and tool failure. Cu t t in g For ce a n d Pow e r . To machine metal at a specified speed, feed, and depth of cut, with a specified lubricant,

cutting tool material, and geometry, generates cutting forces and consumes power. A change in any of the variables alters the forces, but the change is indirect in that the engineer does not specify the forces, only the parameters that generate those forces. Forces are important in that they influence the deflections in the tools, the workpieces, and the workholders, which in turn affect the final part size. Forces also play a roll in chatter and vibration phenomena common in machining. Obviously, the manufacturing engineer would like to be able to predict forces (and power) so that he can safely specify the equipment for a manufacturing operation, including the machine tool, cutting tool, and workholding devices. The basic concepts associated with the modeling and understanding of cutting forces and power are explained in the article "Forces, Power, and Stresses in Machining" in this Volume. Size a n d Pr ope r t ie s of t h e Fin ish e d Pr odu ct . Ultimately, the objective of machining is to obtain a machined

surface of desired size and geometry with the desired mechanical properties. Because machining is a localized, plastic deformation process, every machined surface will have some residual deformation (stresses) left in it. These residual stresses are usually tensile in nature and can interact with surface flaws to produce part failure from fatigue or to cause corrosion. In addition, every process has some inherent process variability (variations about average size) that changes with almost all of the input variables. Thus, the manufacturing engineer must try to select the proper levels of input variables to produce a product that is within the tolerance specified by the designer and has satisfactory surface properties. Su r fa ce Fin ish . The final finish on a machined surface is a function of tool geometry, tool material, workpiece material,

machining process, speed, feed, depth of cut, and cutting fluid. Surface finish is also related to the process variability. Rough surfaces have more variability than smooth surfaces. Often it is necessary to specify multiple cuts, that is, roughing and finish cuts, to achieve the desired surface finish, or it may be necessary to specify multiple processes, such as following turning with cylindrical grinding, in order to obtain the desired finish. The effect of various machining processes on surface finish and on the properties of the final products are described in the article "Surface Finish and Surface Integrity" in this Volume. Tool W e a r a n d Tool Fa ilu r e . The plastic deformation and friction inherent in machining generate considerable heat, which raises the temperature of the tool and lowers its wear resistance. The problem is subtle, but significant. As the tool wears, it changes in both geometry and size. A dull cutting edge and change in geometry can result in increased cutting forces that in turn increase deflections in the workpiece and may create a chatter condition. The increased power consumption causes increased heat generation in the operation, which accelerates the wear rate. The change in the size of the tool changes the size of the workpiece. Again, the engineer has only indirect control over these variables. He can select slow speeds, which produce less heat and lower wear rates, but which decrease the production rates because the metal removal rate is decreased. Alternatively, the feed or depth of cut can be increased to maintain the metal removal rate while reducing the speed. Increasing either the feed or depth of cut directly increases the cutting forces. Therefore, while tool life may be gained, some precision may be lost due to increased deflection and chatter. Wear mechanisms, determination of modes of tool failure, and tool life testing are examined in the article "Tool Wear and Tool Life" in this Volume.

Re la t ion s Be t w e e n I n pu t Va r ia ble s a n d Pr oce ss Be h a vior Understanding the connections between input variables and process behavior is important knowledge for the manufacturing engineer. Unfortunately, this knowledge is difficult to obtain. Machining is a unique plastic deformation

process in that it is constrained only by the cutting tool and operates at very large strains and very high strain rates. The tremendous variety in the input variables results in an almost infinite number of different machining combinations. Basically, there are three ways to deal with such a complex situation. Ex pe r ie n ce requires long-term exposure, because knowledge is basically gained by trial and error, with successful

combinations transferred to other, "similar" situations. This activity goes on in manufacturing every time a new material is introduced into the production facility. It took years for industry to learn how to machine titanium. Unfortunately, the knowledge gained through one process may not transfer well to another even though their input variables appear very similar. Ex pe r im e n t s. Machining experiments are expensive, time consuming, and difficult to carry out. Tool life experiments,

for example, are quite commonly done, yet tool life data for most workpiece/tool material combinations are not available. Even when laboratory data have been published, the results are not necessarily transferable to the particular machine tools and cutting tools on the shop floor. Tool life equations are empirically developed from turning experiments in which all input variables except cutting speed are kept constant. The experimental arrangement may limit the mode of tool failure to wear. Such results are of little value on the shop floor, where tools can and do fail from causes other than wear. Th e or ie s. There have been many attempts to build mathematical models of the metal cutting process. Many of the

theories are extensions of the mechanics presented in the following Section, "Fundamentals of the Machining Process." These theories try to predict the direction of the shearing process of metal cutting. These models range from crude, firstorder approximation to complex, computer-based models using finite-element analysis. Recently, some modest successes have been reported in the literature in which accurate predictions of cutting forces and tool wear were made for certain materials. Clearly such efforts are extremely helpful in understanding how the process behaves. However, the theory of plastic deformation of metals (dislocation theory) has not yet been able to predict values for shear stresses and tool/chip interface from the metallurgy and deformation history of the material. Therefore, it has been necessary to devise two independent experiments to determine the shear strength ( s) of the metal at large strains and high strain rates and the sliding friction situation at the interface between the tool and chip (see the article "Mechanics of Chip Formation" in this Volume).

Fu t u r e Tr e n ds The metal cutting process will continue to evolve, with improvements in cutting tool materials and machine tools leading the evolution. More refined coatings on cutting tools will improve tool life and reliability, as will more robust, rigid machine tools. The challenge for machining will involve dealing with the new types of materials that will need to be machined, including aluminum and titanium alloys, alloy steels, and superalloys. These materials, because of improved processing techniques, are becoming stronger and harder and therefore more difficult to machine. The objective should be to design and build cutting tools that have less variability in their tool lives rather than longer tool lives. The increasing use of structural ceramics, high-strength polymers, composites, and electronic materials will also necessitate the use of nontraditional methods of machining. In addition, grinding will be employed to a greater extent than in the past, with greater attention to creep feed grinding and the use of superabrasives (diamond and cubic boron nitride). As the cutting tools improve, the machine tools will become smarter, with on-board computers providing intelligent algorithms interacting with sensory data from the process. Programmable machine tools, if equipped with the proper sensors, are capable of carrying out measurements of the product as it is being produced. These product data will be fed back to the control program, which is then modified to improve the product or corrected for errors. Thus, the machine will be able to make the adjustments necessary to prevent defective products from being produced. The goal of such control programs should be improved quality (designed not to make a defect), rather than optimum speed or lowest cost. Advancements in computer-aided machining processes are discussed in the Section "Machine Controls and Computer Applications in Machining" in this Volume. Another area in which significant advances will be made is the design of workholders that are capable of holding various parts without any downtime for setups. Included in this search for flexible fixtures will be workholding devices that can be changed over by a robot--the same robot used to load or unload parts from the machine.

M e ch a n ics of Ch ip For m a t ion J.T. Black, Auburn Universit y

I n t r odu ct ion THE BASIC MECHANISM involved in metal cutting is that of a localized shear deformation on the work material immediately ahead of the cutting edge of the tool. The relative motion between the tool and the workpiece during cutting compresses the work material near the tool and induces a shear deformation (called the primary deformation), which forms the chip. The chip passes over the rake face of the cutting tool and receives additional deformation (called the secondary deformation) because of the shearing and sliding of the chip against the tool. These two plastic deformation processes have a mutual dependence. The material element that rubs the rake face has been heated and plastically deformed during its passage through the primary shear process; therefore, the secondary process is influenced by the phenomena on the shear plane. At the same time, the shear direction is directly influenced by the rake face deformation and friction processes. The shear direction influences the heating and straining of the chip in the primary process. In terms of metal cutting theory, this means that shear stress and shear direction must be determined simultaneously. Such theoretical analyses are usually based on the mechanics of the process. This article will review the following: • • • •

The fundamental nature of the deformation process associated with machining The principles of the orthogonal cutting model The effect of workpiece properties on chip formation The mechanics of the machining process

Additional information on the modeling and analysis of chip formation can be found in the article "Forces, Power, and Stresses in Machining," which immediately follows in this Section.

Fu n da m e n t a l M e ch a n ism of M e t a l D e for m a t ion Cu t t in g M ode ls. Before the mechanics of machining are presented, a brief discussion of the fundamental nature of the

deformation processes is helpful in understanding the assumptions that accompany the mechanics. The machining geometry can be simplified from the three-dimensional (oblique) geometry, which typifies most industrial processes, to a two-dimensional (orthogonal) geometry. Figure 1 compares the oblique and the orthogonal cutting geometries. Orthogonal machining can be obtained in practice by: • •

End cutting a tube wall by turning (Fig. 1b) Machining a plate as shown in Fig. 2

Oblique cutting is obtained when the cutting edge and the cutting motion are not perpendicular to each other. Because the orthogonal case is more easily modeled, it will be used in this article to describe the deformation process.

Fig. 1 Com parison of oblique and ort hogonal geom et ry m achining. ( a) Three- force oblique m achining. Fc is t he prim ary cut t ing force, Ff is t he feed force, and Fr is t he radial or t hrust force. ( b) Two- force ort hogonal m achining. Fc is t he m easured cut t ing force, and Ft is t he feed ( t angent ial) force. A t ube- cut t ing applicat ion is shown; t he cut t ing edge of t he t ool is perpendicular t o t he direct ion of m ot ion. ( c) For ort hogonal cut t ing, t he shear area, As occurs for a shear angle

, widt h of cut w, and feed t .

Fig. 2 Developm ent of t he shear front - lam ella st ruct ure. As shown by t his ort hogonal geom et ry, shear deform at ion evolves from a radial com pression zone. See Fig. 5 for an explanat ion of t he effect s of shear deform at ion on area p- q- r- s.

In the orthogonal cutting of a tube (Fig. 1b), the width of the cut is equal to the thickness of the tube wall, w. The direction of shear is specified by the shear angle. The cross-sectional area of the chip is given by tc · wc, where tc is chip thickness and wc is the width of the chip. The cutting edge of the tool is perpendicular to the feed direction. The measured horizontal cutting force, Fc, is the force in the direction of the cutting velocity (or cutting speed). The force in the direction of the feed (vertical or tangential) and perpendicular (orthogonal) to Fc is denoted by Ft. With this twodimensional model of chip formation, the influence of the most critical elements of the tool geometry (rake angle, , and the edge radius of the tool) and the interactions that occur between the tool and the chip can be more easily examined. Sh e a r Zon e . Basically, the chip is formed by a localized shear process that takes place over very narrow regions.

Classically called the shear zone or shear plane, this deformation evolves out of a radial compression zone that travels ahead of the shear process as the tool passes over the workpiece (Fig. 2). Like all plastic deformations, this radial compression zone has an elastic compression region that converts to a plastic compression region as the material

approaches the cutting edge. This plastic compression generates dense dislocation tangles and networks in annealed metals. When this work-hardened material reaches the tool, the material shears in the direction of the free surface. Sh e a r Fr on t - La m e lla St r u ct u r e . The shear process itself is a nonhomogeneous (discontinuous) series of shear fronts (or narrow bands) that produce a lamellar structure in the chips. This fundamental structure occurs on the microscale in all metals when they are machined and accounts for the unique behavior of the machining process.

Individual shear fronts (Fig. 2) coalesce into narrow shear bands. The shear bands are very narrow (20 to 200 nm) compared to the thickness of a lamella (2 to 4 m) and account for the large strain and high strain rates that typify this process. These fundamental structures are difficult to observe in normal metal cutting, but can be readily observed in a scanning electron microscope with specially prepared workpieces. Figure 3 shows micrographs from an orthogonal machining experiment performed inside a scanning electron microscope. The fundamental shear front-lamella structure is readily observed. The side of the workpiece has been given a mirror polish so that the shear fronts can be observed. The shear fronts are produced by the activation of many dislocations traveling in waves from the tool tip to the free surface. The lamella represents heavily deformed material that has been segmented by the shear fronts. When machined, all metals deform by this basic mechanism. The shear fronts relieve the applied stress.

Fig. 3 Chip form at ion process viewed inside a scanning elect ron m icroscope. The workpiece is a rect angular plat e of high- purit y gold t hat was polished on t he sides so t hat t he plast ic deform at ion of t he shear process can be readily observed. The boxed area in ( a) , which is shown at a higher m agnificat ion in ( b) , shows t he shear front s, num bered 1 and 2, advancing from t he t ool t ip region t oward t he free surface of t he workpiece. The let t er D indicat es a defect on t he side of t he chip. The arrows indicat e a scrat ch (S) t hat has been sheared. The t ool has been wit hdrawn from t he workpiece. I n ( c) , t he t ool has been reinsert ed and slight ly advanced. This produced addit ional shear on shear front No. 2 and new shear front No. 3. Not e t he m ovem ent of defect D. These shear front s are difficult t o observe unless t he specim en is polished and exam ined in a scanning elect ron

m icroscope.

Chips are sometimes produced with a sawtooth pattern on the top side--the side that did not rub against the tool. This sawtooth pattern is not produced by the shear front-lamella structure but rather by the unloading of the elastic energy stored in the tool and workpiece, which results in chatter and vibration during cutting. The shear front-lamella structure can and does exist without any vibration of the tool or workpiece. If each sawtooth were to be observed in the scanning electron microscope, many fine shear front-lamella structures would be found in each sawtooth region. The geometry of the sawtooth can be changed (even eliminated) by altering the rigidity of the setup or the machine. The shear frontlamella structure is fundamental to, and characteristic of, the plastic deformation process itself; therefore, it is relatively invariant with respect to cutting parameters and certainly cannot be eliminated.

Or t h ogon a l M a ch in in g Fu n da m e n t a ls Orthogonal machining setups are used to model oblique machining processes. Processes such as turning, drilling, milling, and shaping are all three-force, or oblique, cutting methods. However, the orthogonal model shown in Fig. 4 is an excellent illustration of the behavior of oblique processes without the complications of the third dimension.

Fig. 4 Schem at ics of ort hogonal m et al cut t ing m echanics. ( a) Ort hogonal m odel. t , uncut chip t hickness ( feed or dept h or cut ) ; t c, chip t hickness;

, shear angle;

, back rake angle;

, clearance angle;

, edge angle [

= 90 - ( + ) ] . ( b) Velocit y t riangle. Vs, shear velocit y; Vc, chip velocit y; V, cut t ing velocit y. ( c) Chip freebody diagram . F, frict ion force; N, norm al t o frict ion force; Fs, shear force; Fn , norm al t o shear force; Fc, cut t ing force; Ft , t angent ial force; R, result ant force

Ch ip Ra t io. As described earlier in this article, orthogonal machining can be accomplished by machining a plate or can

be approximated by cutting the end of a tube wall in a turning setup. For the purposes of modeling, the following are assumed: The shear process is a plane, the cutting edge is perfectly sharp, and there is no friction contact between the flank of the tool and the workpiece surface. Because plane-strain conditions are assumed, the chips are assumed to have no side flow (w = wc, Fig. 1c), and the cutting velocity is constant. The shear process occurs at angle for a tool with back rake angle . The chip has velocity Vc and makes contact with the rake face of the tool over length (Fig. 2). Defining the ratio of the uncut chip thickness, t, to the chip thickness, tc, as the chip ratio, r, produces the following:

(Eq 1)

Solving Eq 1 for

yields:

(Eq 2) In practical tests, the average chip thickness can be obtained by carefully measuring the length L and the weight W of a piece of a chip. Then:

(Eq 3)

where is the density of the work material and t is the feed or uncut chip thickness. Chip thickness is usually greater than the depth of cut, t, and is constrained by the rake face of the cutting tool. Sh e a r An gle . There are numerous other ways to measure or compute the shear angle, both during (dynamically) the

cutting process and after (statically) it has been halted. The shear angle can be measured statically by instantaneously interrupting the cut through the use of quick-stop devices. These devices disengage the cutting tool from the workpiece while cutting is in progress, leaving the chip attached to the workpiece. Optical microscopy and scanning electron microscopy are then used to observe the shear angle. High-speed motion pictures have also been used to observe the process at frame rates as high as 30,000 frames per second. More recently, machining stages have been built that allow the process to be performed inside a scanning electron microscope and recorded on video-tape for high-resolution, highmagnification examination of the deformation process. The micrographs shown in Fig. 3 were created in this manner. This technique has been used to measure the velocity of the shear fronts, Vc, during cutting, thus verifying experimentally that the vector sum of V and Vc equals Vs (Fig. 4b). For constancy of volume, it was observed that:

(Eq 4) Equation 4 indicates that the chip ratio (and therefore the shear angle) can be determined dynamically if a reliable means of measuring chip velocity can be found. Thus, one could determine dynamically for a known tool geometry. Therefore, cutting forces can be dynamically predicted, an important consideration in adaptive control machining (see the article "Adaptive Control" in this Volume). Velocities are also important in power calculations, heat determinations, and vibration analyses associated with chip formation. Sh e a r St r a in . When an area of metal (for example, area p-q-r-s in Fig. 2) passes through the shear process, it is

plastically deformed into a new shape, as shown in Fig. 5. The amount of plastic deformation is related to the shear angle, , and the rake angle, .

Fig. 5 St rain on shear plane,

, versus shear plane angle,

Therefore, the chip undergoes a shear strain,

, of:

, for t hree values of rake angle,

(Eq 5) The meaning of shear strain, as well as of the units in which it is measured, is shown in the inset diagram in Fig. 5. A unit displacement of one face of a unit cube is a shear strain of 1 ( = 1). Figure 5 illustrates the relationship between the shear strain in orthogonal cutting and the shear plane angle for three values of the rake angle. For any rake angle, there is a minimum strain at which the mean chip thickness is equal to the feed (tc = t). For zero rake angle, this occurs at = 45°. The change in shape of a unit cube after it passes through the shear plane for different values of the shear plane angle is shown in the lower diagram in Fig. 5 for a tool with a zero rake angle. The minimum strain at = 45° is apparent from the shape change. The shaded region in Fig. 5 shows the typical values of found in practice. At a zero rake angle, the minimum shear strain is 2. The minimum strain occurs when there is no friction at the tool/chip interface. The minimum strain decreases as the rake angle increases. If the rake angle is too large, the tool is weak and will fracture. Rake angles larger than 30° are seldom used in industry. With carbides and ceramics, the tendency has been to decrease the rake angle to make the tools more robust, allowing these harder but less tough tool materials to be used. Therefore, even under optimum cutting conditions, chip formation involves very severe plastic deformation, resulting in considerable work hardening and structural change. Metals and alloys lacking in ductility periodically fracture on the shear plane, producing discontinuous chips (see the section "Effect of Work Material Properties" in this article). In general, metal cutting strains are quite large compared to other plastic deformation processes, being of the order of 2 to 4 mm/mm (2 to 4 in./in.). However, this large strain occurs over very narrow regions (the shear band), which results in extremely high shear strain rates, typically of the order of 104 to 108 mm/mm (104 to 108 in./in.). This strain rate can be estimated from:

(Eq 6) where d is the thickness of the shear bands. This combination of large strains and high strain rates operating within a process constrained only by the workpiece and the tool (actually, the deformation interface at the rake face of the tool) causes great difficulties in theoretical analyses of the process.

Effe ct of W or k M a t e r ia l Pr ope r t ie s Pr in cipa l Ch ip Type s. The properties of the work material control chip formation. Work material properties include

yield strength, shear strength under compressive loading, strain-hardening characteristics, friction behavior, hardness, and ductility. As noted in the section "Shear Strain" in this article, work material ductility is an important factor. Highly ductile materials not only permit extensive plastic deformation of the chip during cutting, which increases work, heat generation, and temperature, but also result in longer, continuous chips that remain in contact longer with the tool face, thus causing more frictional heat. Chips of this type are severely deformed and have a characteristic curl. On the other hand, some materials, such as gray cast iron, lack the ductility necessary for appreciable plastic chip formation. Consequently, the compressed material ahead of the tool can fail in a brittle manner anywhere ahead of the tool, producing small fragments. Such chips are termed discontinuous or segmented (Fig. 6).

Fig. 6 Three charact erist ic t ypes of chips. ( a) Discont inuous. ( b) Cont inuous. ( c) Cont inuous wit h built - up edge

The cutting parameters also influence chip formation. Cutting parameters include tool materials, tool angles, edge geometries (which change due to wear, cutting speed, feed, and depth of cut), and the cutting environment (machine tool deflections, cutting fluids, and so on). Further complications result from the formation of the built-up edge on the cutting tool. A bu ilt - u p e dge is work material that is deposited on the rake face near the cutting edge (Fig. 6c). It is the product of the localized high temperature and extreme pressure at the tool/chip interface. The work material adheres to the cutting edge of the tool (similar to a dead-metal zone in extrusion). Although this material protects the cutting edge, it also modifies the geometry of the tool. Built-up edges are not stable and will slough off periodically, adhering to the chip or passing under the tool and adhering to the machined surface. Built-up edge formation can often be eliminated or minimized by reducing the depth of the cut, increasing the cutting speed, using positive rake tools, or applying a coolant, but these techniques greatly increase the complexity of the chip formation process analysis.

M e ch a n ics of M a ch in in g Orthogonal machining has been defined as a two-component force system, while oblique cutting involves a three-force situation. Figure 4(c) shows a free body diagram of a chip that has been separated at the shear plane. The resultant force R consists of the friction force, F, and the normal force, N, acting on the tool/chip interface contact area (length times width w). The resultant force R' consists of a shear force, Fs, and a normal force, Fn, acting on the shear plane area, As. The forces R and R' are assumed to be equal, opposite, and colinear. Determination of these forces necessitates a third set that can be measured. A dynamometer, mounted in the workholder or the toolholder, can be used to measure Fc and Ft. This set has resultant R'', which is equal in magnitude and colinear to the other resultant forces in the diagram. To express the desired forces (Fs, Fn, F, N) in terms of the dynamometer components Fc and Ft and appropriate angles, a circular force diagram is developed in which all six forces are collected in the same force circle. This is shown in Fig. 7. In Fig. 7, is the angle between the normal force, N, and the resultant force R. It is used to describe the friction coefficient, , on the tool/chip interface area, which is defined as F/N so that:

(Eq 7) The friction force, F, and its normal force, N, can be shown to be:

where

F = Fc sin

+ Ft cos

N = Fc cos

-Ft sin

(Eq 8) (Eq 9)

R=( When the back rake angle,

+

)1/2

(Eq 10)

, is zero, then F = Ft and N = Fc.

Fig. 7 Circular force diagram for ort hogonal chip form at ion

The forces parallel and perpendicular to the shear plane can be shown (from the force circle diagram) to be:

Fs = Fc cos

-Ft sin

(Eq 11)

Fn = Fc sin

-Ft cos

(Eq 12)

The shear force, Fs, is of particular interest because it is used to compute the shear stress on the shear plane. The shear stress, s, is defined as:

(Eq 13)

where As = tw/sin

.

Recalling that t is the depth of the cut and w is the width of the workpiece, the shear stress is:

(Eq 14) For a given polycrystalline metal, this shear stress is a material constant that is not sensitive to variations in cutting parameters, tool material, or the cutting environment. Some researchers are attempting to derive (predict) the shear stress, s, and the shear direction from dislocation theory, but this has not yet been accomplished. Correlations of the shear stress with metallurgical measures, such as hardness or dislocation stacking fault energy, have been useful in these efforts. The cutting force, Fc, is the dominant force in this system, and it is important to understand how it varies with changes in the cutting parameters. As shown in Fig. 8, the cutting forces typically double when the feed or depth of cut is doubled, but remain constant when speed is increased. In addition, the forces will increase (and change direction) when the rake angle is reduced. More detailed information on the determination of cutting forces can be found in the article "Forces, Power, and Stresses in Machining" in this Section.

Fig. 8 General relat ionship of ort hogonal cut t ing forces t o prim ary cut t ing param et ers speed ( a) , feed ( b) , and dept h of cut ( c)

For ce s, Pow e r , M a ch in in g

and

St r e sse s

in

Paul H. Cohen, The Pennsylvania St at e Universit y

I n t r odu ct ion THE MODELING AND ANALYSIS of chip formation has been a continuing exercise over the past century. The metal cutting process is a unique and complex production process distinguished by: • • • • •

Large shear strains, usually of the order of 2 to 5 (Ref 1) Exceptionally high shear strain rates, typically from 103 to 105 s-1 with local variations as high as 107 s-1 (Ref 2, 3) The rubbing of the tool flank over a freshly cut surface that is chemically clean and active Many process and tooling parameters with a wide range of settings that can drastically alter the cutting process A large number of metallurgical parameters in the workpiece that can influence its response to the cutting tool

These factors and others make the modeling of metal cutting a difficult task that continues to evolve over time. The models and the discussion presented in this article will attempt to explain the basic concepts of the many complex factors that influence the forces, power, and stresses in machining.

For ce s a n d En e r gy in Or t h ogon a l M a ch in in g Although most production machining processes are oblique (that is, having three component forces), models of the orthogonal (that is, two force) machining of metals are useful for understanding the basic mechanics of machining and can be extended for modeling of the production processes. For ce s. The classical thin zone mechanics was developed for materials that yield continuous chips with a planar shear process coupled with the following assumptions (Ref 4, 5):

• • • •

The tool tip is sharp, and no rubbing occurs between the tool and the workpiece Plane strain conditions prevail (that is, no side spread occurs) The stresses on the shear plane are uniformly distributed The resultant force, R, on the chip is equal, opposite, and colinear to the force R' at the tool/chip interface (Fig. 1)

Fig. 1 The geom et ry ( a) and forces ( b) in ort hogonal cut t ing

The modeling of the orthogonal cutting process defines two regions of deformation (primary and secondary), each described by its own set of orthogonal forces, as shown in Fig. 1(b). Because these force components cannot be directly measured (except for the forces on the rake face of the tool when = 0°), a dynamometer must be used to measure the primary (horizontal) cutting force, Fc, and the tangential (vertical) force, Ft. Thus, the measured forces can be resolved onto the shear plane through the shear angle, , and onto the rake face through the back rake angle, . The shear angle, , is the angle the primary shear plane makes with respect to the horizontal motion of the tool. Although it is possible to observe and measure this angle in special experiments by using high-speed photography or the machining

stages within a scanning electron microscope, thickness, tc, as follows:

is typically computed by using a ratio of the depth of cut, t, to chip

(Eq 1)

The shear strain necessary to shear the work material at this angle,

= tan (

-

, is:

) + cos

(Eq 2)

Analyzing the primary shearing process in Fig. 1, the shear and normal forces on the shear plane can be written as functions of the measured horizontal and vertical (dynamometer) forces and shear angle, as follows:

Fs = Fc cos

- Ft sin

(Eq 3)

Fn = Fc sin

+ Ft cos

(Eq 4)

Similarly, the forces on the rake face can be written as functions of the same measured force components and the tool back rake angle as:

F = Fc sin

+ Ft cos

(Eq 5)

N = Fc cos

- Ft sin

(Eq 6)

The resultant force, R, which acts on the chip and is shown in Fig. 1(b), can be written as the vector sum of the measured forces, the forces acting on the shear plane, or the forces acting on the rake face of the tool. Therefore:

R=(

+

)1/2

(Eq 7)

R=(

+

)1/2

(Eq 8)

R = (F2 + N2)1/2

(Eq 9)

En e r gy of Ch ip For m a t ion . During the cut, the total energy per unit time (or power) can be calculated simply as the product of the primary cutting force, Fc, and the velocity of cut, V. However, because many parameters can be varied in the cutting process that change the total energy consumed, this energy value is typically normalized by dividing by the rate at which material is removed. The material removal rate is calculated by multiplying the area being cut (t · w for the case of the plate of width w shown in Fig. 1a), by the velocity perpendicular to that area at which the material is removed (V in this case). Thus, the energy per unit time, or specific energy, u, can be calculated as:

(Eq 10) The specific energy can be partitioned into four components (Ref 6, 7): • • • •

Shear energy per unit volume, us Friction energy per unit volume, uf Kinetic (momentum) energy per unit volume, um Surface energy per unit volume, ua

The shear energy per unit volume can be calculated by substituting the energy per unit time necessary to shear the material in place of the total energy per unit time in Eq 8. Thus:

(Eq 11) where Vs is the shear velocity (where Vs = V cos /cos( - ), as defined in the article "Mechanics of Chip Formation" in this Section; see the discussion of shear angle measurement during orthogonal machining). The shear energy per unit volume is the largest of the four components, typically representing more than 75% of the total. The friction energy per unit volume is consumed as the chip slides on the rake face of the tool. This component is very sensitive to cutting velocity and can be written as:

(Eq 12) where Vc is the velocity of the chip as it flows over the tool (Vc = V sin "Mechanics of Chip Formation" in this Section).

/cos(

-

), as defined in Eq 4 of the article

The kinetic (momentum) energy per unit volume required to accelerate the chip is generally neglected but takes on increasing importance with very high speed machining. It can be written as:

(Eq 13) where Fm is the momentum force = strain.

V2tw

sin

, where

is the density of the material being cut and

is the shear

Additional energy is required to produce a new uncut surface. The surface energy per unit volume needed to create this new surface can be written as:

(Eq 14) where T is the surface energy of the material being cut. This component is also generally neglected. Therefore, for most machining applications, the specific energy can be accurately estimated as:

u

us + uf

(Eq 15)

except at high speeds (above 900 to 1200 m/min, or 3000 to 4000 sfm) for which the kinetic specific energy should be included. Specific energies can be used to calculate the power per unit volume per unit time (specific horsepower) and are readily accessible for most engineering materials. They are a good measure of the difficulty involved in machining a particular material.

St r e ss D ist r ibu t ion s in M e t a l Cu t t in g High shear and normal stresses occur both in the primary shear plane and on the rake face of the tool. This region of friction or secondary shear is critical in understanding the process mechanics and the wear of cutting tools. St r e sse s in t h e W or k pie ce . As discussed in the article "Mechanics of Chip Formation" in this section, the

fundamental mechanism of chip formation requires prior work hardening before the workpiece material reaches the shear

plane. Experimental results have shown that the material is elastically deformed at distances sufficiently far from the tool tip. As the material approaches the tool, the compressive stresses will begin to plastically deform the workpiece material as shown in Fig. 2. Behind the tool tip, the stresses will be tensile.

Fig. 2 St resses in t he workpiece

The distance of the elastic-plastic boundary from the tool tip will depend on tooling parameters, cutting parameters, and workpiece material properties. In particular, the amount of prior strain hardening and the ability of the workpiece material to work harden will alter the magnitude of the stresses in the workpiece and will affect the placement of the elastic-plastic boundary. Materials with little prior strain hardening will extend their boundaries farther out from the tool tip. St r e sse s on t h e Sh e a r Pla n e . Consistent with the assumptions in the section "Forces" in this article, the shear plane

is generally modeled to have uniform distributions of both shear and normal forces over its entire area. The shear area, As, is the area of cut (Ac = t · w) inclined at the shear angle . Thus, the shear area is:

(Eq 16) The shear stress on the shear plane can then be calculated as follows:

(Eq 17)

and the normal stress can be computed similarly as:

(Eq 18)

Therefore, the stresses rely only on measured cutting forces (Fc and Ft), the geometry of the cut (t and w), and the deformation geometry ( ). The shear stress, variety of metals.

s,

takes on a constant value for a particular material. Figure 3 and Table 1 provide typical values for a

Table 1 Shear stresses and specific horsepowers of selected engineering materials Material

Magnesium 1100 aluminum alloy 6061-T6 aluminum alloy 2024-T4 aluminum alloy Copper 60-40 brass 65-35 brass 70-30 brass AISI 1020 steel

Shear stress, psi 28,000 16,700 35,722 50,000 44,850 47,000 50,000 56,940 61,500

AISI 1112 steel Type 304 stainless steel Titanium

63,500 105,000 173,500

Specific horsepower, hp/in.3/min 0.17 ... 0.35 0.46 0.78

Hardness, HB ... ... ... ... ...

0.59 0.58 0.67 0.5 1.1-1.9 1.9

... 150-175 176-200 150-175 ... ...

Fig. 3 Shear st ress variat ion wit h Brinell hardness for ferrous and nonferrous m et als. Source: Ref 9

St r e ss D ist r ibu t ion s on t h e Ra k e Fa ce . The nature of the tool/chip interface and the distribution of the shear and normal stresses are critical in understanding the cutting process and the performance of cutting tools. The high stresses, coupled with the high temperatures and large strains in the chip adjacent to the tool face, make the secondary shearing process difficult to model. Un ifor m St r e sse s on t h e Ra k e Fa ce . The classical analysis of the forces and stresses on the rake face assumes that

Coulombic sliding friction is present and that the stresses are uniformly distributed. Therefore, the coefficient of sliding friction is simply the frictional force, F, divided by the normal force, N, acting on the rake face. Thus:

(Eq 19)

The coefficient of friction is velocity dependent, with increasing speeds yielding lower friction. The area of contact on the tool/chip interface is the product of the width of cut, w, and the length of sliding contact, , as illustrated in Fig. 1. Thus, the area of sliding contact on the rake face is:

Af = w · and the shear stress at the interface can be calculated as:

(Eq 20)

(Eq 21)

Analogously, the normal stress on the rake face can be written as:

(Eq 22)

These models have been found to be useful approximations of the behavior of the chip as it slides over the tool. However, there is a large body of experimental evidence to suggest that the stresses are not uniformly distributed on the rake face. N on u n ifor m St r e ss D ist r ibu t ion s on t h e Ra k e Fa ce . The body of experimental evidence indicating the nonuniformity of the stresses on the rake face is extensive, using a wide variety of experimental techniques and observations. Perhaps the simplest observation supporting this conclusion is the transfer of workpiece material to the tool as observed by the unaided eye, light microscope, or scanning electron microscope (Ref 10). The transfer of this material does not occur over the entire contact area, but near the tip of the tool. Experiments utilizing photography through transparent sapphire tools (Ref 11), photoelastic tools (Ref 12), quick-stop devices to observe metal flow in the chip (Ref 13), and other techniques have revealed the nonuniformity of the stresses.

Quick-stop devices that separate the tool from the chip freeze the flow pattern of the material in the chip. Such studies have revealed two major regions on the rake face with respect to flow. When polished and etched, it is clear from the flow lines that the material near the tool tip is seized by the tool. This can be shown by the flow lines in Fig. 4, which run parallel to the tool face. The rest of the contact area exhibits sliding contact.

Fig. 4 Flow lines in a chip

This concept of seizure is quite different from the standard notions of sliding friction. Because of the high interface temperatures and pressures, the material adjacent to the tool surface is almost stationary, and relative shearing takes place in the chip. As originally developed by Zorev (Ref 14) and consistent with the empirical results presented, the stresses on the rake face are inherently nonlinear, as shown in Fig. 5. The normal stress is assumed to take on a maximum value, max, at the tip of the tool; the stress then decreases as a power function of the distance from the tool tip to the point at which the chip leaves the tool (Fig. 5). The shear stress is constant in the region of seizure and then decreases as a power function to the point at which the chip leaves the tool.

Fig. 5 Model of st ress dist ribut ion on t ool during cut t ing. Source: Ref 14

The normal stress on the rake face is defined by: f

=

max(x/

)n

(Eq 23)

where max is the maximum normal stress at the tool tip (or x = ), is the total length of contact of the chip on the tool, x is the distance from the point at which the chip leaves the tool to the point of interest, and n is the exponent. The normal force can be obtained by integrating the normal stress over the area of contact on the tool face:

(Eq 24)

The shear stress is more complicated to evaluate because the behavior of the chip material as it passes over the tool varies along the rake face. The region of seizure close to the tool tip must be modeled differently from the region of sliding (Coulombic) friction. Over the region of seizure ( f x ), the shear stress has a constant value, , because the chip material shears internally, as illustrated by the flow lines in Fig. 4. Over the sliding region, the shear and normal stresses are related by:

= =

f

n max(x/ )

(Eq 25)

Thus, the shear stress over the entire face can be conveniently expressed as:

(Eq 26)

To determine the friction force, F, on the rake face, the shear stress given in Eq 26 must be integrated over the area of contact. This yields:

(Eq 27)

where

s

is the length of seizure (that is, -

f).

Although such models are useful in understanding the process, it is difficult to determine the lengths associated with seizure and sliding. Typically, a tool is ground to restrict the total length of contact and the length of seizure determined by the flow lines, as detailed previously. From this simple orthogonal model, it is clear that the shear strength of the chip material, the relative amount of seizure, and other parameters will significantly alter the machining forces. Increasing shear strengths and lengths of seizure (for constant contact length) will increase the cutting forces. Therefore, the use of tool materials with less propensity for chip seizure or the use of lubricants that decrease s will lower the cutting forces accordingly.

Pow e r Con su m pt ion in Pr odu ct ion Pr oce sse s Although one may wish to describe the energy per unit volume needed to form the chip, machine tools are typically rated in terms of power. Unit (or specific) power values can be calculated by dividing the power input to the process, FcV, by the volumetric rate at which material is removed and then dividing this quantity by 33,000 to convert to horsepower. The specific power, Ps, is a measure of the difficulty involved in machining a particular material and can be used to estimate the total cutting power, P. Typical specific horsepower values are given in Table 1. The specific power is the power required to remove a unit volume per unit time. Therefore, the specific and total powers are related as follows:

P = Ps · MRR

(Eq 28)

where MRR is the material removal rate, or volume of material removed per unit time. The material removal rate can be computed as the uncut area multiplied by the rate at which the tool is moved perpendicular to the uncut area. As previously determined for a plate, the material removal rate is the uncut area, t · w, multiplied by the velocity of the tool, V. Thus, the cutting parameters and machine tool kinematics define the material removal rate. There are many standard sources for specific power values for a variety of materials. Unfortunately, machine tools are not completely efficient. Losses due to component wear, friction, and other sources prevent some power from reaching the tool. Therefore, the gross power, Pg, needed by the motor can be defined as:

(Eq 29) where is the efficiency of the machine.

Pow e r in Tu r n in g. As with the plate, the total power required in a turning operation can be calculated as P = Ps · MRR.

However, the material removal rate must be redefined for turning. Consider the turning operation illustrated in Fig. 6(a), in which a billet of diameter D is turned with depth of cut d to diameter D1. The billet is rotated at N revolutions per minute, while the tool is fed at fr units (millimeters or inches) per revolution, which can be set directly on the machine. Recommended cutting speeds (in meters or feet per minute) are generally available from handbooks and can be converted to rotational speed where V = DN. Suggested feeds are also available.

Fig. 6 Set ups for t urning ( a) , drilling ( b) , and m illing ( c) operat ions

The material removal rate has been defined as the uncut area multiplied by the rate at which the material is removed perpendicular to the area. For turning, the area removed is an annular ring of outside diameter D and inside diameter D1. Thus, the uncut area is (D2 )/4. The rate at which the tool is fed, fm (in unit distance per minute), is fr · N. Therefore, the material removal rate for turning is:

(Eq 30) and the total cutting power is Ps · MRR, which can then be adjusted for machine tool efficiency. The specific power and the material removal rate can also be used to estimate the main cutting force, Fc. The total horsepower can be written as P = Ps · MRR or as the product of the main cutting force multiplied by the velocity, as described for the plate in the section "Energy of Chip Formation" in this article. Equating these two equivalent expressions yields:

(Eq 31) Pow e r in D r illin g. In Fig. 6(b), a drill of diameter D is rotated at N revolutions per minute and fed at fr (unit distance per revolution). The uncut area is the area covered by the drill, or D2/4, while the feed rate perpendicular to the area is fm = fr · N. As with turning, recommended velocities, V, and feeds, fr, are available in handbooks and the literature. The cutting velocity is again related to the rotational velocity by V = DN. The material removal rate is D2/4 · fm, and the power can be calculated as:

P = Ps( D2/4 · fr · N)

(Eq 32)

Thus, the calculation of the material removal rate and the power is quite analogous to the case of turning. Pow e r in M illin g. For the milling operation illustrated in Fig. 6(c), the cutter has diameter D with T teeth. The cutter is

rotated at N revolutions per minute. Unlike turning and drilling, the table (tool) is indexed at a feed rate of fm (millimeters or inches per minute), which is set directly on the machine. However, the feed rate is not given in handbooks. Instead, the

feed (in unit distance) per tooth, ft, is specified because too large a cut with any tooth may damage the tooth and because different cutters may contain varying numbers of teeth. The feed rate is related to the feed per tooth by:

fm = fr · T · N

(Eq 33)

and the material removal rate is specified by:

MRR = Uncut area · Feed rate MRR = (d · w)fm MRR = dwfrTN

(Eq 34)

The power required is therefore Ps · MRR. Fa ct or s Affe ct in g Spe cific Pow e r . The specific power is a material-related property that can be significantly altered

by cutting parameters, prior strain hardening of the workpiece material, and the material and geometry of the cutting tool. The cutting velocity has a significant effect on the specific power and/or energy because the coefficient of friction on the rake face is speed dependent. Increasing speeds decrease the friction (up to a point), thus decreasing the specific power through the frictional component of specific power (Fig. 7).

Fig. 7 I nfluence of speed, t ool geom et ry, and prior st rain hardening on t he specific energy of brass

The cutting tool material will also change the specific power due to changes in friction. For example, under similar conditions, the coefficient of friction and therefore the frictional component of the specific power will be lower for carbide tools than for high-speed steel tools. Therefore, the specific power will be higher, resulting in higher cutting forces. The tool geometry will also play a role in determining the specific power. In particular, the back rake angle, ,

will influence the friction and therefore the specific power. Larger back rake angles will result in lower power and/or energy consumed per unit volume per unit time, provided the tool maintains its integrity. Figure 7 illustates this in the machining of brass. In addition, cutting tools (inserts) are available in a variety of nose radii. The radius of the tool has been shown to change the specific power and cutting forces significantly with larger radii yielding higher specific power. Research has shown that ductile materials such as aluminum, brass, or copper will behave quite differently during machining, depending on the amount of strain hardening from prior processing. These materials have a great propensity for seizing the tool when in an annealed state. Cutting forces and specific powers are greater than for the same materials with significantly more work hardening. Thus, the processing history of a metal is an important determining factor for the specific power requirements.

Re fe r e n ce s 1. S. Ramalingam and J T. Black, An Electron Microscopy Study of Chip Formation, Metall. Trans., Vol 4, 1973, p 1103 2. J T. Black, Flow Stress Model in Metal Cutting, J. Eng. Ind., Vol 101, 1979, p 403 3. B.F.von Turkovich, Dislocation Theory of Shear Stress and Strain Rate in Metal Cutting, in Advances in Machine Tool Design and Research, Pergamon Press, 1967, p 531 4. M.E. Merchant, Mechanics of the Metal Cutting Process--I, J. Appl. Phys., Vol 16 (No. 5), 1945, p 267-275 5. M.E. Merchant, Mechanics of the Metal Cutting Process--II, J. Appl. Phys., Vol. 16 (No. 6), 1945, p 318325 6. E. Vidosic, Metal Machining and Forming Technology, The Ronald Press, 1964 7. N. Cook, Manufacturing Analysis, Addison-Wesley, 1966 8. E.P. DeGarmo, J T. Black, and R.A. Kosher, Materials and Processes in Manufacturing, 7th ed., Macmillan, 1988 9. S. Ramalingham and K.J. Trigger, in Advances in Machine Tool Design and Research, Pergamon Press, 1971 10. E.M. Trent, Metal Cutting, Butterworths, 1984, p 24 11. P.K. Wright, Friction Interactions In Machining: Comparisons Between Transparent Sapphire Tools and Steel Cutting Tools, Met. Technol., Vol 8, 1981, p 150 12. E. Amini, J. Strain Anal., Vol 3, 1968, p 206 13. P.K. Wright and J.L. Robinson, Material Behaviour in Deformation Zones of Machining Operation, Met. Technol., May 1977, p 240 14. N.N. Zorev, Interrelationship Between Shear Process Occurring Along Tool Face and on Shear Plane in Metal Cutting, International Research in Production Engineering, Presented at International Production Engineering Research Conference (Pittsburgh), 1963, p 42

Su r fa ce Fin ish a n d Su r fa ce I n t e gr it y Michael Field, John F. Kahles, and William P. Kost er, Met cut Research Associat es I nc.

I n t r odu ct ion A PART SURFACE has two important aspects that must be defined and controlled. The first concerns the geometric irregularities of the surface, and the second concerns the metallurgical alterations of the surface and the surface layer. This second aspect has been termed surface integrity. Both surface finish and surface integrity must be defined, measured, and maintained within specified limits in the processing of any product. Standards have been adopted for surface finish and are available in ANSI/ASME B46.1-1985 (Ref 1). A companion standard for surface texture symbols is ANSI Y 14.36-1978 (Ref 2). The standard for surface integrity is ANSI B211.1-1986 (Ref 3).

Su r fa ce Fin ish or Su r fa ce Te x t u r e Surface finish as described in Ref 1 is concerned with only the geometric irregularities of surfaces of solid materials and the characteristics of instruments for measuring roughness. Surface texture is defined in terms of roughness, waviness, lay, and flaws (Fig. 1): • • • •

Surface roughness consists of fine irregularities in the surface texture, usually including those resulting from the inherent action of the production process, such as feed marks produced during machining Waviness is a more widely spaced component of surface texture and may result from such factors as machine or work deflections, vibration, or chatter Lay is the direction of the predominant surface pattern Flaws are unintentional, unexpected, and unwanted interruptions in the surface--for example, cracks, nicks, scratches, and ridges

Fig. 1 Schem at ic of roughness and waviness on a surface wit h unidirect ional lay and one flaw. See Fig. 2 for definit ion of Ra and waviness height . Source: Ref 1

Both surface roughness and waviness can be measured by a variety of instruments, including both surface contact and noncontact types. By far the most universal technique is to measure surface roughness with a stylus contact-type instrument that provides a numerical value for surface roughness. Such instruments can usually provide an indication of roughness in terms of the arithmetic average, Ra (Fig. 2a), or the root mean square (rms) value, Rq (Fig. 2b).

Fig. 2 Som e com m only used designat ions of surface t ext ure. ( a) Ra. ( b) Rq . ( c) Ry or Rm ax . ( d) Rz. ( e) W. Source: Ref 1

Su r fa ce t e x t u r e sym bols used for illustrations and specifications are shown in Fig. 3. Symbols for defining lay and its direction are shown in Fig. 4.

Fig. 3 Surface t ext ure sym bols used for drawings or specificat ions. I n t his exam ple, all values are in inches except Ra values, which are in m icroinches. Met ric values ( m illim et ers and m icrom et ers) are used on m et ric drawings. Source: Ref 2

Fig. 4 Sym bols used t o define lay and it s direct ion. Source: Ref 2

D e sign a t ion s of Su r fa ce Rou gh n e ss. Figure 2 illustrates some of the designations of surface roughness. The most

common method of designating surface roughness in the United States is the arithmetical average Ra, although the rms value Rq is also used. The ratio between Rq and Ra varies with the manufacturing process producing the surface (Table 1). A preferred series of roughness values is given in Table 2.

Table 1 Ratio of root mean square roughness to arithmetic average roughness Root mean square roughness Arithmetic average roughness Theoretical ratio of sine waves, Rq/Ra Actual ratios of Rq/Ra for various processes Turning Milling Surface grinding Plunge grinding Soft honing Hard honing Electrical discharge machining Shot peening Practical first approximation of Rq/Ra For most processes For honing

Rq Ra 1.11 1.17 to 1.26 1.16 to 1.40 1.22 to 1.27 1.26 to 1.28 1.29 to 1.48 1.50 to 2.10 1.24 to 1.27 1.24 to 1.28 1.25 1.45

Source: Ref 4

Table 2 Preferred series of roughness average values (Ra) m 0.012 0.025(a) 0.050(a) 0.075 0.10(a) 0.125 0.15 0.20(a) 0.25 0.32 0.40(a) 0.50 0.63 0.80(a) 1.00 1.25 1.60(a) 2.0 2.5 3.2(a) 4.0 5.0 6.3(a) 8.0 10.0 12.5(a) 15 20 25(a)

in. 0.5 1(a) 2(a) 3 4(a) 5 6 8(a) 10 13 16(a) 20 25 32(a) 40 50 63(a) 80 100 125(a) 160 200 250(a) 320 400 500(a) 600 800 1000(a)

Source: Ref 2

(a)

Recommended.

Su r fa ce Rou gh n e ss Pr odu ce d in M a n u fa ct u r in g Pr oce sse s. The predominant method of producing engineering

surfaces is by a machining process, although some finished surfaces result from primary techniques such as casting, extruding, or forging. Each surface-producing method has a characteristic surface roughness range, some of which are

shown in Fig. 5. The finer finishes are generally produced by machining techniques. Traditional machining techniques include chip removal processes (such as turning, milling, and reaming) and abrasive processes (such as grinding, polishing, buffing, and superfinishing). A variety of surface finishes can also be produced by nontraditional machining techniques such as electrical discharge machining, electrochemical machining, or laser beam machining. Surface finish requirements for representative machine tool components and aircraft engine components are given in Tables 3 and 4, respectively.

Table 3 Typical surface finish requirements for machine tool components Part name (material) Quill (4145 H) End face Outside diameter Holes Inside diameter Cam (1018) Key (1018) Holder (1018) Bracket (1018) Plate (1018) Block (1018) Junction block (1018) Ball screw (4150) Keyways Outside diameter Thread diameter Ball nut (8617) Slots Diameters Holes

Machining process

Surface finish required, Ra m in.

Mill Lathe Drill Grind Grind Mill Mill Mill Mill Mill Grind

1.60 1.60 1.60 0.40 0.40-0.80 3.2 3.2 3.2 3.2 3.2 1.60

63 63 63 16 16-32 125 125 125 125 125 63

Mill Turn Grind

3.2 3.2 0.80

125 125 32

Mill Grind Drill

3.2 1.60 3.2

125 63 125

Table 4 Typical surface finish requirements for aircraft engine components Part name and material

Operation

Fan disk (Ti-6Al-4V) and turbine disk (Inconel 718) Turned Ultrasonic envelope Turned General surfaces Reamed Bolt holes Broached Dovetails Mass media finished Corner breaks Compressor casing (M-152 stainless steel) Turned Flowpath (inside diameter) Milled Outside (outside diameter) Bored Vane bores Turned Flange faces High-pressure turbine blade (René·80) Tumbled Airfoil Ground Dovetail form Ground General surfaces High-pressure turbine vane X40 (cobalt base) Tumbled Airfoil, convex Tumbled Airfoil, concave Tumbled Flowpath Ground, tumbled General surfaces Turbine shaft (Inconel 718) Journals (chromium plated) Ground Reamed Bolt holes Turned General surfaces

Surface finish requirements, Ra m in. 1.60 1.60-3.2 0.80-1.60 0.80-1.60 0.80

63 63-125 32-63 32-63 32

1.60-3.2 3.2 1.60 1.60

63-125 125 63 63

0.80 0.80 1.1

32 32 45

0.61-0.80 1.1 0.80 0.80

24-32 45 32 32

0.40 0.80 1.60-3.2

16 32 63-125

Fan blade (Ti-6Al-4V) Airfoil, convex Airfoil, concave Dovetail

Belt ground, tumbled Belt ground, tumbled Broached

0.61-0.80 0.80-1.1 0.80

24-32 32-45 32

Fig. 5 Surface roughness produced by com m on product ion m et hods. The ranges shown are t ypical of t he processes list ed. Higher or lower values can be obt ained under special condit ions. Source: Ref 1

Su r fa ce Rou gh n e ss a n d D im e n sion a l Tole r a n ce s. Surface roughness is closely tied to the accuracy or tolerance of a machine component (Table 5). A close-tolerance dimension requires a very fine finish, and the finishing of a component to a very low roughness value may require multiple machining operations. For example, a 3.2 m (125 in.) surface roughness can be produced by milling or turning, while a very fine (low roughness value) surface would require grinding or additional subsequent operations, such as honing, superfinishing, buffing, or abrasive flow. Therefore, specifying very fine finishes will normally result in increased costs (Table 5.)

Table 5 Classification of machined surface finishes Class

Roughness, Ra

Super finish Polish Ground Smooth Fine Semifine Medium Semirough Rough Cleanup

m 0.10 0.20 0.40 0.80 1.60 3.2 6.3 12.5 25 50

in. 4 8 16 32 63 125 250 500 1000 2000

Suitable for tolerance of plus or minus mm in. 0.0125 0.0005 0.0125 0.0005 0.025 0.001 0.050 0.002 0.075 0.003 0.100 0.004 0.175 0.007 0.330 0.013 0.635 0.025 1.25 0.050

Typical method of producing finish

Approximate relative cost to produce

Ground, microhoned, lapped Ground, honed, lapped Ground, lapped Ground, milled Milled, ground, reamed, broached Ground, broached, milled, turned Shaped, milled, turned Milled, turned Turned Turned

40 35 25 18 13 9 6 4 2 1

Th e or e t ica l Su r fa ce Rou gh n e ss Pr odu ce d by M illin g a n d Tu r n in g Tools. It is possible to calculate the theoretical surface roughness profile produced by milling cutters and lathe tools. Surface roughness calculations have been made for three of the most common cutting tool shapes. These tool shapes (Fig. 6(a) and 6(b)) are designated as Type A, a sharp-nose milling tooth, and Types B and C, which are a round tool and a tool with a nose radius, respectively. Types B and C can be used for either milling (Fig. 6(a)) or turning (Fig. 6(b)). Calculations have been made of the theoretical surface roughness as a function of the feed, the tool radius, the end cutting edge angle (ECEA), and the side cutting edge angle (SCEA) (Fig. 6(a) and 6(b)). The theoretical surface roughness obtained from these calculations represents the best finish commonly produced by that particular milling or turning tool and thus provides an indication of the minimum surface roughness possible with a designated tool shape and feed rate. The actual surface roughness may be poorer, because the surface is further degraded by a built-up edge that is usually formed as a characteristic of the machining process. Under some less common occurrences, a finer surface finish than the theoretical is produced because of the wear of the cutting edge that produces the finished surface; the worn tool develops a wearland that provides the tool with a wiping action, which tends to smooth out the theoretical surface irregularities. Surface roughness is sometimes improved in milling by providing the milling cutter with one additional finishing or wiper tooth designed to produce a broad finished machining path following the cutting action of the regular chip producing teeth in the cutter.

Fig. 6 ( a ) Theoret ical surfaces produced in m odels of face m illing wit h a sharp- nose m illing t ool ( Type A) , a round t ool ( Type B) , and a round- nose t ool ( Type C) . Source: Ref 5

Fig. 6 ( b) Theoret ical surfaces produced by t urning wit h a round t ool ( Type B) and a round- nose t ool ( Type C) . Source: Ref 5

The theoretical surface roughness produced by a face milling cutter containing teeth with a zero nose radius is plotted in Fig. 7. The theoretical surface for turning or face milling with round cutting edges is illustrated in Fig. 8, and the theoretical surface roughness for turning or face milling with a radius of 0.396 mm (0.0156 in.) and various end cutting edge angles is shown in Fig. 9. Numerical tables of the theoretical roughness produced by milling and turning tools as a function of the feed, end cutting edge angle, and side cutting edge angle are provided in Ref 5. Table 6 gives the arithmetic roughness average, Ra, and the maximum peak-to-valley roughness height, Ry, for turning and milling.

Table 6 Theoretical values for arithmetic roughness average, Ra, and maximum peak-to-valley roughness height, Ry, for turning or milling with a Type C tool Tool nose radius = 0.40 mm (0.031 in.) Feed/rev (turning) or feed/tooth (milling) mm (in.) 0.020 (0.001) 0.040 (0.002) 0.060 (0.003) 0.080 (0.004) 0.100 (0.005) 0.120 (0.006) 0.140 (0.007) 0.160 (0.008) 0.180 (0.009) 0.200 (0.010) 0.250 (0.012) 0.300 (0.014) 0.350 (0.016) 0.400 (0.018) 0.450 (0.020) 0.500 (0.025) 0.600 (0.030)

End cutting edge angle 3° 5° Ra Ry Ra Ry Surface roughness, m ( in.)

6° Ra

Ry

10° Ra

Ry

15° Ra

Ry

30° Ra

Ry

40° Ra

Ry

45° Ra

Ry

0.03 (1.0) 0.13 (4.1) 0.28 (9.3) 0.46 (16.0) 0.65 (25.0) 0.85 (34.0) 1.1 (43.0) 1.3 (53.0) 1.5 (63.0) 1.7 (73.0) 2.2 (94.0) 2.8 (115.0) 3.4 (136.0) 3.9 (158.0) 4.5 (180.0) 5.1 (236.0) 6.3 (293.0)

0.03 (1.0) 0.13 (4.1) 0.29 (9.3) 0.51 (16.0) 0.80 (26.0) 1.1 (37.0) 1.5 (50.0) 1.9 (66.0) 2.2 (82.0) 2.6 (100.0) 3.6 (136.0) 4.7 (174.0) 5.7 (213.0) 6.8 (253.0) 7.9 (294.0) 9.0 (398.0) 11.0 (506.0)

0.13 (4.0) 0.50 (16.0) 1.1 (36.0) 2.0 (64.0) 3.1 (100.0) 4.3 (145.0) 5.5 (196.0) 6.9 (253.0) 8.2 (312.0) 9.6 (373.0) 13.0 (503.0) 17.0 (639.0) 21.0 (781.0) 25.0 (927.0) 29.0 (1076.0) 34.0 (1463.0) 42.0 (1865.0)

0.03 (1.0) 0.13 (4.1) 0.29 (9.3) 0.51 (16.0) 0.80 (26.0) 1.2 (37.0) 1.6 (51.0) 2.1 (66.0) 2.6 (84.0) 3.2 (103.0) 4.7 (149.0) 6.3 (202.0) 7.9 (259.0) 9.6 (319.0) 11.0 (381.0) 13.0 (543.0) 17.0 (711.0)

0.13 (4.0) 0.50 (16.0) 1.1 (36.0) 2.0 (64.0) 3.1 (100.0) 4.5 (145.0) 6.2 (197.0) 8.0 (257.0) 9.9 (326.0) 12.0 (403.0) 17.0 (579.0) 23.0 (771.0) 29.0 (975.0) 35.0 (1189.0) 42.0 (1410.0) 48.0 (1992.0) 62.0 (2607.0)

0.03 (1.0) 0.13 (4.1) 0.29 (9.3) 0.51 (16.0) 0.80 (26.0) 1.2 (37.0) 1.6 (51.0) 2.1 (66.0) 2.6 (84.0) 3.2 (103.0) 5.1 (149.0) 7.2 (204.0) 9.5 (267.0) 12.0 (339.0) 14.0 (417.0) 17.0 (636.0) 22.0 (873.0)

0.13 (4.0) 0.50 (16.0) 1.1 (36.0) 2.0 (64.0) 3.1 (100.0) 4.5 (145.0) 6.2 (197.0) 8.1 (257.0) 10.0 (326.0) 13.0 (403.0) 20.0 (582.0) 27.0 (795.0) 36.0 (1043.0) 44.0 (1318.0) 53.0 (1609.0) 62.0 (2389.0) 82.0 (3232.0)

0.03 (1.0) 0.13 (4.1) 0.29 (9.3) 0.51 (16.0) 0.80 (26.0) 1.2 (37.0) 1.6 (51.0) 2.1 (66.0) 2.6 (84.0) 3.2 (103.0) 5.1 (149.0) 7.4 (204.0) 10.0 (267.0) 14.0 (339.0) 17.0 (420.0) 22.0 (665.0) 31.0 (972.0)

0.13 (4.0) 0.50 (16.0) 1.1 (36.0) 2.0 (64.0) 3.1 (100.0) 4.5 (145.0) 6.2 (197.0) 8.1 (257.0) 10.0 (326.0) 13.0 (403.0) 20.0 (582.0) 29.0 (795.0) 40.0 (1043.0) 54.0 (1326.0) 69.0 (1646.0) 85.0 (2613.0) 120.0 (3842.0)

0.03 (1.0) 0.13 (4.1) 0.29 (9.3) 0.51 (16.0) 0.80 (26.0) 1.2 (37.0) 1.6 (51.0) 2.1 (66.0) 2.6 (84.0) 3.2 (103.0) 5.1 (149.0) 7.4 (204.0) 10.0 (267.0) 14.0 (339.0) 17.0 (420.0) 22.0 (665.0) 33.0 (972.0)

0.13 (4.0) 0.50 (16.0) 1.1 (36.0) 2.0 (64.0) 3.1 (100.0) 4.5 (145.0) 6.2 (197.0) 8.1 (257.0) 10.0 (326.0) 13.0 (403.0) 20.0 (582.0) 29.0 (795.0) 40.0 (1043.0) 54.0 (1326.0) 69.0 (1646.0) 88.0 (2613.0) 132.0 (3842.0)

0.03 (1.0) 0.13 (4.1) 0.29 (9.3) 0.51 (16.0) 0.80 (26.0) 1.2 (37.0) 1.6 (51.0) 2.1 (66.0) 2.6 (84.0) 3.2 (103.0) 5.1 (149.0) 7.4 (204.0) 10.0 (267.0) 14.0 (339.0) 17.0 (420.0) 22.0 (665.0) 33.0 (972.0)

0.13 (4.0) 0.50 (16.0) 1.1 (36.0) 2.0 (64.0) 3.1 (100.0) 4.5 (145.0) 6.2 (197.0) 8.1 (257.0) 10.0 (326.0) 13.0 (403.0) 20.0 (582.0) 29.0 (795.0) 40.0 (1043.0) 54.0 (1326.0) 69.0 (1646.0) 88.0 (2613.0) 135.0 (3842.0)

0.13 (4.0) 0.50 (16.0) 1.1 (36.0) 1.7 (63.0) 2.4 (93.0) 3.1 (125.0) 3.9 (159.0) 4.6 (194.0) 5.4 (230.0) 6.2 (268.0) 8.3 (344.0) 10.0 (423.0) 13.0 (503.0) 15.0 (585.0) 17.0 (669.0) 19.0 (881.0) 24.0 (1099.0)

0.03 (1.0) 0.13 (4.1) 0.29 (9.3) 0.51 (16.0) 0.79 (26.0) 1.1 (37.0) 1.4 (50.0) 1.7 (64.0) 2.0 (79.0) 2.4 (94.0) 3.2 (126.0) 4.1 (159.0) 5.0 (192.0) 5.9 (226.0) 6.9 (261.0) 7.8 (350.0) 9.7 (441.0)

0.13 (4.0) 0.50 (16.0) 1.1 (36.0) 2.0 (64.0) 3.0 (100.0) 4.0 (144.0) 5.1 (191.0) 6.3 (242.0) 7.5 (294.0) 8.7 (349.0) 12.0 (462.0) 15.0 (581.0) 19.0 (704.0) 22.0 (830.0) 26.0 (959.0) 29.0 (1291.0) 36.0 (1634.0)

0.700 (0.035) (0.040) (0.045) (0.050) (0.060)

Source: Ref 5

7.5 (351.0) (409.0) (468.0) (527.0) (646.0)

28.0 (1321.0) (1545.0) (1772.0) (2001.0) (2465.0)

12.0 (534.0) (628.0) (723.0) (819.0) (1013.0)

44.0 (1985.0) (2343.0) (2706.0) (3074.0) (3821.0)

14.0 (615.0) (726.0) (839.0) (953.0) (1185.0)

51.0 (2277.0) (2698.0) (3127.0) (3562.0) (4447.0)

21.0 (885.0) (1063.0) (1244.0) (1429.0) (1805.0)

76.0 (3247.0) (3906.0) (4583.0) (5274.0) (6693.0)

28.0 (1120.0) (1377.0) (1640.0) (1910.0) (2466.0)

102.0 (4124.0) (5055.0) (6020.0) (7015.0) (9080.0)

42.0 (1348.0) (1783.0) (2263.0) (2778.0) (3889.0)

158.0 (5320.0) (6940.0) (8684.0) (10542.0) (14572.0)

46.0 (1350.0) (1808.0) (2355.0) (2983.0) (4447.0)

183.0 (5370.0) (7253.0) (9456.0) (11874.0) (17356.0)

48.0 (1350.0) (1808.0) (2360.0) (3018.0) (4629.0)

192.0 (5370.0) (7253.0) (9581.0) (12273.0) (18554.0)

Fig. 7 Theoret ical surface roughness for a face m illing cut t er cont aining t eet h wit h a zero nose radius. Source: Ref 5

Fig. 8 Theoret ical surface roughness for t urning or face m illing t ools wit h round cut t ing edges. Source: Ref 5

Fig. 9 Theoret ical surface roughness for t urning or face m illing t ools wit h a radius of 0.39 m m ( 0.0156 in.) and various ECEAs. Source: Ref 5

Su r fa ce I n t e gr it y The specification and manufacture of unimpaired or enhanced surfaces require an understanding of the interrelationship among metallurgy, machinability, and mechanical testing. To satisfy this requirement, an encompassing discipline known as surface integrity was introduced, and it has gained worldwide acceptance. Surface integrity technology describes and controls the many possible alterations produced in a surface layer during manufacture, including their effects on material properties and the performance of the surface in service. Surface integrity is achieved by the selection and control of manufacturing processes, estimating their effects on the significant engineering properties of work materials.

Surface integrity involves the study and control of both surface roughness or surface topography, and surface metallurgy. Both of these factors influence the quality of the machined surface and subsurface, and they become extremely significant when manufacturing structural components that have to withstand high static and dynamic stresses. For example, when dynamic loading is a principal factor in a design, useful strength is frequently limited by the fatigue characteristics of materials. Fatigue failures almost always nucleate at or near the surface of a component; similarly, stress corrosion is also a surface phenomena. Therefore, the nature of the surface from both a topographical and a metallurgical point of view is important in the design and manufacture of critical hardware. The importance of surface integrity is further heightened when high stresses occur in the presence of extreme environments. Heat-resistant, corrosion-resistant, and high-strength alloys are used in a wide variety of such applications. Typical alloys used in these applications include alloy steels with hardnesses of 50 to over 60 HRC and heat-treated alloys with strength levels as high as 2070 MPa (300 ksi). Additional materials include stainless steels, titanium alloys, and high-temperature nickel-base alloys developed for high-temperature and corrosion-resistant applications. Unfortunately, the alloys suitable for high-strength applications are frequently difficult to machine. The hard steels and high-temperature alloys, for example, must be turned and milled at low speeds, which tend to produce a built-up edge and poor surface finish. The machining of these alloys tends to produce undesirable metallurgical surface alterations, which have been found to reduce fatigue strength. Typical surface integrity problems created in metal removal operations include: • • • • • •

Grinding burns on high-strength steel aircraft landing gear components Untempered martensite in drilled holes in steel components Grinding cracks in the root sections of cast nickel-base gas turbine buckets Lowering of fatigue strength of parts processed by electrical discharge machining Distortion of thin components Residual stress induced in machining and its effect on distortion, fatigue, and stress corrosion

Su r fa ce Alt e r a t ion s The types of surface alterations associated with metal removal practices are (Ref 6):

Mechanical • • • • • • • •

Plastic deformations (as result of hot or cold working) Tears and laps and crevicelike defects (associated with built-up edge produced-in-machining) Hardness alterations Cracks (macroscopic and microscopic) Residual stress distribution in surface layer Processing inclusions introduced Plastically deformed debris as a result of grinding Voids, pits, burrs, or foreign material inclusions in surface

Metallurgical • • • • • • • •

Transformation of phases Grain size and distribution Precipitate size and distribution Foreign inclusions in material Twinning Recrystallization Untempered martensite or overtempered martensite Resolutioning or austenite reversion

Chemical

• • • • • • • • •

Intergranular attack Intergranular corrosion Intergranular oxidation Preferential dissolution of microconstituents Contamination Embrittlement by the chemical absorption of elements such as hydrogen and chlorine Pits or selective etch Corrosion Stress corrosion

Thermal • • • •

Heat-affected zone Recast or redeposited material Resolidified material Splattered particles or remelted metal deposited on surface

Electrical • • •

Conductivity change Magnetic change Resistive heating or overheating

The principal causes of surface alterations produced by machining processes are: • • • • •

High temperature or high-temperature gradients developed in the machining process Plastic deformation Chemical reactions and subsequent absorption into the nacent machined surface Excessive electrical currents Excessive energy densities during processing

Virtually all material removal methods produce altered surface and subsurface conditions. The possible surface alterations resulting from various processes are summarized in Table 7. The mechanically and metallurgically altered zones produced by material removal processes may also extend into the surface to a considerable depth as a function of whether roughing or finishing conditions are used in the material removal process (Table 8).

Table 7 Summary of possible surface alterations resulting from various material removal processes R, roughness of surface; PD, plastic deformation and plastically deformed debris; L & T, laps and tears and crevicelike defects; MCK, microcracks; SE, selective etch; IGA, intergranular attack; UTM, untempered martensite; OTM, overtempered martensite; OA, overaging; RS, resolution or austenite reversion; RC, recast, respattered metal, or vapor-deposited metal; HAZ, heat-affected zone Material

Nonhardenable 1018 steel

Hardenable 4340 and D6ac steels

D2 tool steel

Type 410 stainless steel (martensitic)

Type 302 stainless steel (austenitic)

17-4 PH steel

250 grade maraging (18% Ni) steel

Nickel and cobalt-base alloys Inconel alloy 718 René 41 HS 31 IN-100 Ti-6Al-4V

Refractory alloy TZM

Tungsten (pressed and sintered)

Source: Ref 7

Process Conventional Milling, Grinding drilling, or turning R R PD PD L&T R R PD PD L&T MCK MCK UTM UTM OTM OTM R R PD PD L&T MCK MCK UTM UTM OTM OTM R R PD PD L&T MCK MCK UTM UTM OTM OTM R R PD PD L&T R R PD PD L&T OA OA R R PD PD L&T RS RS OA OA HAZ HAZ R R PD PD L&T MCK MCK HAZ HAZ R R PD PD L&T MCK R R L&T MCK MCK R R L&T MCK MCK

Nontraditional Electrical Electrochemical discharge machining machining R R MCK SE RC IGA R R MCK SE RC IGA UTM OTM OTM

Chemical machining R SE IGA R SE IGA

R MCK RC UTM OTM

R SE IGA

R SE IGA

R MCK RC UTM OTM

R SE IGA

R SE IGA

R MCK RC R MCK RC OA R RC RS OA

R SE IGA R SE IGA

R SE IGA R SE IGA

R SE IGA

R SE IGA

R MCK RC

R SE IGA

R SE IGA

R MCK RC R MCK

R SE IGA R SE IGA R SE MCK IGA

R SE

R MCK

R SE R SE MCK IGA

Table 8 Comparison of depth of surface integrity effects observed in material removal processes Property and type of effect

Condition

Mechanically altered material zones Finishing(b) Plastic deformation Roughing(d) Finishing Plastically deformed debris Roughing Finishing Hardness alteration(e) Roughing Finishing Microcracks or macrocracks Roughing Finishing Residual stress(f) Roughing Metallurgically altered material zones Finishing Recrystallization Roughing Finishing Intergranular attack Roughing Finishing Selective etch, pits, protuberances Roughing Finishing Metallurgical transformations Roughing Finishing Heat-affected zone or recast layers Roughing

Maximum observed depth of effect(a) Turning or Drilling Grinding milling

Chemical machining

Electrochemical machining

Electrochemical grinding

mm

in.

mm

in.

mm

in.

mm

in.

mm

in.

mm

0.043 0.076

0.0008 0.0047

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

0.0003 0.0035 0.0005 0.0013 0.0015 0.0100 0.0005 0.0090 0.0005 0.0125

(c)

(c)

0.008 0.089 0.013 0.033 0.038 0.254 0.013 0.229 0.013 0.318

(c)

(c)

0.020 0.119

(c)

(c)

0.0017 0.0030

(c)

(c)

(c)

0.013

(c)

(c)

0.013 0.127 0.013 0.038 0.152 0.356

0.0005 0.0050 0.0005 0.0015 0.0060 0.0140

0.025 0.508 0.013 0.038

0.0010 0.0200 0.0005 0.0015

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

0.0004 0.0010 0.0004 0.0030 0.0001 0.0010

(c)

(c)

0.010 0.025 0.010 0.076 0.003 0.025

(c)

0.076 0.038 0.508 (c)

0.076

(c)

0.3000 0.0015 0.0200 (c)

0.0030

in.

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

0.0010 0.0080 0.0005 0.0070 0.0020 0.0030

(c)

(c)

(c)

(c)

0.015 0.102 0.005

0.0006 0.0040 0.0002

(c)

(c)

(c)

(c)

0.025 0.025

0.0010 0.0010

0.0005

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

0.0003 0.0060 0.0006 0.0015

0.008 0.038 0.010 0.064 0.000 0.005

0.0003 0.0015 0.0004 0.0025 0.0000 0.0002

0.000

0.000

(c)

(c)

(c)

(c)

0.008 0.152 0.015 0.038

(c)

(c)

(c)

(c)

(c)

(c)

0.003 0.013 0.003 0.008

0.0001 0.0005 0.0001 0.0003

(c)

(c)

(c)

(c)

(c)

(c)

(c)

0.0005 0.0016 0.0006 0.0050 0.0006 0.0050

(c)

0.015 0.038 0.015 0.038

0.0006 0.0015 0.0006 0.0015

0.005 0.010 0.013 0.152 0.018 0.318

0.0002 0.0004 0.0005 0.0060 0.0007 0.0125

0.0010 0.0031

mm

0.036 0.051 0.008 0.038 0.000 0.000

(c)

0.025 0.079

Laser beam machining

in.

Electrical discharge machining mm in.

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

(c)

0.0014 0.0020 0.0003 0.0015 0.0000 0.0000

0.018 0.038 0.000 0.025 0.000 0.000

0.0007 0.0015 0.0000 0.0010 0.0000 0.0000

0.025 0.203 0.013 0.178 0.051 0.076

(c)

(c)

(c)

(c)

(c)

0.013 0.041 0.015 0.127 0.015 0.127

(c)

(c)

Source: Ref 8

(a) (b) (c) (d) (e) (f)

Normal to the surface. Finishing, gentle, or low-stress conditions. No occurrences or not expected. Roughing, off-standard, or abusive conditions. Depth to point at which hardness becomes less than ±2 points HRC (or equivalent) of bulk material hardness (hardness converted from Knoop microhardness measurements). Depth to point at which residual stress becomes and remains less than 140 MPa (20 ksi) or 10%, of tensile strength, whichever is greater.

Su r fa ce Alt e r a t ion s Pr odu ce d in Tr a dit ion a l M a ch in in g Ope r a t ion s. Traditional machining methods, which

are the principal means of metal removal, include chip cutting (such as milling, drilling, turning, broaching, reaming, and tapping) and abrasive machining methods (such as grinding, sanding, and polishing). These machining operations can produce surface layer alterations when abusive machining conditions are used (Fig. 10 and 11). In general, abusive machining promotes higher temperatures and excessive plastic deformation. Gentle machining operations occur when a sharp tool is employed and when machining conditions result in reduced machining forces.

Fig. 1 0 Surface alt erat ions produced from drilling wit h dull t ools. ( a) Sect ion perpendicular t o t he drill-hole axis in a 4340 st eel ( 48 HRC) . Abusive drilling produced a cracked unt em pered m art ensit e surface alt erat ion. Also not e t he soft er overt em pered zone below t he unt em pered m art ensit e layer. ( b) Cross sect ion t hrough a hole drilled in t ype 410 st ainless wit h a dull drill ( 1.5 m m , or 0.060 in., wearland) . The drill broke down at t he corner during t he t est and frict ion welded a port ion of t he high-speed st eel drill bit t o t he workpiece. The base m et al exhibit s a rehardened and subsequent ly overt em pered zone as a result of t he high localized heat ing. 20× . Source: Ref 9

Fig. 1 1 Surfaces produced by t he face m illing of Ti- 6Al- 4V ( aged, 35 HRC) . ( a) Wit h gent le m achining condit ions, a slight whit e layer is visible, but changes in m icrohardness are undet ect ed. 1000× . ( b) Wit h abusive condit ions, an overheat ed whit e layer about 0.01 m m ( 0.0004 in.) deep and a plast ically deform ed layer t ot alling 0.04 m m ( 0.0015 in.) deep are visible. 1000× . ( c) Microhardness m easurem ent s show a t ot al affect ed zone 0.08 m m ( 0.003 in.) deep from abusive condit ions. Source: Ref 9

Su r fa ce Effe ct s in Gr in din g. In grinding, there are five important parameters that determine gentle or abusive conditions: wheel grade, wheel speed, downfeed or infeed, wheel dressing, and grinding fluid. As grinding parameters become more aggressive (that is, harder wheels, higher wheel speeds, increased infeed, and so on), the grinding process becomes more abusive and therefore more likely to produce surface damage. Gentle, or low-stress, grinding conditions for a variety of alloys are summarized in Table 9.

Table 9 Low-stress grinding procedures Grinding parameters Surface grinding Wheel Wheel speed, m/s (sfm)(b) Downfeed per pass, mm (in.) Table speed, m/min (sfm)(c) Crossfeed per pass, mm (in.) Grinding fluid Traverse cylindrical grinding Wheel Wheel speed, m/s (sfm)(b) Infeed per pass, mm (in.) Work speed, m/min (sfm)(c) Grinding fluid

Steels and nickel-base high-temperature alloys(a)

Titanium

A46HV 13-15 (2500-3000) 0.005-0.013 (0.0002-0.0005) 12-30 (40-100) 1-1.25 (0.040-0.050) Highly sulfurized oil

C60HV 10-15 (2000-3000) 0.005-0.013 (0.0002-0.0005) 12-30 (40-100) 1-1.25 (0.040-0.050)

A60IV 13-15 (2500-3000) 0.005-0.013 (0.0002-0.0005) 20-30 (70-100) Highly sulfurized oil

C60HV 10-15 (2000-3000) 0.005-0.013 (0.0002-0.0005) 20-30 (70-100)

(d)

(d)

Source: Ref 10

(a)

(b)

(c) (d)

For a wide variety of metals (including high-strength steels, hightemperature alloys, titanium, and refractory alloys), low-stress grinding practices develop very low residual tensile stresses. In some materials, the residual stress produced near the surface is actually in compression instead of tension. Low-stress grinding requires wheel speeds lower than the conventional 30 m/s (6000 sfm). To apply low-stress grinding, it would be preferable to have a variable-speed grinder. Because most grinding machines do not have wheel-speed control, it is necessary to add a variable-speed drive or to make pulley modifications. Increased work speeds even above those indicated are considered to be advantageous for improving surface integrity. Cutting fluids should be nitrate free for health reasons. Some manufacturers also prohibit cutting fluids with chlorine when machining

titanium.

Figure 12(a) illustrates the surface characteristics produced by the low-stress grinding of AISI 4340 steel quenched and tempered to 50 HRC. The low-stress condition produced no visible surface alterations, while the abusive grinding condition (Fig. 12b) produced an untempered martensite layer 0.03 to 0.13 mm (0.001 to 0.005 in.) deep with a hardness of 65 HRC. Below this white layer there was an overtempered martensitic zone with a hardness of 46 HRC. The hardness returned to its normal value at a depth of 0.30 mm (0.012 in.) below the surface (Fig. 12c).

Fig. 1 2 Surface charact erist ics produced by t he low- st ress and abusive grinding of AI SI 4340 st eel. ( a) Lowst ress grinding produced no visible surface alt erat ions. ( b) The whit e layer shown from abusive condit ions has a hardness of 65 HRC and is approxim at ely 0.025 t o 0.050 m m ( 0.001 t o 0. 002 in.) deep. ( c) Plot of m icrohardness alt erat ions showing a t ot al heat - affect ed zone of 0.33 m m ( 0.013 in.) from abusive condit ions. ( d) Plot of residual st ress. ( e) Effect on fat igue st rengt h. Source: Ref 11

Abusive grinding also tends to produce a residual stress within the altered surface layer. A residual stress profile can be readily obtained by using x-ray diffraction techniques both at the surface and then by stepping into the surface with repeated x-ray diffraction readings after successive surface etching. The abusive grinding produced high tensile stresses in the altered zone, while low-stress grinding produced a surface with small compressive stresses (Fig. 12d). Fatigue tests conducted on flat specimens indicate that abusive grinding may seriously reduce the fatigue strength, as shown in Fig. 12(e). In this example, the abusive grinding depressed the endurance limit from 760 MPa (110 ksi) for lowstress grinding to 520 MPa (75 ksi). Su r fa ce Alt e r a t ion s Pr odu ce d in N on t r a dit ion a l M a ch in in g Ope r a t ion s. Nontraditional machining includes a

variety of methods for removing and finishing materials. Examples of nontraditional operations are electrical discharge machining, laser beam machining, electrochemical machining, electropolishing, and chemical machining. Ele ct r ica l disch a r ge m a ch in in g ( ED M ) tends to produce a surface with a layer of recast spattered metal that is usually hard and cracked and porous to some degree (Fig. 13). Below the spattered and recast metal, it is possible to have some of the same types of surface alterations that occur in abusive or traditional machining practices.

Fig. 1 3 Surface charact erist ics of cast I nconel 718 ( aged, 40 HRC) produced by EDM. ( a) Finishing condit ions produced a variable recast layer 0.005 m m ( 0.0002 in.) t hick. 860× . ( b) Roughing condit ions produced an ext ensively cracked variable recast layer up t o 0.05 m m ( 0.002 in.) . The random acicular st ruct ure is not relat ed t o t he surface phenom ena. 660× . ( c) The recast st ruct ure has a hardness of about 53 HRC, and slight overaging due t o localized surface overheat ing was also not ed. Source: Ref 9

The effect is more pronounced when EDM is used under abusive or roughing conditions. Figure 14 shows a surface produced by EDM under finishing and roughing conditions. Under roughing conditions, globs of recast metal are spattered onto a white layer of rehardened martensite. An overtempered zone up to 46 HRC is also found beneath the surface. The surface produced under finishing conditions contains discontinuous patches of recast metal plus a thin layer of rehardened martensite 0.003 mm (0.0001 in.) deep.

Fig. 1 4 Surface charact erist ics of AI SI ( quenched and t em pered, 50 HRC) produced by EDM under finishing and roughing condit ions. ( a) Finishing condit ions produce discont inuous pat ches of recast m et al plus a t hin layer ( 0.0025 m m , or 0.0001 in.) of rehardened m art ensit e. 620× . ( b) Roughing condit ions produce globs of recast m et al and a whit e layer of rehardened m art ensit e 0.075 m m ( 0.003 in.) deep. 620× . ( c) Microhardness m easurem ent s show a t ot al heat - affect ed zone approaching 0.25 m m ( 0.010 in.) . Globs of recast and t he whit e layer are at 62 HRC. An overt em pered zone as soft as 46 HRC is found beneat h t he surface. Source: Ref 9

La se r be a m m a ch in in g ( LBM ) tends to produce the same types of surface alterations as EDM. Figure 15 illustrates the heat-affected zone produced by LBM on Inconel 718. The intense heat generated by the laser beam resulted in a recast surface layer at the entrance and exit of the hole produced by the laser beam.

Fig. 1 5 Surfaces from t he laser beam drilling of I nconel 718 shown at m agnificat ions of 185× ( a) and 1160× ( b) . Not e t he grain st ruct ure in t he heat - affect ed zones of t he ent rance ( A) and t he exit ( B) . Source: Ref 9

Ele ct r och e m ica l m a ch in in g ( ECM ) is capable of producing a surface that is essentially free of metallurgical surface

layer alterations. However, when ECM is not properly controlled, selective etching or intergranular attack may occur (Fig. 16). Abusive ECM conditions also tend to degrade surface roughness (Fig. 17).

Fig. 1 6 Surface charact erist ics of Waspaloy ( aged, 40 HRC) produced by ECM. ( a) Gent le condit ions produce a slight roughening of t he surface and som e int ergranular at t ack. ( b) Abusive condit ions produce severe int ergranular at t ack. ( c) Microhardness is unaffect ed by t he abusive condit ions. Source: Ref 9

Fig. 1 7 Surface charact erist ics of 4340 st eel ( quenched and t em pered, 30 HRC) produced by ECM. ( a) Gent le condit ions produce slight surface pit t ing but no ot her visible changes. ( b) Abusive condit ions produce surface roughening but no ot her visible effect on m icrost ruct ure. ( c) Gent le and abusive condit ions bot h produce a slight hardness loss at t he surface. Source: Ref 12

Ele ct r opolish in g ( ELP) a n d Ch e m ica l M a ch in in g ( CM ) . Surface softening is produced on most materials by

electrochemical machining as well as by electropolishing and chemical machining, also referred to as chemical milling. Figure 17 illustrates the surface softening produced by both gentle and abusive ECM on 4340 steel. The surface is about 5 HRC points lower in hardness than the interior to approximately 0.05 mm (0.002 in.) in depth. With CM, the same steel had its surface softened by about 5 HRC points to a depth of about 0.05 mm (0.002 in.) (Fig. 18). This softening may be severe enough and deep enough to affect the fatigue strength and other mechanical properties of metals and may necessitate postprocessing.

Fig. 1 8 Surface charact erist ics of 4340 st eel ( annealed, 31 t o 36 HRC) produced by CM. ( a) Gent le condit ions produce no visible surface effect s and a surface finish of 0.9 m ( 35 in.) . ( b) Abusive condit ions produce a slight roughening and a surface finish of 3 m ( 120 in.) . ( c) Gent le and abusive condit ions bot h produce a soft ening at t he surface. Source: Ref 9

Re sidu a l St r e ss. Machining processes impart a residual stress in the surface layer. In grinding, the residual stress tends

to be tensile when more abusive conditions are used (Fig. 19). By using gentle grinding conditions, the stress can be reduced in magnitude and can even become compressive. In milling, the residual stress tends to be compressive (Fig. 20). In facing milling 4340 steel (Fig. 20), the stresses are tensile at the surface but go into compression below the surface.

Wheel Wheel speed, m/min (ft/min) Cross feed, mm/pass (in./pass) Table speed, m/min (ft/min) Depth of grind, mm (in.) Grinding fluid Specimen size, mm (in.)

A46K8V 1800 (6000) 1.25 (0.050) 12 (40) 0.25 (0.010) Soluble oil (1:20) 1.5 × 19 × 108 (0.060 ×

×4

)

Fig. 1 9 Residual st ress from surface grinding of D6AC st eel ( 56 HRC) .

Tool End cutting edge angle Peripheral clearance Cutting speed, m/min (ft/min) Feed, mm/tooth (in./tooth) Depth of cut, mm (in.) Width of cut, mm (in.) Cutting fluid Specimen size, mm (in.)

100 mm (4 in.) diam single-tooth face mill with Carboloy 370 (C-6) carbide 5° 8° 55 (180) 0.125 (0.005) 1.0 (0.040) 1.9 (0.750) Dry 1.5 × 19 × 108 (0.060 ×

×4

)

Fig. 2 0 Residual st ress from surface m illing 4340 st eel ( quenched and t em pered t o 52 HRC) .

D ist or t ion a n d r e sidu a l st r e ss are two related results from abusive machining conditions. In grinding, for example,

a more abusive machining condition (such as an increase in downfeed) increases the distortion of the workpiece (Fig. 21) and creates more residual stress at the surface (Fig. 19). Figures 22 and 20 illustrate a similar relation in milling. As the machining condition becomes more abusive with a duller tool, the distortion (Fig. 22) and residual stress (Fig. 20) become greater. Residual stress and distortion thus exhibit the following relation: The greater the area under the residual stress curve, the greater the distortion of the workpiece.

Wheel grade Cross feed, mm/pass (in./pass) Table speed, m/min (ft/min) Depth of cut, mm (in.) Specimen size, mm (in.) Grinding fluid

A46K8V 1.25 (0.050) 12 (40) 0.25 (0.010) 1.8 × 19 × 108 (0.070 × Soluble oil (1:20)

×4

)

Fig. 2 1 Change in deflect ion versus wheel speed and down feed in t he surface grinding of D6AC st eel ( 56 HRC)

Tool End cutting edge angle Peripheral clearance Cutting speed, m/min (ft/min) Feed, mm/tooth (in./tooth) Cutting fluid

100 mm (4 in.) diam single-tooth face mill with Carboloy 370 (C-6) carbide 5° 8° 55 (180) 0.125 (0.005) Dry

Fig. 2 2 Change in deflect ion versus t ool wearland for t he face m illing of 4340 st eel ( quenched and t em pered t o 52 HRC)

M e ch a n ica l Pr ope r t y Effe ct s The surface alterations produced by abusive metal removal conditions generally have a minor effect on the static mechanical properties of materials. A major exception to this, however, has been a marked decrease in ductility and tensile strength on materials that have been processed using EDM, followed by stress-relief heat treatment. D u ct ilit y a n d Te n sile St r e n gt h . Figure 23 illustrates the change in ductility and tensile strength from an EDM

application. It was found that a carbon deposit produced on the surface during the EDM operation was diffused into the grain boundaries during a subsequent stress-relief treatment and caused an excessive grain-boundary precipitation of carbides. These precipitates were responsible for reductions in ductility and strength, the extent of which was found to be a function of surface roughness (Fig. 23). Reductions in ductility as high as 80% were noted after the heat treatment of Inconel specimens that had been EDM-machined to a surface finish of 16.5 m (650 in.) Rq.

Fig. 2 3 Loss of t ensile st rengt h, duct ilit y, and yield st rengt h versus surface finish of I nconel 718 aft er EDM. Source: Ref 13

Fa t igu e St r e n gt h . Surface alterations produced in machining are known to affect the fatigue and stress-corrosion properties of many materials. Extensive investigations have been performed on high-strength steels, and data illustrating the effect of some machining methods on fatigue strength are given in Table 10. The electropolishing of 4340 steel resulted in a 12% decrease in fatigue strength compared to gentle grinding. Chemical machining of Ti-6Al-4V resulted in an 18% drop in fatigue strength compared to gentle grinding, while the electrochemical machining and electrical discharge machining of Inconel 718 produced a 35% and a 63% drop, respectively, in endurance limit compared to gentle grinding. Generally, ECM, ELP, and CM produce a stress-free surface. These surfaces exhibit reduced fatigue strength when compared to a gently ground surface because the fatigue enhancement by the compressive stress associated with low-stress grinding is not present.

Table 10 Effect of method of machining on fatigue strength Alloy

Machining operation

4340 steel, 50 HRC

Gentle grinding Electropolishing Abusive grinding Gentle milling Gentle grinding Chemical milling Abusive milling Abusive grinding Gentle grinding Electrochemical machining Conventional grinding Electrical discharge machining

Ti-6Al-4V, 32 HRC

Inconel 718, aged, 44 HRC

Endurance limit in bending, 107 cycles MPa ksi 703 102 620 90 430 62 480 70 430 62 350 51 220 32 90 13 410 60 270 39 165 24 150 22

Gentle grinding, %

100 88 61 113 100 82 52 21 100 65 40 37

Source: Ref 9 Effe ct of Gr in din g. Endurance limits vary with selected changes in grinding parameters (Fig. 24). When abusively grinding, there is a tendency to form patches or streaks of untempered martensite (UTM) or overtempered martensite (OTM) on the surface. It has been found that the presence of either one of these two microconstituents is usually associated with a significant drop in fatigue strength. For example, the presence of a depth of UTM as little as 0.01 mm (0.0005 in.) or as large as 0.09 mm (0.0035 in.) produces a decrease in endurance limit from 760 MPa (110 ksi) down to 480 to 520 MPa (70 or 75 ksi) (Fig. 24). Typically, residual tensile stresses are involved in this condition.

Fig. 2 4 Loss of fat igue st rengt h from t he abusive grinding of 4340 st eel ( quenched and t em pered t o 50 HRC) .

Fat igue t est s involved cant ilever bending at room t em perat ure and zero m ean st ress. Source: Ref 9

Retempering of the workpiece does not correct the problem. Although it tempers the UTM and reduces its brittleness, it does not restore the softening of the OTM to its prior condition. In addition, the tempering cycle does not reduce the residual tensile stresses formed in the abusive grinding operation. The effect on fatigue strength from EDM and grinding of aged Inconel 718 is illustrated in Fig. 25. After gentle grinding, the Inconel 718 had an endurance limit of 410 MPa (60 ksi). With either gentle or abusive EDM, the endurance limit dropped to 150 MPa (22 ksi).

Fig. 2 5 Effect of EDM and grinding on t he fat igue st rengt h of I nconel 718. Fat igue t est s involved cant ilever bending at room t em perat ure and zero m ean st ress. Source: Ref 9

The effect of ECM on the fatigue properties of Ti-6Al-4V is shown in Fig. 26. The endurance limit after low-stress grinding was 460 MPa (67 ksi). The tests were conducted on flat specimens that were longitudinally surface ground to a surface finish of 0.35 m (14 in.). Specimens that were electrochemically machined by a frontal or pocketing operation had an endurance limit of 410 MPa (60 ksi) for the same surface finish. When the titanium specimens were electrochemically machined by a trepanning operation, the endurance limit was reduced to 280 MPa (40 ksi), a 40% drop with respect to the low-stress grinding conditions.

Fig. 2 6 Effect of ECM on t he fat igue st rengt h of Ti- 6Al-4V. Fat igue t est s involved cant ilever bending at room t em perat ure and zero m ean st ress. Source: Ref 9

I m pr ovin g or Pr e se r vin g M e ch a n ica l Pr ope r t ie s W it h Sh ot Pe e n in g. Shot peening provides the most practical technique for enhancing the mechanical properties affected by surface alterations. It is frequently used in industry to improve the fatigue strength and stress-corrosion properties of most structural alloys subjected to harsh environments and high stress. Shot peening not only enhances the properties of gently machined materials but also improves the properties of metals that have been processed by machining techniques that tend to abuse or degrade fatigue strength and other mechanical properties. Shot peening parameters should be selected to meet surface finish specifications.

Fatigue properties can also be improved by other cold-work finishing processes, such as the burnishing of holes and the roll burnishing of fillets. The advantage of shot peening is that it can be applied to large surfaces and to contoured or irregular surfaces. With shot peening, the magnitude and depth of the compressive stress layer can be regulated by controlling the type and size of the shot, the intensity, and the coverage. However, the compressive stress produced by shot peening can be reduced in service by prolonged exposure to elevated temperature. I m pr ovin g Fa t igu e St r e n gt h . Shot peening can improve the fatigue strength of machined metals (Table 11). On

4340 steel at 50 HRC, for example, shot peening improved the fatigue strength of a gently ground surface by 10% (Table 11). Shot peening after roughing EDM increased the fatigue strength to 130% of the gently ground level. Similar improvements in fatigue strength by shot peening occur on surfaces produced by ECM and ELP (Table 11).

Table 11 Effect of shot peening on fatigue strength Alloy

Machining operation

4340 steel, 50 HRC

Gentle surface ground Gentle surface ground and shot peened Abusive surface ground Abusive surface ground and shot peened Electropolished Electropolished and shot peened Gentle surface ground EDM roughing EDM roughing and shot peened EDM finishing EDM finishing and shot peened ECM ECM and shot peened ELP ELP and shot peened

Inconel 718, solution treated and aged, 44 HRC

Endurance limit in bending, 107 cycles MPa ksi 703 102 772 112 430 62 630 92 620 90 660 96 410 60 170 25 540 78 170 25 480 70 285 41 560 81 290 42 540 79

Gentle grinding, %

100 110 61 90 88 94 100 42 130 42 117 68 135 70 132

Note: Shot size: S110; shot hardness: 50-55 HRC; coverage: 300%. Source: Ref 14 Re t a r din g St r e ss- Cor r osion Cr a ck in g. It has been shown that the compressive stresses introduced by shot peening

retard stress-corrosion cracking (Ref 14). For example, shot peening increased the stress-corrosion life of 4340 steel specimens that had been prestressed to 75% of the yield strength and exposed to 3.5% sodium chloride solution at room temperature. Shot peening increased the life of gently ground specimens from 400 to 850 h (Fig. 27). On abusively ground specimens, shot peening increased the life from 200 to over 1000 h, while on an electropolished specimen, the life was increased from 300 to 400 h by shot peening.

Fig. 2 7 Effect of shot peening on t he st ress- corrosion resist ance of AI SI 4340 st eel ( 50 HRC) . Source: Ref 14

Abr a sive t u m blin g is also an effective technique for improving fatigue and surface properties. Abrasive tumbling can

be used to reverse unfavorable tensile stress by inducing a compressional stress in the surface layer. M e a su r in g Su r fa ce I n t e gr it y Effe ct s In a systematic study of surface integrity, there are two recommended sets of data: a minimum data set and a standard data set. Th e m in im u m da t a se t is used for initial screening and consists of measurements of surface texture, macrostructure, microstructure, and microhardness alterations. The minimum recommended data set consists of the following (Ref 12):

• • • • • • • •

Surface roughness measurement according to Ref 1 or a microtopographic map Lay designation or photograph Macroscopic examination at a magnification of 10× or less Macroetch indications from die penetrants or magnetic flux Chemical etches Microscopic examination of a cross section of the altered surface (cross section examination at 1000× preferred) Indications of heat-affected zones and microhardness alterations (Optional) A scanning electron microscope can occasionally supplement surface texture measurements with a series of photographs at increasing magnifications (preferably 20, 200, 1000, and 2000×)

Th e st a n da r d da t a se t in measuring surface integrity consists of the above minimum data set plus the following

additional tests (Ref 12): • •

Residual stress profile or distortion measurements Fatigue tests (screening only)

Residual stresses in surfaces can be determined with x-ray diffraction techniques. A residual stress profile requires a determination of stress not only at the surface but also into the surface. This can be done by taking multiple x-ray readings after successive layers are removed by etching. Full-reverse bending at room temperature using tapered flat specimens taken to 107 cycles is recommended for fatigue testing. M a cr oscopic e x a m in a t ion involves some type of visual inspection for detecting visible cracks and other surface

defects. Die penetrants are generally applied to nonmagnetic materials, while magnetic particle inspection can detect small cracks in magnetic materials. Parts manufactured from martensitic high-strength steels can be visually inspected after a macroetch for evidence of untempered or overtempered martensite resulting from grinding or machining. A dilute solution of nitric acid is usually used to etch the damaged areas so that they can be visually detected. Typically, untempered martensite appears white, and the overtempered areas appear darker than the background material. A specific etching technique for detecting grinding damage in hardened steel is given in Table 12.

Table 12 Etching techniques for the detection of grinding injury in hardened steel Operation Solution used Description, time, or function Double-etch method 4-5% nitric acid in water Until black, 5-10 s; do not over etch. 1. Etch No. 1 Warm water To remove acid 2. Rinse Methanol or acetone(a) To remove water 3. Rinse 5-10% hydrochloric acid in methanol or acetone(a) Until black smut is removed, 5-10 s 4. Etch No. 2 Running warm water To remove acid 5. Rinse 2% sodium carbonate + phenolpthalien indicator in water To neutralize any remaining acid 6. Neutralize Methanol To remove water 7. Rinse Warm air blast 8. Dry Low-viscosity mineral oil with rust inhibitor To enhance contrast, prevent corrosion 9. Oil dip Nital etch method 5-10% nitric acid in ethanol or methanol Until contrast is evident 1. Etch 2. Repeat steps 5-9 above. Dark areas are overtempered, light areas are rehardened, uniform gray indicates no injury.

Source: Ref 15 4% nitric acid in water for etch No. 1 used with 2% hydrochloric acid in acetone for etch No. 2 sometimes gives greater sensitivity on high-carbon hardened steel. It is important that appropriate precautions be taken to avoid fire hazards, and good ventilation must be provided.

(a)

M icr oscopic e x a m in a t ion of a cross section of an altered surface is used to detect the following microstructural

alterations: • • • • • • • • • •

Microcracks Plastic deformation (a cross section parallel to the lay is suggested) Phase transformations Intergranular attack Micro defects such as laps and inclusions Built-up edges or deposits of debris Recast layers Selective etching Metallurgical transformations Heat-affected zones

These surface alterations can be detected by magnifying a cross section of an altered surface. A metallurgical mount holds the test specimen, and a cross section of the altered surface is ground and polished. The edge of the altered surface is then examined under a magnifying instrument with a magnification factor of 500 to 1000×. When polishing the cross section of an altered surface, special techniques for mounting and polishing are used to reduce edge rounding. Edge retention is important because surface alterations are only about 0.025 to 0.075 mm (0.001 to 0.003 in.) deep and because depth of field decreases as magnification factors increase. One technique is to vacuum cast the metallurgical mount and to use special procedures for polishing. Polishing units with a horizontal specimen holder, which is rotated with a preselected force against various grinding and polishing surfaces, have also been successfully used for edge retention. An additional mounting and preparation technique for edge retention is as follows (Ref 7): •

Samples are sectioned from the workpiece in a manner that leads to the least possible distortion or burring. Band sawing or hacksawing is preferred. A minimum of 0.75 mm (0.030 in.) is then removed from the cut surface using a 120-grit silicon carbide paper on a low-speed polisher

•

Copper molds (or tubes), 31.8 mm (1 in.) in inside diameter by 70 mm (2 in.) high, are placed on a pallet approximately 125 mm (5 in.) in diameter. The inner surface of the molds and the surface of the pallet are previously sprayed with a silicone releasing agent After placing a metallurgical specimen in a mold, a mixture of epoxy resin, hardener, and pelletized

•

•

•

• •

• •

•

aluminum oxide sufficient to produce a layer 6.4 to 9.5 mm ( to in.) is poured over the specimen. The ratio of resin to hardener is 4 to 1. The amount of pellets added is in the range of 10 to 15 g (0.35 to 0.5 oz). The hardness or abrasive level of the pelletized material used (low, medium, or high fired) is strictly a function of the alloy to be prepared and its hardness characteristics The pallet containing the molds is placed in a vacuum chamber at 10 to 100 Pa (1 × 10 -2 to 1 × 10-3 torr) to degas the mixture, thus improving the adherence of the epoxy and pellets to the surface of the specimen. When vigorous bubbling of the mixture decreases after vacuum impregnation, sufficient resin and hardener (4 to 1 ratio) is added to produce a mount approximately 25 mm (1 in.) high The mounts are cured at a temperature not greater than 20 °C (70 °F) for approximately 10 h. Casting of the mounts is accomplished during the latter portion of the laboratory workday so that curing occurs overnight After curing, the mounts are placed in an oven at a temperature of 65 °C (150 °F) for 1 h, after which they are removed from the molds Approximately 0.50 mm (0.020 in.) of stock is then removed from the as-mounted metal surface on a positive-positioning automatic polishing unit, using the side of a 25 × 330 mm (1 × 13 in.) aluminum oxide 320-grit grinding wheel as the grinding medium. Water is used as a coolant Subsequent rough grinding is performed wet on silicon carbide papers or equivalent ranging from 240 to 600 grit For steels and nickel- and cobalt-base superalloys, the intermediate polish is performed on an automatic polisher using a polishing cloth with a soft nap texture and 6 m diamond paste. The final polish is achieved using deep nap or pile cloth similar to billiard cloth with a suspension of 0.1 m or finer aluminum oxide in water. Titanium and refractory alloys require an etch-polish cycle (using a slurry of hydrogen peroxide, water, and 0.1 m or finer aluminum oxide), which is accomplished between a diamond polish and a final polish procedure. The final polish for titanium and refractory alloys is accomplished on a vibratory polisher using a deep pile cloth with a suspension of 0.1 m or finer aluminum oxide in water Samples are etched by swabbing. Examples of some typical etchants used are given below:

Material Steels Nickel-base alloys Titanium alloys

Etchant 2% HNO3 and 98% denatured anhydrous alcohol 100 mL HCl, 5 g CuCl2·2H2O, and 100 mL denatured anhydrous alcohol 2% HF and 98% H2O or 2% HF, 3% HNO3, and 95% H2O

Gu ide lin e s for M a t e r ia l Re m ova l, Post pr oce ssin g, a n d I n spe ct ion The following guidelines are meant to serve only as general or starting recommendations for the machining of structural components. More detailed information on guidelines for surface integrity is given in Ref 16. Data and experience indicate that these practices will lead to increased surface integrity in some important applications. However, the current state of knowledge of surface alterations is such that general recommendations are not always applicable for all specific surface integrity situations. For highly critical parts, it is mandatory to make individual, specific evaluations. It is also important to note that these guidelines were formulated based on experience with structural surfaces as opposed to mating or bearing surfaces. However, many of these guidelines are also applicable to the mating surfaces of bearings, cams, gears, and other similar parts. Surface integrity controls often result in increased manufacturing costs and decreased production rates. Therefore, surface integrity practices should not be implemented unless the need exists. Process parameters that provide surface integrity should be applied selectively to critical parts or to critical areas of given parts to help minimize cost increases. These surface integrity guidelines are primarily intended for application to metal removal processes used for final surface generation rather than roughing cuts. It is important, however, to know the type and depth of surface alterations produced

during roughing so that adequate provisions can be made for establishing surface integrity during finishing operations by removing damaged surface layers. Abr a sive Pr oce sse s. Grinding distortion and surface damage can be reduced by using low-stress grinding conditions.

Low-stress grinding conditions (Table 9) differ from conventional practices by employing softer-grade grinding wheels, reduced grinding wheel speed, reduced infeed rates, chemically active cutting fluids, and coarse wheel dressing procedures. Low-stress grinding and other guidelines should be used as follows: • • • • • • • • • •

Low-stress grinding should be used for removing the last 0.25 mm (0.010 in.) of material If low-stress grinding is required for finish grinding, then conventional grinding can be used to within 0.25 mm (0.010 in.) of finish size if the materials being ground are not sensitive to cracking Alloys for high-stress applications, such as titanium and high-temperature nickel and cobalt alloys, should be finished with low-stress grinding instead of conventional grinding Frequent dressing of grinding wheels can reduce surface damage by keeping the wheels open and sharp, thus helping to reduce temperatures at the wheel/workpiece interface Cutting fluids, properly applied, help promote surface integrity. As a general rule, at least 10 L/kW (2 gal./hp) per minute is needed Hand wheel grinding of sensitive alloys should be discouraged When abrasive cutoff is used, steps should be taken to determine the extent of the disturbed layer and to make proper stock allowances for subsequent cleanup by suitable machining Controls for hand power sanders should be maintained The relatively new high-speed grinding processes should not be used for finishing highly stressed structural parts unless a standard data set is developed Abrasive processing and especially finish grinding must be accomplished under strict process control when employed for the manufacture of aerospace components

Ch ip Re m ova l Ope r a t ion s. For turning and milling, there are at least two very important steps that will improve

surface integrity. First, machining conditions should be selected that will provide long tool life and good surface finish. Second, all machining should be done with sharp tools. Sharp tools minimize distortion and generally lead to better control during machining. The maximum flank wear when turning or milling should be 0.13 to 0.20 mm (0.005 to 0.008 in.). A good rule of thumb is to remove the tool when the wearland becomes visible to the naked eye. Guidelines for other chip removal operations are as follows: • • • • •

Rigid, high-quality machine tools are essential All hand feeding during drilling should be avoided The wearland on drills should be limited to 0.13 to 0.20 mm (0.005 to 0.008 in.) The finishing of drilled holes is imperative. The entrance and exit of all holes should be carefully deburred and chamfered or radiused Special precautions should be taken when reaming holes in sensitive alloys. The should be employed for the reaming of straight holes. On tapered holes (using power-driven machines), hand feeding is permissible, but power feeding is preferred

•

All straight holes 7.9 mm (

• •

1.2 mm ( in.) on the diameter. The operator should visually check the reamer for sharpness after each operation. At the first sign of chipping, localized wear, or average flank wear beyond specification, the reamer should be replaced, and the hole should be inspected. In addition, regardless of the hole and reamer condition, a maximum number of holes should be specified for reamer replacement Deburring and chamfering or radiusing should be used to remove all sharp edges Honing is an excellent finishing operation for developing surface integrity. A multistone head is preferred; heads with steel shoes and/or steel wipers are not recommended Boring can be used as a finish machining operation if roughness is within the manufacturing engineering limits. The tool wearland in finish boring should be limited to 0.13 mm (0.005 in.), but it should often be far less than this in order to achieve the desired accuracy and surface finish

•

in.) or larger should be double reamed, with a minimum metal removal of

Ele ct r ica l, Ch e m ica l, a n d Th e r m a l M a t e r ia l Re m ova l Pr oce sse s. Guidelines for these processes are as follows:

•

• •

•

•

Whenever EDM is used in the manufacture of highly stressed structural parts, the heat-affected layer produced should be removed. Generally, during EDM roughing, the layer showing microstructural changes, including a melted and resolidified layer, is less than 0.13 mm (0.005 in.) deep. During EDM finishing, it is less than 0.025 mm (0.001 in.) deep Surface integrity evaluations should be made when CM and ECM processes are used for finishing critical parts When substituting ECM for other machining processes, it may be necessary to add postprocessing operations such as steel shot or glass bead peening or mechanical polishing. Some companies require the peening of all electrochemically machined surfaces of highly stressed structural parts Special cognizance should be taken of the surface softening that occurs in the chemical and electrochemical machining of aerospace materials. Hardness reduction for CM and ECM range from 3 to 6 HRC points to a depth of 0.025 mm (0.001 in.) for CM and 0.05 mm (0.002 in.) for ECM. Shot peening or other suitable postprocessing should be used on such surfaces to restore mechanical properties Laser beam machining develops surfaces showing the effects of melting and vaporization. It is suggested that critical parts made by LBM be tested to determine if surface alterations lower the critical mechanical properties

Post pr oce ssin g guidelines are given below:

•

• •

•

•

•

Cracks, heat-affected layers, and other detrimental layers created during material removal processing should be removed from critically stressed parts. In addition to conventional machining processes, chemical machining and abrasive flow machining are successfully used for this purpose Steel shot and glass bead peening as well as burnishing can be used to improve surface integrity The use of controlled shot peening practices to restore the fatigue life of components processed by electrical, chemical, and thermal removal processes should be evaluated. Shot peening has been shown to be extremely effective in improving the fatigue life of specimens processed by ECM, EDM, and ELP (Table 11). Component tests are recommended to confirm the favorable trends shown in laboratory tests Postheat treatments following material removal are of limited utility. Stress-relief treatments, used to soften the hardened layers produced during the grinding of steels, do not restore the hardness of overtempered layers present immediately below the damaged surface layer. In addition, heat treatment does not repair any cracks produced during material removal Abrasive tumbling is an effective technique for improving surface properties, including fatigue. Abrasive tumbling can be used to reverse unfavorable tensile stresses by inducing a compressionstressed surface layer Washing procedures should be employed for critical parts and assemblies to remove all traces of cutting fluids, which may cause stress corrosion. Typical suspect compounds are sulfur compounds on aluminum- and nickel-base alloys and chlorine compounds on titanium alloys. Some companies use this precaution only for parts subjected to temperatures over 260 °C (500 °F). For applications at less than 260 °C (500 °F), carefully controlled washing procedures are often used to remove the chlorinated and sulfurized cutting oils

I n spe ct ion pr a ct ice s should be reviewed and amplified to meet surface integrity requirements. Some inspection

practices include the following: •

•

Microscopic examination, including microhardness testing, is used on a sampling basis to determine the type of surface layer being produced and its depth. This method can be used to check for such defects as microcracks, pits, folds, tears, laps, built-up edge, intergranular attack, and sparking The white layer or overtempered martensite produced during the grinding of steels can be detected by immersion etching using a 3 to 5% aqueous nitric acid solution

• •

Magnetic particle, penetrant inspection, ultrasonic testing, and eddy current techniques are recommended for detecting macrocracks X-ray diffraction methods are available for detecting residual surface stresses

Re fe r e n ce s 1. "Surface Texture (Surface Roughness, Waviness and Lay)," ANSI/ASME B46.1-1985, American Society of Mechanical Engineers, 1985 2. "Surface Texture Symbols," ANSI Y14.36-1978, American Society of Mechanical Engineers, 1978 3. "Surface Integrity," ANSI B211.1-1986, Society of Manufacturing Engineers, 1986 4. J. Peters, P. Vanherck, and M. Sastrodinoto, Assessment of Surface Typology Analysis Techniques, Annals of the CIRP, Vol 28 (No. 2), 1979 5. Machining Data Handbook, Vol 2, 3rd ed., Metcut Research Associates, 1980, p 18-15 to 18-37 6. G. Bellows and D.N. Tishler, "Introduction to Surface Integrity," Technical Report TM 70-974, General Electric Company, 1970, p 3 7. L.R. Gatto and T.D. DiLullo, "Metallographic Techniques for Determining Surface Alterations in Machining," Technical Paper IQ 71-225, Society for Manufacturing Engineers, 1971 8. Machining Data Handbook, Vol 2, 3rd ed., Metcut Research Associates, 1980, p 18-58 9. M. Field and J.F. Kahles, Review of Surface Integrity of Machined Components, Annals of the CIRP, Vol 20 (No.2), 1971 10. Machining Data Handbook, Vol 2, 3rd ed., Metcut Research Associates, 1980, p 18-87 11. M. Field, J.F. Kahles, and J.T. Cammett, A Review of Measuring Methods for Surface Integrity, Annals of the CIRP, Vol 2 (No. 1), 1972 12. W.P. Koster et al., "Surface Integrity of Machined Structural Components," AFML-TR-70-11, Metcut Research Associates, March 1970, p 2 13. A.R. Werner and P.C. Olson, Paper MR 68-710, Society of Manufacturing Engineers, 1968 14. W.P. Koster, L.R. Gatto, and J.T.Cammett, Influence of Shot Peening on Surface Integrity of Some Machined Aerospace Materials, in Proceedings of the First International Conference on Shot Peening (Paris, France), Sept 1981, Pergamon Press, p 287-293 15. W.E. Littman, "The Influence of Grinding on Workpiece Technology," Paper MR 67-593, Society of Manufacturing Engineers, 1967 16. J.F. Kahles, G. Bellows, and M. Field, "Surface Integrity Guidelines for Machining," Paper 69-730, Society of Manufacturing Engineers, 1969

Tool W e a r a n d Tool Life L. Alden Kendall, Washingt on St at e Universit y

I n t r odu ct ion CUTTING TOOLS WEAR because normal loads on the wear surfaces are high and the cutting chips and workpiece that apply these loads are moving rapidly over the wear surfaces. The cutting action and friction at these contact surfaces increase the temperature of the tool material, which further accelerates the physical and chemical processes associated with tool wear. In order to remove the unwanted material as chips, these forces and motions are necessary; therefore, cutting tool wear is a production management problem for manufacturing industries. To successfully manage machining processes, production engineers and managers need to establish a system that: • • • • • •

Selects the proper machine tools and cutting tools to produce the geometric features in a part being machined from a particular material Ensures that the tool distribution system provides quality tools having the required geometry Specifies the correct cutting velocity, tool feed rate, and tool engagements with the workpiece Establishes on-line or off-line procedures to monitor the condition of the cutting tool and the quality of the surfaces machined by the tool Has maintenance procedures that ensure consistent machine tool operation Takes into account the cost of machine tool operation and tool use, permitting a clear idea of the economic objective for the machining system

Such a system can provide the necessary information to determine when a tool should be changed. Unfortunately, there are many variables to consider; thus, it is not surprising that tool wear assessment and tool change decisions are difficult problems.

Th e W e a r En vir on m e n t Cutting tool wear occurs along the cutting edge and on adjacent surfaces. Figure 1 shows a view of the cutting process in which the rake and clearance surfaces intersect to define the cutting edge. Figure 2 in the article "Mechanics of Chip Formation" in this Volume shows a similar view but adds the stress and strain states that are present in the material and chip; Fig. 4 from the same article also shows the machining forces and velocities that produce this state of stress and strain. These are average forces and velocities. Cutting tool wear is localized on specific surfaces where stress, strain, velocity, and temperature are above critical levels. It is important to understand where these critical conditions exist and how they interact to cause tool wear.

Fig. 1 Chip, workpiece, and t ool relat ionship

W e a r Su r fa ce s. Figure 2 shows how the sharp tool of Fig. 1 may wear. Along the rake surface, the chip motion and

high normal stress have produced a wear scar called crater wear. Along the clearance surface, the tool motion and high normal stress have increased the area of contact between the tool and work, producing flank wear. Lastly, the cutting edge radius has increased. Figure 3 shows the characteristic wear surfaces on a turning tool insert, end mill, form tool, and drill. The cutting edge view shown in Fig. 1 and 2 is identified as section A--A in Fig. 3.

Fig. 2 Typical wear surfaces

Fig. 3 Wear surfaces on com m on t ools due t o t he t ool m ot ion, V

These figures show how the wear process changes the geometry of these different types of cutting tools. Flank wear decreases the diameter of the end mill as well as the depth of cut for a lathe tool. Both of these changes in the geometry of the cutting tool could produce out-of-tolerance dimensions on machined parts. The edge wear and crater wear on the rake surface alter the state of stress and strain in the cutting region, thereby changing cutting forces and the mechanics associated with the chip-making process. Severe geometric changes that decrease the angle between the rake and clearance surfaces can weaken the tool so that the edge may suddenly fracture. It should be apparent that the location and size of these wear surfaces play an important role in determining the useful life of the cutting tool. Localized stresses on cutting tool surfaces are a major contributing factor in regard to location and size of wear surfaces. St r e sse s on Tool W e a r Su r fa ce s. Figure 4 shows the approximate distribution of normal and shear stresses on the

tool wear surfaces. The normal stresses, n, are caused by normal forces acting along the rake surface, the cutting edge surface, and the clearance surface. In addition to the normal stresses, Fig. 4 shows the shear stresses, , that act along the surface of the tool and are associated with sticky and sliding shear processes. For the sticky zone, the normal force has a magnitude that results in a shear stress component that equals the shear yield strength, y, of the strain-hardened workpiece material. Rather than sliding along the surface, the chip tends to adhere and periodically separate along shear fracture planes. The existence and size of this sticky zone is dependent upon the magnitude of the normal force and the friction coefficient along these surfaces. The sliding zones have friction forces and associated surface shear stresses that vary according to the normal force and coefficient of friction. The state of strain created by the shear stresses associated with the sticky and sliding zones on the rake surface is generally called the secondary shear zone and is identified in Fig. 1. Tool surface roughness and lubrication conditions affect the magnitude of these surface shear stresses.

Fig. 4 Wear surface st resses

The primary shear zone shown in Fig. 1 extends from the cutting edge to the surface and is the zone where the chip material is plastically deformed and sheared from the work material. The complex state of stress along the cutting edge is caused by the strain associated with the chip separating from the work and moving along the rake surface of the tool and the strained material that remains a part of the work material. This conflict between chip motion strain and stationary workpiece strain produces a plowing action by the cutting edge. Normal stresses can become very high and exceed the strength of the tool material, causing plastic deformation or fracture of the cutting edge. A sticky zone may not exist for certain cutting conditions; however, the plowing action along the cutting edge always exists to some degree because it is impossible to create a cutting edge with no radius and with a primary shear zone that is a perfect plane. The magnitude of the state of stress in the cutting region also varies with time and creates a potential fatigue failure environment. M ot ion Alon g t h e W e a r Su r fa ce s. One way to increase machining productivity is to increase the volumetric chip removal rate. Volumetric chip removal rate is the product of the engagement area of the tool with the work material times the cutting velocity, V; thus, one option the process planner has is to increase the cutting velocity. This productivity gain must be balanced against increased tool wear caused by higher cutting velocities.

In the article "Mechanics of Chip Formation" in this Volume, the chip velocity, Vc, and the shear velocity, Vs, were shown to be functionally related to the primary shear angle, , the rake angle, , and the cutting velocity, V. The cutting velocity is the relative velocity between the clearance surface of the tool and the work, while the chip velocity is the relative velocity between the chip and the rake surface of the tool. The magnitude of these two velocities and the related shear stresses at these interfaces determine the amount of thermal energy released per unit of contact area. The magnitude of the shear velocity causes a high strain rate in the primary shear zone and in the sticky zone. The volume of material strained by this strain rate releases additional thermal energy. This thermal energy is the heat source that causes the temperature of the workpiece, cutting tool, and chip to increase. The productivity gain from increasing the cutting velocity proportionately increases wear surface velocities, strain rates, and release of thermal energy, thus increasing the wear environment temperature. Te m pe r a t u r e s in t h e W e a r Zon e s. The difference between the thermal-energy release rate and the thermal-energy

dissipation rate determines the temperature of the material in these wear zones. Thermal-energy dissipation is a function

of the thermal-conductivity properties of the tool material and workpiece material. Additionally, the workpiece size and specific heat determine the workpiece heat capacity; to a lesser extent, the surface area plays a role in convective-heat transfer to the surrounding air. If a cutting fluid is used, its conductivity and convective-heat transfer boundary coefficient with the hot tool and workpiece play an important role. The development and ultimate selection of tool materials are based on the ability of the tool to maintain hardness, toughness, and chemical stability at high temperatures. Even with the best of tools, these properties eventually adversely change with increasing temperatures. Figure 5 shows the change in yield strength as the temperature changes for three common cutting tool materials.

Fig. 5 Cut t ing t ool m at erials yield st rengt h as a funct ion of t em perat ure. Lower curve is high-speed st eel. Upper t wo curves are t ungst en carbide. Source: Ref 1

There have been many studies of temperatures in the cutting tool. Most commonly, average temperature conditions are approximated using the following type of relationship (Ref 2):

where T is the mean temperature (°F) of tool-chip interface, u is the specific cutting energy, which is the energy used per unit volume of material removed, V is the cutting speed, h is the undeformed chip thickness (see Fig. 1), and k, , c are the conductivity, density, and specific heat, respectively, of the workpiece material.

Figure 6 shows a typical temperature distribution. The high temperature gradient can cause portions of the tool to be at dangerously high temperature levels, well above the mean temperature. The mean temperature relationship identifies the important variables and their effect on temperature. High-strength work materials have high values of specific cutting energy. Some of these materials, such as titanium, are also poor conductors and are of low density. This combination of conditions results in higher tool temperatures, and the cutting tool material used must maintain its hardness and strength at these temperature levels if it is to have a useful service life.

Fig. 6 Generalized t em perat ure dist ribut ions in t he cut t ing region

Coolants are a mixture of lubricating and heat dissipation constituents; however, their primary function is to cool the tool. The lubricating component becomes more important at lower cutting speeds where there is less heat generated but a greater tendency for the chip and work material to adhere to the surfaces of the tool. More detailed information on machining coolants can be found in the article "Metal Cutting and Grinding Fluids" in this Volume. Combining these temperature effects with the state of stress, strain rates, and material motion makes the wear zones of the tool a complex battleground in which these conditions interact to trigger the mechanisms that cause wear.

W e a r M e ch a n ism s Material scientists and engineers have documented many volumes of information on wear. One important area of concern in these studies has been the identification of wear mechanisms. What has emerged is that a particular wear mechanism is dependent upon the contact stress, relative velocities at the wear interface, temperature, and physical properties of the materials in contact.

For a particular set of contacting materials, wear mechanism maps have been used to identify the ranges of normal pressure and velocity that result in a particular wear mechanism (Ref 3). Figure 7 shows the general structure of such a map when the area of the contacting wear surface remains constant. Temperature is not a mapped variable; however, it is a variable that is dependent upon the pressure (normal stress), velocity, and size of the wear surface, as discussed in the previous section. The four broad classes of mechanisms shown in Fig. 7 are: • • • •

Seizure Melt wear Oxidation/diffusion-dominated wear Plasticity-dominated wear

Fig. 7 Wear m echanism m aps and safe operat ing regions for cut t ing t ools. Source: Ref 3

Cutting tool wear research has established similar maps for feed rate, which is related to pressure for a given engagement condition, and velocity (Ref 1). The dashed lines in Fig. 7 show some possible boundaries that could be used to define a safe operating zone. This overlay of the two maps provides a framework to discuss the progressive wear and failure conditions for cutting tools. I n it ia l W e a r M e ch a n ism s. The two materials in contact have surface roughness irregularities in the form of

protrusions or asperities. At the interface, asperities from the two materials touch, creating tiny contact areas. The total area from these contact points is a fraction of the projected area of the contact surface. The stresses and heat are

intensified in the asperities, and partial removal may occur due to seizure accompanied by fracture of the asperity or melting in the asperity. As these asperities are removed, the initial surface roughness is altered and the contact area increases. If the force conditions remain unchanged, pressure decreases and the active wear mechanisms change to plasticity and/or mild oxidation/diffusion dominated wear. This initial wear period creates small, visible wear surfaces. St e a dy- St a t e W e a r M e ch a n ism s. A velocity and normal stress condition that would continue to cause seizure and melt should be avoided because it would soon cause complete failure of the cutting tool. Assuming that such conditions do not exist, the wear surfaces get progressively larger. If the wear surfaces are plasticity dominated, small particles of material are mechanically deformed and fractured away from the wear surface. Generally called abrasion, this is the most common wear process along the clearance surfaces of most tools.

As discussed earlier, normal stress and temperature vary over the wear surfaces so that a plasticity mechanism that dominates in one wear zone may not dominate in another. Figure 6 indicates that the maximum tool surface temperature occurs on the rake surface a small distance from the cutting edge. This is where the crater wear condition occurs, with diffusion wear often being the dominant mechanism (Ref 4). The high temperature and pressure cause atoms to move between the contacting materials, and this diffusion process locally aids in the removal of tool material to form the crater. In very hard cutting tool materials, such as ceramics, for which high cutting velocities are commonly used, these oxidation and diffusion mechanisms may be responsible for a majority of the wear. A built-up edge (BUE) condition, which is indicated by a dashed line in Fig. 7, affects the process in two ways. Near the cutting edge, the higher pressures could cause particles of work material to adhere to the cutting tool in the sticky zone. If the shear forces due to chip movement are high enough, the bond will be temporary and the adhered material will fracture away from the tool surface. When it fractures, small particles of tool material may be removed with the previously adhered material. This then is a wear process and would be associated with conditions in the safe zone just outside the built-up edge line. The second effect on the process occurs when the BUE is not fractured away by the chip motion and remains to alter the geometry of the cutting edge. The presence of the BUE changes the shear angle, causing instabilities in the chip-forming process and damage to the machined surface. The lubricating characteristics of cutting fluids are helpful in eliminating this BUE condition. A discussion of the BUE condition can also be found in the article "Mechanics of Chip Formation" in this Volume. Wear can also occur as chipping along the cutting edge. Such chipping more commonly occurs when the cutting edge intermittently removes chips. This results in cyclic impact and thermal loading of the cutting edge. These two cyclic loading states can initiate small cracks and then propagate these cracks or other residual cracks to form the chips. These abrasion, oxidation, diffusion, and chipping wear mechanisms that occur at operating conditions within the safe zone shown in Fig. 7 cause the initial wear surfaces to enlarge over time. This surface life period is often referred to as the steady-state wear period. Te r t ia r y W e a r M e ch a n ism s. The steady-state period of wear eventually enlarges the wear surfaces to a critical size

that triggers accelerated wear. In tools that have a hard, wear-resistant coating such as titanium nitride, wear through this coating or separation of small volumes of the coating (Fig. 8a) exposes the less-resistant core material (Fig. 8b), resulting in accelerated wear. The pressures and velocities on these enlarged surfaces begin increasing the temperature so that the rapid oxidation/diffusion and local seizure or melting conditions shown in Fig. 7 exist, causing rapid destruction of the tool. A similar vulnerability exists for the softer, core material of coated tools. A tool change must be made before this point is reached. Figure 9 shows these three zones in terms of amount of wear over time.

Fig. 8 Delam inat ion of t it anium nit ride coat ing from HSS end m ill. ( a) 940× . ( b) 4700× . Source: Ref 5

Fig. 9 Tool wear curves for different cut t ing velocit ies

As a final factor complicating wear mechanisms, there is a growing trend toward high-speed machining. This means cutting velocities five to ten times normal speeds are used. New tool materials and machine tool designs have made this possible; however, these changes alter the wear environment. Proportionally more thermal energy is removed by the chip due to the decrease in contact time between the tool and the chip. The higher velocities increase absolute temperatures on tool wear surfaces. Abrasive wear may become less important, and the diffusion and oxidation processes dominate in creating and enlarging the wear surfaces. See the Section "High-Productivity Machining" in this Volume for information on high-speed machining practice.

M a ch in e , Cu t t in g Tool, a n d Tool W e a r I n t e r a ct ion s Tool wear, too, is influenced by the construction of the tool; the rigidity of the tool, workpiece, and machine tool system; and the proper positioning of the tool. Other Sections of this Volume provide information on tool materials, different machining processes, and the mechanics of the cutting process. The information and suggestions given in these Sections must be followed in order to avoid some undesirable tool wear side effects. A number of these effects are emphasized in this article. Cutting tools may be made of high-speed steel; high-speed steel coated with a thin, hard surface; or cemented carbide materials. As mentioned earlier, the coated material provides good wear resistance until the hard coating is removed, exposing the softer core material. The cemented carbides are very hard, but brittle, and have a lower thermal conductivity than high-speed steel. Inserts may easily fracture if the mechanical attachment to the insert toolholder does not firmly hold the insert. An insert cannot easily dissipate heat through the attachment interfaces to the more massive toolholder; this situation could cause unacceptably high temperatures in the insert. Carbide inserts that are brazed in place are better supported and dissipate heat better; however, they are still susceptible to fracture failures caused by impact or cyclic loads. At times, the application requires small cutting tools and tool inserts. In such instances, thermal overload and fracture failures can be avoided by decreasing cutting velocities and feed rates, reducing the size of cut, and/or increasing the coolant flow rate. Sometimes tool wear is excessive even when the recommended cutting conditions are used. The cause of the problem may be the rigidity of the machine tool and workholding fixture. An older machine tool or one that has not been properly maintained can cause vibrations at the tool that accelerate the breakdown and wear of the cutting edge. Even with a well-maintained rigid machine and workpiece system, improper positioning of the tool can cause tool wear problems. Excessive overhang of the tool can cause dynamic instabilities similar to those mentioned above. In the cutting zone, the cutting edge must be positioned so that clearance surfaces do not rub against the machined surface. For a singlepoint cutting tool, this requires correct alignment of the point of the tool with the rotating workpiece centerline and careful positioning of the rake and flank surfaces so that correct angles exist between the tool surfaces and the workpiece surfaces. Each type of machining operation has such specifications, and the tool engineer should ensure that they are followed when tooling setups are made. Understanding the tool service environment and the conditions that cause wear is the first step in planning and controlling of tool use in machining. Unfortunately, research and development efforts in tool wear have not progressed to the point that the life of a cutting tool can be predicted using basic tool and work material properties plus fundamental science and engineering principles. Tool replacement decisions are dependent on predicting or sensing the magnitude of the wear damage caused by the active wear mechanisms on critical wear surfaces.

Tool Re pla ce m e n t Tools are replaced to minimize the probable consequences of a failure event. The tool wear environment is so complex that even with the utmost care it is possible that the tool may fail in service. A very conservative strategy may replace the tool frequently to reduce the probability of an in-process failure, but, as a consequence, interrupt the process so frequently that productivity decreases and tool costs increase. To establish the strategy objectively, a clear understanding of the failure event must be agreed upon, and suitable data concerning tool performance and operating costs must be available. Tool Fa ilu r e Eve n t s. The steady-state wear process removes material gradually until damage reaches a critical level.

Wear beyond this level is considered damaging to the process and represents a failed state for the tool based on wear. This is called a wear failure event. The wear environment also weakens the tool so that sudden variations in tool thermal and/or mechanical loading can cause rapid destruction of the tool cutting surfaces. These sudden variations can come from interrupted cuts, hard spots in the work material, or sudden loss of coolant in the cutting region. Sudden destruction can also be caused by residual defects in new tools. Even a new tool with no residual defects can be destroyed if the abrupt load changes are large enough. This second failure state is called a catastrophic-failure event. Fa ilu r e e ve n t con se qu e n ce s are dependent on the type of feature being machined, the quality specifications associated with the feature, and the production economics. The type of feature determines the type of machine and tool used for processing. A feature such as a small-diameter hole has a potential risk for a catastrophic failure event that could

have severe consequences. A sudden increase in torsional loading could fracture the drill, leaving the fractured part in the hole and requiring expensive rework to remove the broken portion and resize the hole. In contrast, a pocket being milled out to a rough size using a large end mill has less severe consequences. A wear or catastrophic-failure event could mean loss of the tool, but little damage to the part. However, a final finish cut for this same pocket will change the failure event consequences: Excessive tool wear may result in an unacceptable surface finish or out-of-tolerance dimensions. Economic analysis can be used to establish a tool replacement strategy that optimizes the performance objective associated with the making of each feature. Common objectives are minimum cost, maximum production rate, and maximum profit; however, minimum acceptable reliability or minimum downtime could also be used. The tool replacement strategy is dependent on the mechanism by which wear and catastrophic-failure events influence the objective. To access failure event consequences and achieve feature production objectives, tool wear must be mathematically modeled and/or measured in a timely way in order to change tools strategically. Real time measurement of tool wear is an emerging but promising technology (Ref 6). Until it matures and is commercially available, mathematical models of tool wear will be primarily used. M ode lin g Tool W e a r . One of the earliest applications of science to production management was done by F.W. Taylor

(Ref 7). Recognizing that tool wear was dependent on cutting velocity, he developed the following equation using data from tool life tests:

Vtn = C This became known as Taylor's tool life equation, in which the tool lifetime, t, was related to cutting velocity, V, by means of the constants n and C. These constants were obtained by testing cutting tools at different cutting velocities and using a "tool life criterion" to establish the point at which the useful life of the cutting tool had ended. This criterion was a wear limit that could not be exceeded if a wear failure event was to be avoided. Figure 9 shows typical tool wear curves for different cutting velocities. The wear limit or failure criterion, W1, shows that the elapsed time before tool replacement increases with a decrease in cutting velocity. Taylor's relationship models this behavior (Fig. 10) and has been used by industry as originally conceived or in modified forms to the present time. In fact, the constant C, which equals the cutting velocity for 1 min of elapsed time before reaching the wear limit of the tool, has been widely used as a measure of machinability for a particular material being machined using a particular cutting tool and cutting condition. The cutting condition in this context is the feed rate and depth of cut (engagement) used to collect the data. Table 1 contains other tool life models that add feed rate and/or depth of cut as variables that influence tool life. In all of these, life testing of tools must be conducted.

Table 1 Tool life models proposed in the literature Linearized Taylor model Linearized Taylor model with interaction term Partial second-order model

ln t = b0 + b1x1 + b2x2 ln t = b0 + b1x1 + b2x2 + b12x1x2 ln t = b0 + b1x1 + b2x2 + b12x1x2 + b11

Full second-order model

+ b22 ln t = b0 + b1x1 + b2x2 + b3x3 + b12x1x2 + b13x1x3 + b23x2x3 + b11

+ b22

+ b33 Konig and DePiereux (Ref 8)

ln t = b0 + b1V

+ b3f

t = tool life; x1 = ln V, where V is the cutting velocity; x2 = ln f, where f is the feed rate; x3 = ln D, where D is the depth of cut; b0, b1, b2, . . ., b33 are experimentally determined coefficients.

Fig. 1 0 Taylor t ool life m odel using dat a from Fig. 9 and a ln- ln plot

The wear limit or failure criterion is an important decision to be made in the process of developing the tool life model. The size and characteristics of particular wear surfaces are often used. Figure 11 shows the variety of wear and fracture surfaces that may be present on a worn or failed cutting tool. This figure also shows how to measure the amount of wear on one or more of these wear surfaces. Flank wear has been commonly used, and Table 2 presents some typical values for some common tools.

Table 2 Typical flank wear limits Operation and material

Turning HSS Carbide Face milling HSS Carbide End milling-slotting HSS Carbide End milling-peripheral HSS Carbide Drilling HSS Carbide Reaming HSS Carbide Tapping HSS

Source: Ref 10

Flank wear land Average wear Maximum local wear mm in. mm in. 1.5 0.45

0.060 0.015

1.5 0.9

0.060 0.030

1.5 0.45

0.060 0.015

1.5 0.9

0.060 0.030

0.30 0.30

0.012 0.012

0.50 0.50

0.020 0.020

0.30 0.30

0.012 0.012

0.50 0.50

0.020 0.020

0.45 0.45

0.015 0.015

0.45 0.45

0.015 0.015

0.15 0.15

0.006 0.006

0.15 0.15

0.006 0.006

Go-no go gage

Tap fracture

Fig. 1 1 Cut t ing t ool failure m odes. ( a) Charact erist ic wear and fract ure surfaces on cut t ing t ools. ( b) Cat ast rophic failure. ( c) Typical wear m easurem ent s for a t urning t ool. VB = flank wear. Source: Ref 9

To use these wear limits, the tool life test must be stopped and a measurement made using optical instruments at suitable magnification levels. Generally, a toolholding fixture must be made to position the tool properly and consistently in the field of view. Typical flank wear scars are shown in Fig. 12(a) to (d). Each of these figures shows that the size of the scar varies. The end mill in Fig. 12(a) has enlarged flank wear toward the end where the flank wear from the end cutting edge and flute cutting edge overlap. The drill in Fig. 12(b) shows that the width of the flank wear scar is roughly proportional to the distance from the center of the drill. Velocity increases with distance from the center, and such increased wear is consistent with the dependence of wear on velocity.

Fig. 1 2 Typical flank wear scars. ( a) Flank wear on an end m ill. ( b) Flank wear on t he cut t ing edge of a drill. ( c) and ( d) Flank wear on a t urning t ool used for an ort hogonal cut on a ring. ( c) High-speed st eel t ool. ( d) Coat ed high- speed st eel t ool. Court esy of A.E. Bayoum i, Washingt on St at e Universit y

Figures 12(c) and 12(d) show the flank wear on a turning tool that was used for an orthogonal cut on the edge of a ring. In these figures, the nature and size of the scar reflect the difference between coated and uncoated high-speed steel tools. These figures also illustrate the importance of establishing where the flank measurements are to be made. It is apparent, too, that it is a good practice to take multiple measurements along the cutting edge; with multiple-tooth cutters, such measurements should be made on all of the cutting edges. If crater wear is part of the wear limit criterion, the topography can be analyzed by the depth of the crater or the projected area of the crater. For chipping, the size and number of such chips can be determined during examination. A wear limit criterion for chipping may have to be based on the number and size that detrimentally affect surface finish. Often it is necessary to use more detailed metallographic examination techniques. For external topographical examination of wear surfaces, a portion of the tool needs to be appropriately prepared for viewing under varying levels of magnification. A scanning electron microscope is sometimes used for this purpose. To evaluate internal damage, sectional views perpendicular to the cutting edge are generally needed. Near-surface thermal softening and phase changes are of interest, and microhardness measurements can be used to quantify the magnitude of these changes. Figure 8 is an example of the way high magnifications can be used to examine a failed tool and provide insight concerning the wear mechanisms that caused failure.

Tool Life Te st in g

A successful tool life testing program depends on establishing the test plan objective, designing the test, conducting the test, analyzing the results, and applying the results. These points are described below. Additional information on tool life testing can also be found in the article "Machinability Test Methods" in this Volume. Te st in g Pr ogr a m Ex pe r t ise Metallurgical, manufacturing-process, and experimentation knowledge is needed for a testing program. The first step is to find an individual or assemble a team that can provide this expertise. Metallurgical assistance is needed to: • • •

Provide the material information concerning the cutting tool and the workpiece Provide metallographic knowledge during examination of worn tools Interpret evidence concerning wear mechanisms and tool fracture

Manufacturing-process help is required to: • • • •

Select the machine tool, cutting tool, and geometric feature to be machined for the cutting test Ensure that the machine is properly set up and calibrated for the test Establish the correct operating procedure for the test equipment Help relate the results to productivity and economic gains

Experimentation assistance is needed to: • • •

Design the test plan including the statistically designed experiment if one is to be used Conduct the calibration tests and determine what assignable errors exist in the test equipment Conduct the data analysis and compute any parameters that are needed in the wear models

With this expertise in place, the testing program can begin. Te st Pla n Obj e ct ive s There are global and local objectives that must be dealt with. The global concern often comes from management, who have decisions to make that require tool wear data. Only two objectives will be defined and expanded upon using examples. (However, the tool wear problems of industry can provide many, diverse reasons to conduct tool wear tests.) The first is the problem of selecting tools to buy. This will be referred to as the acceptance objective. The second concerns the development of cutting condition data for a new tool-operation-material combination. This will be referred to as the operating objective. Many aspects of the test plan will be the same for these objectives; however, there are a number of important differences, which will be identified. These two objectives might be stated as: • •

The acceptance objective: To establish which HSS steel end mills to purchase for use in a machining center that machines parts from an AISI 4340 steel The operating objective: To develop a tool life model for an HSS end mill used in a machining center to cut AISI 4340 alloy steel

The local objectives are the testing conditions and wear criterion that should be used for these global objectives. These target more specifically the intended use of the tool and how its performance will be judged. For example, the following local objectives will be established: •

Acceptance local objectives: A sizable rough cut will be used at feed and speed conditions that are at the maximum generally used to machine this material on this particular machine. A limiting flank wear value will be used as a failure criterion

•

Operating local objective: A statistically designed experiment will be used to develop the tool life model for the range of operating conditions normally used for this tool-machine-material combination. A limiting flank wear value will be used as a failure criterion

It should be noted that these objectives are still broadly stated. During the test plan design stage, specific values will be established that appropriately define the bounds for these objectives. D e sign in g t h e Te st The establishment of test conditions, a sampling plan, and test plans are all integral to the design of the test. Est a blish m e n t of Te st Con dit ion s. A machine that represents the type of machining center that will be used in production must be identified and made available. It is important to examine this machine closely and perform all maintenance that is necessary to ensure that it will operate in a nominal way. The speed and feed settings must be checked to see whether they produce the correct motion values. The runout of the table and spindle must be checked.

It is desirable to add instrumentation to the machine and connect it to an appropriate data acquisition system. Power, spindle torque, and force components on the tool are useful measurements to obtain (Ref 5). Acoustic emission sensors are also proving to be of value (Ref 11). These measurements will not be used in the procedure described in this article; however, they should be obtained if the wear study results are to be used for process monitoring and control using on-line sensors. Size of the workpiece should be considered next. As discussed earlier, tool temperature influences the type of wear mechanism and the wear rate, and the mass of the workpiece is a variable heat sink as it is cut. For the current example, this temperature condition will be handled as follows for the use of acceptance and operating tests. The acceptance test part size will be large plates. The tools to be tested will be sequenced so that each vendor's tool will have a balanced exposure to the heat sink mass effects of the plates as they are reduced in size during machining. Coolant flow will be constant and will be measured to verify. The operating test part size is more difficult to establish. In a production environment, the goal would be to use coolant flow rates and cutting conditions that would maintain a tool temperature level below a safe upper bound. In the test environment, this same upper-bound temperature constraint is desirable, but it is also necessary to allow normal temperature changes below this upper bound caused by changes in the operating conditions. The recommendation is to use the largest workpiece for the machine being used and to monitor the surface temperature of the material a fixed distance from the cut. An infrared sensor could be used. When this temperature exceeds the upper bound value, a larger piece of material is called for. Cutting conditions for the acceptance objective must severely test each vendor's tools, whereas operating objective tests require conditions that represent the range of values used for the operating environment. Acceptance cutting conditions will be the same for every test cut. Some analysis will have to be made of the maximum speed and feed conditions normally used on that machine for the rough-cut size selected. Some sample cuts will most likely be necessary in order to establish that these conditions can be adequately handled by the machine tool. Operating cutting conditions will be systematically varied over the range of values that are anticipated for use in the production environment. A statistically designed experiment will be used to provide data for estimation of the coefficients in the partial second order model described in Ref 11 and given in Table 1. The cutting conditions to be varied will be feed rate and cutting velocity. A representative depth of cut will be used for all tests. A full 22 factorial design will be used with center point replication for a total of six tests. A fractional replication of this full design requiring fewer tests is another option, and more recent optimal designs have been proposed for tool life models (Ref 12). Sa m plin g Pla n . The work material for each of these test programs should come from the same manufacturer and carry

the same heat number. Verification of chemical composition and condition should be obtained. The tool sampling should be:

• •

Acceptance tools: Randomly selected from a batch of tools from each vendor. In this case, five tools should be used Operating tools: Randomly selected from a batch of the particular tool type being tested. In this case, six would be needed

Te st Poin t s. For both the acceptance and operating test plans, the test data are used to generate a wear curve for each

tool. This means that the test must be interrupted periodically to measure the flank wear on the tool. Equal intervals of time should be used. Con du ct in g t h e Te st An unneeded tool should be used to machine a starting surface on the edge of the workpiece. The test series can now be started and should be conducted in the same manner for each test plan. The end mills should be removed and placed in the measurement fixture. Each flute of these mills will have multiple measurements made along the length of the flank wear surface. Averages for each mill must be used when the results are analyzed. The test can be interrupted and measurements made until the wear limit has been exceeded. At the end of the test, metallographic examinations are conducted using sectioned portions of tools. These examinations will provide data concerning wear mechanisms and wear surface topography. An a lyzin g t h e Re su lt s The first step is the same for both testing plans; the wear curves are plotted. If all the test data fall within the steady-state wear region, a linear wear trend may exist, and such a line could be fitted to the data. If one assumes such data exist for the acceptance test, as plotted in Fig. 13, this data would provide a set of five tool life values for each tool vendor. The operating test plan would produce similar plots and six tool life values, one for each test condition. Further analysis will differ for each of the plans.

Fig. 1 3 Accept ance t est linearized wear curves for t he t ools of one vendor. The relat ionship of t he Weibull probabilit y densit y funct ion t o t he t ool life variabilit y is shown, as well as t he reliabilit y probabilit y.

Acce pt a n ce da t a a n a lysis seeks a quality measure for the tool of each vendor. Reliability is the measure and can be computed using the five tool life values for each vendor. The method of median ranks and a graphical technique are used to establish the parameters for the Weibull probability distribution for each vendor (Ref 13). This is shown in Fig. 14 for some hypothetical data. To compare one vendor with another, a tool life value, t, was selected, and the reliability computed using the values of (characteristic life parameter) and (shape parameter) for each vendor's tools. For the parameter values = 100 and = 6 calculated in Fig. 14 for t = 70 minutes, R(t) = 0.889. This number is the probability that the vendor's tools will survive this service life. Higher reliability values imply that the tools are of higher quality. A statistician can provide other tests that compare tool life values; however, this reliability measure is used to provide background for application of this test information for tool change decisions.

Fig. 1 4 Graphical m et hod t o develop t he Weibull param et ers

Ope r a t in g da t a a n a lysis establishes the value of the significant coefficients in the tool life model. Any coefficients that are statistically insignificant will be set to zero, which eliminates the product term in which the coefficient appears. The terms that remain define the tool life equation. Any combination of cutting condition variables within the span of the test cutting conditions can be substituted into the equation, and the tool life can then be computed. The method can also provide a confidence interval for this tool life estimate.

Applyin g t h e Re su lt s The results for these two tests will be used to indicate the economic impact of the wear information on the production process. Before addressing each of these test plans individually, some general background concerning such analyses is presented below. The primary objective is to select cutting conditions to obtain a tool service life that meets the process management objectives. The decision-making process must take into account the possibility that a wear or catastrophic-failure event may occur before replacement is made. These will be called premature failures. The tool life models in Table 1 predict average performance. If a statistical technique was used to develop the equation, this is an accurate assumption. A simple way to use these models is to select a replacement interval and then choose cutting conditions that satisfy the model. At these cutting conditions, about half of the tools will prematurely fail. Because this is generally unacceptable, the cutting conditions must be changed or the interval must be shortened to ensure fewer premature failures. The lower confidence bound on a statistically developed model can be used to make this margin of safety more precise. These approaches ignore some very important economic consequences. First, the decision does not take into account the gains to be made by increased material removal rates and tool usage penalties associated with such productivity gains. Figure 15 shows this trade-off for costs, but such relationships could also be based on production rates or profits. The recommended operating point is obtained by computing the operating condition that minimizes the total cost function

shown in Fig. 15. Second, the method incorrectly assumes that replacing premature failures costs the same as scheduled replacements. Premature failure not only stops the process, it can damage the part and add rework or scrapped-part expense to the process.

Fig. 1 5 Minim um cost operat ing point for a cut t ing process

Acce pt a n ce t e st pla n e va lu a t ion s are conducted to measure the impact of poor tool quality. The wear data provide

a probability distribution for tool life for each vendor's tools. Table 3 presents two common fixed interval replacement policies that can be used. The scheduled replacement policy is used in this example, and the expected replacement cost per unit of operating time computed, for different scheduled replacement intervals, ts. There are formal optimization approaches that can compute this interval directly (Ref 14); however, good, useful information can be easily obtained using a spreadsheet program with plotting capabilities. The and parameters of the Weibull distribution are used to obtain renewal function values from statistical tables (Ref 15). For certain conditions, an approximation for this function can be obtained directly by using the reliability function (Ref 16). Figure 16 shows typical plots for three vendors and indicates that the expected cost for the higher-quality tools was less sensitive to the choice of the scheduled replacement interval. This knowledge would certainly be of comfort to a process manager who otherwise would not know that by using the higher-quality tools, more flexibility was available concerning the choice of the replacement interval.

Table 3 Expected cost functions for tool replacement policies

Fig. 1 6 The expect ed replacem ent cost for t hree qualit y levels of t ools

Ope r a t in g t e st pla n e va lu a t ion s are used to establish operating conditions that minimize the variable costs to

machine this particular feature. The tool life model that was developed has two independent variables. It is possible to establish the combination of values for these variables that will minimize this cost; however, there are often other machining considerations that will fix one of these. As shown in Fig. 15, this variable cost, C, equals the sum of the production costs and the tool usage costs for the feature. The functional expression is:

C = c0t0 + nt(cott + Ct) where c0 is the cost per unit time for machine operation; t0 is the time to produce the feature; nt is the fraction of tool life used for the feature, which is equal to t0/t; tt is the time required for a tool change; and Ct is the cost of the cutting tool. The proper expression for t0 depends on the type of tool and feature machined by the tool. The value of t for any combination of operating conditions comes from the tool life model, and an evaluation of the cost C could be accomplished using a spreadsheet or formal optimization procedures (Ref 17). This procedure works well for a tool that is used to produce this same feature again and again until the tool is replaced. A typical application of this type is a transfer line. When the tool is used to produce a variety of features such as in a machining center, the analysis becomes more complex. A system that estimates the proportion of the life of the tool used for each feature must be used and then these values must be added together to establish when the tool needs to be changed.

Fu t u r e Tr e n ds The complex machining wear environment will continue to be studied and better understood by tribologists, material scientists, and manufacturing engineers. Its importance deserves such attention, and the research efforts by these investigators needs continuing support. The use of this wear knowledge for tool change decisions, process monitoring, and control strategies is even more important. Dependence on off-line laboratory testing and model development will eventually become too costly and time consuming for the current and future automated machining systems. The process will become the direct source of wear information through the use of new monitoring techniques and sensors. Local data bases that store performance information concerning the production of each feature will be able to access the progressive wear of the tool more precisely. Some of the motivation to develop such systems will come from changes in controller technology that will include the intelligent decision-making capabilities currently done by human experts.

Re fe r e n ce s 1. P.K. Wright, Physical Models of Tool Wear for Adaptive Control in Flexible Machining Cells, in Proceedings of the Symposium on Computer Integrated Manufacturing, Vol 8, American Society of Mechanical Engineers, 1986 2. M.C. Shaw, Metal Cutting Principles, Oxford University Press, 1984 3. S.C. Lim and M.F. Ashby, Overview No. 55, Wear-Mechanism Maps, Acta Metall., Vol 35 (No. 1), 1987, p 1-24 4. B.M. Kramer and N.P. Sub, Tool Wear by Solution: A Quantitative Understanding, J. Eng. Ind. Trans. ASME, Vol 102, 1980, p 303 5. A.E. Bayoumi and L.A. Kendall, Modeling and Measurement of Wear of Coated and Uncoated High Speed Steel End Mills, J. Mater. Shap. Technol., to be published 6. S. Jetly, Measuring Cutting Tool Wear On-Line: Some Practical Considerations, Manuf. Eng., July 1984, p 55 7. F.W. Taylor, On the Art of Cutting Tools, Trans. ASME, Vol 28, 1907 8. B.N. Colding and W. Konig, Validity of the Taylor Equation in Metal Cutting, Ann. CIRP, Vol 19 (No. 4), 1971, p 793 9. S. Kalpakjian, Manufacturing Processes for Engineering Materials, Addison-Wesley, 1984, p 489 10. D.A. Dornfeld, The Role of Acoustic Emission in Manufacturing Process Monitoring, in Proceedings of the 13th NAMRC, Society of Manufacturing Engineers, May 1985, p 69-74 11. S.M. Wu, Tool Life Testing by Response Surface Methodology, Parts I and II, J. Eng. Ind. Trans. ASME, Series B, Vol 86, 1964, p 105

12. W.J. Zdeblick and R.E. Devor, An Experimental Strategy for Designing Tool Life Experiments, J. Eng. Ind. Trans. ASME, Vol 100, 1978, p 441 13. K.C. Kapur, chapter 11 in Reliability in Engineering Design, John Wiley & Sons, 1977 14. A.K. Sheikh, L.A. Kendall, and S.M. Pandit, Probabilistic Optimization of Multitool Operations, J. Eng. Ind. Trans. ASME, Vol 102, 1980 15. J.S. White, "Weibull Renewal Analysis," Paper presented at the 3rd Annual Aerospace Reliability Maintainability Conference (Washington, D.C.), June/July 1964 16. S.B. Billatos and L.A. Kendall, Approximate Renewal Functions for Optimization, in Proceedings of the 40th Meeting of the Mechanical Failures Prevention Group, T.R. Shives, Ed., Cambridge University Press, 1985 17. G. Boothroyd, chapter 6 in Fundamentals of Metal Machining and Machine Tools, McGraw-Hill, 1975

H igh - Spe e d Tool St e e ls Alan M. Bayer and Bruce A. Becherer, Teledyne Vasco

I n t r odu ct ion HIGH-SPEED TOOL STEELS and their requirements are defined by The American Society for Testing and Materials in Specification A600-79 as follows: High-speed tool steels are so named primarily because of their ability to machine materials at high cutting speeds. They are complex iron-base alloys of carbon, chromium, vanadium, molybdenum, or tungsten, or combinations thereof, and in some cases substantial amounts of cobalt. The carbon and alloy contents are balanced at levels to give high attainable hardening response, high wear resistance, high resistance to the softening effect of heat, and good toughness for effective use in industrial cutting operations. Commercial practice has developed two groups of cutting materials: • •

The recognized standard high-speed tool steel, which serves almost all applications under mild to severe metal-cutting conditions A smaller group of intermediate steels, which are satisfactory for limited applications under mild to moderate metal-cutting conditions

The minimum requirements that must be met to be classed as a standard high-speed tool steel, and those for an intermediate high-speed tool steel, are listed in Table 1. To be acceptable for either group, an alloy must meet all of the requirements shown for that group.

Table 1 Requirements for high-speed tool steels per ASTM A 600 Requirement Chemical requirements Minimum alloy content by major elements Carbon Chromium Vanadium Tungsten + 1.8% molybdenum Minimum total alloy content based on tungsten equivalents ( Cr + 6.2 V + W + 1.8 Mo) Grades containing less than 5% cobalt Grades containing 5% or more cobalt Hardening response requirements Ability to be austenitized, and tempered at a temperature not less than 510 °C (950 °F) with a fine-grain structure (Snyder-Graff grain size 8 min) to

Standard

Intermediate

0.65 3.50 0.80 11.75

0.70 3.25 0.80 6.50

22.50 21.00

13.00 12.00

63 HRC

62 HRC

A chronology of some of the significant developments in high-speed tool steels is given in Table 2. The research work in 1903 on a 14% tungsten alloy led to the development of the first high-speed tool steel, which is now designated T1.

Table 2 Significant dates in the development of high-speed tool steels Date 1903 1904 1906 1910 1912 1923 1939 1940-1952 1953 1961 1970 1973 1980 1982

Development 0.70% C, 14% W, 4% Cr prototype of modern high-speed tool steels 0.30% V addition Introduction of electric furnace melting Introduction of first 18-4-1 composition (AISI T1) 3 to 5% Co addition for improved hot hardness 12% Co addition for increased cutting speeds Introduction of high-carbon high-vanadium super high-speed tool steels (M4 and T15) Increasing substitution of molybdenum for tungsten Introduction of sulfurized free-machining high-speed tool steel Introduction of high-carbon high-cobalt super hard high-speed tool steels (M40 series) Introduction of powdered metal high-speed tool steels Addition of higher silicon/nitrogen content to M-7 to increase hardness Development of cobalt-free super high-speed tool steels Introduction of aluminum-modified high-speed tool steels for cutting tools

Ack n ow le dge m e n t s The authors wish to express their thanks to the following individuals for their useful contributions to this article: Gene Bistrich, National Broach and Machine Division, Lear Siegler, Inc.; George Kiss, The Cleveland Twist Drill Company; Jim Kucynski, The Weldon Tool Company; Ron Oakes, Niagara Cutter Inc.; Robert Phillips, Pfauter Maag; Ira Redinger, Thurston Manufacturing Company; and Scott Schneier, Regal-Beloit Corporation.

M a n d T Cla ssifica t ion There are presently more than 40 individual classifications of high-speed tool steels, according to the American Iron and Steel Institute (AISI). When these are compounded by the number of domestic manufacturers, the total number of individual steels in the high-speed tool steel category exceeds 150. The AISI established a classification system for the high-speed tool steels many years ago. That system consists of a T for those steels that have tungsten as one of their primary alloying elements and an M for those steels that have molybdenum additions as one of their primary alloying elements. In addition, there is a number that follows either the M or the T. Thus, there are high-speed tool steels designated M1, M2, M41, T1, T15, and so on. That number does not have any special significance other than to distinguish one from another. For example, M1 does not mean that it is more highly alloyed than M2 or has greater hardenability or poorer wear resistance, and so on. It merely separates the types and attempts to simplify selection for the user. Table 3 lists the nominal analyses of the common M and T types.

Table 3 Composition of high-speed tool steels AISI type UNS designation Molybdenum high-speed tool steels T11301 M1 M2 Regular C T11302 ... High C M3 T11313 Class 1 T11323 Class 2 T11304 M4 T11306 M6 T11307 M7 M10 Regular C T11310 ... High C T11315 M15 T11330 M30 T11333 M33 T11334 M34 T11335 M35 T11336 M36 T11341 M41 T11342 M42 T11346 M46 T11348 M48 T11350 M50(a) T11352 M52(a) T11362 M62 Tungsten high-speed tool steels T12001 T1 T12004 T4 T12005 T5 T12006 T6 T12008 T8 T12015 T15

(a)

C

Si

Cr

V

W

Mo

Co

0.83

0.35

3.75

1.18

1.75

8.70

...

0.83 1.00

0.33 0.33

4.13 4.13

1.98 1.98

6.13 6.13

5.00 5.00

... ...

1.05 1.20 1.33 0.80 1.01

0.33 0.33 0.33 0.33 0.38

4.13 4.13 4.25 4.13 3.75

2.50 3.00 4.13 1.50 2.00

5.88 5.88 5.88 4.25 1.75

5.63 5.63 4.88 5.00 8.70

... ... ... 12.00 ...

0.89 1.00 1.50 0.80 0.89 0.89 0.80 0.85 1.10 1.10 1.26 1.50 0.80 0.90 1.30

0.33 0.33 0.33 0.33 0.33 0.33 0.33 0.33 0.33 0.40 0.53 0.33 0.40 0.40 0.28

4.13 4.13 4.00 4.00 3.75 3.75 4.00 4.13 4.13 3.88 3.95 3.88 4.13 4.00 3.88

2.00 2.00 5.00 1.25 1.18 2.10 2.00 2.00 2.00 1.15 3.15 3.00 1.00 1.93 2.00

... ... 6.50 2.00 1.70 1.75 6.00 6.00 6.63 1.50 2.05 10.00 ... 1.25 6.25

8.13 8.13 3.50 8.00 9.50 8.48 5.00 5.00 3.75 9.50 8.25 5.13 4.25 4.45 10.50

... ... 5.00 5.00 8.25 8.25 5.00 8.25 8.25 8.25 8.30 9.00 ... ... ...

0.73 0.75 0.80 0.80 0.80 1.55

0.30 0.30 0.30 0.30 0.30 0.28

4.13 4.13 4.38 4.38 4.13 4.38

1.10 1.00 2.10 1.80 2.10 4.88

18.00 18.25 18.25 19.75 14.00 12.38

... 0.70 0.88 0.70 0.70 1.00

... 5.00 8.25 12.00 5.00 5.00

Intermediate high-speed tool steel

Effe ct of Alloyin g Ele m e n t s The T series contains 12 to 20% tungsten, with chromium, vanadium, and cobalt as the other major alloying elements. The M series contains approximately 3.5 to 10% molybdenum, with chromium, vanadium, tungsten, and cobalt as the other alloying elements. All types, whether molybdenum or tungsten, contain about 4% chromium; the carbon and vanadium contents vary. As a general rule, when the vanadium content is increased, the carbon content is usually increased (Ref 1). The tungsten type T1 does not contain molybdenum or cobalt. Cobalt-base tungsten types range from T4 through T15 and contain various amounts of cobalt. Molybdenum types M1 through M10 (except M6) contain no cobalt, but most contain some tungsten. The cobalt-base, molybdenum-tungsten, premium types are generally classified in the M30 and M40 series. Super high-speed steels normally range from M40 upward; they are capable of being heat treated to high hardnesses. The M series steels generally have higher abrasion resistance than the T series steels and less distortion in heat treatment; also, they are less expensive (Ref 2). Tools made of high-speed tool steel can also be coated with titanium nitride, titanium carbide, and numerous other coatings by physical vapor deposition technique for improved performance and increased tool life.

Various elements are added to M and T series high-speed tool steels to impart certain properties to the tool steels. These elements and their effects are discussed below. Ca r bon is by far the most important of the elements and is very closely controlled. While the carbon content of any one

high-speed tool steel is usually fixed within narrow limits, variations within these limits can cause important changes in the mechanical properties and the cutting ability. As the carbon concentration is increased, the working hardness also rises; the elevated temperature hardness is higher; and the number of hard, stable, complex carbides increases. The latter contribute much to the wear resistance and other properties of the high-speed tool steels. Silicon . The influence of silicon on high-speed tool steel, up to about 1.00%, is slight. Increasing the silicon content

from 0.15 to 0.45% gives a slight increase in maximum attainable tempered hardness and has some influence on carbide morphology, although there seems to be a concurrent slight decrease in toughness. Some manufacturers produce at least one grade with silicon up to 0.65%, but this level of silicon content requires a lower maximum austenitizing temperature than does a lower silicon level in the same grade, if overheating is to be avoided. In general, however, the silicon content is kept below 0.45% on most grades. M a n ga n e se . Generally, manganese is not high in concentration in high-speed tool steels. This is because of its marked

effect in increasing brittleness and the danger of cracking upon quenching. Ph osph or u s has no effect on any of the desired properties of high-speed tool steels, but because of its well-known effect in causing cold shortness, or room-temperature brittleness, the concentration of phosphorus is kept to a minimum. Ch r om iu m is always present in high-speed tool steels in amounts ranging from 3 to 5% and is mainly responsible for the

hardenability. Generally, the addition is 4% because it appears that this concentration gives the best compromise between hardness and toughness. In addition, chromium reduces oxidation and scaling during heat treatment. Tu n gst e n . In the high-speed tool steels, tungsten is of vital importance. It is found in all T-type steels and in all but two

of the M-type steels. The complex carbide of iron, tungsten, and carbon that is found in high-speed tool steels is very hard and significantly contributes to wear resistance. Tungsten improves hot hardness, causes secondary hardening, and imparts marked resistance to tempering. When the tungsten concentration is lowered in high-speed tool steels, molybdenum is usually added to make up for its loss. M olybde n u m forms the same double carbide with iron and carbon as tungsten does but has half the atomic weight of

tungsten. As a consequence, molybdenum can be substituted for tungsten on the basis of approximately one part of molybdenum, by weight, for two parts of tungsten. The melting point of the molybdenum steels is somewhat lower than that of the tungsten grades, and thus they require a lower hardening temperature and have a narrower hardening range. The M-type high-speed tool steels are tougher than the T-type high-speed tool steels, but the hot hardness is slightly lower. Compensation for this reduced hot hardness is partially accomplished by the addition of tungsten and, to a lesser extent, vanadium to the plain molybdenum grades. This is one important reason for the popularity of the tungsten-molybdenum grades (like M2, M3, M4): they afford good hot hardness, which is so desirable in high-speed tool steels. Va n a diu m was first added to high-speed tool steels as a scavenger to remove slag impurities and to reduce nitrogen

levels in the melting operation, but it was soon found that the element materially increased the cutting efficiency of tools. The addition of vanadium promotes the formation of very hard, stable carbides, which significantly increase wear resistance and, to a lesser extent, hot hardness. An increase in vanadium, when properly balanced by carbon additions, has relatively little effect on the toughness. For this reason, vanadium-bearing grades are a very good choice when very fast cutting operations are demanded, as in finishing cuts, or when the surface of the material is hard and scaly. The special characteristics of the high-speed tool steels that are due to high vanadium additions have given rise to several specially developed steels for very severe service requiring high toughness as well as exceptional hot hardness and wear resistance. The T15, M4, and M15 grades are in this category; their vanadium content is 4.88, 4.13, and 5.00%, respectively. Coba lt . The main effect of cobalt in high-speed tool steel is to increase the hot hardness and thus to increase the cutting

efficiency when high tool temperatures are attained during the cutting operation. Cobalt raises the heat-treating temperatures because it elevates the melting point. Hardening temperatures for cobalt high-speed tool steels can be 14 to 28 °C (25 to 50 °F) higher than would be normal for similar grades without cobalt. Cobalt additions slightly increase the brittleness of high-speed tool steels.

Cobalt steels are especially effective on rough or hogging cuts, but they are not usually suited to finishing cuts that do not involve high temperatures. Cobalt types usually perform quite well when cutting materials that have discontinuous chips such as cast iron or nonferrous metals. The necessity of using deep cuts and fast speeds or of cutting hard and scaly materials justifies the use of cobalt high-speed tool steels. Su lfu r , in normal concentrations of 0.03% or less, has no effect on the properties of high-speed tool steels. However,

sulfur is added to certain high-speed tool steels to contribute free-machining qualities, as it does in low-alloy steels. The amount of free-machining high-speed tool steels is a small but significant percentage of the total consumption of highspeed tool steels. One of the major areas for free-machining high-speed tool steels is in larger-diameter tools such as hobs, broaches, and so on. Sulfur forms complex sulfides, containing chromium, vanadium, and manganese, which are distributed throughout the steel as stringer-type inclusions. The stringers interrupt the steel structure and act as notches, which aid the metalremoving action of a cutting tool when machining the high-speed steel, because the resulting chip is discontinuous, a characteristic of free-machining steels. Very high sulfur additions (up to 0.30%) are made to some powder metallurgy (P/M) high-speed tool steels for improved machinability/grindability by forming globular sulfides rather than stringers (see the article "P/M High-Speed Tool Steels" in this Volume). N it r oge n is generally present in air-melted high-speed tool steel in amounts varying from approximately 0.02 to 0.03%.

The nitrogen content of some high-speed tool steels is deliberately increased to about 0.04 to 0.05%, and this addition, when combined with higher than usual amounts of silicon, results in a slight increase of maximum attainable tempered hardness and some change of carbide morphology.

Pr ope r t ie s of H igh - Spe e d Tool St e e ls High-speed tool steels, regardless of whether they are an AISI M-type or T-type, have a rather striking similarity in their physical makeup: • • • •

They all possess a high-alloy content They usually contain sufficient carbon to permit hardening to 64 HRC They harden so deeply that almost any section encountered commercially will have a uniform hardness from center to surface They are all hardened at high temperatures, and their rate of transformation is such that small sections can be cooled in still air and be near maximum hardness

All high-speed tool steels possess excess carbide particles, which in the annealed state contain a high proportion of the alloying elements. These carbide particles contribute materially to the wear resistance of hardened high-speed tool steel. By partially dissolving during heat treatment, these carbides provide the matrix of the steel with the necessary alloy and carbon content for hardenability, hot hardness, and resistance to tempering. While all high-speed tool steels have many similar mechanical and physical characteristics, the properties may vary widely due to changes in chemical composition. Basically, the most important property of a high-speed tool steel is its cutting ability. Cutting ability depends on a combination of properties, the four most important of these being: • • • •

Hardness: Resistance to penetration by diamond-hard indenter, measured at room temperature Hot hardness: The ability to retain high hardness at elevated temperatures Wear resistance: Resistance to abrasion, often measured by grindability, metal-to-metal, or various other types of tests to indicate a relative rating Toughness: Ability to absorb (impact) energy

The relative importance of these properties varies with every application. High machining speeds require a composition with a high initial hardness and a maximum resistance to softening at high temperatures. Certain materials may abrade the cutting edge of the tool excessively; hence, the wear resistance of the tool material may well be more important than its resistance to high cutting temperatures.

Hardness is necessary for cutting harder materials and generally gives increased tool life, but it must be balanced against the toughness required for the application. The desired combination of properties in a high-speed tool steel may be obtained, first, by selection of the proper grade and, second, by the proper heat treatment, two equally important decisions. H a r dn e ss. Hardness is the most commonly stipulated requirement of a high-speed tool steel and is used as an acceptance check of a heat-treated tool. All high-speed tool steels can be hardened to room temperature hardness of 64 HRC, while the M40 series, some of the M30 series, and T15 can reach nearly 69 HRC. H ot H a r dn e ss. A related and important component of cutting ability is hot hardness. It is simply the ability to retain

hardness at elevated temperatures. This property is important because room temperature hardness values are not the same values that exist at the elevated temperature produced by friction between the tool and workpiece. Hot hardness values of some representative grades are plotted in Fig. 1. It is noteworthy that the cobalt-base types as a group exhibit higher hot hardness than non-cobalt-base types. For a comparison of the hot hardness of other metals, alloys, carbides, and ceramics to that of high-speed tool steels, see Fig. 1 in the article "Cast Cobalt Alloys" in this Volume.

Fig. 1 Com parison of t he hot hardness of cobalt - base ( M33, M36, M4, and T15) t ype versus noncobalt -base ( M1, M2, M4, M7, and T1) t ype high- speed t ool st eels

W e a r Re sist a n ce . The third component of cutting ability is resistance to wear. Wear resistance of high-speed tool

steels is affected by the matrix hardness and composition, by precipitated M2C and MC carbides responsible for secondary hardness, by the volume of excess alloy carbides, and by the nature of these excess carbides. In practically any given high-speed tool steels, wear resistance strongly depends on hardness of the steel, and higher hardness, however achieved, is an aim when highly abrasive cutting conditions are encountered (Fig. 2).

Fig. 2 Effect of hardness on wear rat e for high- speed t ool st eels, each having been double t em pered t o t he indicat ed hardness

For the ultimate in wear resistance, carbon content has been increased simultaneously with vanadium content, to permit the introduction of a greater quantity of total carbide and a greater percentage of extremely hard vanadium carbide in high-speed tool steel. Examples of this effect are given when discussing the effect of vanadium on the properties of highspeed tool steels. Steels T15, M3 (class 2), M4, and M15 are in this category, and all have extremely high wear resistance. Laboratory tests for wear resistance are diverse, making comparisons between different procedures difficult. Therefore, production tests on actual tools are used to a great extent. However, laboratory tests can produce valuable data on the relative wear resistance for these steels (Fig. 3). The data given in Fig. 3 were generated by measuring the volume loss of a high-speed tool steel sample against the volume loss of a known vitrified abrasive wheel after a predetermined grinding procedure.

Fig. 3 Com parison of relat ive abrasion resist ance at t ypical working hardness for high- speed t ool st eels

Tou gh n e ss. The fourth component of cutting ability mentioned above is toughness, which is defined as a combination

of two factors: • •

The ability to deform before breaking (ductility) The ability to resist permanent deformation (elastic strength)

If either of these factors is to be used to describe toughness (a practice not to be condoned), the second appears more practical for high-steel tool steel because rarely are large degrees of flow or deformation permissible with fine-edge tools. The first, however, cannot be ignored, as frequently the stress applied to a tool (through overloads, shock, notches, and sharp corners) exceeds the elastic strength. Toughness tests on high-steel tool steel are usually conducted at room temperature. Tool failures that occur from spalling of the tool edge generally occur during the initial contact of the tool with the work, and once the tool becomes heated, its performance in this respect is much superior. Therefore, room temperature tests are perhaps of greater value when toughness is considered than when hardness is in question. Laboratory tests for the measurement of toughness of hardened high-speed tool steel include bend, unnotched or C-notch impact, static torsion, and torsion impact tests. Figures 4 and 5 compare relative unnotched impact values for representative high-speed tool steels. Modest improvements in toughness (within a grade) can be made by lowering the tempered hardness. Lower austenitizing temperatures enhance the toughness for a given hardness and grade.

Fig. 4 Plot of im pact t oughness versus hardness for high- speed t ool st eels

Fig. 5 Relat ive t oughness of high- speed t ool st eels at t ypical working hardness

H e a t Tr e a t m e n t of H igh - Spe e d Tool St e e ls Proper heat treatment is as critical to the success of the cutting tool as material selection itself. Often the highest-quality steel made into the most precise tools does not perform because of improper heat treatment. The object of the heat treating or hardening operation is to transform a fully annealed high-speed tool steel consisting mainly of ferrite (iron) and alloy carbides into a hardened and tempered martensitic structure having carbides that provide the cutting tool properties (see Fig. 6 and 7).

Fig. 6 Microst ruct ure of fully annealed high- speed t ool st eel consist ing of ferrit e ( iron) and alloy carbides. 1000×

Fig. 7 Microst ruct ure of hardened, t em pered high- speed t ool st eel having m art ensit ic st ruct ure wit h carbides. 1000×

The heat treatment process can be divided into four primary areas, preheating, austenitizing, quenching, and tempering. Figure 8 outlines graphically these four heat treatment steps.

Fig. 8 Tim e versus t em perat ure plot illust rat ing sequences required t o properly heat t reat high-speed t ool st eels

Pr e h e a t in g. From a metallurgical standpoint, preheating plays no part in the hardening reaction; however, it performs three important functions. The first of these is to reduce thermal shock, which always results when a cold tool is placed into a warm or hot furnace. Minimizing thermal shock reduces the danger of excessive distortion or cracking. It also relieves some of the stresses developed during machining and/or forming, although conventional stress relieving is more effective.

The second major benefit of preheating is to increase equipment productivity by decreasing the amount of time required in the high-heat furnace. Thirdly, if the high-heat furnace is not neutral to the surface of the tool or part, preheating will reduce the amount of carburization and decarburization that would result if no preheat were employed. In commercial salt bath hardening, a two-step preheat is typically used for high-speed tool steels. The first preheat is carried out between 650 and 760 ° (1200 and 1400 °F); the second preheat cycle is carried out between 815 and 900 °C (1500 and 1650 °F). In atmosphere or vacuum heat treating, the furnace is usually heated slowly to a single preheat of 790 to 845 °C (1450 to 1550 °F). Preheat duration is of little importance as long as the part is heated throughout its cross section. Au st e n it izin g ( h a r de n in g) is the second step of the heat treatment operation. Austenitizing is a time/temperature

dependent reaction. High-speed tool steels depend upon the dissolving of various complex alloy carbides during austenitizing to develop their properties. These alloy carbides do not dissolve to any appreciable extent unless the steel is heated to a temperature within 28 to 56 °C (50 to 100 °F) of their melting point. This temperature is dependent upon the particular high-speed tool steel being treated and is in the range of 1150 to 1290 °C (2100 to 2350 °F). The generally recommended hold time for high-speed tool steel is approximately 2 to 6 min, depending upon high-speed tool steel type, tool configuration, and cross-sectional size. Lowering the hardening temperature (underhardening) generally improves the impact toughness while lowering the hot hardness. Raising the hardening temperature increases heat-treated room-temperature hardness and also increases the hot hardness. Qu e n ch in g. The quenching or cooling of the workpiece from the austenitizing temperature is designed to transform the austenite that forms at the high temperature to a hard martensitic structure. The rate of cooling, which must be controlled, is dictated by the analysis of the particular steel. Sometimes high-speed steels are two-step quenched, initially in a molten

salt bath maintained at approximately 540 to 595 °C (1000 to 1100 °F) or an oil quench, followed by air cooling to near ambient temperature. The least drastic form of quenching is cooling in air, although only in the smaller and/or thinner cross sections would high-speed tool steels air quench rapidly enough to transform the majority of the structure into the desirable martensitic condition. The austenite-martensite transformation is exemplified in Fig. 9 illustrating a timetemperature-transformation curve.

Fig. 9 Tim e- t em perat ure- t ransform at ion diagram for M2 high- speed t ool st eel t hat was annealed prior t o quenching. Aust enit izing t em perat ure was 1230 °C ( 2 250 °F) , and crit ical t em perat ure was 830 °C ( 1530 °F) .

Te m pe r in g. Following austenitizing and quenching, the steel is in a highly stressed state and therefore is very

susceptible to cracking. Tempering (or drawing) increases the toughness of the steel and also provides secondary hardness, as illustrated by the peak on the right of the tempering curve in Fig. 10. Tempering involves reheating the steel to an intermediate temperature range (always below the critical transformation temperature), soaking, and air cooling.

Fig. 1 0 Tem pering curve for M2 high- speed t ool st eel. To opt im ize t he t ransform at ion of ret ained aust enit e t o fresh m art ensit e during t he t em pering sequence, t he high ( right ) side of t he secondary hardness peak curve is preferred, and t he low ( left ) side should be avoided.

Tempering serves to stress relieve and to transform retained austenite from the quenching step to fresh martensite. Some precipitation of complex carbide also occurs, further enhancing secondary hardness. It is this process of transforming retained austenite and tempering of newly formed martensite that dictates a multiple tempering procedure. High-speed tool steels require 2 to 4 tempers at a soak time of 2 to 4 h each. As with austenitizing temperatures and quenching rates, the number of tempers is dictated by the specific grade. High-speed tool steels should be multiple tempered at 540 °C (1000 °F) minimum for most grades. It is essential to favor the right (high) side of the secondary hardness peak of the tempering curve in order to optimize the above-described transformations. Subzero treatments are sometimes used in conjunction with tempering in order to continue the transformation of austenite to martensite. Numerous tests have been run on the effect of cold treatments, and the findings generally prove that cold treatments used after quenching and first temper enhance the transformation to martensite, in much the same way that multiple tempering causes transformation. Cold treatments administered to high-speed tool steels immediately after quenching can result in cracking or distortion because the accompanying size change is not accommodated by the newly formed, brittle martensite. It is generally accepted that subzero treatments are not necessary if the steel is properly hardened and tempered.

Su r fa ce Tr e a t m e n t s Tools made of high-speed tool steel are available with either a bright, black oxide or nitride finish or they can be coated with titanium nitride and other coatings using a vapor disposition process that greatly increases tool life. Br igh t Fin ish . Most tools are finished with a ground or mechanically polished surface that would be categorized as a bright finish. Bright finished tools are often preferred to tools with an oxide finish for machining nonferrous work material. The smooth or bright finish tends to resist galling, a type of welding or buildup associated with many nonferrous alloys. However, work materials of ferrous alloys tend to adhere to similar, iron-base tools having a bright finish. This buildup on the cutting edges leads to increased frictional heat, poor surface finish, and increased load at the cutting edge. Bla ck Ox ide Fin ish . This characteristic black finish is typically applied to drills and other cutting tools by oxidizing in a steam atmosphere at approximately 540 °C (1000 °F). The black oxide surface has little or no effect on hardness, but serves as a partial barrier to galling of similar ferrous metals. The surface texture also permits retention of lubricant. N it r ide Fin ish . Nitriding is a method of introducing nitrogen to the surface of high-speed tool steels at a typical temperature of 480 to 595 °C (900 to 1100 °F) and is accomplished either by the dissociation of ammonia gas, exposure to sodium cyanide salt mixtures, or bombardment with nitrogen ions in order to liberate nascent nitrogen, which combines with the steel to form a hard iron nitride. Nitriding improves wear resistance of high-speed steel, at the expense of notch toughness.

Coa t e d H igh - Spe e d Tool St e e ls. The addition of wear-resistant coatings to high-speed tool steel cutting tools lagged

behind the coating of carbide inserts by approximately 10 years until the development of the low-temperature physical vapor deposition (PVD) process, an innovation, which is much more suitable for coating high-speed tool steels than is the older chemical vapor deposition (CVD) process, and which also eliminates the need for subsequent heat treatment (Ref 3). As described in Ref 4, titanium nitride is the most commonly used and most durable coating available, although substitutes such as other nitrides (hafnium nitride and zirconium nitride) and carbides (titanium carbide, zirconium carbide, and hafnium carbide) are being developed. These other coatings are expected to equal or surpass the desired properties of titanium nitrides in future years. The hard thin (2 to 5 m, or 80 to 200 in. thick) deposit of high-density titanium nitride, which has 2500 HV hardness and imparts a characteristic gold color to high-speed tool steels, provides excellent wear resistance, minimizes heat buildup, and prevents welding of the workpiece material, while improving the surface finish of high-speed tool steels (Ref 5). The initial use, in 1980, of titanium nitride coatings was to coat gear cutting tools. Subsequent applications include the coating of both single-point and multipoint tools such as lathe tools, drills, reamers, taps, milling cutters, end mills, and broaches (Ref 3). Today, titanium nitride coated hobs and shapers dominate high-production applications in the automotive industry to such an extent that 80% of such tools use this coating. As described in Ref 6, significant cost savings are possible because the titanium nitride coating improves tool life up to 400% and increases feed and speed rates by 30%. This is primarily attributable to the increased lubricity of the coating because its coefficient of friction is one-third that of the bare metal surface of a tool. Examples of increased tool life obtained when using coated versus uncoated single-point and multipoint cutting tools are listed in Table 4. The increased production obtained with a coated tool justifies the application of the coating despite the resulting 20 to 300% increase in the base price of the tool (Ref 3).

Table 4 Increased tool life attained with coated cutting tools Coating

Workpiece material

High-speed tool steel, AISI type M7 M7 M3 M2 M3

Workpieces machined before resharpening Uncoated Coated

TiN TiN TiN TiN TiN

1022 steel, 35 HRC 6061-T6 aluminum alloy 7075T aluminum alloy 8620 steel Type 303 stainless steel

325 166 9 40 100,000

1,200 1,500 53 80 300,000

M2 M2

TiN TiN

48% nickel alloy Type 410 stainless steel

3,400 31,000

Pipe tap Tap

M2 M2

TiN TiN

Gray iron 1050 steel, 30-33 HRC

200 10,00012,000 3,000 60-70

Form tool Form tool Cutoff tool

T15 T15 M2

1045 steel Type 303 stainless steel Low-carbon steel

5,000 1,840 150

Drill Drill

M7 M7

TiC TiN TiCTiN TiN TiN

9,000 750800 23,000 5,890 1,000

Low-carbon steel Titanium alloy 662 layered with D6AC tool steel, 48-50 HRC

1,000 9

4,000 86

Cutting tool Type

End mill End mill End mill Gear hob Broach insert Broach Broach

Coated tools can meet close-tolerance requirements and significantly improve the machining of carbon and alloy steels, stainless steels (especially the 300 series, where galling can be a problem), and aluminum alloys (especially aircraft grades). Coated high-speed tool steels are less of a factor in the machining of certain titanium alloys and some high-nickel alloys because of chemical reactions between the coatings and the workpiece materials (Ref 3).

H igh - Spe e d Tool St e e l Applica t ion s High-speed tool steels are used for most of the common types of cutting tools including single-point lathe tools, drills, reamers, taps, milling cutters, end mills, hobs, saws, and broaches. Sin gle - Poin t Cu t t in g Tools The simplest cutting tools are single-point cutting tools, which are often referred to as tool bits, lathe tools, cutoff tools, or inserts. They have only one cutting surface or edge in contact with the work material at any given time. Such tools are used for turning, threading, boring, planing, or shaping, and most are mounted in a toolholder that is made of some type of tough alloy steel. The performance of such tools is dependent on the tool material as well as factors such as the material being cut, the speeds and feeds, the cutting fluid, and fixturing. Following is a discussion of material characteristics and recommendations for the most popular lathe tools. M1, M2, and T1 are suitable for all-purpose tool bits. They offer excellent strength and toughness and are suitable for both roughing and finishing and can be used for machining wrought steel, cast steel, cast iron, brass, bronze, copper, aluminum, and so on (see the Section "Machining of Specific Metals and Alloys" in this Volume). These are good economical grades for general shop purposes. M3 class 2 and M4 high-speed tool steels have high-carbon and high-vanadium contents. The wear resistance is several times that of standard high-speed steels. These bits are hard and tough, withstanding intermittent cuts even under heavy feeds. They are useful for general applications and especially recommended for cast steels, cast iron, plastics, brass, and heat-treated steels. On tool bit applications where failure occurs from rapid wearing of the cutting edge, M3 class 2 and M4 will be found to surpass the performance of regular tool bits. T4, T5, and T8 combine wear resistance resulting from the higher carbon and vanadium contents together with a higher hot hardness, resulting from a cobalt content. Because of the good resistance to abrasion and high hot hardness, these steels should be applied to the cutting of hard, scaly, or gritty materials. They are well adapted for making hogging cuts, for the cutting of hard materials, and for the cutting of materials that throw a discontinuous chip, such as cast iron and nonferrous materials. The high degree of hot hardness permits T4, T5, and T8 to cut at greater speeds and feeds than most high-speed tool steels. They are much more widely used for single-point cutting tools, such as lathe, shaper, and planer tools, than for multiple-edge tools. Superhard tool bits made from the M40 series offer the highest hardness available for high-speed tool steels. The M40 steels are economical cobalt alloys that can be treated to reach a hardness as high as 69 HRC. Tool bits made from them are easy to grind and offer top efficiency on the difficult-to-machine space-age materials (titanium and nickel-base alloys, for example) and heat-treated high-strength steels requiring high hot hardness. T15 tool bits are made from a steel capable of being treated to a high hardness, with outstanding hot hardness and wear resistance. The exceptional wear resistance of T15 has made it the most popular high-speed tool steel for lathe tools. It has higher hardness than most other steels, and wear resistance surpassing that of all other conventional high-speed tool steels as well as certain cast cutting tool materials. It has ample toughness for most types of cutting tool applications, and will withstand intermittent cuts. These bits are especially adapted for machining materials of high-tensile strength such as heat-treated steels and for resisting abrasion encountered with hard cast iron, cast steel, brass, aluminum, and plastics. Tool bits of T15 can cut ordinary materials at speeds 15 to 100% higher than average. Often an engineer will specify a grade that is not necessary for a given application. For example, selecting M42 for a general application that could be satisfied with M2 does not always prove to be beneficial. The logic is that the tool can be run faster and therefore generate a higher production rate. What happens many times is that the M42 will chip because of its lower toughness level, whereas the M2 will not. M u lt ipoin t Cu t t in g Tools Applications of high-speed tool steels for other cutting tool applications such as drills, end mills, reamers, taps, threading dies, milling cutters, circular saws, broaches, and hobs are based on the same parameters of hot hardness, wear resistance, toughness, and economics of manufacture. Some of the cutting tools that require extensive grinding have been produced of P/M high-speed tool steels (see the article "P/M High-Speed Tool Steels" in this Volume).

Ge n e r a l- pu r pose dr ills, other than those made from low-alloy steels for low production on wood or soft materials, are

made from high-speed tool steels, typically M1, M2, M7, and M10. For lower cost hardware quality drills, intermediate high-speed tool steels M50 and M52 are sometimes used although they cannot be expected to perform as well as standard high-speed tool steels in production work. For high hot hardness required in the drilling of the more difficult-to-machine alloys such as nickel-base or titanium product, M42, M33, or T15 are used. High-speed tool steel twist drills are not currently being coated as extensively as gear cutting tools because many drills are not used for production applications. Also, the cost of coating (predominantly with titanium nitride) is prohibitive because it represents a higher percentage of the total tool cost. Drills coated with titanium nitride reduce cutting forces (thrust and torque) and improve the surface finishes to the point that they eliminate the need for prior core drilling and/or subsequent reaming. Coated drills have been found especially suitable for cutting highly abrasive materials, hard nonferrous alloys, and difficult-to-machine materials such as heatresistant alloys. These tools are not recommended for drilling titanium alloys because of possible chemical bonding of the coating to the workplace material. When drilling gummy materials (1018 and 1020 steels, for example) with coated tools, it may be necessary to provide for chipbreaking capabilities in the tool design (Ref 3). En d m ills are produced in a variety of sizes and designs, usually with two, four, or six cutting edges on the periphery. This shank-type milling cutter is typically made from general-purpose high-speed tool steels M1, M2, M7, and M10. For workpieces made from hardened materials (over 300 HB), a grade such as T15, M42, or M33 is more effective. Increased cutting speeds can be used with these cobalt-containing high-speed tool steels because of their improved hot hardness.

One manufacturer realized a fourfold increase in the tool life of end mill wear lands when he switched to a titaniumnitride coated tool (Fig. 11). Titanium nitride coated end mills also outperform uncoated solid carbide tools. When machining valves made from type 304 stainless steel, a switch from solid carbide end mills to titanium nitride coated end mills resulted in a fivefold increase in tool life, that is, 150 parts compared to 30 finished with the carbide tools (Ref 3). Furthermore, the cost of the coated high-speed steel end mills was only one-sixth that of the carbide tools. Both types of 19 mm ( in.) fluted end mills were used to machine a 1.6 mm ( 51 mm/min (2 in./min).

in.) deep slot at a speed of 300 rev/min and a feed of

Fig. 1 1 Wear lands developed wit h uncoat ed and t it anium nit ride coat ed end m ills show a 4: 1 increase in t ool life wit h coat ed t ools. The crosshat ched area at left ( ext ending from 0 t o 20 part s) indicat es t he num ber of pieces produced by uncoat ed end m ill aft er 0. 25 m m ( 0.010 in.) wear land on t he t ool; t he crosshat ched area at right represent s quant it y produced by t it anium nit ride coat ed end m ill aft er 0.25 m m ( 0.010 in.) wear land on t ool. Source: Ref 3

Re a m e r s are designed to remove only small amounts of metal and therefore require very little flute depth for the

removal of chips. For this reason, reamers are designed as rigid tools, requiring less toughness from the high-speed tool steel than a deeply fluted drill. The general-purpose grades M1, M2, M7, M10, and T1 are typically used at maximum hardness levels. For applications requiring greater wear resistance, grades such as M3, M4, and T15 are appropriate. M illin g Cu t t e r s. The size, style, configuration, complexity, and capacity of milling cutters is almost limitless. There are staggered-tooth and straight-tooth, form-relieved and formed milling cutters with sizes that range from 51 to 305 mm (2

to 12 in.) and are used to machine slots, grooves, racks, sprockets, gears, splines, and so on. They cut a wide variety of materials, including plastics, aluminum, steel, cast iron, superalloys, titanium, and graphite structures. The generalpurpose high-speed tool steel used for more than 70% of milling cutter applications is M2, usually the free-machining type. It has a good balance of wear resistance, hot hardness, toughness, and strength and works well on carbon, alloy, and stainless steels, aluminum, cast iron, and some plastics (generally any material that is under 30 HRC in hardness). When higher hardness materials or more wear-resistant materials need to be milled, M3 or M4 are selected. The higher carbon and vanadium content in those materials improves wear resistance nd allows for the machining of materials greater than 35 HRC in hardness. For workpiece hardness levels above that and as high as 50 HRC, either M42 with its high hardness and high hot hardness properties or T15 with its high wear resistance and high hardness characteristics are desirable. The powder metallurgy grades in M4 and T15 are increasing in popularity for milling cutters because of their ease of grinding and regrinding. H obs are a type of milling cutter that operates by cutting a repeated form about a center, such as gear teeth. The hob cuts by meshing and rotating about the workpiece, forming a helical pattern. This type of metal cutting creates less force at the cutting edge (less chip load on the teeth) than do ordinary milling cutters. Accordingly, less toughness and edge strength is required of hob materials; wear is more commonly a mode of failure. Most hobs are made from a high-carbon version of M2, although normal carbon levels are also used. M2 with a sulfur addition or P/M product for improved machinability and surface finish is often used for hobs. Sa w s are quite similar to milling cutters in style and application, but they are usually thinner and tend to be smaller in

diameter. Sizes range from 0.076 mm (0.003 in.) thick by 13 mm ( in.) outside diameter to more than 6.4 mm ( in.) thick by 203 mm (8 in.) outside diameter. Used for cutting, slitting, and slotting, saws are available with straight-tooth, staggered-tooth, and side-tooth configurations and are made from alloys similar to those used for milling cutters. Again, M2 high-speed tool steel is the general-purpose saw material, but, because of the typical thinness of these products, toughness is optimized with lower hardness. There are relatively few saws that are made from M3 or M4 high-speed tool steel because generally T15 and M42 are the two alternative materials to the standard M2 steel. M42 is often used to machine stainless steels, aluminum, and brass because it increases saw production life and can be run at considerably higher speeds. T15 is used for very specialized applications. Saws made of high-speed tool steel are used to cut, slit, and slot everything from steel, aluminum, brass, pipe, and titanium to gold jewelry, fish, frozen foods, plastics, rubber, and paper. Br oa ch e s. M2 high-speed tool steel is the most frequently used material for broaches. This includes the large or circular

broaches that are made in large quantities as well as the smaller keyway and shape broaches. Sometimes the highercarbon material is used, but generally free-machining M2 is used because it results in a better surface finish. P/M products are very popular for broaches in both M2 as well as M3 class 2 and M4 when they are used to improve wear resistance. M4 is probably the second most widely used material for this application. M42 and T15 are often used for difficult-tomachine materials such as the nickel-base alloys and other aerospace-type alloys. A high-nickel (48%) alloy magnet manufacturer using a 3.2 × 13 × 305 mm ( × × 12 in.) flat broach made of M2 increased tool life from 200 pieces to 3400 pieces when a titanium nitride coating was added, and also obtained a smoother surface finish. Replacing the flat broach with an uncoated 11.99 mm (0.472 in.) diam, by 660 mm (26 in.) long round broach increased the production to about 7000 pieces, and coating the round broach with titanium nitride further increased the magnet production to about 19,000 pieces (Ref 3). Thus, going from an uncoated flat broach to a coated round broach increased production by a factor of 95.

Fa ct or s I n Se le ct in g H igh - Spe e d Tool St e e ls No one composition of high-speed tool steel can meet all cutting tool requirements. The general-purpose molybdenum steels such as M1, M2, and M7 and tungsten steel T1 are more commonly used than other high-speed tool steels. They have the highest toughness and good cutting ability, but they possess the lowest hot hardness and wear resistance of all the high-speed tool steels. The addition of vanadium offers the advantage of greater wear resistance and hot hardness, and steels with intermediate vanadium contents are suited for fine and roughing cuts on both hard and soft materials. The 5% V steel (T15) is especially suited for cutting hard metals and alloys or high-strength steels, and is particularly suitable for the machining of aluminum, stainless steels, austenitic alloys, and refractory metals. Wrought high-vanadium high-speed tool steels are more difficult to grind than their P/M product counterparts. The addition of cobalt in various amounts allows still higher hot hardness, the degree of hot hardness being proportional to the cobalt content. Although cobalt steels

are more brittle than the noncobalt types, they give better performance on hard, scaly materials that are machined with deep cuts at high speeds. High-speed tool steels have continued to be of importance in industrial commerce for 70 to 80 years despite the inroads made by competitive cutting tool materials such as cast cobalt alloys, cemented carbides, ceramics, and cermets. The superior toughness of high-speed tool steels guarantees its niche in the cutting tool materials marketplace.

Re fe r e n ce s 1. Machining, Vol 1, Tool and Manufacturing Engineers Handbook, Society of Manufacturing Engineers, 1983, p 3-6 2. S. Kalpakjian, Manufacturing Processes for Engineering Materials, Addison-Wesley, 1984, p 524 3. C. Wick, HSS Cutting Tools Gain a Productivity Edge, Manufacturing Engineering, May 1987, p 38 4. W.D. Sproul, Turning Tests of High Rate Reactively Sputter-Coated T-15 HSS Inserts, Surf. Coat. Tech., Vol 33, 1987, p 133 5. TiN Coatings Continue to Revolutionize the Metalworking Industry, Machining Source Book, ASM INTERNATIONAL, 1988, p 98 6. How Cutting Tools Can Help You Make a Quick Buck, Machining Source Book, ASM INTERNATIONAL, 1988, p 20

P/ M H igh - Spe e d Tool St e e ls Revised by Kennet h E. Pinnow and William St asko, Crucible Mat erials Corporat ion

I n t r odu ct ion POWDER METALLURGY (P/M) high-speed tool steels are used extensively for drills, taps, end mills, reamers, broaches, and other cutting tools because of their excellent manufacturing and performance characteristics. For most applications, they offer distinct advantages over conventional high-speed tool steels which, as a result of pronounced ingot segregation, often contain a coarse, nonuniform microstructure, accompanied by poor toughness and grind-ability, and also present problems of size control and hardness uniformity in heat treatment. Rapid solidification of the atomized powders used in the production of P/M high-speed tool steels eliminates such segregation and produces a very fine microstructure with a uniform distribution of carbides and nonmetallic inclusions. As a result, a number of important end properties of high-speed tool steels have been improved by powder processing, notably toughness, dimensional control during heat treatment, grindability, and cutting performance under difficult conditions when good toughness is essential (Ref 1). Further, powder processing allows the production of high-speed tool steels with much greater alloy contents than are practical or possible by conventional ingot methods. Two examples of such highly alloyed high-speed tool steels are CPM Rex 76 and ASP 60. Since the early 1970s, several P/M methods for producing high-speed tool steels have been developed, including controlled spray deposition (CSD), the Osprey process, rapid omnidirectional compaction, consolidation at atmospheric pressure (CAP process), the STAMP process, and injection molding. These processes are discussed in Powder Metal Technologies and Applications, Volume 7 of the ASM Handbook. The present discussion describes procedures for producing tool steel powder by inert-gas atomization, followed by compaction by hot isostatic pressing (HIP). These processes include the Anti-Segregation Process (ASP), developed in Sweden by Stora Kopparberg and ASEA, and the Crucible Particle Metallurgy process, developed in the United States by the Crucible Materials Corporation. The FULDENS process, which uses water-atomized powders compacted by vacuum sintering, is also discussed. It was developed in the United States by Consolidated Metallurgical Industries, Inc.

For additional data concerning the classification, composition, heat treatment, and properties of conventionally processed and P/M processed high-speed tool steel materials, see the articles "High-Speed Tool Steels" in this Volume; "Wrought Tool Steels" and "Powder Metallurgy Tool Steels" in Properties and Selection: Irons, Steels, and High-Performance Alloys, Volume 1; and "Particle Metallurgy Tool Steels" in Powder Metal Technologies and Applications, Volume 7 of the ASM Handbook.

Th e An t i- Se gr e ga t ion Pr oce ss The Anti-Segregation Process, or ASEA-STORA process, is used to produce high-speed tool steels by powder metallurgy. In this process, an alloy steel melt is atomized in an inert gas to form spherical powder particles. These are poured into cylindrical sheet steel capsules (cans), which are vibrated to pack the particles as tightly as possible. A cover is then welded onto the capsule and the air inside is evacuated. The capsule and its contents are cold isostatically pressed at 400 MPa (58 ksi). The capsule is hot isostatically pressed at 100 MPa (14.5 ksi) at 1150 °C (2100 °F) to full density. After compaction, the steel is conventionally hot worked by forging and rolling to the desired dimensions. Figure 1 compares the processing of conventional (wrought) high-speed tool steels with that of ASP high-speed tool steels.

Fig. 1 Com parison of convent ionally ( wrought ) processed high- speed t ool and P/ M processed ASP high- speed t ool st eel

This processing results in a fine-grain material with a uniform distribution of small carbides. The homogeneous material, free from segregation, has a uniform structure, regardless of bar size and alloy content. Figure 2 compares the microstructures of conventional high-speed tool steel and P/M processed ASP high-speed tool steel.

Fig. 2 Com parison of m icrost ruct ures of convent ional high- speed t ool st eel and P/ M high-speed t ool st eel. ( a) Convent ional high- speed t ool st eel m icrost ruct ure showing carbide segregat ion. ( b) Microst ruct ure of P/ M processed ASP st eel showing sm all, uniform ly dist ribut ed carbide part icles. Court esy of Speedst eel, I nc.

Pr ope r t ie s of ASP St e e ls ( Re f 2 ) The primary benefits of ASP techniques include improved toughness and ultimate strength due to uniform carbide distribution and the absence of metallurgical defects. Improved grindability due to the small carbide size and improved dimensional stability in heat treatment caused by the absence of segregation are also benefits. Additionally, wear resistance can be improved by increasing alloy content, without sacrificing toughness or grindability. Currently, ASP high-speed tool steel is available in three grades: ASP 23, 30, and 60 (ASP 60 can be made only by the powder metallurgy process). The compositions and recommended applications of these grades are given in Table 1. Additional information on applications of ASP steels can be found in the section "Applications of P/M High-Speed Tool Steels" in this article.

Table 1 ASP steel grades, compositions, hardnesses, and applications ASP grade

Composition, % C Cr Mo

W

V

Co

23

1.28

4.2

5.0

6.4

3.1

...

Typical hardness, HRC 65-67

30

1.28

4.2

5.0

6.4

3.1

8.5

66-68

60

2.30

4.0

7.0

6.5

6.5

10.5

67-69

Recommended applications For ordinary applications of most cutting tools when hot hardness is not of primary concern. Also for tools used in cold-working applications For cutting tool applications when hot hardness is important. Suitable for cutting most stainless steels and superalloys, and for cutting at higher speeds. Also for cold work-tools when wear resistance is critical For cutting tools when wear resistance and hot hardness are critical. Particularly suitable for extratough applications (cutting titanium, highhardness materials, and iron forgings). Also for cold-work tools requiring highest wear resistance

W e a r r e sist a n ce is generally a function of the hardness of the tool and the specific alloy content or type of carbide.

The higher hardness that is possible with P/M high-speed tool steels, plus the higher carbon and vanadium contents, promote better wear resistance. Tou gh n e ss of a tool or high-speed tool steel is usually defined as a combination of strength and ductility or as resistance

to breaking or chipping. A tool that deforms from lack of strength is useless, and one that lacks adequate ductility will fail prematurely. The importance of toughness of high-speed tool steel is illustrated in Fig. 3. A cutting edge may suffer from repeated microchipping. As shown in Fig. 3, the ASP 23 cutting edge shows minimal wear. The M2 cutting edge, however, shows microchipping under the same service conditions. Microchipping blunts the cutting edge, increases stress, and accelerates other wear factors.

Fig. 3 Com parison of cut t ing edge wear of a convent ional high- speed t ool st eel and a P/ M high-st eel t ool st eel. ( a) Cut t ing edge of t ool m ade of convent ional AI SI M2 m at erial, showing severe m icrochipping. ( b) Cut t ing edge of t ool m ade of P/ M- processed ASP 23 m at erial, showing no m icrochipping under t he sam e service condit ions. Court esy of Speedst eel, I nc.

One method of measuring toughness of high-speed tool steel after heat treatment is bend testing. Bend yield strength, ultimate bend strength, and deflection are measured on 5 mm (0.2 in.) diam test bars on which a load is exerted. The results of these laboratory tests correlate well with shop experience. As shown in Fig. 4, toughness and hardness can be controlled by varying the hardening temperature. A low hardening temperature produces good toughness. Raising the hardening temperature increases hardness, but lowers toughness.

Fig. 4 Bend t est result s t o det erm ine t oughness of PM/ processed ASP high- speed t ool st eels. A, ult im at e bend st rengt h; B, bend yield st rengt h; C, hardness ( HRC) . ( a) Bend st rengt h of a t est bar of ASP 23 st eel aft er hardening and t em pering at 560 °C ( 1040 °F) ( t hree t im es for 1 h) . ( b) Bend st rengt h of a t est bar of ASP 30 st eel aft er hardening and t em pering at 560 °C ( 1040 °F) ( t hree t im es for 1 h) . ( c) Bend st rengt h of a t est bar of ASP 60 st eel aft er hardening and t em pering at 560 °C ( 1040 °F) ( t hree t im es for 1 h) . Ult im at e bend st rengt h m ay vary ± 10% ; bend yield st rengt h m ay vary ± 5% ; hardness values m ay vary ± 1% . Court esy of Speedst eel, I nc.

Gr in da bilit y of ASP steel is superior to that of conventional high-speed tool steel of the same chemical composition.

This is due to the small carbide size and the uniform distribution of carbides, regardless of bar size. Figure 5 compares the grindability of several tool steels. These data are based on laboratory measurements, but results are confirmed by shop experience.

Fig. 5 Grindabilit y of P/ M high- speed t ool st eel and convent ional high- speed t ool st eel m at erials. Grindabilit y index is t he rat io of t he volum e of m at erial rem oved t o t he volum e of grinding wheel wear.

H e a t Tr e a t m e n t of ASP H igh - Spe e d Tool St e e ls Only with proper heat treatment can optimum mechanical properties of tools and dies be obtained. Improper heat treatment may result in a tool with greatly reduced productivity or even an unusable tool. Heat treatment consists of four stages: preheating, austenitizing, quenching, and tempering. The heat treatment procedure for ASP high-speed tool steels is essentially the same as for wrought high-speed tool steels. Optimum heat-treating temperatures may vary, however, even if chemical compositions are identical. The following procedures should be used to heat treat ASP high-speed tool steels as well as all P/M high-speed tool steels: • • •

•

Annealing: Heat to 850 to 900 °C (1560 to 1650 °F). Slow cool 10 °C/h (18 °F/h) to 700 °C (1290 °F). Hardness values are 260 HB maximum for ASP 23, 300 HB for ASP 30, and 340 HB for ASP 60 Stress relieving: Hold for approximately 2 h at 600 to 700 °C (1110 to 1290 °F). Slow cool to 500 °C (930 °F) in furnace Hardening: Preheat in two steps, first at 450 to 500 °C (840 to 930 °F) and then at 850 to 900 °C (1560 to 1650 °F). Austenitize at 1050 to 1180 °C (1920 to 2155 °F) and quench, preferably in a neutral salt bath. Cool to hand warmth. See Table 2 for recommended temperatures Tempering: Raise temperature to 560 °C (1040 °F) or higher three times for at least 1 h at full temperature. Cool to room temperature between tempers

Hardness of ASP high-speed tool steel after hardening and tempering is shown in Fig. 6.

Table 2 Austenitizing temperatures of ASP 23 steel Hardness(a), HRC 58 60 62 64 66

(a)

Temperature °C °F 1000 1830 1050 1925 1100 2010 1140 2085 1180 2155

Salt bath(b) min/mm 0.59 0.47 0.39 0.31 0.24

min/in. 15 12 10 8 6

Other furnace(c), min 30 25 20 15 10

After triple temper at 560 °C (1040 °F); hardness values may vary by ±1%.

(b) (c)

Total immersion time after preheating. Holding time in minutes after tool has reached full temperature

Fig. 6 Hardness of ASP st eels aft er hardening and t em pering a 25 m m ( 1 in.) diam specim en t hree t im es for 1 h. ( a) ASP 23. ( b) ASP 30. ( c) ASP 60, cooled in st ep bat h. Hardening t em perat ure for curves is: A, 1180 °C ( 2155 °F) ; B, 1150 °C ( 2100 °F) ; C, 1100 °C ( 2010 ° F) ; D, 1050 °C ( 1920 °F) .

D im e n sion a l St a bilit y in H e a t Tr e a t m e n t . Three types of distortion are experienced metallurgically during heat

treatment: • • •

Normal volume change due to phase transformations in the steel Variations in volume change in different parts of the tool due to the segregation in the steel Distortion due to residual stress caused by machining or nonuniform heating and cooling during heat treatment

P/M grades, however, differ significantly from conventionally manufactured high-speed tool steels. Dimensional changes are more uniform in all directions. Because P/M high-speed tool steels are segregation free, variations in dimensional change are smaller. As a result, dimensional change occurring during hardening can be predicted more accurately. Conventionally processed high-speed tool steels go out-of-round in a four-sided pattern. The extent of distortion during heat treatment depends on the type and degree of segregation. In P/M high-speed tool steels, anisotropy is smaller, and out-of-roundness occurs in a close, circular pattern. Figure 7 shows typical results of measuring 102 mm (4 in.) diam disks after hardening and tempering. With P/M high-speed tool steels, cracking and variation of hardness are minimized because of their fine-grain, uniform structure.

Fig. 7 Out - of- roundness m easurem ent s on t est disks aft er hardening and t em pering. Test disks m achined from 102 m m ( 4 in.) diam bars. ( a) AI SI M2. ( b) ASP 30

The same precautions must be taken to control distortion due to residual stresses during heat treating. Mechanical stresses from rough machining can be eliminated by stress relieving prior to finish machining and heat treating.

Cr u cible Pa r t icle M e t a llu r gy Pr oce ss Since 1970, Crucible Materials Corporation has been producing powder metal tool steels commercially by the Crucible Particle Metallurgy (CPM) process. The process consists of induction melting and inert-gas atomizing, screening, and containerizing the prealloyed particles, followed by hot isostatic pressing to full density. See Fig. 8 for a schematic of the process elements. The desired chemical composition is melted, and the molten stream is poured into an atomizing chamber where high-pressure gas jets disperse it into spheroidal droplets that are rapidly quenched to ambient temperature. Powder is removed from the atomizing chamber, dried, and screened to obtain the desired size fraction. It is then poured into cylindrical steel cans that are evacuated and sealed. The cans are subsequently heated to a specific temperature and hot isostatically compacted to achieve a fully dense product. Compacts are processed to the desired billet and bar sizes by conventional hot rolling and forging (Ref 3).

Fig. 8 Schem at ic of CPM processing

As stated earlier, the most detrimental tendency of conventionally produced high-alloy high-speed tool steels is the high degree of alloy and carbide segregation that occurs during ingot solidification. This segregation not only reduces the hot workability and machinability of these alloys, but also results in reduced mechanical properties and tool performance. An increase in the carbon and alloy content results in increased segregation and low product yield after hot working of conventional ingot products. The CPM process was developed to minimize alloy segregation in standard high-alloy high-speed tool steel grades. Additionally, the CPM process is used to produce more highly alloyed grades than can be made by conventional practices. Pr ope r t ie s of CPM H igh - Spe e d Tool St e e ls A variety of CPM high-speed tool steels are available, as is shown in Table 3. Some of these grades are standard AISI steels such as M2, M3, M35, and M42, which normally are produced by conventional means, but when made by the CPM process offer notable advantages in toughness, out-of-roundness after heat treatment, and grindability. Others such as M4 and T15 are very difficult to produce by conventional means, but are readily producible by the CPM process with an improvement in properties. Still others, such as CPM Rex 20 and CPM Rex 76, are superhigh-speed tool steels that are very difficult or impossible to produce by conventional means and can only be made by the CPM route.

Table 3 Commercial CPM high-speed tool steel compositions Steel designation Trade name

AISI

Composition, % C Cr W

Mo

V

Co

S

CPM Rex M2HCHS CPM Rex M3HCHS CPM Rex M4 CPM Rex M4 HS CPM Rex M35 HCHS CPM Rex M42 CPM Rex 45 CPM Rex 45 HS CPM Rex 20 CPM Rex T15 CPM Rex T15 HS CPM Rex 76 CPM Rex 76 HS

M2 M3 M4 M4 M35 M42 ... ... M62 T15 T15 M48 M48

1.00 1.30 1.35 1.35 1.00 1.10 1.30 1.30 1.30 1.55 1.55 1.50 1.50

5.00 5.00 4.50 4.50 5.00 9.50 5.00 5.00 10.50 ... ... 5.25 5.25

2.00 3.00 4.00 4.00 2.00 1.15 3.00 3.00 2.00 5.00 5.00 3.10 3.10

... ... ... ... 5.00 8.00 8.25 8.25 ... 5.00 5.00 9.00 9.00

0.27 0.27 0.06 0.22 0.27 ... 0.03 0.22 0.06 0.06 0.22 0.06 0.22

4.15 4.00 4.25 4.25 4.15 3.75 4.00 4.00 3.75 4.00 4.00 3.75 3.75

6.40 6.25 5.75 5.75 6.00 1.50 6.25 6.25 6.25 12.25 12.25 10.00 10.00

Typical hardness, HRC 64-66 65-67 64-66 64-66 65-67 66-68 66-68 66-68 66-68 65-67 65-67 67-69 67-69

AISI T15 (Fe-12.25W-5Co-5.0V-4Cr-1.55C) demonstrates the advantages of the CPM process. This high-speed tool steel is one of the most wear- and heat-resistant grades of the standard American Iron and Steel Institute (AISI) high-speed tool steel materials. However, T15 usage has been limited because this highly alloyed, carbide-rich high-speed tool steel is difficult to produce by conventional production methods. The CPM process makes it possible to produce such difficult compositions in high volume. The size distributions of the primary carbides in CPM and conventional T15 are shown in Fig. 9. Most carbides in CPM high-speed tool steel are less than about 3 m (120 in.), whereas those in the conventional product cover the entire size range to approximately 34 m (1360 in.) with a median size of 6 m (240 in.). The microstructures of CPM and conventionally processed T15 are compared in Fig. 10.

Fig. 9 Prim ary carbide size dist ribut ions in CPM and convent ionally produced T15 high- speed t ool st eel

Fig. 1 0 Microst ruct ures of high- speed t ool st eels. Left : CPM T15. Right : Convent ional T15. Carbide segregat ion and it s det rim ent al effect s are elim inat ed wit h t he CPM process, regardless of t he size of t he product s. Court esy of Crucible Mat erials Corporat ion

Figure 11 shows a high-speed tool steel graph that summarizes the relative toughness, wear resistance, and red (hot) hardness characteristics of CPM and conventional versions of various AISI high-speed tool steel grades. As shown by the graph, the wear resistance and red hardness of a given grade of high-speed tool steel are equal for both its CPM and conventional versions. However, the wear resistance and red hardness generally increase with increasing alloy content,

and it is the very highly alloyed grades, such as CPM M4, CPM T15, CPM Rex 20, CPM Rex 45, and CPM Rex 76, that are best produced or can only be produced by the CPM method. The toughness comparison shows that the CPM version of each grade is notably tougher than the conventional high-speed tool steel.

Fig. 1 1 Com paragraph showing wear resist ance, red ( hot ) hardness, and t oughness of CPM and convent ional high- speed t ool st eels

The alloyed grade CPM Rex 76 is an example of a high-speed tool steel designed for production by the CPM process. This grade is a cobalt-rich high-speed tool steel with exceptional hot hardness and wear resistance and greatly increased tool life in difficult cutting operations. The high alloy content (32.5% compared to 27.8% for T15 and 25% for M42) renders this alloy unforgeable if produced by conventional processing. Two prominent high-speed steel cutting tool grades used in machining difficult-to-machine superalloys and titanium alloys used by the aircraft industry are T15 and M42, which contain 5 and 8% Co, respectively (Table 3). Cobalt increases the solidus temperature in high-speed tool steels, thereby permitting the use of high austenitizing temperatures to achieve greater homogeneous mixing of alloying elements. Cobalt also enhances the secondary hardening reaction, which results in a 1 to 2 HRC hardness advantage in the fully heat-treated condition. It also enhances hot hardness and temper resistance, thus allowing a tool to retain a sharp cutting edge at higher machining speeds that generate heat. Despite the advantages of cobalt additions, the high cost and occasional lack of availability of cobalt have necessitated the development of cobalt-free alternatives. CPM Rex 20 was developed as a cobalt-free P/M equivalent of M42. The chemical composition of CPM Rex 20 is listed in Table 3, and property comparisons of CPM Rex 20 with those of CPM and conventional M42 are shown in Tables 4, 5, and 6.

Table 4 Temper resistance of CPM alloys Alloy grade

CPM Rex 20 CPM M42 Conventional M42

Hardness, HRC As heat treated at 1190 °C (2175 °F) + 550 °C (1025 °F) 3 times/2 h 67.5 67.5 67.5

As heat treated + 595 °C (1100 °F)/ 2h

595 °C (1100 °F)/ 2 + 2h

As heat treated + 650 °C (1200 °F)/ 2h

650 °C (1200 °F)/ 2+2h

66 65.5 65

65.5 65 65

60 59 59

57 55.5 55

Table 5 Hot hardness of CPM alloys Alloy grade

CPM Rex 20 CPM M42 Conventional M42

Hot hardness, HRC At room temperature before test 67.5 67.0 66.5

At 540 °C (1000 °F) 58.0 58.5 58.5

At 595 °C (1100 °F) 56.0 56.0 56.0

At 650 °C (1200 °F) 47.5 48.0 48.0

At room temperature after test 64.0 63.0 62.0

Source: Ref 3

Table 6 Charpy C-notch impact energy and bend fracture strengths of CPM alloys and conventional alloy Alloy

CPM Rex 20 CPM M42 Conventional M42

(a)

Austenitizing temperature(a) °C °F 1190 2175 1190 2175 1190 2175

Hardness, HRC 67.5 67.5 67.5

Charpy C-notch Impact energy J ft·lbf 16 12 16 12 7 5

Bend fracture strength MPa ksi 4006 581 4006 581 2565 372

4 min soak in salt bath and oil quenched. Tempered at 550 °C (1025 °F) three times for 2 h.

Tables 4 and 5 show the results of temper resistance and hot hardness comparisons for specimens that were heat treated to full hardness. These results show that CPM Rex 20 and CPM and conventional M42 are equivalent in both temper resistance and hot hardness. Table 6 compares the Charpy C-notch impact energy and bend fracture strength values obtained for CPM Rex 20 with those obtained for CPM and conventional M42. Both the Charpy C-notch impact energy and the bend fracture strength of CPM Rex 20 are equal to those of CPM M42, but are notably higher than those of conventional M42. Table 7 shows the results of laboratory lathe tool tests in single-point turning on H13 and P/M René 95 superalloy. The overall performance of CPM Rex 20 is comparable to that of CPM M42.

Table 7 Lathe tool test results on CPM alloys Alloy grade

Austenitizing temperature °C °F

CPM Rex 20 1190 2175 1190 2175 CPM M42 Test conditions Speed, m/s (sfm) Feed, mm/rev (in./rev) Depth of cut, mm (in.) Coolant

Hardness, HRC

67.5 67

Tool life, minutes to 0.038 mm (0.015 in.) flank wear Intermittent cut on H13 steel at 33 HRC 8.5 8

Continuous cut on H13 steel at 33 HRC 14 16

Continuous cut on P/M René 95 at 33 HRC 31 27

0.20 (40) 0.10 (0.004) 1.57 (0.062) None

0.20 (40) 0.14 (0.0055) 1.57 (0.062) None

0.06 (12) 0.18 (0.007) 1.57 (0.062) None

Th e FULD EN S Pr oce ss Another method for the consolidation of P/M high-speed tool steels is the FULDENS process. This process differs from the others discussed in this article in that it uses water-atomized powders compacted either mechanically or by cold isostatic pressing and sintered in a vacuum to full density. For a discussion of mechanical properties of water-atomized powders that have been compacted and sintered, see the article "Production of Steel Powders" in Powder Metal Technologies and Applications, Volume 7 of the ASM Handbook. The FULDENS process allows close-tolerance complex shapes to be made with mechanical properties and performance characteristics comparable to those of equivalent parts made by conventional machining, with considerable material and labor savings. Pr oce ssin g St e ps. Figure 12 is a flowchart for the FULDENS process. The powders, usually water atomized, are

specially prepared in regard to composition as well as particle shape and size distribution. The powder is then annealed and pressed into green compacts, either by conventional mechanical pressing or by cold isostatic pressing. When part geometry allows, the part is compacted by filling a closed die with annealed powder and compacting with pressure ranging from 414 to 690 MPa (60 to 100 ksi). Cold isostatic pressing is more suitable for parts of relatively low volume, high complexity, and liberal tolerance; mechanical pressing is more suitable for parts of relatively high volume, low complexity, and close tolerance. Examples of isostatically pressed and mechanically pressed parts are shown in Fig. 13. For more information on cold isostatic pressing, see the article "Cold Isostatic Pressing" in Powder Metal Technologies and Applications, Volume 7 of the ASM Handbook.

Fig. 1 2 The FULDENS process for producing t ool st eel powder. Source: Ref 4

Fig. 1 3 Exam ples of part s m anufact ured by t he FULDENS process. Not e t he com plexit y of shapes at t ainable by t his process. ( a) Using cold isost at ic pressing. ( b) Using m echanical pressing

The compact is then sintered in a specialty built vacuum sintering furnace. After sintering, it emerges from the furnace fully dense. For more information on vacuum sintering, see the article "Production Sintering Practices" in Powder Metal Technologies and Applications, Volume 7 of the ASM Handbook. Adva n t a ge s. Hardenability, elongation, and impact strength of materials drop significantly if small amounts of porosity

(even 1 or 2%) are present. This has often limited the applicability of pressed and sintered powdered metals in demanding applications. Because the FULDENS process provides a finished part of nearly 100% density, good properties and product performance can be expected. In the FULDENS process, only the net weight of materials in the preform is actually used. In conventional manufacturing, much material is wasted in the form of chips. With higher-alloy high-speed tool steels that contain cobalt, molybdenum, and tungsten (materials often in short supply), savings in scrap reduction by using P/M processing can be substantial. The FULDENS process also eliminates much preheat-treat machining labor. Applica t ion s. Components produced by the FULDENS process should be considered for applications in which the

following conditions exist: • • • • •

Low net-to-gross weight ratios Relatively high strength-to-ductility ratios High wear environments High temperature environments High Hertz stresses

Applications in which fully dense P/M high-speed tool steel parts are currently in use include screw machine tooling, gear cutting tools, high-speed tool steel indexable inserts, and forming tools.

Applica t ion s of P/ M H igh - Spe e d Tool St e e ls M illin g. Milling cutters, such as those shown in Fig. 14, are emerging as a major application of P/M high-speed tool steels. Stock removal rates can usually be increased by raising the cutting speed and/or feed rate. In general, the feed per cutter tooth is increased in roughing operations, and the cutting speed is increased for finishing operations.

Fig. 1 4 Typical m illing cut t ers m ade from P/ M high- speed t ool st eels. Court esy of Speed- st eel I nc.

The performances of conventionally processed and P/M high-speed tool steel end mills in milling Ti-6Al-4V have been evaluated. In these tests, ASP 30 and ASP 60 were compared to M42. The cutting conditions used for this evaluation are given in Fig. 15, which shows tool life versus cutting speed. Both feed per tooth (0.203 mm, or 0.008 in.) and cutting speeds (>45.7 m/min, or 150 sfm) are higher than those used in production practice for machining aircraft parts. At a

constant metal removal rate that corresponds to a cutting speed of 53.3 m/min (175 sfm), ASP 60 and ASP 30 lasted eight times and 4.5 times longer, respectively, than the M42 end mill.

Cutter Feed Radial depth of cut Axial depth of cut Cutting fluid Tool life end point

25 mm (1 in.) diam end mills 0.203 mm/tooth (0.008 in./tooth) 6.35 mm (0.250 in.) 25.4 mm (1.000 in.) Soluble oil (1:20) 0.5 mm (0.020 in.) wear

Fig. 1 5 End m ill t est on Ti- 6Al- 4V. Hardness: 321 HB

Ot h e r m a t e r ia ls machined by P/M high-speed tool steel milling cutters include tough, hardened steels such as 4340, austenitic stainless steels such as AISI type 316, and nickel-base superalloys such as Nimonic 80. H ole M a ch in in g. Reamers, taps, and drills (Fig. 16) are also made from P/M high-speed tool steels. In one application,

the tool life of -20 GH3 four-flute plug taps made from CPM M4 and conventional M1, M7, and M42 were compared. The operation consisted of tapping a reamed 5.18 mm (0.204 in.) diam, 12.7 mm (0.500 in.) deep through-hole in AISI 52100 steel at 32 to 34 HRC using a speed of 7.8 m/min (26 sfm) and chlorinated tapping oil. Eight to thirteen taps of each grade were tested. The CPM M4 tap had an average tool life of 157 holes tapped before tool failure, compared to 35 holes for M1, 18 holes for M7, and 32 holes for M42. The tool life of the CPM M4 in this application was about five times the life of conventional M42.

Fig. 1 6 Ream ers, t aps, and drills m ade from P/ M high- speed t ool st eels. Court esy of Crucible Mat erials Corporat ion

Br oa ch in g. These tools constitute another major application for P/M high-speed tool steels because tool life is often

improved when P/M steels are used to broach difficult-to-cut materials such as case-hardened steels and superalloys. One application required broaching six ball tracks that are used in front-wheel-drive automobiles in constant-velocity joint hubs made of a case-hardening steel. Figure 17 shows the joint hubs and broaching tool used in this application.

Fig. 1 7 Broaching applicat ion. ( a) Tool m ade of P/ M high- speed t ool st eel t hat was used t o produce ball t racks on j oint hub. ( b) ASP 30 t ools produced 20,000 part s com pared t o 5600 part s by t ools m ade from convent ional high- speed t ool st eel. Court esy of Speedst eel I nc.

In this broaching application, surface finish and form tolerance requirements are high because subsequent machining is not performed on the ball tracks. In broaching with tools 18.0 mm (0.709 in.) in diameter and 185.0 mm (7.283 in.) in length made from low-carbon M35 steel (similar to M41 in chemical composition), the total number of hubs machined per tool was 5600. The M35 tools experienced severe flank wear and developed a large built-up edge, which produced poor surface finishes. With an ASP 30 tool, 20,000 parts were produced. Large broaching tools, such as those shown in Fig. 18, are also being made from P/M high-speed tool steels, such as P/M M3 and M4, to upgrade the broach material. In general, large rounds for broaches are not available in conventional highspeed tool steels in sizes above about 254 mm (10 in.), but larger sizes are available in P/M high-speed tool steels. One

application for these tools is the broaching of involute splines in bores of truck transmission gear blanks. Bores up to 305 mm (12 in.) in diameter by 1380 mm (54

in.) long have been cut using such tools.

Fig. 1 8 Large broaching t ool m ade from P/ M high- speed t ool st eel t hat was used for broaching involut e splines in bores of t ruck t ransm ission gear blanks. Court esy of Crucible Mat erials Corporat ion

Ge a r M a n u fa ct u r in g. Gear hobs (Fig. 19) made from P/M high-speed tool steels can also provide substantial cost reductions by increasing machining rates. One application called for hobbing of rear axle gears for heavy-duty trucks and tractor differentials. Hobs made of conventionally processed AISI M35 (65 HRC) and ASP 30 (67 HRC) were compared. Test parameters for both materials were:

• • • • • • • •

Hob dimensions: 152 mm (6 in.) diam × 50 mm (2 in.) diam × 205 mm (8 in.) length Work material: Case-hardening steel (Fe-1.2Ni-1Cr-0.2C-0.12Mo); hardness: 160 to 180 HB Cutting speed: 70 m/min (230 ft/min) Spindle speed: 150 rev/min Roughing: Feed, 4.24 mm (0.167 in.); depth of cut, 15.0 mm (0.591 in.) Finishing: Feed, 5.92 mm (0.233 in.); depth of cut, 0.81 mm (0.032 in.) Coolant: Cutting oil Number of parts per resharpening: 20

Fig. 1 9 Gear hobs m ade from P/ M high- speed t ool st eels. Court esy of Speedst eel I nc.

Production results showed that the flank wear land on the hobs made of ASP 30 (0.44 mm, or 0.017 in.) was much less than hobs made of M35 (0.71 mm, or 0.028 in.). Chipping of the edges was infrequent on the ASP 30 hobs, while the M35 hobs frequently displayed such damage. ASP 30 is now the standard grade for hobs used by one automotive manufacturer.

Tool Bit s. Figure 20 shows tool bits that have been produced using P/M high-speed tool steels. Typical applications

include machining of turbine blades made from superalloys and turning of hardened steels, such as AISI 4340 (1.9Ni0.75Cr-0.4C).

Fig. 2 0 Tool bit s m ade from P/ M high- speed t ool st eels. Court esy of Crucible Mat erials Corporat ion

Re fe r e n ce s 1. R. Riedl et al., Developments in High Speed Tool Steels, Steel Res., Vol 58 (No. 8), 1987, p 339-352 2. O. Siegwarth, "Higher Productivity with ASP Tooling Material," Technical Paper MF 81-137, Society of Manufacturing Engineers, 1981, p 1-22 3. F.R. Dax, W.T. Haswell, and W. Stasko, Cobalt-Free CPM High Speed Steels, in Processing and Properties of High Speed Tool Steels, The Metallurgical Society of the American Institute of Mining, Metallurgical, and Petroleum Engineers, 1980, p 148-158 4. M.T. Podob and R.P. Harvey, Advantages and Applications of CMI's FULDENS Process, in Processing and Properties of High Speed Tool Steels, The Metallurgical Society of the American Institute of Mining, Metallurgical, and Petroleum Engineers, 1980, p 181-195 Se le ct e d Re fe r e n ce s • • •

•

• • • • • •

B. Alvelid, and H. Wisell, Wear of High Speed Steel in Orthogonal Milling, Scand. J. Metall., Vol 9, 1980, p 59-67 J.T. Berry, "High Performance High Hardness High Speed Steels," Climax Molybdenum Company, 1970 E.A. Carlson, J.E. Hansen, and J.C. Lynn, Characteristics of Full-Density P/M Tool Steel and Stainless Steel Parts, in Modern Developments in Powder Metallurgy, Vol 13, Metal Powder Industries Federation, Princeton, NJ, 1980 B.-A. Cehlin, "Improving Productivity With High Strength P/M High Speed Steel Cutting Tools," Technical Paper MR82-948, Society of Manufacturing Engineers, presented at Increasing Productivity With Advanced Machining Concepts Clinic (Los Angeles), 1982 "Crucible CPM Rex-High Speed Steel for Superior Cutting Tools," Colt Industries E.J. Dulis and T.A. Neumeyer, Particle-Metallurgy of High-Speed Tool Steel, in Materials for Metal Cutting, Publication 126, The Iron and Steel Institute, 1970, p 112-118 P. Hellman, Wear Mechanism and Cutting Performance of Conventional and High-Strength P/M High-Speed Steels, Powder Metall., Vol 25 (No. 2), 1982 P. Hellman et al., The ASEA-STORA-Process, Modern Developments in Powder Metallurgy, Vol 4, Plenum Press, 1970, p 573-582 P. Hellman and H. Wisell, "Effect of Structure on Toughness and Grindability of High Speed Steels," Paper presented at Colloquium on High Speed Steels (Saint-Étienne, France), Nov 1975 W.E. Henderer and B.F. von Turkovich, "The Influence of Heat and Surface Treatments on the Performance of M1 HSS Taps," Paper presented at American Society for Metals Symposium on Hole

•

• • • • • • •

Making Operations (Boston), May 1977 W.J. Huppmann and P. Beiss, "Sintering of P/M Tool Steels to Full Density," Paper presented at Fully Dense P/M Materials for High Performance Applications, and Metal Powder Industries Federation Short Course, Sintermetallwerk Krebsoge (New Orleans), Feb 1982 A. Kasak and E.J. Dulis, Powder-Metallurgy Tool Steels, Powder Metall., Vol 21 (No. 2), 1978, p 114-123 A. Kasak, G. Steven, and T.A. Neumeyer, "High-Speed Tool Steels by Particle Metallurgy," Paper 720182, Society of Automotive Engineers, 1972 Properties and Selection: Stainless Steels, Tool Materials and Special-Purpose Metals, Vol 3, 9th ed., Metals Handbook, American Society for Metals, 1980 Powder Metallurgy, Vol 7, 9th ed., Metals Handbook, American Society for Metals, 1984 G.A. Roberts, J.C. Hamaker, Jr., and A.R. Johnson, Tool Steels, American Society for Metals, 1962, p 710-713 S. Söderberg, S. Hogmark, H. Haag, and H. Wissel, "Wear Resistance of High Speed Steel Milling Tools," Report 821 OR., Uppsala University, Institute of Technology "Tool Steels Today Newsletter," Committee on Tool Steel Producers, American Institute of Mining, Metallurgical, and Petroleum Engineers, Washington, D.C., Jan 1976

Ca st Coba lt Alloys I n t r odu ct ion CAST COBALT ALLOYS were developed to bridge the gap between high-speed steels and carbides. Although comparable in room-temperature hardness to high-speed steel tools, cast cobalt alloy tools retain their hardness to a much higher temperature (Fig. 1) and can be used at higher ( 20%) cutting speeds than high-speed steel tools. Unlike the highspeed steels that can be heat treated to obtain the desired hardness, cast cobalt alloys are hard in the as-cast condition and cannot be softened or hardened by heat treatment.

Fig. 1 Com parison of hot hardness values of cast cobalt alloys wit h alt ernat e cut t ing t ool m at erials. ( a) Hot hardness as a funct ion of t em perat ure. ( b) Recovery hardness as a funct ion of t em perat ure. Source: Ref 1

Pr oce ssin g, Pr ope r t ie s, a n d Applica t ion s Pr oce ssin g. Cast cobalt alloys are produced by electric or induction melting under a protective atmosphere, and for cutting tool applications they are preferably cast in permanent graphite molds. However, they can be cast in investment, shell, or sand molds to produce special and intricate shapes. Each of the melting and casting processes mentioned above is discussed in Casting, Volume 15 of ASM Handbook, formerly 9th Edition Metals Handbook. Pr ope r t ie s a n d Applica t ion s. Cast cobalt alloys contain a primary phase of cobalt-rich solid solution that is

strengthened by chromium and tungsten and is dispersion strengthened by complex, hard, refractory carbides of tungsten and chromium (Ref 2, 3). Nominal compositions for two commercially available grades are as follows:

Element Cobalt Chromium Tungsten Carbon Tantalum or niobium Manganese Iron Nickel

(a)

Alloy Tantung G, % 42-47 27-32 14-19 2-4 2-7 1-3 2-5 7(a)

Tantung 144, % 40-45 25-30 16-21 2-4 3-8 1-3 2-5 7(a)

Maximum

The typical microstructure of a permanent graphite mold cast Tantung G alloy is shown in Fig. 2. Properties of Tantung G are given in Table 1. For comparison, permanent mold cast Tantung 144 has a hardness of 61 to 65 HRC, a transverse strength of 2070 MPa (300 ksi), and an elastic modulus of 295 GPa (43 × 106 psi).

Table 1 Typical properties of cast Tantung G Property Melting temperature, °C (°F) Casting temperature, °C (°F) Density, g/cm3 (lb/in.3) Thermal expansion,

m/m · °C ( in./in. · °F)

Thermal conductivity, W/m · K (Btu/ft · h · °F) Hardness, HRC Transverse strength, MPa (ksi) Modulus of elasticity, GPa (106 psi) Tensile strength, MPa (ksi) Compressive strength, MPa (ksi) Impact strength, J (ft · lb)

Permanent mold cast 1150-1200 (2100-2200) 1370 (2500) 8.3 (0.30) 4.2 (2.3) 26.8 (15.5) 60-63 2240 (325) 265 (41) 585-620 (85-90) 2760 (400) 6.1 (4.5)

Refractory mold cast 1150-1200 (2100-2200) 1370 (2500) 8.3 (0.30) 4.2 (2.3) 26.8 (15.5) 53-58 1030-1200 (150-175) ... 450 (65) 2930 (425) 6.1 (4.5)

Fig. 2 Microst ruct ure of cast Tant ung G alloy. Et ched wit h Murakam i's reagent ( st andard m ix: 10 g sodium

hydroxide, 10 g pot assium ferricyanide, 100 m L H2 O) . 400× . Court esy of G.F. Vander Voort , Carpent er Technology

Tantung G is recommended for general-purpose cutting tools and parts for wear applications, and it is more widely used than Tantung 144. Tantung 144 has higher hardness than Tantung G and was developed for use where resistance to abrasion is paramount and where there is little or no shock or impact. The good resistance of Tantung G and Tantung 144 to foods, particularly those containing acetic acid, makes them highly suitable for food-processing equipment, especially parts requiring good resistance to abrasion and corrosion. Water containing chlorine and hypochlorites may produce some corrosion and pitting of alloys G and 144. In addition, these alloys are attacked by strong acid solutions, alkalies, and solutions of some heavy-metal salts such as ferric chloride, ferric sulfate, and cupric chloride. When heated in air, Tantung G and Tantung 144 both tarnish on short-time exposures at 400 °C (750 °F) and lose appreciable weight at 750 °C (1380 °F) or higher. Scaling may be progressive above 1000 °C (1830 °F). Hot hardness data for cast cobalt alloys are shown in Fig. 1. The use of cast cobalt cutting tools should be considered: • • • • • •

Where relatively low surface speeds cause buildup with cemented carbides Where machines lack the power or rigidity to use cemented carbides effectively Where higher production is desired than is possible with high-speed steel tools For multiple-tool operations in which the surface speed of one or more operations falls between the recommended speeds for high-speed steel and carbide tools For short runs on automatic equipment in which the form grinding of carbide tools is excessively costly For machining rough surfaces of castings where the surfaces contain abrasive material, such as residual sand, surface oxides, slag, or refractory particles

Tools made of cast cobalt alloys usually are not recommended for light, very fast finishing cuts. Typical wear applications for these alloys include wear strips for belt sanders; dies for extruding copper, for extruding molybdenum tubing, and for hot swaging tungsten rod; burnishing rolls; internal chuck jaws; drill bushings; and knives for slicing fruits, vegetables, and meat.

Re fe r e n ce s 1. R. Komanduri, Tool Materials, in Encyclopedia of Chemical Technology, Vol 23, 3rd ed., John Wiley & Sons, 1983 2. K.J.A. Brooks, World Directory and Handbook of Hard Metals, Engineer's Digest Ltd., 1976 3. R.L. Hatschek, Am. Mach., Vol 733 (No. 165), May 1981

Ce m e n t e d Ca r bide s A.T. Sant hanam and P. Tierney, Kennam et al I nc.

I n t r odu ct ion CEMENTED CARBIDES belong to a class of hard, wear-resistant, refractory materials in which the hard carbide particles are bound together, or cemented, by a soft and ductile metal binder. These materials were first developed in Germany in the early 1920s in response to demands for a die material having sufficient wear resistance for drawing tungsten incandescent filament wires to replace the expensive diamond dies then in use. The first cemented carbide to be produced was tungsten carbide (WC) with a cobalt binder. Tungsten carbide was first synthesized by the French chemist Henri Moissan in the 1890s (Ref 1). There are two types of tungsten carbide: WC, which directly decomposes at 2800 °C (5070 °F), and W2C, which melts at 2750 °C (4980 °F) (Ref 2, 3). Early attempts to produce drawing dies from a eutectic alloy of WC and W2C were unsuccessful, because the material had many flaws and fractured easily. The use of powder metallurgy techniques by Schroeter in 1923 paved the way for obtaining a fully consolidated product (Ref 4). Schroeter blended fine WC powders with a small amount of iron, nickel, or cobalt powders and pressed the powders into compacts, which were then sintered at approximately 1300 °C (2400 °F). Cobalt was soon found to be the best bonding material. Over the years, the basic WC-Co material has been modified to produce a variety of cemented carbides, which are used in a wide range of applications, including metal cutting, mining, construction, rock drilling, metal forming, structural components, and wear parts. Approximately 50% of all carbide production is used for metal cutting applications. Although the term cemented carbide is widely used in the United States, these materials are better known as hard metals internationally. This article will discuss the manufacture and composition of cemented carbides and their microstructure, classifications, applications, and physical and mechanical properties. New tool geometries, tailored substrates, and the application of thin, hard coatings to cemented carbides by chemical vapor deposition and physical vapor deposition will also be discussed. This article is limited to tungsten carbide cobalt-base materials. Information on titanium carbide nickel-base materials is given in the article "Cermets" in this Volume. Extensive reviews of the scientific and industrial aspects of cemented carbides are available in Ref 5, 6, 7, and 8.

M a n u fa ct u r e of Ce m e n t e d Ca r bide s Cemented carbides are manufactured by a powder metallurgy process consisting of a sequence of steps in which each step must be carefully controlled to obtain a final product with the desired properties, microstructure, and performance. The steps include: • • • • • •

Processing of the ore and the preparation of the tungsten carbide powder Preparation of the other carbide powders Production of the grade powders Compacting or powder consolidation Sintering Postsinter forming

The sintered product can be directly used or can be ground, polished, and coated to suit a given application. Pr e pa r a t ion of Tu n gst e n Ca r bide Pow de r . There are two methods by which tungsten carbide powders are produced from the tungsten-bearing ores. Traditionally, tungsten ore is chemically processed to ammonium paratungstate and tungsten oxides. These compounds are then hydrogen-reduced to tungsten metal powder. The fine tungsten powders are blended with carbon and heated in a hydrogen atmosphere between 1400 and 1500 °C (2500 and 2700 °F) to produce tungsten carbide particles with sizes varying from 0.5 to 30 m (Fig. 1). Each particle is composed of numerous tungsten

carbide crystals. Small amounts of vanadium, chromium, or tantalum are sometimes added to tungsten and carbon powders before carburization to produce very fine (50

50 max

>50

135-275 HB

Condition

Hot rolled, normalized, annealed, cold drawn, or quenched and tempered Carburized and/or quenched and tempered

Hot rolled, normalized, annealed, cold drawn, or quenched and tempered Carburized and/or quenched and tempered

Annealed or quenched and tempered

Quenched and tempered

Annealed or cold drawn

Thread designation Pitch, Threads mm per 25 mm (1 in.)

Wheel identification ANSI

Wheel speed

Work speed

m/s

sfm

m/min

sfm

6-2.5 2-1 0.80.45 0.40.35 6-2.5 2-1 0.80.45 0.40.35

4-12 13-27 28-63

A90RB A120SB A180TB

43 46 48

8,500 9,000 9,500

1.2

4.0

64-80

A220UB

51

10,000

4-12 13-27 28-63

A100KV A150JV A220MV

36 38 43

7,000 7,500 8,500

1.8

6.0

64-80

A240PV

46

9,000

6-2.5 2-1 0.80.45 0.40.35 6-2.5 2-1 0.80.45 0.40.35

4-12 13-27 28-63

A90RB A120SB A180TB

43 46 48

8,500 9,000 9,500

1.2

4.0

64-80

A220UB

51

10,000

4-12 13-27 28-63

A100KV A150JV A220MV

36 38 43

7,000 7,500 8,500

1.8

6.0

64-80

A240PV

46

9,000

6-2.5 2-1 0.80.45 0.40.35 6-2.5 2-1 0.80.45 0.40.35

4-12 13-27 28-63

A90RB A120SB A180TB

43 46 48

8,500 9,000 9,500

1.2

4.0

64-80

A220UB

51

10,000

4-12 13-27 28-63

A100KV A150JV A220MV

36 38 43

7,000 7,500 8,500

1.8

6.0

64-80

A240PV

46

9,000

6-2.5 2-1 0.80.45 0.4-

4-12 13-27 28-63

A90RB A120SB A180TB

43 46 48

8,500 9,000 9,500

1.2

4.0

64-80

A220UB

51

10,000

0.35 Gray cast irons 45 max Ferritic: ASTM A48, class 20; SAE J431c, grade G1800 Pearlitic-ferritic: ASTM A48, class 25; SAE J431c, grade G2500 Pearlitic: ASTM A48, classes 30, 35, 40; SAE J431c, grade 45-52 G3000 Pearlitic + free carbides: ASTM A48, classes 45, 50; SAE J431c, grades G3500, G4000 Pearlitic or acicular + free carbides: ASTM A48, classes 55, 60 Wrought high-temperature nickel-base alloys 200-390 AF2-1DA, Astroloy, Haynes HB 263, IN-102, Incoloy 901, Incoloy 903, Inconel 617, Inconel 625, Inconel 700, Inconel 702, Inconel 706, Inconel 718, Inconel 721, Inconel 722, Inconel X-750, 300-475 Inconel 751, M252, Nimonic 75, HB Nimonic 80, Nimonic 90, Nimonic 95, René 41, René 63, René 77, René 95, Udimet 500, Udimet 700, Udimet 710, Unitemp 1753, Waspaloy Carbides 89-94 Titanium carbide, tungsten HRA carbide

As-cast, annealed, or quenched and tempered

As-cast, annealed, or quenched and tempered

Annealed or solution treated

Solution treated and aged

...

6-2.5 2-1 0.80.45 0.40.35 6-2.5 2-1 0.80.45 0.40.35

4-12 13-27 28-63

A90RB A120SB A180TB

43 46 48

8,500 9,000 9,500

64-80

A220UB

51

10,000

4-12 13-27 26-63

A90RB A120SB A180TB

43 46 48

8,500 9,000 9,500

64-80

A220UB

51

10,000

6-2.5 2-1 0.80.45 0.40.35 6-2.5 2-1 0.80.45 0.40.35

4-12 13-27 28-63

A90SB A100TB A180TB

46

64-80

A280TB

4-12 13-27 28-63

A90SB A100TB A180TB

64-80

A280TB

6-2.5

4-12

2-1

13-27

0.80.45 0.40.35

28-63 64-80

D240R100B D320T100B D400V100B D400V100B

1.2

4.0

1.2

4.0

9,000

0.451.2

1.54.0

46

9,000

0.451.2

1.54.0

23

4,500

0.45

1.5

25

5,000

28

5,500

30

6,000

Source: Ref 3

As indicated in Table 4, in most thread grinding, the recommended wheel speed increases as the number of threads per 25 mm (1 in.) increases. Grinding of heat-resistant alloys, however, is an exception; the recommended speed is the same regardless of thread pitch.

Gr in din g Flu ids Grinding oils are always used in thread grinding. Water-base fluids have never been successful, although many attempts have been made to replace oils with water-base fluids. Mineral-base sulfochlorinated grinding oils such as G1 or G2 are the most widely used, especially when finish and accuracy of the threads are important:

G1 G2

Transparent sulfochlorinated grinding oil containing fats, 4% S, and 2% Cl (both active); viscosity: 230 SUS at 40 °C (100 °F) Dark sulfochlorinated grinding oil containing fats, 3% S, and 0.5% Cl (both active); viscosity: 190 SUS at 40 °C (100 °F)

Such oils are generally satisfactory for all steels, as well as for other work metals, and many shops have standardized on them for all thread grinding. Other formulations prepared especially for thread grinding are available for specific applications where the conventional sulfochlorinated oils have proved unsatisfactory.

Cylin dr ica l Gr in din g of Th r e a ds Five basic methods, identified by the design of the wheel, are employed for cylindrical thread grinding. Four of these methods use multiribbed wheels: • • • • •

Single-rib wheel traverse grinding Multirib wheel traverse grinding Multirib wheel plunge grinding Multirib wheel skip-rib, or alternate-rib, grinding Multirib wheel three-rib grinding

Sin gle - Rib W h e e l Tr a ve r se Gr in din g M a ch in in g Pr oce ss. In this method (Fig. 4a), the edge of the wheel is trued to the form of the thread to be ground; the

profile of the cutting edge works like that of a single-point metal tool for cutting threads on a lathe. Because the grinding wheel is frequently 457 to 508 mm (18 to 20 in.) in diameter, the arc of contact at full depth is appreciable. Thus, for accuracy, the wheel should be inclined to the helix angle; this adjustment is imperative for helices of 4° or more.

Fig. 4 Cylindrical grinding of t hreads using ( a) single- rib, ( b) m ult irib t raverse, ( c) m ult irib plunge, and ( d) skip- rib m ult iribbed grinding wheels

Single-rib wheel traverse grinding is the original and most versatile method of thread grinding, and it is still applied in the majority of operations. It can also be considered the method by which the highest degree of dimensional accuracy of the ground thread can be attained. The capacity limits of single-rib wheel thread grinding are the widest available, extending from the finest to the coarsest thread types. The single-rib wheel is also adaptable, with the aid of appropriate truing devices, to many different profile configurations, including those that differ from the conventional thread shapes, such as hobs, worms, and translation screws with special ball tracks.

Re ce n t D e ve lopm e n t s t o I n cr e a se Pr odu ct ivit y. One respect in which single-rib wheel grinding is inferior to

methods employing multirib wheels is productivity. Several design improvements, however, have been incorporated into different models of thread grinding machines for increasing the productivity of the single-rib wheel method. These design characteristics, although found only in certain models of thread grinders, include the following: • •

• •

•

The wheel diameter has been increased to 508 mm (20 in.) to extend the length of the operating periphery, thus removing a higher volume of stock before retruing is necessary The peripheral speed of the grinding wheel has been increased, often to a multiple of the conventional speed. This preceded by many years the application of high-speed grinding, which more recently is used in other precision grinding methods Wheel truing systems and devices are used that substantially reduce or even eliminate the loss in productive grinding time due to wheel truing For multipass grinding--a procedure that is generally needed for the single-rib wheel method--the accelerated table return has long been provided, but in certain models of modern production grinders it is being superseded by the even more productive two-way grinding. In this latter system, also referred to as two-directional, the grinding is carried out during the table movements in both directions, but it requires specially designed grinders equipped with dependable backlash-eliminator devices. Some manufacturers are building thread grinders that can operate in an automatic cycle using both methods subsequently: the two-way grinding (in roughing) for increased productivity and the one-way grinding (in finishing) for higher work accuracy The wheel and work speeds are varied automatically in different phases of the processes. High wheel speeds are used during the roughing passes for increased rate of stock removal. The wheel speed is then decreased, and the work speed increased, during the finishing passes, thus obtaining a better finish on the ground surface

M u lt ir ib W h e e l Gr in din g Multiribbed wheels have two or more parallel grooves or ribs around the periphery of the wheel. Each rib is trued to the form of the thread to be ground. The thread form is imparted to the wheel by diamond or crush truing. Tr a ve r se Gr in din g. If the grinding wheel is as thick as, or thicker than, the required length of the thread (Fig. 4b), the

thread can be completed in one revolution of the work plus a half revolution for feeding-in of the wheel. Traverse grinding operates with a wheel that is wide enough to contain ribs corresponding to six to eight threads of the workpiece. The leading edge of the wheel is usually tapered, with the first two or three ribs truncated in order to cause the wheel to penetrate gradually into the work material as it is traversed axially. The wheel is advanced to the intended radial position of the cut, which may be the full depth of the thread for single-pass grinding or only a part of the total thread depth for multipass grinding. After the wheel has reached its advance position, the traversing and rotating workpiece contacts the leading edge of the wheel and continues the traverse movement across the face of the wheel until the end of the threaded part section has passed the wheel. In the case of multipass grinding, the process of wheel retraction and work return is the same as in single-rib wheel grinding, but a longer return distance must be covered because of the greater wheel width. Traverse grinding can also be operated in the two-way system. In this case, both sides of the wheel function alternately as leading edges, and the starting ribs are truncated at both ends of the wheel face. Multirib wheel traverse grinding is more productive than single-rib wheel grinding because of the higher rate of stock removal per pass, and it is frequently used on thread grinding machines that are adapted to this method by the use of wider wheels and the appropriate multirib wheel-truing devices. The application of multirib wheel grinding, however, has certain limitations: • •

Thread pitch should not exceed one-eighth of the available wheel width There should be no shoulder on the workpiece that could interfere with its passing across the total width of the wheel face

Regarding the accuracy attainable, multirib wheel traverse grinding is successfully applied to threads even with tight tolerances, except for such extreme accuracy requirements as specified for thread gages.

Plu n ge Gr in din g. For this method of grinding, wheels that are about 1

thread pitches wider than the entire threaded

section of the workpiece must be used, but the complete thread is produced in less than 1 turns of the workpiece. The wheel, which has ring-shaped thread profile ribs of equal depth around its entire periphery, is advanced gradually into the rotating and traversing workpiece during about turn of the latter. The wheel then stays in this position while the workpiece continues making somewhat more than an additional full turn (Fig. 4c). At this position, the entire thread has been ground in a single plunged pass; the wheel retracts and the machine table returns into its starting position, ready to accept the next workpiece. Multirib wheel plunge grinding is the most productive thread grinding method, and it is successfully used where the dimensions of the workpiece and of the thread permit its application. The accuracy attainable satisfies most general requirements, with the exception of parts having very critical thread tolerances. The truing of the multirib wheel is commonly performed by crushing, a truing method that was adopted only for certain types of thread grinding machines, including the special machine types built expressly for plunge grinding with crushdressed wheels. For high-volume production, diamond-impregnated form rolls are also used for the very efficient truing of multirib thread grinding wheels. Because of the relatively high cost of the required special tooling, the crushing rolls, and the diamond rolls (even much more expensive), multirib wheel thread grinding by plunge advance must be used for the production of parts in substantial quantities in order to be economical in comparison to other thread grinding methods. Sk ip- Rib Gr in din g. A type of multirib wheel plunge grinding, designated as skip-rib, with rib spacing on the wheel twice that of the thread pitch, is occasionally used for threads of very fine pitch or for resinoid-bonded wheels where the truing of ribs corresponding to the actual thread pitch would be difficult to carry out. The grinding wheel can be formed by crush or diamond truing. Diamond truing is used on resinoid wheels or to dress fine, accurate pitches. In the first pass, the wheel grinds every other thread of the workpiece; in a subsequent pass, after the work is traversed by a distance equal to the thread pitch, the previously untouched threads are ground over a distance comprising a small circumferential

overlap for bending (Fig. 4d). This method requires about 2

total turns of the workpiece for the entire operation.

Th r e e - Rib Gr in din g. Another variation of the multirib thread grinding wheel is the three-rib wheel (Fig. 5). This wheel has a roughing rib (A, Fig. 5) that removes about two-thirds of the material, and an intermediate rib (B, Fig. 5) that takes the remainder of the material except for about 0.13 mm (0.005 in.), which is left to be cleaned up by the finishing rib (C, Fig. 5). If desired, the wheel can be dressed so that a flattened area (D, Fig. 5) will finish grind the crest of the thread.

Fig. 5 Schem at ic of a t hree- rib t hread grinding wheel. A, roughing rib; B, int erm ediate rib; C, finishing rib. The

flat t ened area ( D) is opt ional and can be used t o finish grind t he crest of t he t hread.

The three-rib wheel will produce threads whose accuracy equals or exceeds that of threads produced by the single-rib wheel. If necessary, this type of wheel can be inclined to the helix angle, provided an allowance is made for the radius of curvature of the work. The three-rib wheel is designed to traverse the work rather than to plunge cut. I n fe e d When a single-rib wheel is used, a roughing cut of 0.51 to 1.02 mm (0.020 to 0.040 in.) per pass can be made without burning the work metal. An infeed of 0.038 to 0.10 mm (0.0015 to 0.004 in.) per pass is used for close dimensional accuracy, while infeed for larger dimensional tolerances common to many commercial products ranges from 0.10 to 0.25 mm (0.004 to 0.010 in.) per pass. The work can be ground to final size in one cut when the depth of cut is less than 1.02 mm (0.040 in.). Work speed for finish grinding in one cut is slow; for example, a work speed of 0.46 to 0.61 m/min (1 to 2 sfm) is used for grinding 20 threads per 25 mm (1 in.) with a thread depth of 0.81 mm (0.032 in.). For depths greater than 1.02 mm (0.040 in.), roughing cuts of 0.51 to 1.02 mm (0.020 to 0.040 in.) or more per pass at work speeds of 0.9 to 1.2 m/min (3 to 4 sfm) are required. Work speed for the finish cut of 0.038 to 0.10 mm (0.0015 to 0.004 in.) is 1.8 to 2.4 m/min (6 to 8 sfm). In multicut operations, infeed should be controlled so that the cutting edge of the grinding wheel will not break down before the finishing cut. For example, an infeed of 0.91 mm (0.036 in.) per pass for roughing and 0.46 mm (0.018 in.) per pass for finishing is satisfactory for grinding 20 screw threads per 25 mm (1 in.). Thread grinding can be done on the up cut, usually with a better finish than that possible when grinding is done on the climb cut (Fig. 6). With up cut, the materials removed from the work and the wheel do not have to pass over the finished surface. Figure 7 shows a tube held in a chuck while being threaded internally using a cylindrical grinding setup.

Fig. 6 Direct ion of cut in grinding t hreads. ( a) Up cut . ( b) Clim b cut

Fig. 7 Schem at ic of a cylindrical grinding set up in which a grinding wheel is generat ing int ernal t hreads on a piece of t ubing

Ce n t e r le ss Gr in din g of Th r e a ds Centerless grinding is the most productive method of grinding screw threads. This method is capable of grinding threads of classes 2, 2A, 3, 3A, and 7 at high production rates; threads of class 4 and 5A are produced at slower rates. Wheels for centerless thread grinding can also be either single-ribbed or multiribbed. Grain size ranges from 100 to 320, depending on the number of threads per 25 mm (1 in.), as indicated in Tables 1 and 2. Harder wheels are used in centerless thread grinding than for other centerless grinding operations on the same materials. Grade (hardness) ranges from M to R. Wheels of grade Q are generally used for such products as headless setscrews made of steel at 40 HRC. Wheel structure is usually fairly open, ranging from 8 to 11, depending on the wheel manufacturer. Wheel life depends on the type and hardness of the material being ground, thread tolerance, and production rate. Because maximum wear occurs at the thread crests of the grinding wheel, the wheel must be dressed to maintain the tolerance of the root of the product thread. Centerless thread grinding wheels are dressed by crushing. Re gu la t in g w h e e ls for centerless thread grinding are similar to those used for other centerless grinding operations. For precision thread grinding, however, regulating wheels of finer grain size are used.

Regulating wheels usually revolve in the same direction as the grinding wheel (down grinding, Fig. 8a). However, for coarse threads (1 to 8 coarser), the regulating wheel rotates in the direction opposite to that of the grinding wheel (up grinding, Fig. 8b).

Fig. 8 Two m et hods of cent erless t hread grinding. ( a) Down grinding. ( b) Up grinding

Regulating wheels are dressed with a single-point diamond dressing tool. The frequency of dressing depends on the depth of the grooves that develop in the wheel face during thread grinding. In general, the regulating wheel is dressed each time the grinding wheel is dressed, but in some operations the regulating wheel is dressed only once for every two or three dressings of the grinding wheel. Scr e w Th r e a d Pr odu ct ion . Screw threads can be ground from the solid by the centerless method. Adjustments are

provided to accommodate work of different sizes and for varying the rates of feed. The grinding wheel is a multirib type, which consists of a series of annular ridges across the face. These ridges conform in pitch and profile with the thread to be ground. The grinding wheel is inclined to suit the helix or lead angle of the thread. In grinding threads on such work as socket-type setscrews, the blanks are fed automatically and passed between the grinding and regulating wheels in a continuous stream. To illustrate production possibilities, hardened socket setscrews of -20 size can be ground from the solid at the rate of 60 to 70 per min and with the wheel operating continuously for 8 h without redressing (Fig. 9). The lead errors of centerless ground screw threads can be limited to 0.0005 mm/mm (0.0005 in./in.) or even less by reducing the production rate. The pitch diameter tolerances are within 0.0051 to 0.0076 mm (0.0002 to 0.0003 in.) of the basic size. The grain size for the wheel is selected with reference to the pitch of the thread. The following sizes are recommended: •

For 11 to 13 threads per 25 mm (1 in.), 150

• • • •

For 16 threads per 25 mm (1 in.), 180 For 18 to 20 threads per 25 mm (1 in.), 220 For 24 to 28 threads per 25 mm (1 in.), 320 For 40 threads per 25 mm (1 in.), 400

Fig. 9 Principle of cent erless t hread grinding. This m et hod is used t o produce headless set screws. The blanks are hopperfed t o posit ion A. The regulat ing wheel causes t hem t o t raverse t he grinding wheel face, from which t hey em erge at posit ion B in com plet ed form . A product ion rat e of 60 t o 70 screws of 12.7 m m ( per m in is possible.

in.) lengt h

Pr odu ct ion Pr a ct ice Pitches as fine as 80 threads per 25 mm (1 in.) can be ground with conventional thread grinding procedures. Special techniques have been developed for grinding threads as fine as 400 per 25 mm (1 in.). Most threads ground in regular production are considerably coarser than 80 per 25 mm (1 in.). A hard thread can be obtained by thread grinding because the workpiece can be ground after it has been fully hardened. This eliminates distortion, which is commonly found in workpieces threaded before heat treating, and produces true thread forms with the lead and pitch diameter held to close tolerances. Limits of 0.005 mm (0.0002 in.) or less can be held on pitch diameter, and less than 0.0002 mm/mm (0.0002 in./in.) on lead tolerances. Thread grinding can be used on surfaces containing drilled holes, keyways, or other irregularities. Shell-type parts are more amenable to thread grinding than to other methods of producing threads, in which pressure from the tool can cause distortion. Small worms can be roughed faster by thread grinding than by conventional thread milling. Thread grinding is applicable to virtually all materials and all hardnesses. On the other hand, thread cutting is superior from the standpoint of material handling and production rate, especially when the thread tolerances are broad. However, these advantages of thread cutting are quickly offset when close tolerances are required. When threads of fine finish and close tolerances are needed, thread grinding is more economical than other methods, including thread cutting and thread milling. Internal threads can also be produced by grinding, using a machine especially designed for internal threading or a universal machine. H igh - Volu m e Applica t ion s of Th r e a d Gr in din g In production quantities, thread grinding can be used to produce diverse types of threaded parts, as well as other types of parts that have helical profiles of particular shapes. Table 5 lists commonly used mechanical parts whose threading and similar features are often produced by thread grinding.

Table 5 Mechanical parts often produced with ground threads or similar features Designation Cutting tools Thread cutting taps

Thread cutting dies Thread cutting chasers

Thread milling cutters

Gear-generating hobs Forming tools Crush rolls for thread forms Thread rolling dies

Discussion Taps were the first types of workpieces manufactured in high-volume production with ground threads, and they still represent one of the principal applications of the method. Thread grinding ensures superior accuracy of all the important parameters and can produce relieved thread profiles. By grinding threads in solid blanks, thread grinding also has a productivity rate exceeding that of only approximately equivalent alternative processes. Universal thread grinding machines with internal grinding attachments are commonly used for grinding the threaded surfaces of these tools and will produce threads with excellent dimensional accuracy and the unique advantage of relieved thread profile. Both the radial and the tangential types, used in threading heads, can be produced on thread grinding machines equipped with special devices imparting a straight-line reciprocating motion to the workpieces, which are usually mounted in sets. Similar methods can be used for multirib thread chasers used in lathe work. Round thread chasers are ground on basic thread grinding machine equipment. Multirib thread milling cutters are generally manufactured with ground threads that have relieved form on the entire peripheral sections between the gashes. Producing these relieved profiles by thread grinding results in superior accuracy, and thread grinding is also faster than the earlier applied relief-form turning. The relief grinding of thread milling cutters with helical gashes generally requires a special attachment. Hobs used as the tools of continuously operating gear milling machines generally have ground profiles, particularly for ensuring a high accuracy of tooth form and spacing, but also for higher cutting ability of tool edges that are free of decarburization. Special attachments or thread grinding machines, operating with small-diameter wheels, are needed. Multirib wheels are the means of very efficient thread grinding methods, and crush dressing is the most commonly used process for forming the wheels. The crush rolls are ground on thread grinding machines equipped with devices to grind no-lead (annular) thread forms. Both types, the flat dies and the round dies, also known as thread forming rolls, are ground on thread grinding machines. The flat dies require a reciprocating attachment. Higher thread accuracy and better service life of the dies are obtained by thread grinding than by milling followed by heat treatment.

Measuring instruments The accuracy level that thread gages for internal and external surfaces must satisfy, with respect to all the Thread plug and significant parameters, is most effectively and dependably obtained by thread grinding. Plug gages and (with the ring gages aid of an internal grinding attachment) ring gages are ground with all types of profiles, both straight and tapered. Micrometer screws Screw threads are a reliable and sensitive mechanical means of imparting accurately controlled linear displacement, combined with the ability to display such motions at a high rate of dependably proportional amplification. Micrometer screws are commonly manufactured with ground threads for the highest degree of lead accuracy. Racks, in combination with pinions or worms, have the potential to produce very sensitively specific linear Gage and displacements, whose accuracy is controlled by the shape and spacing of the meshing teeth. Thread grinders with instrument racks special attachments for work reciprocation and incremental table traverse are suitable for grinding such precise racks. Translation elements The potential advantages of screws serving as translation elements, such as the favorable ratio of input torque to Leadscrews and the resulting translational force, with sensitive and precise motion control, can best be implemented by screws feed screws with ground threads that can be produced even on very long lead screws. Screws and nuts connected by recirculating balls represent translation elements with excellent force efficiency; Ball screws and they operate with practically no play and have very long service lives. Precision-type ball screws and nuts, with a ball nuts circular or gothic-arch ball-track profile, are produced with high form and lead accuracy on thread grinding machines. The efficiency of worm gear drives is particularly sensitive to the accuracy of the driving member. Therefore, Worms ground worms, with flank contours correct either with respect to the axis or the helix, are often required and produced on thread grinding machines adapted to those operations that may need special wheel profiles and very high helix angles. Fasteners High-strength threaded fasteners are preferably made by thread grinding because of the favorable stress Screws and bolts distribution that results from the maximum flank contact ensured by accurate flank angles and lead. High production rates can be achieved on conventional or centerless automatic thread grinders. Hardened steel machine elements having screw threads only on limited sections but in a very precise relation to Machine elements other ground portions of the surface are best finished by thread grinding. Preferably, the part is held on the same with threaded mounting surfaces that were used in the preceding operations for producing the locationally related surface sections elements. Special helical profiles When these must be produced with specific contours, precise lead, and excellent surface texture, thread grinding Machine elements may be the best suited or only method. Standard machines or special equipment is used. Examples include with helical

features

metering feed screws and steel regulating wheels for special centerless grinding processes.

Source: Ref 2

Thread grinding as a metalworking method is not limited to parts whose quality specifications warrant or even require its use. In many cases, thread grinding can be the most productive method of producing certain part configurations. In such applications, the accuracy, finish, and other quality-related properties of the ground surface constitute additional benefits. With respect to productivity, the advantages of thread grinding over those of alternative methods may arise from various conditions, as follows. I n cr e a se d Ra t e of St ock Re m ova l. The limiting conditions of chip disposal from a confined work area can be

overcome more easily in grinding than in other metalworking methods. W h e e l Tr u in g. Instead of frequent tool changes, only the regular truing of the installed grinding wheel needs to be

applied, often in a manner not interfering with the productive cycle time of the operation. This property of thread grinding may prove particularly valuable in the production of long threaded parts, such as leadscrews, where the volume of stock removed may exceed the edge-holding limits of a single-point cutting tool. Rou gh in g a n d Fin ish in g Ope r a t ion s. Roughing, even from the solid, and subsequent finishing can be combined

into a single operation, using the same tool, by changing the infeed rate. The condition of the work produced in a single pass is usually adequate, thus eliminating the need for finishing in a second operation. W or k pie ce Con figu r a t ion . The repetitive shape of the screw thread frequently permits the use of one-pass plunge

grinding, producing (by an essentially single turn of the workpiece) the entire threaded section. In principle, this is similar to the operation of the multirib thread milling cutter, but it is generally accomplished with a much higher productivity and with superior accuracy. Su bse qu e n t M a ch in in g of t h e W or k pie ce . Threads or similar profiles can be produced along a helical path on a

basically cylindrical part with ribs having eccentric surfaces. The resulting surfaces are thus relieved with respect to the leading edges, which are located along the individual rib boundaries, and the configuration produced is highly beneficial to the effective operation of thread cutting tools, especially taps and multirib thread milling cutters. Th r e a d Gr in din g M a ch in e s The classification of thread grinding machines can be based on the degree of automation, distinguishing manually controlled, semiautomatic, automatic, and fully automatic thread grinders. The fully automatic classification comprises equipment that can operate essentially without attendance, performing even the loading and unloading of the work by automatic controls. Although the degree of automation is an important consideration in the selection of equipment intended for particular production purposes, thread grinding machines of widely different systems, capacities, and application purposes are available in designs providing different degrees of automation. A better method of classification is based on the principal application of the machine. The principal application is usually the governing aspect in the design of the various types of thread grinders. A survey of this classification is presented in Table 6. This classification is also a flexible one; overlaps will occasionally occur because of the expanded adaptability of the basic machine and particularly because of the use of optional accessories.

Table 6 Principal categories of thread grinding machines Category Universal thread grinding machines

Field of application Toolroom work for single pieces or limited volumes; general production of a wide range of workpieces, including special helical, annular, or even straight-line element shapes

Production-type thread grinding machines

Construction parts for precision machinery with threads or worms; thread plug gages and thread cutting tools; high-strength threaded fasteners, and so on, produced in medium lots or large volumes

Fully automatic tap grinders

Taps are typical, but not exclusive, examples of small-size threaded parts that, by their overall dimensions, permit automatic work handling and are required in volumes warranting equipment for completely automatic operation.

Thread grinding machines for long workpieces

Single-setup grinding of very long, threaded machine components, particularly leadscrews for machine tools, feed and actuator screws, and ball screws with gothic-arch thread profiles

Thread grinding machines for internal threads

Production grinding of threads on internal surfaces; principal applications are thread ring gages, nuts of precision-type translation elements (including those with ball tracks and Acme threads), threading dies, and so on

Special-purpose thread grinding machines

Workpieces that, because of extreme size or a particular configuration, cannot be accommodated on standard thread grinding machines. Typical examples are parts requiring the grinding of special profiles or particular feature-generating movements.

General characteristics A particularly adaptable machine type, several models also operating with multirib wheels; wide tilt range of the wheelhead; capable of relief grinding. Can be supplied with many different accessories for extending the scope of application to include various profile types and work surface configurations This category comprises several models with different work-size capacities, generally ranging from about 150 to 300 mm (6 to 12 in.) maximum diameter and 200 to 1000 mm (8 to 40 in.) maximum length. Some of these models have change gears; others have interchangeable thread-generating elements, considering the infrequent pitch change needed in high-volume production. Many models can operate with single-rib or multirib wheels, and depending on the design, most or all phases of the grinding cycle are automated, with the variables adjustable to suit the operational requirements. The thread grinders for this type of application are manufactured in models using single-rib wheels in one-way or two-way grinding at multiple cuts and high work speeds and in models using multirib wheels usually operating in a plunge grinding system, with wheel dressing by crush rolls or diamond rolls. In either case, the operation is continuously automatic, including all the elements of the grinding cycle as well as the loading and positioning of the workpiece. Manufactured in different sizes for maximum grinding lengths of the order of 4 m (12 ft). Some models have hollow work spindles to accommodate parts longer than the grinding capacity; usually operated in a two-way grinding cycle; designed to reduce the extent of, or to compensate for, thermal effects; equipped with various types of wheel-dressing devices; accept special steady rests for providing intermediate support to long and thin workpieces which are prone to deflection Usually designed with entirely automatic grinding cycle, operating in different modes, such as one-way or two-way grinding. Lead generation by interchangeable master screws and nuts; supplied with interchangeable grinding spindles for various speed ranges. Grinds right- or left-hand threads, as well as annular profiles, and can be used for multistart threads Overall design and capacities adapted to the special purpose, although the operating principles and the methods of wheel dressing are often similar to those of general-purpose thread grinders. This category also includes such unconventional designs as the centerless thread grinding machines.

Source: Ref 2

Re fe r e n ce s 1. Machining, Vol 1, 4th ed., Tool and Manufacturing Engineers Handbook, Society of Manufacturing Engineers, 1983 2. F.T. Farago, Abrasive Methods Engineering, Vol 2, Industrial Press, 1980 3. Machining Data Handbook, Vol 2, 3rd ed., Metcut Research Associates, 1980

Th r e a d Rollin g I n t r odu ct ion THREAD ROLLING (also known as roll threading) is a cold forming process for producing threads or other helical or annular forms by rolling the impression of hardened steel dies into the surface of a cylindrical or conical blank. Polygonal blanks are also thread rolled for the purpose of fabricating thread-forming and self-locking screws. The preferred polygonal shape is trilobular and is produced in flat die machines. In contrast to thread cutting and thread grinding, thread rolling does not remove metal from the work blank. Rather, thread rolling dies displace the surface metal of the blank to form the roots and crests of a thread. Dies for threading rolling may be either flat or cylindrical (Fig. 1). Flat dies operate by a traversing motion. Methods that use cylindrical dies are classified as radial infeed, tangential feed, through feed, planetary, and internal. Each method is discussed in a separate section of this article.

Fig. 1 Two com m on t ypes of t hread rolling dies, flat and cylindrical

Ca pa bilit ie s a n d Lim it a t ion s Most thread rolling is done on blanks having a hardness of 32 HRC or less. However, threads on fasteners used for hightemperature service are rolled in metal as hard as 52 HRC. Some metal products such as gray iron castings and sintered metal pieces cannot be thread rolled because of their low ductility. These materials crumble rather than conform to the contour of the die. All of the commonly used straight and tapered thread forms can be rolled. These include Unified, International Standard (the same as UNR), metric, Whitworth, Acme, worm, buttress, screw shell, wood screw, tapping screw, lag screw, and drive screw. Thread diameters vary from less than 1.25 mm (0.050 in.), for instrument threads, to 380 mm (15 in.); the larger threads can be as long as 6 m (20 ft). Thread forms of 60° roll easily. In more blunt forms, metal flows with greater difficulty. Threads with fully rounded roots are more easily rolled than threads with wide, flat roots. Flank angles of less than 10° included angle and thread depths exceeding one-sixth of the major diameter should be avoided, except with the most ductile metals. For multiple threads, a thread depth of one-fourth the major diameter for double and quadruple lead threads, and one-fifth the major diameter for triple lead threads, is generally acceptable.

The rolling process can also accomplish many nonthreading operations, such as the rolling of splines, helical and annular grooves, knurls, and involute teeth. Rolling may also be used for burnishing and for displacing metal to form flanges and similar cylindrical shapes. Su r fa ce Fin ish . When properly made, thread rolling dies impart smooth, burnished roots and flanks to threads. Rolled

threads are free of tears, chatter marks, or cutting-tool marks common to cut threads. Such imperfections nucleate wear and can serve as starting points for fatigue failure. Surface roughness on rolled threads is usually from 0.20 to 0.60 m (8 to 24 in.), whereas on cut threads it is often 1.63 to 3.18 m (64 to 125 in.) with small ridges or unevenness along the flanks of the thread. However, because the surface finish of thread flanks is extremely difficult to check, it is rarely specified on drawings. In general, the coefficient of friction of a rolled thread surface in sliding contact is considerably lower than that of a comparable cut thread surface. The coefficient of friction between the thread and its mating nut determines how effectively a bolt can be tightened or a moving screw can transmit power. Therefore, the relatively low coefficient of friction of rolled thread surfaces provides more uniform and consistent tightening of fasteners and less loss of power in overcoming friction when a load is moved by a screw. The reduced and more uniform friction of rolled threads also contributes to the torque control of threaded connections with self-locking features. The smoother finish of rolled threads also retards corrosion. St r e n gt h of Rolle d Th r e a ds. Thread rolling deforms the blank plastically as it is forced to flow along the contour

imposed by the dies. The worked metal is appreciably harder and stronger than the blank prior to rolling. Thus, fasteners with rolled threads are harder and stronger than those with cut threads, as indicated in Table 1.

Table 1 Average mechanical properties of hexagonal-head capscrews with rolled and cut threads Data are based on 15 pieces of each size and each threading method, made of 4027 steel. Screw size and pitch

-9 1-8 1

(a) (b)

Type of thread

Core hardness(a), HRB Shank Thread area

Tensile strength MPa ksi

Rolled Cut Rolled Cut Rolled

82 82 91 91 91

631.2 489.2 678.1 630.5 710.9

92 82 94.5 91 96.5

91.55 70.95 98.35 91.45 103.1

Fatigue life, cycles to failure(b),×103 71.8 14.3 51.8 21.3 68.5

-7 Core hardness refers to the hardness at the area of the centerline of the thread below the pitch diameter. Converted from Knoop readings. Fatigue test was tension-tension, at 415 MPa (60 ksi) using a preload equal to 10% of maximum.

Heating thread rolled fasteners made of steel above the transformation range during a heat-treating process completely relieves the favorable compressive stresses induced by rolling. Therefore, steel fasteners produced by the thread rolling of blanks already quenched and tempered to a given hardness usually have higher fatigue strength than fasteners quenched and tempered to the same hardness after thread rolling. Figure 2 shows this difference for the most common range of hardness. The different effects of hardness on fatigue strength of bolts that were roll threaded before and after heat treatment is illustrated in Fig. 3.

Fig. 2 Effect of rolling/ heat - t reat ing sequence on t he fat igue st rengt h of 16 m m ( in.) diam 50B40 st eel bolt s. ( a) Four different lot s t hat were rolled before being t em pered t o an average hardness of 22.7, 26.6, 27.6, and 32.6 HRC. ( b) Five different lot s t hat were rolled aft er being t em pered t o an average hardness of 23.3, 27.4, 29.6, 31.7, and 33 HRC. Harder bolt s in bot h ( a) and ( b) had fat igue st rengt h t oward t he high side of t he ranges shown.

Fig. 3 Effect of hardness on fat igue st rengt h for bolt s wit h t hreads rolled before or aft er heat t reat m ent

Eva lu a t ion of M e t a ls for Th r e a d Rollin g The three characteristics that are important in evaluating and selecting metals for thread rolling are rollability, flaking, and seaming.

Rolla bilit y involves ductility and the resistance of a metal to flow when subjected to cold forming in thread rolling dies.

Rollability indexes for 17 steels and for 6 nonferrous alloys commonly threaded are given in Table 2. The power required to form a given thread shape at a given rate in various metals is inversely proportional to the rollability indexes of the metals. If material in an operation is changed for one of lower rollability index, the production rate per horsepower for a rolled form in that operation decreases. For example, if a through-feed machine using its full 7.5 kW (10 hp) output produces -13 UNC-2A threads at a rate of 11.5 m/min (450 in./min) in a solid bar of steel with an index of 1.00, the rate will be only 6.9 m/min (270 in./min) when a steel of index 0.60 is threaded in the same machine under the same operating conditions.

Table 2 Rollability of alloys Metal Carbon and low-alloy steels 1010 1018 1020 1095 1095 1112 1117 1144 4140 4140 4140 4340 8620 Stainless steels 303 316 416 430 Nonferrous alloys Aluminum, 2017 and 2024 Brass Cartridge Naval Phosphor bronze Monel

(a) (b) (c)

Hardness, HB

Rollability index(a)

Flaking tendency(b)

Seaming tendency(c)

137 148 156 260 320 198 173 225 205 234 300 235 215

1.11 1.08 0.96 0.47 0.42 1.00 1.03 0.78 0.93 0.57 0.42 0.45 0.60

B B B B B C C B B B B B B

C C C B B C C C C C B B C

174 150 221 225

0.46 0.45 0.58 0.56

C B C C

B B B B

135

1.40

B

C

190 155 130 235

1.55 1.00 1.28 0.93

B C C B

B B B B

Index applies to metals rolled at room temperature. B, minor susceptibility; C, strong susceptibility. B, negligible susceptibility; C, moderate susceptibility. See Fig. 5, which also indicates the two extremes of seaming tendency (A and D).

The rollability index also provides a means of comparing radial die loads and expected die life in rolling two materials under identical operating conditions. The radial die loads required for roll threading various metals are approximately inversely proportional to the rollability indexes of the metals. Die life is approximately proportional to the third or fourth power of the indexes if final die failure is due to crumbling of thread crests of the die. Thus, if a die life of 30,000 m (100,000 ft) of threaded rod is obtained when rolling steel having an index of 1.00, a die life of about 1800 to 3700 m (6000 to 12,000 ft) can be expected when steel with an index of 0.5 is threaded under the same conditions. Fla k in g is related to the shear strength of the metal being rolled. Lead and sulfur in brass and steel increase susceptibility

to flaking during rolling. An increase in the carbon content of steel decreases susceptibility to flaking. In general, flaking increases directly with the amount of previous cold working of the blank material. This is true of almost all rollable metals, and especially of the work-hardening alloys such as series 300 stainless steel, copper, and some aluminum alloys. Annealing prior to rolling reduces flaking.

Work metals may be classified into four groups with respect to susceptibility to flaking: • • • •

Group A: Little or no susceptibility, regardless of whether or not the material was previously cold worked, or regardless of the bluntness of form to be rolled Group B: Minor susceptibility Group C: Strong susceptibility Group D: Excessive susceptibility, which precludes the rolling of all but the most simple, shallow forms

Table 2 indicates susceptibility to flaking for metals most commonly thread rolled. As indicated, all metals listed in Table 2 have either minor or strong susceptibility to flaking. Not many metals can be classified as having little or no susceptibility to flaking. Copper and some of the extremely ductile copper alloys when in the annealed condition are sometimes given the A rating. Metals that contain excessive amounts of free-machining additives, such as the specially prepared screw-machine steels, are likely to fall within the D class. Also, some metals that work harden at an excessive rate (some of the stainless steels and the less-ductile heat-resisting alloys) are likely to fall in the D class. Se a m in g. If, during rolling, the work metal flows up the flanks of the die threads faster than it does at the center of the thread form, the displaced metal may fold together to form a seam as the metal fills the full crest of the thread form, as shown in Fig. 4. The formation of seams, or folds, depends first on the metal being rolled, and second on the shape of the thread form.

Fig. 4 Seam at crest of t hread caused by fast er m et al flow along flanks of die t hread

Open seams in the thread crests may occur when undersize blanks are rolled. The open seams can shorten the service life of the thread in a corrosive environment, although they are not usually detrimental in normal service, in which corrosion is less important. The softer, more ductile metals usually form deeper seams than the harder, less ductile metals. Figure 5 shows the types of metal flow associated with four degrees of seaming. Table 2 indicates susceptibility to seaming for specific alloys commonly used for thread rolling stock.

Fig. 5 Four degrees of suscept ibilit y t o seam ing during t hread rolling. ( A) Negat ive suscept ibilit y t o form seam s. Met al flow adj acent t o t he die surface is slower t han in t he m iddle of t he roll form . This is charact erist ic of m et als having a relat ively high coefficient of frict ion wit h t he die st eel. ( B) Negligible suscept ibilit y t o form seam s. Met al flows up wit h an alm ost flat t op during rolling in convent ional t hread form s. ( C) Moderat e suscept ibilit y t o form seam s, t ypical of low- carbon st eels. ( D) Excessive suscept ibilit y t o form seam s. Cavit y is likely t o be form ed under crest of t hread.

Pr e pa r a t ion a n d Fe e din g of W or k Bla n k s The diameter of the blank to be threaded is between the major and minor diameters of the thread to be rolled, as shown in Fig. 6. It is common practice to produce blanks for rolling Unified Standard threads with tolerances greater than 0.05 mm (0.002 in.) by extruding, by cold heading, or by shaving on automatic machines. Most of the class 3A threads in the sizes generally used have pitch diameter tolerances greater than 0.05 mm (0.002 in.). Therefore, they can be rolled on extruded, cold-headed, or shaved blanks. Some of the smaller class 3A threads have tolerances closer than 0.05 mm (0.002 in.). Blanks for these threads should be ground.

Fig. 6 Relat ion of blank diam et er t o m aj or and m inor diam et ers of t hreads

To produce threads having a pitch diameter within a tolerance of 0.05 mm (0.002 in.), the tolerance of the blank diameter should be within 0.013 mm (0.0005 in.). Closer thread tolerances can increase the cost of blank preparation, sometimes far beyond the usual costs. Blanks with close-tolerance diameters are ground. Blanks of material not suitable for extruding or cold heading, such as titanium and some stainless steels, are also ground.

Blank diameter must be within the tolerance required for the particular size and class of thread specified. It is not practical to roll threads to a close tolerance except on blanks held to appropriate diameter. Over-rolling an undersize blank to provide a screw of correct size causes premature die failure. Maximum die life is obtained when the crest is not rolled full. When rolling a class 2A thread, for example, the most economical procedure is to use a blank with a diameter such that the thread can be rolled to the mean class 2A pitch diameter (halfway between the high and low limit of pitch diameter) and to maintain the major diameter of the thread just above the lower tolerance limit. Re la t ion of Bla n k a n d Pit ch D ia m e t e r s. The thread form is said to be balanced when the volume of cavity below the pitch line (Fig. 7) is equal to the volume of metal above the pitch line. (Pitch line is defined as the location at which the widths of the thread ridge and the thread groove are equal.) In such threads, the correct blank diameter is substantially the same as the pitch diameter. With metals that are commonly cold headed, the maximum blank diameter is generally equal to the mean pitch diameter. The optimum blank diameter varies with different work metals; some adjustment may be needed to get the desired crest formation. Additional dimensional allowance is necessary if subsequent plating is planned. If the form to be rolled is unbalanced (Fig. 7), the blank diameter is not the same as the pitch diameter.

Fig. 7 Relat ion of blank diam et er t o pit ch line for balanced and unbalanced t hread form s.

Be ve l on Bla n k s. Ground and extruded blanks are made with a bevel at the end or ends of the section to be threaded.

This bevel ranges from 15° to 45°. Blanks for the majority of thread rolled products have a bevel angle of 30°. Thread rolling, however, increases the bevel by 15° to 30°. Therefore, the bevel angle on the blank must be less than the angle desired on the finished part. In addition to the bevel angle, some blanks have a bevel in the area that connects the section to be threaded with the section that will remain unthreaded. This bevel, sometimes called the extruding angle, is usually 30°. Fe e din g. The various thread rolling methods employ three basic techniques for feeding the blank into the dies:

• • •

Radial infeed: The die, usually cylindrical, moves in a radial line directly toward the axis of rotation of the workpiece Tangential feed: The die, either cylindrical or flat, moves past the workpiece on a path that brings the pitch line of the thread form tangent to the work surface End feed or through feed: The cylindrical die "tracks" on the workpiece causing the workpiece to move axially as it rotates

Fla t D ie Rollin g

Flat traversing dies are the type most commonly used for rolling threads in commercial fasteners and similar parts. One technique involves the use of two flat, rectangular dies; one is stationary, and the other traverses in a plane parallel to the stationary die and separated from it by a distance equal to the minor diameter of the thread to be rolled. In another technique, both dies traverse the workpiece. Figure 8 shows the most common arrangement of dies and workpiece. As the blank is forced into the space between the dies by a feed finger, it is engaged by the forward motion of the traversing die and caused to roll between the threaded faces of the dies. This action forms the thread.

Fig. 8 Operat ing principle of flat t raversing die t hread rolling

A thread is rolled on one blank at a time during the forward stroke of the machine. There is no appreciable axial movement of the blank during rolling. The diameter of the finished thread is controlled by the diameter of the blank and the distance between the faces of the dies at the end of the stroke. M a ch in e s. The generic term flat die rollers encompasses a large and varied family of machines that are made in a

number of sizes, each for a limited diameter range and with a specified die length. As a general rule, the correct machine size accommodates a die length that allows the blank to complete from six to eight full revolutions. Most flat die rollers have the dies in the side-by-side position shown in Fig. 8. Most of the newer machines also have dies that traverse in the horizontal plane, but with the faces of the dies at an angle to the vertical. The feed track is at an angle, thereby permitting gravity feed down an incline into the inserting mechanism. Specialized types of vertical and horizontal rolling machines also are available. One horizontal traversing unit has dies placed one above the other rather than in the usual side-by-side position. It is used to roll splines and related forms as well as threads. Another variation of the horizontal traversing machine uses an inclined feed chute at right angles to, rather than parallel with, the axis of traverse. This type of machine incorporates a different method of die alignment. Dies are adjusted longitudinally to match the threads of the two dies, rather than vertically by shims, as in most other machines. D ie s used in flat die rollers consist of matching pairs of rectangular plates, with each of the opposed faces having a

reverse image of the form to be produced on the part. Dies are made in various widths and are used to roll screws of any thread length up to the maximum die capacity.

The lead angle of flat dies can theoretically vary from 0° to 45°, but for producing most standard screw threads, it is less than 5°. (Lead angle of the die is defined as the angle between the thread form and the longitudinal axis of the die.) In flat die design, penetration rate is primarily governed by the length of the die. Best practice calls for complete penetration prior to the last revolution of the workpiece in the die. The last complete revolution of the workpiece should only iron out small irregularities. It is important that the die be long enough to prevent an excessive penetration rate. The above principles can be used to roll more than one form on a part by use of a multiple stack of dies, which are inserted one atop the other in the machine and held together with a clamp. Thus, many combinations of forms that would be impossible to generate on a one-piece die can be easily produced on a multiple die, provided that the spread between diameters of the individual forms to be rolled on the same part is minimal. In addition to the dies, secondary tooling is required for feeding, sorting, orienting, and inserting the parts between the dies. Ca pa bilit ie s. The flat-die process is commonly used for all types of straight- and taper-threaded commercial fasteners.

Flat die rolling can produce more than one form in one operation, such as two entirely different types of threads at opposite ends of a part, knurling and a thread, or knurling and an annular groove, on the same part. Duplex face dies can be used for rolling straight threads. Such dies have threads on both the front and back sides so that they provide two rolling surfaces. When the screw length is less than half the die width, the die can be reversed, top for bottom, so that four rolling edges are available for still greater economy. Production rates vary widely and usually are inversely proportional to the size of the product. The small machines are capable of producing parts at a rate of 10,000 to 36,000 per hour. Larger units, producing 9.5 or 13 mm ( can roll 3000 to 12,000 pieces per hour. Products such as 32 mm (1 rates, ranging from 900 to 3000 pieces per hour.

or

in.) bolts,

in.) diam bolts are thread rolled at much slower

Lim it a t ion s. In general, flat dies are used for threading metals no harder than 32 HRC before rolling, although steel as

hard as 52 HRC can be roll threaded. Thread diameters are commonly limited to 25 mm (1 in.), although a few machines can roll up to 38 mm (1 in.) diam threads. Thread lengths up to 265 mm (10 in.) are rolled. These limitations-hardness, diameter, and length--are interrelated so that a workpiece having more than one or two of these measurements near maximum value may not be rollable. The flat-die method is also limited to parts of an overall size that can be accommodated in the machine. Because of interference between the part and elements of the machine and die, part size as well as thread dimensions must be considered before the flat die method is selected for a specific piece.

Ra dia l- I n fe e d Rollin g Radial-infeed thread rolling consists of moving a rotating cylindrical die or dies radially toward the center of the rotating workpiece. The operating principle of this method is shown in Fig. 9.

Fig. 9 Operat ing principle of radial- infeed cylindrical- die t hread rolling

A minimum amount of axial movement between the dies and the workpiece occurs during the rolling cycle. This characteristic distinguishes infeed rolling from the through-feed method of cylindrical-die thread rolling. Axial movement is canceled by designing the die with an effective lead angle equal in magnitude but opposite in direction to that on the -32 double-lead worm thread, movement would be as much as 4.8 mm ( in.) during the rolling work. In rolling a cycle if there were no compensating lead on the die. Axial movement does not affect thread quality, but it may restrict the ability to produce a full thread close to a shoulder and will reduce the amount of full thread that can be produced with a special die face. The effective lead angle varies slightly during die penetration so that some axial movement does occur. The amount of movement is usually insignificant when rolling standard threads, but can be considerable when rolling blunt or very deep thread forms, or those with large lead angles. Dies can be designed to give slight axial movement to the blank to increase die life or to simplify regrinding of the dies. Movement of 3.2 to 6.4 mm ( applications.

to

in.) is common; movement of up to 13 mm (

in.) has advantages in some

When two or three dies are used, they must be matched so that the helical path produced by one die is a continuation of that produced by the other die or dies; otherwise, there will be steps in the product thread (Fig. 10). Dies are matched by rotating one or more dies in relation to the others, or by moving one or more dies axially, to produce a continuous helix on the work.

Fig. 1 0 Effect of m ism at ched and correct ly m at ched dies on t hread im pression m ade by cylindrical dies

Cylin dr ica l- die m a ch in e s capable of infeed thread rolling are equipped with either two or three dies (Fig. 9). Two-die

machines are usually of the horizontal type; that is, the workpiece is horizontal during rolling. Three-die machines are available in both horizontal and vertical models. In two-die machines, either one or both dies can move radially. If one die does not move radially, the work must move radially during die penetration. If both dies move equally, the work can stay in place. Most three-die machines provide equal radial movement of all dies so that the work position does not change during rolling. A few three-die machines have one or two stationary dies, and radial movement of the work must be allowed for. Two-die machines require a work support to position the centerline of the work slightly below (usually about 0.25 mm, or 0.010 in.) the same plane as the centerline of the dies. This offset prevents the workpiece from rising out of the dies. A work rest, as shown in the two-die machine of Fig. 9, can be used for short, manually loaded parts. Larger parts may require additional supports. In many instances, the parts can be inserted in a tube or bushing, or held between centers, to provide proper positioning for rolling. A spring-loaded work stop can be used for positioning the part in proper axial location. The dies in a three-die machine serve to locate the part so that often the only fixture required for manual loading is a spring-loaded work stop to provide correct axial position. La t h e s a n d Au t om a t ic M a ch in e s. Radial infeed rolling can be done in lathes and automatic bar machines equipped

with single-roll or double-roll radial threading attachments. In these attachments, the cylindrical dies are commonly called thread rolls. Figure 11 shows the operating principle. The single-roll attachment is a simple roll or knurl holder mounted on a cross slide. As the cross slide advances, the roll is pressed into the workpiece so that it rotates with the work and thus forms the thread. The travel of the cross slide is controlled so that the thread roll in its final position produces a thread of correct size. After the cross slide completes its full length of travel, it is rapidly retracted.

Fig. 1 1 Operat ing principle of t wo t ypes of radial at t achm ent s for t hread rolling on lat hes and aut om at ic bar m achines

Double-roll radial attachments operate by means of a toggle arrangement that causes the rolls to close approximately radially to contact the rotating work, at which time the rolls begin to turn to form the thread. After reaching full depth, the rolls and attachment are rapidly retracted. D ie s ( Rolls) . A little axial movement occurs between the work and the cylindrical dies, and it is necessary that the

threaded length of the dies be about two to three threads per product lead longer than the length of the thread to be rolled. In practice, it is desirable to provide a bevel at each end of the die (Fig. 12); thus, the width of the die face must exceed the length of the thread to be rolled by twice the width of a bevel.

Fig. 1 2 Bevel on a cylindrical t hread rolling die

Long bevels with a small bevel angle increase die life by reducing breakage at the edge of the die. Usually 30° bevels are recommended, although workpiece requirements may necessitate 45°. For the harder workpieces, bevel angles of less than 30° are desirable; they may need to be as small as 15° for the hardest rollable metals. D ie Size . For infeed rolling, the pitch diameter of the die must be a multiple of the pitch diameter of the finished workpiece. For single-lead threads, the number of thread starts is equal to the ratio of die-to-work pitch diameters. For

rolling multiple-lead threads, the number of die thread starts equals the number of thread starts on the work multiplied by the ratio of die-to-work pitch diameters. The diameter of dies for two-die machines is limited by the size of the machine and fixtures and is approximately constant regardless of work diameter. For instance, a typical two-die machine having the capacity for 1.6 to 38 mm ( to 1 in.) workpieces uses dies that are 125 to 150 mm (5 to 6 in.) in diameter for all sizes of work. In general, but not necessarily, two-die machines use dies of larger diameter than three-die machines for threading the same size of workpiece. Three-die machines generally cannot use dies larger than about five times the work diameter, because larger dies will contact each other before reaching full thread depth in the work. Slightly larger dies can be used for rolling multiple-lead threads. The size of thread rolls for radial infeed lathe attachments is determined by the dimensions of the attachment being used, rather than by the work diameter. Su ppor t s for D ie Spin dle s. Because of the limitation on the die-to-work diameter, three-die machines require a series

of spindles and spindle supports, graduated in size, to accommodate the entire work-diameter capacity range. For smalldiameter work, the spindles are necessarily quite slender and may not be strong enough to roll hard alloys or long thread lengths. Ca pa bilit ie s. The minimum practical diameter of a workpiece for rolling in two-die machines or attachments is 1.3 mm

(0.050 in.). The maximum diameter is limited only by the capacity of available equipment. Two-die machines capable of rolling threads 380 mm (15 in.) in diameter and 400 mm (16 in.) long are in use. Three-die machines roll threads from 6.4 to 115 mm ( seldom practical for rolling threads smaller than 6.4 mm (

to 4

in.) in diameter and up to 125 mm (5 in.) long. They are

in.) in diameter.

The rate of die penetration into the work is adjustable; thus metals of various hardnesses can be threaded. Most metals threaded by cylindrical dies or attachments are no harder than about 32 HRC. However, work metals as hard as 52 HRC have been thread rolled on cylindrical dies. Die life does deteriorate rapidly when material harder than 32 HRC is rolled. The versatility of the radial infeed method makes possible the rolling of thin-wall parts such as tubing or stampings and also some metals harder than 48 HRC that would be difficult or impossible on other types of machine. Rolling of thinwall parts is discussed in a subsequent section of this article. The radial infeed method permits rolling of threads close to shoulders, with a minimum of imperfect threads. Also, threads can be rolled between two sections of larger diameter, as is common when rolling worms on large transmission shafts. Threading close to shoulders is considered in the section "Rolling Threads Close to Shoulders" in this article. The three-point support provided by three-die machines is advantageous for rolling parts with irregular or unbalanced overhangs and parts requiring a thread length considerably shorter than one diameter. Often such parts can be rolled while being supported only by the three dies, whereas cylindrical two-die machines or other types of threading equipment may require expensive or unwieldy fixturing. Radial-feeding single-roll or double-roll attachments are used to best advantage for rolling threads at the collet end of pieces being machined in a lathe or automatic bar machine (Fig. 13). Such threads are usually behind a shoulder making radial infeed and tangential feed rolling (discussed below) the only practical thread rolling methods.

Fig. 1 3 Use of a double- roll at t achm ent for t hread rolling near t he collet of an aut om at ic bar m achine

Because double-roll attachments exert a minimum of transverse pressure on the work, small diameters of considerable length can be rolled, and at a much greater distance from the collet than with the single-roll type of attachment. For example, a double-roll attachment can roll a -20 UNF, 19 mm ( in.) long thread on a 13 mm ( in.) diam bar at a distance of 25 mm (1 in.) from the collet to the first thread. With a single-roll attachment, maximum distance from the collet is 6.4 mm (

in.).

Production capabilities of cylindrical-die infeed machines vary according to the method and equipment for feeding the work and the type of work involved. In general, hand feeding is not practical at rates above 25 pieces per minute. Automatic feeding equipment can be installed on two-die or three-die machines to produce up to about 90 pieces per minute. Typical applications of cylindrical-die infeed thread rolling are shown in Table 3.

Table 3 Typical applications of cylindrical-die thread rolling Product

Steel

Infeed rolling Adjusting worm(a)

8620

Armature shaft (worm)(b) Armature shaft (worm)(c) Double-end stud

1045

Feed screw(e)

410

Feed screw(a)

410

Worm Worm

4140 8620

Through-feed rolling Threaded rod

1018

Jackscrew(g)

1018

Setscrew

1040 1018

(h)

Double-end stud

(a) (b) (c) (d) (e) (f) (g) (h) (i) (j) (k)

-10 1.7 mm (0.065-in.) pitch 0.2805-32 1 -8 UN-3A 0.330-56 Acme

1335 or 1041(i) 3135(i)

Length mm in.

Insertion

Die speed, rev/min

Rate, pieces/min

Die life, pieces

29

...

85

8-10

...

Manual

45

...

6,000

-33 Acme LH Buttress(f) -5; 4 starts

...

...

Manual

...

6

...

51

2

Manual

...

4(d)

...

...

170

8-10

...

...

...

10

...

Manual Manual

85 ...

8-10 8-10

... 20,000

Automatic

500

6

260,000

56 ... 38 1800

-16 UNC -10 Acme stub

-20 UNF-2A

1

19

95

-19 BSP F(h)

Automatic continuous rolling 8115 Automotive stud Double-end stud

Thread

210

3 2 ... 1 72

Manual

260

...

25,000

30

8 1.2

Hopper

...

300

...

16.3

0.640

Hopper

...

125

500,000

(j)

Hopper

...

200(d)

...

(k)

Hopper

...

80(d)

...

-24 UNF-3A(j) -11 UNC

(k)

Pressure angle, 14 °. Pressure angle, 20°. Pressure angle, 25°; two leads on thread. Both ends. Pressure angle, 29°. Modified buttress; 2 starts; 11 threads per inch; flanks 10° and 30°. Infeed and through-feed rolling. British Standard Parallel Fastener; made of resulfurized screw stock. Cold drawn, cut to length, and ends extruded to blank diameter. Thread length, 14 mm (

in.) on each end of 117 mm (4

One end class 3A, 29 mm (1 70 mm (2

in.) long stud.

in.) long; the other end, class 5A, 25 mm (1 in.) long; total length of stud,

in.)

Lim it a t ion s. The axial travel caused by thread rolling machines when they are producing deep, blunt thread forms or

high lead angles is not objectionable except when it causes interference during rolling of a thread close to a shoulder. In single-roll or double-roll lathe attachments, the work cannot move axially, and roll movement is usually limited to approximately one-third of the pitch so that rolls must be designed to cause a minimum of axial travel. The rolling of deep, blunt forms or high lead angles requires careful attention to the roll design and also to the rate of penetration of the rolls, the accuracy of the setup, and the condition of the attachment.

The accuracy of threads rolled by single-roll infeed attachments depends on the accuracy of the cross-slide travel. Thread accuracy is limited also because of the bending action developed by the force of radial-infeed rolling. In some applications, back-up rolls, bearing on a plain cylindrical surface outside the threaded area, can be used to reduce the bending. If a plain surface is not available or machine tooling does not permit the use of backup rolls, single-roll attachments are restricted to thread rolling near the collet and to short thread lengths in the softer materials.

Ta n ge n t ia l Rollin g Tangential thread rolling is similar to infeed rolling except that the dies (rolls) are fed past the blank on a path parallel to the radial path at a distance such that when the axis of the roll is opposite the axis of the blank the pitchline of the thread form is tangent with the surface of the blank. Figure 14 shows the operating principle of this process. As the rolls advance, they reach maximum penetration when the centerline of the rolls is directly opposite the centerline of the work. The total depth of penetration is determined by the amount the rolls are offset in relation to the work. As in radial-infeed rolling, only a slight axial movement occurs between the rolls and the work.

Fig. 1 4 Operat ing principle of t angent ial- feed t hread rolling

M a ch in e s. Tangential rolling is done in lathes or automatic bar machines equipped with one-roll or two-roll attachments

mounted on a cross slide of the machine. The rolls are rotated by their contact with the rotating work. Two-roll attachments are the most common for tangential rolling. They are available in various sizes; each size has a capacity for a range of work diameters. Capacity up to a work diameter of 65 mm (2 obtainable for special applications.

in.) is commonly available, and larger sizes are

Ca pa bilit ie s. Tangential feeding attachments have essentially the same capabilities as radial feeding attachments. Some

advantage is gained with tangential attachments because the adjustments and control for a given size of work are made within the attachment rather than by the travel of the machine slide. With a two-roll tangential attachment, no radial movement occurs between the roll spindles during rolling, and rolling pressures are greater than in radial-infeed rolling. Threads can be rolled at spindle speeds compatible with other machining operations; therefore, speed changes for threading are unnecessary. Two-roll tangential rolling produces bending loads somewhat higher than the two-roll radialfeeding attachment, but low enough to allow the rolling of threads on relatively hard work metals at a considerable distance from the collet. For instance, a 19 mm ( hardened to 30 HRC, at a distance of 16 mm (

in.) long

-20 UNF thread can be rolled on a 13 mm (

in.) diam bar

in.) from the collet to the first thread.

Lim it a t ion s. Thread rolling with a two-roll tangential attachment is limited primarily by the size and capacity of the

equipment. Rolling threads with high lead angles or with deep, blunt forms can be trouble-some because of axial travel. However, with proper attention to roll design and setup, using only attachments in very good condition, the problem can usually be overcome. Single-roll tangential rolling produces transverse loads capable of bending the workpiece. Therefore, this procedure is limited to rolling threads of short length near the collet in soft materials such as nonferrous alloys or soft steel (generally

no greater than 197 HB). In rolling a

-20 UNF thread next to the collet on a 13 mm (

maximum practical thread length would be approximately 22 mm ( rolls can sometimes be used to increase this capacity.

in.) diam bar of 1112 steel, the

in.). As with radial single-roll attachments, support

Th r ou gh - Fe e d Rollin g In through-feed rolling, the work moves axially through the dies. Through-feed dies are designed with a lead angle generally different from that of the work, so that the part can feed. The dies are made with a starting taper, so that the thread is formed progressively as the blank feeds through the dies. The finish end of the dies also is tapered slightly so that rolling pressure is released gradually without marking the work. Fe e d r a t e , in terms of feed per work revolution, is proportional to the ratio of the difference in lead angles of the dies to

the lead angle of the work. The die lead angle can be either greater or less than that of the work and may be zero (annular form dies). Also, the die lead angle can be the same hand as the work (right-hand lead dies to produce right-hand threads). M a ch in e s. Any cylindrical-die thread rolling machine is capable of through-feed rolling, but the capacity may be

restricted. Vertical three-die machines, for example, can feed only short lengths because the gearbox or other equipment located a short distance below the dies prevents passage of longer work. In three-die horizontal machines, a passage through the gearbox usually allows unlimited length, but restricts the diameter of long work. Obstructions behind the dies of some two-die machines limit the length or diameter of the work. Machines having no provision for skewing the spindles are limited to parallel-axis through-feed rolling, for which the feed rates are low. Most machines, of both two-die and three-die types, have infinitely variable die tilt adjustments up to 10°. This permits the use of annular dies, which are less expensive and more productive than helical dies. Most conventional cylindrical thread rolling machines have one or all dies mounted on slides or pivoting arms for infeed rolling. For through-feed roll threading of the full length of a blank, the dies are held in the closed position by a hydraulic or mechanical system. Such machines can also be used for rolling threads on only a portion of the blank, either by feeding through until the desired length has been threaded, and then opening the dies, or by inserting the work between the open dies to the correct position, and then closing the dies so that the work feeds out of the dies. Some machines designed specifically for through-feed rolling of the entire workpiece are not equipped with die advance-retract mechanisms that permit partial-length rolling. En d- Fe e din g At t a ch m e n t s. Cylindrical through-feed rolling can be done with two-die and three-die end-feeding

heads. Two-die heads are used for small work. The rolling heads can be mounted on the rotating headstock of a boltthreading machine in which the blank is clamped to a slide that advances the work toward the head. When used in an automatic bar or chucking machine, the head is mounted on the tool slide, which advances it toward the rotating workpiece. The dies are made with annular grooves, and the axes of the dies are set at an angle with the work axis equal to the required lead angle of the product thread. The rate of feed per revolution is equal to the lead of the thread being rolled. Some heads have interchangeable frontplate units that can accommodate a range of lead angles. In other heads, lead angles are varied by interchangeable bushings. End-feeding attachments are adjustable to produce correct thread size. With the dies locked in the closed position, they engage the blank and roll the thread as they pass over the blank or the blank passes between the dies. When the desired length of thread has been rolled, the head opens, and the work is withdrawn from the head. When used for continuous rolling, the head remains in the closed position at all times. D ie s for through-feed rolling are usually relieved at both ends. Through-feed dies for Acme, worm, or other wide threads

often have a modified, pointed thread form at the starting end for efficient penetration into the blank. The number of thread starts, which, together with the diameter of the die, determines the lead angle, is different for through-feed rolling on cylindrical machines than for infeed rolling a similar size thread. Compared to infeed dies of similar diameter for a specific thread, the through-feed dies may have more or fewer starts for parallel-axis rolling, or no thread starts (annular form) for skewed-axis rolling.

Ca pa bilit ie s. Through-feed rolling is applicable to threading the full length of a cylindrical part of uniform diameter and

to threading one or more sections of the largest diameter of a multiple-diameter cylindrical part. End-feeding heads are used for straight and tapered threads. Annular rings can be produced by through-feed rolling in skewed-axis dies or rolls. Typical examples of parts made by through-feed rolling are commercially threaded rod, high-strength studs, headless set screws, threaded mounting tubes for electrical fittings, pole line hardware, recirculating ball screws, and jackscrews of all types. Threads are through-feed rolled to partial length on parts such as compressor studs, large-diameter cap screws, clamp and jackscrews, finned heat-exchanger tubing, and reinforcing rods. Three specific applications of through-feed rolling are given in Table 3. Most thread rolling machines and heads except the three-die vertical types are virtually unrestricted in the length of bar that can be threaded. Mill length bars or tubes 3 to 5 m (10 to 16 ft) long are commonly threaded. Heat exchanger tubing is through-feed rolled to produce integral fins in lengths up to 15 m (50 ft). Sometimes it is economical to through feed parts of short length. For example, blanks for socket set screws and various types of studs are thread rolled at high production rates. Rods or parts up to 6.4 mm ( in.) in diameter can be through feed rolled at speeds up to 20 m/min (800 in./min) depending on the type of machine, available horse-power, and hardness of the blanks. Larger sizes will feed slower; 300 mm/min (12 in./min) is a typical speed for 75 mm (3 in.) diam steel parts at a hardness of 22 HRC. In many applications, a larger thread can be rolled in a given machine by partial-length through-feed rolling than by infeed rolling. Threads of very small diameter, such as 0-80 UNF, can be through-feed rolled on two-die machines or endfeeding heads. Three-die machines usually cannot roll threads smaller than about 6.4 mm ( die interference. Machines are available for 115 mm (4

in.) in diameter because of

in.) diam work up to about 0.6 m (2 ft) long.

Standard end-feeding attachments are available for threads up to 230 mm (9 in.) in diameter. Feed rates depend on the maximum speed at which the head or work can be rotated with the available horsepower. Feed rates of 7.6 m/min (300 in./min) are common for thread sizes up to 16 mm ( in.) in diameter. Most standard heads have clearance holes through the center so that long thread lengths can be rolled. The maximum length depends on the ability of the equipment to grip the workpiece to prevent rotation or excessive torsional windup of the piece. All of the common thread forms can be through-feed rolled, including blunt forms such as Acme and worm threads and ball-screw forms. Threads of very blunt form are produced with less difficulty by through-feed rolling than by infeed rolling. Lim it a t ion s. Like other types of cylindrical rolling, the process is limited primarily by the characteristics of the

equipment being used.

Pla n e t a r y Th r e a d Rollin g Planetary thread rolling machines have one central rotating die on a fixed axis and one or more stationary concave segment dies located near the outside of the rotary die, as shown in Fig. 15. One or several blanks may be rolled on a segment die at one time, depending on the gearing of the starting mechanism.

Fig. 1 5 Operat ing principle of planet ary t hread rolling. More t han one segm ent die m ay be used.

The starting end of the segment is adjusted so that the blank contacts both dies. As the rotary die revolves, the blank is rolled between the dies until it traverses the full arc of the segment die, after which it drops out of the threading area. For most applications, the finish end of the segment die is adjusted to produce the desired thread size. However, when rolling easily work-hardened parts, the dies may have to be adjusted so that the starting end of the die does most of the work and the finishing end does little, the part being completely formed by the time it is halfway through the die. Final thread size, however, does not depend entirely on correct adjustment of the die; hardness and size of the blank can cause variations. The effect of blank hardness is illustrated by the example that follows.

Ex a m ple 1 : Effe ct of H a r dn e ss of Bla n k on Th r e a d D im e n sion s. The effect of blank hardness on thread dimensions (primarily pitch diameter and major diameter) was investigated in an effort to reduce the rejection rate of 4-40 UNC-2A machine screws. (Rejection was for undersize pitch diameter and major diameter.) The screws, which were made of 1038 steel cold heading wire, were produced in an automated planetary die threader. Blank diameter was 2.35 to 2.36 mm (0.0925 to 0.0930 in.). A rejection rate of not more than 0.1% was accepted. Blanks of four different hardness levels were threaded: 60 HRB, 95 HRB, 28 HRC, and 32 HRC. Blank diameter was held within specified limits, and the same type of lubricant was used for all tests. Results are shown in Fig. 16. Some endwise stretching in the softer blanks was observed, probably because the part was short and of small diameter. The harder blanks work hardened perceptibly, resulting in undersize thread dimensions and shorter die life. Other details of the threading operation are listed in the table that accompanies Fig. 16.

Thread Blank diameter Machine Operating conditions Die speed, rev/min Feed rate(a), pieces/rev Cutting fluid Production rate, pieces/min Cost breakdown Material, % Manufacture(b), % Finish(c), %

(a) (b) (c)

4-40 UNC-2A by 6.4 mm ( in.) 2.35-2.36 mm (0.0925-0.0930 in.) No. 10 planetary die threader 90 4 Soluble oil 360 15 65 20

Material was fed into dies by a four-lobe cam. Includes overhead. Zinc chromate finish by the barrel method

Fig. 1 6 Effect of blank hardness on ( a) m aj or and ( b) pit ch diam et ers of m achine screws m ade of 1038 st eel

M a ch in e s. Planetary machines are made in several sizes, each having a different maximum rolling capacity. Although

machines are usually rated on the basis of nominal work diameter, the blank hardness and length of thread rolled have considerable influence on the practical capacity of a machine.

The basic planetary machine is comparatively simple, consisting of a spindle for the rotary die; a mounting block for the segment die, which also provides size adjustment; and a starting mechanism for inserting the blanks. The starting finger is adjusted by means of gearing or an adjustable cam, so that the workpiece is inserted at the exact point on the rotary die where it is in match with the segment. There is one such match point for each thread start on the rotary die. The number of thread starts varies from 10 to more than 100; however, it is seldom possible to feed a blank into the die at every thread start. Planetary machines usually can feed from three to eight parts per die revolution, depending on size; five pieces per revolution is common for this mass-production process. To make use of the high production capacity of planetary machines, automatic feeders are essential and are generally supplied as an integral part of the machine. D ie s. Lead angles are similar to those of radial infeed and flat dies; therefore, axial travel is at a minimum. Bevels are

similar to those used on other types of dies. Planetary dies vary in diameter from 100 to 350 mm (4 to 14 in.); the most commonly used machines have dies approximately 180 mm (7 in.) in diameter. The segment die has an inside radius (IR) equal to or slightly greater than the sum of the rotary die radius and the minor diameter of the threaded part. The width of the dies can be the maximum accepted by the machine or can be much narrower when short-length threads are rolled. Dies for straight threads, such as those on machine screws can be reversed so that both the upper and lower portions can be used for screws that have a thread length of less than half the die width. The maximum length of the segment die is limited by the length of the pocket in which it is held. Shorter dies are used for small-diameter work so that the work receives the proper number of revolutions during rolling. Because the surface of the rotary die is longer than that of the segment die, the rotary die generally has a longer life. Planetary dies for gimlet-pointed screws, such as type A and AB sheet metal screws, are very similar to flat dies for the same screws. The length of the segment die is equivalent to one flat die, and the circumference of the rotary die is usually equivalent to five segments and is capable of threading at least five pieces during each revolution of the die. Ca pa bilit ie s. Planetary thread rolling machines can roll most of the smaller parts that can be rolled on flat die machines. Production rates of planetary rollers are higher than those of other types of thread rolling equipment. Rates of 3000 pieces per minute can be reached on small pieces. The practical limit of production rate depends on the ability to feed the blanks rather than on speed of rolling. Planetary rollers lose their economic advantage over other equipment when it is necessary to reduce the production rate in order to feed difficult parts properly, or when quantities are too low (for instance, less

than 1 million 6.4 mm (

in.) screws).

Because they are produced in large quantities and can be fed easily, headed parts comprise most of the production on planetary machines. Typical products are machine screws; types A, AB, and B sheet metal screws; and drive screws. Size of product ranges in the popular machine sizes are No. 4 machine screws to 16 mm ( machines with capacities up to 29 mm (1

in.) diam screws, although

in.) have been built.

Lim it a t ion s. The cost of rotary dies is in proportion to die size. Die cost per piece produced is generally competitive

with other types of dies, but unless volume is high enough to use up the available die life, large inventories of dies will accumulate.

Con t in u ou s Rollin g Continuous rolling is a high-production method suited to cylindrical-die machines. The method uses two cam-type segmental dies, maintained at a predetermined center distance, as required for the desired pitch diameter. Workpieces are fed from a hopper or magazine to a revolving cage-type workrest that indexes them into and away from the rolling position. Depending on the number of die segments, one, two, or three workpieces are threaded for each revolution of the dies. Continuous rolling in a two-roll cylindrical-die machine provides the highest rate of production for headed workpieces. This method is applicable to the threading of double-end studs and similar parts. Threads of different diameters, pitches,

and tolerances, as well as those of identical specifications, can be produced in a single pass. Parts can be rolled to produce threads on one end and knurling or splines on the other. When thread rolling double-end studs, two sets of cam-type dies are used. Each spindle contains two dies maintained a fixed distance apart, depending on stud size. Both ends of the stud can be rolled simultaneously when thread diameters are the same. Except for studs larger than about 19 mm ( in.) in diameter, segmented dies can be used to roll different pitches on the two ends of a stud in one spindle revolution. Studs and similar parts up to 19 mm ( in.) in diameter and 345 mm (13 in.) long have been produced by this method. Production rate varies with size and shape of the part. Three applications of continuous rolling are shown in Table 3. Threads on double-end studs 8.0 MM ( in.) in diameter and 120 mm (4 complete parts per minute by this method.

in.) long can be produced at a rate of 240

I n t e r n a l Th r e a d Rollin g The rolling of internal threads requires dies of comparatively small diameter, which greatly limits die life and the loadcarrying capabilities of die bearings and spindles. As a result, internal thread rolling is limited in its use. However, it has been used successfully for forming helical fins on the internal surfaces of heat-exchanger tubes and for rolling internal threads in pipe couplings. Internal threads can be rolled by impressing the inside surface of a workpiece shell of suitable wall thickness onto a closefitting threaded mandrel, as shown in Fig. 17(a). Pressure is provided by three or four rotating plain external dies. The workpiece and mandrel may be clamped in a stationary position with the dies mounted in a rotating die head, or the dies may be stationary while the work and mandrel rotate. It is necessary to unscrew the part from the mandrel after threading has been completed.

Fig. 1 7 I nt ernal t hread rolling. ( a) Wit h a close- fit t ing t hreaded m andrel. ( b) Wit h a t hreading die t hat is considerably sm aller t han t he inside diam et er being rolled

To be thread rolled internally, the workpiece must be made of highly ductile metal such as aluminum, brass, or lowcarbon steel. The wall must be thick enough to provide adequate material to fill the die thread, but not so thick that the external dies are prevented from creating sufficient load to cause the threaded mandrel to penetrate the workpiece. Parts have also been threaded internally by using a smaller, threaded mandrel or die and a single, plain external die or support roll, as shown in Fig. 17(b). In this procedure, the part feeds axially when a single-start thread is being rolled. Dies can roll a multiple-start thread with a minimum of axial movement. Cold form tapping, discussed in the article "Tapping" in this Volume, is another method of producing internal threads by metal displacement without the production of chips.

Se le ct ion of Rollin g M e t h od Table 4 shows approximate ranges of production rates for different types of thread rolling equipment. Actual production rates depend on the condition of the machine being used, the work-handling equipment, the type of workpiece, and the metal being rolled. The quantity of pieces to be rolled is also an important factor in the selection of a machine.

Table 4 Approximate range of production rates of thread rolling equipment

3.2

Infeed rolling, threads per minute Flat-die Cylindrical die traversing Single Multiple revolution(a) revolution 40-500 75-300 20-90

450-2000

Through-feed or end-feed rolling, threads per minute ParallelSkewedaxis dies axis dies 20-40 140-280

6.4

40-400

60-150

20-90

250-1200

20-40

200-450

13

25-90

50-100

15-70

100-400

25-55

110-300

19

20-60

...

10-50

...

25-65

80-300

15-50 ...

... ...

8-40 6-30

... ...

20-50 15-30

70-300 50-200

... ...

... ...

4-25 3-20

... ...

10-20 6-15

30-140 20-90

... ... ...

... ... ...

2-15 1-5

... ... ...

4-10 1-3 ...

15-40 5-10 ...

Thread diameter mm

25 38 51 64 76 100 125

(a)

in.

1 1 2 2 3 4 5

-1

Rotary planetary

Two threads can be rolled on double-end studs in one die revolution in some machines.

In general, a low volume of identical parts (up to several thousand pieces) can be produced most economically on a handfed flat die machine. Hopper feeding is usually most economical for more than 10,000 pieces. Depending on the size of part, a production run of more than 100,000 pieces can often be produced most economically on a planetary-die machine.

In one operation involving standard hexagon-head screws ranging from 6.4 to 9.5 mm ( to in.) in diameter, the optimum run for a hand-fed flat die machine was approximately 10,000 pieces. The hopper-fed flat die machine was best for quantities between 10,000 and 400,000 pieces, and the planetary-die machine proved to be the most practical for runs of more than 400,000 pieces. The quantity of pieces economically producible on cylindrical-die machines is difficult to assess. These machines can roll special thread forms, and the parts produced are frequently of superior quality, including greater precision. They also can roll a greater range of diameters than either flat die or planetary-die machines. Selection of a single machine that is suitable for a variety of applications is discussed in the example that follows.

Ex a m ple 2 : Equ ipm e n t for Va r ie d Pr odu ct M ix in Qu a n t it ie s fr om 5 0 t o 1 0 0 0 Pie ce s. Various products produced in one plant were originally threaded by cutting tools in secondary-operation equipment. Thread sizes ranged from 6.4 to 25 mm ( to 1 in.) in diameter and from two diameters to 300 mm (12 in.) long. Production quantities ranged from 50 to 1000 pieces. Although thread cutting produced a quality product, the company decided to purchase thread rolling equipment capable of rolling class 2A and 3A threads and of handling the varied product mix. The various types of thread rolling equipment were considered, and selection was based on the factors discussed below. Tr a ve r sin g Fla t D ie M a ch in e . Die and setup costs for this equipment were favorable, and a high production rate

could have been obtained for most of the thread lengths involved. The limited diameter capacity, however, would have necessitated the purchase of more than one size machine in order to include the total range of thread diameters of the product mix, and this in turn would have been unsuitable for the long thread lengths required and the variety of workpiece configurations. Pla n e t a r y- D ie M a ch in e . The quantity of parts to be threaded was insufficient to warrant the purchase of this highproduction type of equipment. Also, die cost per piece would have been excessive, and the machine would have been unsuitable for the long thread lengths required. Cylin dr ica l- D ie M a ch in e s. This type of equipment would have been satisfactory from the aspect of versatility in

application to the product mix. Infeed rolling could have been used for the short thread lengths, and through-feed rolling could have been used for the long thread lengths. Two-die and three-die machines were available, and setup time would have been about the same as for other machines of similar capacity. However, because these machines rotate the workpiece, it was doubtful that the equipment could have been used for all sizes and shapes of product. En d- Fe e din g H e a ds. Thread-rolling machines equipped with end-feeding heads (thread rolls) could have rolled both

short and long threads. Because each head could have rolled only a limited range of thread diameters, several sizes would have been required for the product mix involved. Setup of the machine and changing of the head would have been done easily and quickly. Cost of thread rolls was low because of their small size. Because end-feeding heads rotate around the workpiece, simple work holding fixtures could have been employed. Se le ct ion . The equipment selected was a thread rolling machine equipped with end-feeding heads. Although a cylindrical-die machine would have been suitable for the work required in this plant, it was more expensive. Final selection of the machine with end-feeding heads was based on the low cost of heads and thread rolling dies and the simplicity of fixtures needed for holding the work.

Fa ct or s Affe ct in g D ie Life The life of thread rolling dies is determined primarily by the rate of deterioration of the profile of the die threads. Rolling imposes severe stress on the dies from pressure and bending and sliding action. Dies usually fail by spalling and crumbling of the thread crests, which roughens the minor diameter of the product thread and causes the screw thread to go out of tolerance. Failure is probably caused by fatigue from the stresses imposed in rolling. The best products and maximum die life can be obtained only when the dies are properly set up and the correct

die speed and number of blank revolutions are used. The surface of the material being worked should be relatively free of oxide and scale. Spalling or chipping may be the direct result of endwise extrusion or stretching of the blank during rolling, over-rolling, improper blank design, or inferior design or quality of tools. When spalling occurs, relatively large pieces are broken out of the threads of the tools. This failure usually occurs near the edges of the tools or in the tool area where the ends of the blanks are rolled. Spalling is frequently the result of having improper bevels on the blanks and tools, or of excessive tool hardness. Excessive variation or irregularity of blank diameter and hardness increases susceptibility to spalling, as does contact between the edge thread on the tools and the shoulders and fillets on the workpiece. Mismatching of the tools during rolling imposes transverse loading on the tool threads and causes spalling. When crumbling takes place on the crests of die threads, failure usually starts in the more highly stressed rolling areas and gradually spreads over other portions of the die threads. Excessively sharp die crests are subject to greater initial crumbling action. Useful tool life is greatly affected by the hardness and work hardening characteristics of the metal being rolled. Premature failures may be minimized by preventing over-rolling and by using clean blanks of correct, uniform size and hardness. The wear of thread rolling dies seldom need be considered. When wear does occur, it is primarily associated with the abrasive action of scale or dirt on the surface of the blanks, or the use of contaminated coolant. Deep threads are generally subject to the most rapid abrasive wear because of the greater amount of sliding action between workpiece thread and die thread. The rolling of deep threads of comparatively short length and small minor diameter may result in excessive extrusion of the blank, which causes permanent bending or chipping of the die thread profile at the end of the blank. Die life ranges from millions of pieces for soft work metals to a few hundred pieces for hardened steels. Table 5 shows the approximate relationship between die life and method of rolling different thread sizes in steel at 85 HRB and 30 HRC.

Table 5 Nominal die life for three thread rolling methods Thread diameter

Die life Flat die pieces

Cylindrical dies Infeed, pieces Through-feed km ft × 103

mm in. Threading steel at 85 HRB (550 MPa, or 80 ksi, tensile strength) 700,000 350,000 6.4

21

70

625,000

315,000

19

63

450,000 ...

245,000 190,000

15 9.1

50 30

...

125,000

5.3

17.5

...

70,000

2.1

7

... 18,000 4 Threading steel at 30 HRC (965 MPa, or 140 ksi, tensile strength) 35,000 20,000 6.4

...

...

4.3

14

32,000

18,000

3.8

12.5

24,000 ...

14,000 11,000

3.4 1.8

11 6

...

7,000

1.1

3.5

...

4,000

0.4

1.4

13 25 38 64 89

1 1 2 3

115

13 25 38 64 89

1 1 2 3

115

4

...

1,000

...

...

Even when production conditions are as nearly constant as is feasible, die life can vary greatly, as illustrated in Fig. 18. Seventy-two identical thread rolling dies, made of D2 steel by hobbing before hardening, failed with the die life distribution shown in Fig. 18. These data for dies of the same design and tool steel, all used for rolling the same threads on the same blank material, represented about the narrowest variation of factors affecting die life possible in normal production.

Fig. 1 8 Dist ribut ion of die life of 72 dies used for t hreading low- carbon st eel bolt s in a 13 m m ( bolt m aker m achine

in.)

Fin e Th r e a ds. Dies for rolling 40 or more threads per inch, especially with class 3A fits, require greater precision than

do dies for rolling coarse threads, but they impose less cold work on the blanks during rolling and have longer life. Fasteners with the closest tolerances and finest pitches require grinding of dies after hardening to achieve required accuracy. D ie Life Ve r su s H a r dn e ss of Bla n k . Increasing hardness of the blanks being rolled shortens die life. Figure 19(a)

shows the approximate average life of a number of flat dies made of D2 tool steel in rolling blanks of various hardnesses.

-20 threads on 1022 steel

Fig. 1 9 Relat ion of die life t o hardness and diam et er of blank. ( a) Die life versus hardness of blanks of 1022 st eel t hreaded wit h flat dies of D2 t ool st eel. ( b) Die life versus hardness of blanks of alum inum or brass t hreaded wit h cylindrical dies of A2 t ool st eel. ( c) Die life versus diam et er of st eel blanks t hreaded wit h st andard flat dies or wit h special-cont our dies in bolt m akers, and of alum inum and brass blanks t hreaded wit h cylindrical dies

Some of the tests were run on hardened blanks, including the two hardest specimens plotted. On blanks harder than 94 HRB, die life was low and inconsistent. The shaded portion of the curve indicates both the spread and inconsistency in die life. A blank hardness of about 32 HRC is the limit for normal thread rolling. Figure 19(b) shows minimum and maximum die life versus hardness of the screw blank for a large number of circular dies made of A2 tool steel. These dies were used mainly in rolling fine-pitch threads on parts made mostly from freecutting brass and aluminum alloys. The data include some high-production runs in which conditions of setup and blank material were nearly ideal. Therefore, the average life of cylindrical dies would be less than the mean between the two curves shown. Also, dies that roll fine threads on small parts have longer average life than those that roll coarser threads on larger parts. Figure 19(c) shows the relationship between die life and diameter of threads rolled on aluminum and brass with circular dies and on steel with flat dies. The upper and lower curves relate to A2 circular dies, most of which were ground after hardening and used primarily for rolling fine-pitch threads, the majority of them being special threads. The original data indicated no detectable difference in die life between aluminum and brass at a hardness of about 88 HRB. The curve relating to 1016 and 1020 steel gives the average life in a single setting of several hundred D2 dies with special contour on entering threads and used as hardened after machining. The other curve gives similar data for 1010 and 1038 steels on machines of the traversing type. Unusual variables can alter die life considerably. For example, rolling 16 mm ( in.) diam hot-rolled 1040 steel blanks that were nonuniform in diameter, scaled, and unannealed, has reduced die life to only 45,000 pieces.

Effe ct of Th r e a d For m on Pr oce ssin g The form of the thread affects the radial die load, seam formation, surface finish, and thread dimensions.

Ra dia l D ie Loa d. Greater loads are required as the form becomes more blunt. For example, if a radial load of 45 kN

(10,000 lbf) is required to produce a

-13 UNC thread 50 mm (2 in.) long, a radial load 40 to 70% greater would be

required to roll a -10 Acme thread under the same conditions. The amount of increase in load is determined by the blend radii on the crest of the die thread and by the degree of fullness of thread required. Se a m for m a t ion at the crest of the thread is greatly affected by the shape of the thread form, as indicated in Fig. 20.

Sharp crests on the dies cause sharper seams than those produced when a wider form is rolled more deeply.

Fig. 2 0 Effect of t hree different t hread form s on seam form at ion at t hread crest , and t he rolled shape of t he workpiece at t hree levels of die penet rat ion ( 30, 60, and 90% )

Su r fa ce fin ish is usually independent of the form for 60° threads. However, as the thread form becomes more blunt, or

if it has sharp root corners, the normal, smooth flow pattern of the metal being rolled is altered. Restriction of metal flow may cause localized subsurface shear failures. During subsequent die contacts, these shear failures form small flakes, which downgrade the surface finish.

Th r e a d D im e n sion s. Table 6 shows proper specifications for thread rolling high-quality threads in ground and

extruded blanks. It is worth noting that no specification for chamfer angle on the end thread is required. In addition to the dimensional details tabulated and shown in the inset illustration, principal dimensions such as the major and minor diameters should, of course, be specified.

Table 6 Specifications for high-quality rolled threads on ground and extruded blanks Thread size

p (pitch) mm

0.25p (min) in.

mm

in.

R runout radius) mm in.

(min

T root radius)(a) mm in.

(min

10-32

0.7938 0.9070

0.03125 0.03571

0.20 0.23

0.008 0.009

0.30 0.33

0.012 0.013

0.08 0.10

0.003 0.004

1.0584

0.04167

0.25

0.010

0.38

0.015

0.13

0.005

1.0584

0.04167

0.25

0.010

0.38

0.015

0.13

0.005

1.2700

0.05000

0.30

0.012

0.46

0.018

0.15

0.006

1.2700

0.05000

0.30

0.012

0.46

0.018

0.15

0.006

1.4112

0.05556

0.36

0.014

0.51

0.020

0.15

0.006

1.4112

0.05556

0.36

0.014

0.51

0.020

0.15

0.006

1.5875

0.06250

0.41

0.016

0.58

0.023

0.18

0.007

-28 -24 -24 -20 -20 -18 -18 -16

(a) (b)

Maximum root radius is limited by requirement that threads accept a "go" thread gage. Radius at root of runout thread must not be less than minimum radius at root of full thread.

Changes in the thread form affect the relationship between pitch (p) diameter and outside diameter of the thread. Figure 21 shows that if a standard 60° thread form with a typical sharp thread crest is not quite fully rolled, an additional die penetration of 0.025 mm (0.001 in.) on the diameter, with a commensurate reduction in the pitch diameter, results in an increase of 0.08 mm (0.003 in.) in the outside diameter until the thread form is filled.

Fig. 2 1 Effect of t hread form and die penet rat ion on t he relat ion bet ween pit ch diam et er and out side diam et er of t he t hread

For the more blunt Acme thread, the outside diameter increases at only about 1 diameter until the material fills the form through the crest radii.

times the rate of decrease of the pitch

The class of thread affects the size of the blank diameter and the major and minor diameters for a given thread. Classes 1A, 2A, and 3A threads of the same size and pitch in the American Standard series can be rolled with the same die. However, a different die may be required for rolling a given thread of a class other than 1A, 2A, or 3A.

Su r fa ce Spe e d Permissible surface speeds for thread rolling are governed by the mechanical power limitations of the threading equipment and, when the workpiece is rotated, by the speed (rev/min) of the workpiece or of the holding equipment used with end-feeding heads. Table 7 compares die surface speeds in modern thread rolling equipment with the surface speeds of thread cutting tools. Table 8 shows the approximate thread rolling time for different spindle speeds used with tangential-infeed double-roll attachments.

Table 7 Operating speeds of thread rolling and thread cutting tools Tool Thread rolling tools

Surface speed m/min sfm

Flat traversing die Cylindrical die Lathe attachments Thread cutting tools High speed steel tools Carbide tools

(a)

30-100 20-180 20-90

100-325 70-600 70-300

3-45 75

10-150(a) 250

Speed range for threading aluminum, copper, steel, and Monel

Table 8 Thread rolling time for tangential infeed double-roll attachments Threads per 25 mm (per inch) 32 24 18 14 10

(a)

Approximate work revolutions(a) 11-27 14-31 17-35 20-39 23-43

Thread rolling time (s) for spindle speed (rev/min) of: 500 1000 1500 2500 5000 1.3-3.2 0.7-1.6 0.4-1.1 0.3-0.6 0.1-0.3 1.7-3.7 0.8-1.9 0.6-1.2 0.3-0.7 0.2-0.4 2.0-4.2 1.0-2.1 0.7-1.4 0.4-0.8 0.2-0.4 2.4-4.7 1.2-2.3 0.8-1.6 0.5-0.9 0.2-0.5 2.8-5.2 1.4-2.6 0.9-1.7 0.6-1.0 0.3-0.5

The actual number of work revolutions used within the ranges shown depends on the material rolled and the size of the thread rolling attachment.

Pe n e t r a t ion Ra t e a n d Loa d Re qu ir e m e n t s The total die penetration per revolution of the blank varies with different machines, the kind of work, and the type and hardness of the metal being rolled. Low penetration rates are necessary for hard metals, hollow workpieces, and workpieces of nonrigid cross section. Higher penetration rates are used for rolling metals that work harden at an excessive rate. The rate of die penetration is normally limited by the rigidity of the workpiece and the machine and by the hardness of the metal being rolled. Penetration rates for infeed rolling range from 0.0125 to 0.15 mm (0.0005 to 0.0060 in.) per revolution of the workpiece. Penetration rates for through-feed and end-feed rolling are governed by the lead and pitch of the thread and the length of the entrance threads on the dies. If the pitch and other rolling variables are held constant and only the diameter of the work is increased, the rolling load will increase about one-third as much as the increase in diameter. For example, if a load of 17 kN (3800 lbf) is required to roll a 25 mm (1 in.) length of -20 thread, the initial die load for rolling a -20 thread under the same conditions would be about 22 kN (5000 lbf). However, because the thread pitch normally increases proportionately with the diameter, a proportional change in the rolling conditions causes the die load to increase about half to three-fourths as much as the increase in the workpiece diameter. Many rolling machines have insufficient power to provide proportionately rapid penetration rates on larger work; thus, an increase in workpiece size presents a problem when rolling threads of large diameter in a metal that work hardens rapidly. In such applications, the inability of the equipment to produce adequate absolute penetration rates increases the number of die contacts necessary and causes too rapid an increase in the hardness of the work metal. As a result, disproportionately large radial die loads are required to complete the thread form. This is hard on machines and dies. More powerful machinery must be used if work hardening of the metal being rolled causes this kind of load escalation. Rated motor horsepower of flat die and cylindrical die infeed thread rolling machines is not a direct measure of peak power consumption on specific jobs. The motor horse-power of these intermittently loaded machines is supplemented by the stored energy of the drive system momentum. Motor horse-power of through-feed machines can be rated more directly, because power is applied continuously. Each unit of power delivered to the through-feed die rolls threads in material with a rollability index of 1.00 at approximately the rates shown in Table 9.

Table 9 Approximate rates of through-feed thread rolling for metals with a rollability index of 1.00

Thread diameter mm in. 6.4 7.9 11 13 14 16 19 22 25

1

Threads per 25 mm (per inch) 28 20 24 18 20 14 20 13 18 12 18 11 16 10 14 9 14 8

Rolling rate(a) mm/min/kW 5930 5650 5310 4120 2180 2080 1600 1500 1460 1330 1360 1190 1160 950 890 750 610 540

in./min/hp 174 166 156 121 64 61 47 44 43 39 40 35 34 28 26 22 18 16

Note: To determine the through-feed rate for a metal of different rollability index, multiply the rolling rate by the rollability index.

(a)

Power is actual power delivered to the dies.

In addition to the power used for threading the workpiece, the rolling process develops various power losses. Because of these losses, power available for actual threading ranges from 30 to 90% of rated power of the machine. Energy delivered to the thread rolling dies is dissipated in forming the metal and also because of friction between the work and the die. On shallow threads, die friction accounts for about 10% of the energy delivered. For deep threads, losses due to die friction may be as much as 25% of the energy delivered; the remainder is used to flow the metal.

W a r m Rollin g Low-strength materials, such as aluminum, brass, and low-carbon steel, are always thread rolled at room temperature. Because of their relatively low yield strength, these metals are easily penetrated by the threading die. However, some metals, such as high-strength steels and heat-resisting alloys, offer considerable resistance to die penetration. To facilitate thread rolling, blanks of these materials are often heated and rolled while warm. For example, warm rolling is used to thread fasteners made of 5% Cr-Mo-V steel (H11) that has been quenched and tempered. The H11 steel is induction heated to about 480 °C (900 °F) immediately before threading. Heating temporarily decreases the yield strength so that threads can be formed without great difficulty. However, for optimum life, the dies should not be allowed to reach too high a temperature. An ideal ambient die temperature is 90 °C (200 °F). Dies can be cooled by spraying, as needed, with a conventional die-cooling oil or a soluble-oil emulsion. Workpieces that have been heated under controlled conditions and warm rolled on dies with a positive lead have a tensile strength equal to, and a fatigue strength greater than, similar parts rolled cold. In addition, die life in warm rolling is greater than in cold rolling of high-strength materials.

Th r e a din g of Th in - W a ll Pa r t s Three-die machines and end-feeding heads are better suited to the threading of hollow work than are two-die machines, because the application of rolling forces at three points on the circumference has much less tendency to collapse the work. Table 10 shows the minimum wall thickness for thread rolling tubular workpieces on three-die machines. For satisfactory thread rolling of tubular workpieces in a two-die machine, a minimum wall thickness of twice the thickness shown in Table 10 for any given set of conditions is the recommended practice.

Table 10 Preferred minimum wall thickness for thread rolling tubular sections Threads per 25 mm (per inch) 32 24 20 18 16 14 12 10 8

Minimum wall thickness, mm (in.), for thread pitch diameter, 25 mm (in.), of: 25-50 (1-2) 50-75 (2-3) 75-100 (3-4) 100-125 (4-5) 0.75 (>30)

500

320

220

180

30-91

A or B

Q

500

500

320

220

180

27-84

45-52 HRC

A or B

N

500

400

280

220

120

24-76

52-58 HRC

A or B

K

500

320

280

120

120

23-69

58-60 HRC

B or A

I

400

280

220

120

120

20-69

>60 HRC

B or A

I

320

220

120

120

120

18-56

Tool steels, wrought A or B 150-300

Q

600

500

320

220

180

23-64

300-400

A or B

N

500

500

320

220

180

21-61

45-52 HRC

A or B

N

400

320

220

150

120

20-55

52-58 HRC

A or B

K

400

280

220

150

120

15-46

>60 HRC

B or A

I

320

220

180

120

120

15-46

600

500

320

280

180

30-76

315-420

A or B

N

500

500

320

220

150

26-72

45-52 HRC

A or B

N

500

400

280

220

150

21-67

52-58 HRC

A or B

I

500

320

220

150

120

15-61

Gray, ductile, and malleable cast iron C or D Q 500 110-315

500

320

220

180

315-420

C or D

Q

500

500

320

220

180

36.5100 26-91

45-52 HRC

C or D

N

500

400

320

220

120

20-67

Austenitic (Ni-Resist), gray, and malleable cast irons C or D J to S 500 400 280 100-270

220

120

3868.5

Aluminum alloys, wrought and cast

Cutting fluid(c)

Rotation speed m/min sfm

300-400

Stainless steels, wrought and cast A or B O 135-315

Working pressure

4.9-30 4.3-27 3.9-25 3.4-23 3.0-21

3.7-22 3.4-20 2.716.5 2.716.5

4.6-26 3.7-24 2.7-23

C or D L 30-450 (500 kg, or 1100 lb) Titanium alloys, wrought C or D N 110-315 315-440

C or D

N

600

500

400

280

220

15-64

50210

2.7-23

9-75

276

40

A

600

500

400

280

180

15-61

2.7-23

9-75

276

40

A

500

500

400

280

180

9-46

50200 30150

1.516.5

5-55

414

60

A

Copper alloys, wrought and cast C or D I 40-200 (500 kg, or 1100 lb) Nickel alloys, wrought and cast A or B J 80-315

600

500

500

320

180

15-76

50250

2.7-27

9-90

276

40

A

600

500

400

280

180

30-76

100250

5.5-27

1890

276

40

A

Nickel alloys, wrought and cast A or B J 315-420

500

500

400

220

180

27-69

4.9-24

40

A

400

320

220

180

24-61

1680 1475

276

500

90225 80200

517

75

A

500

320

280

180

24-61

80200 50160

4.3-23

414

60

A

2.7-18

1475 9-60

414

60

A

45-52 HRC

A or B

J

High-temperature alloys, wrought and cast C or D J 500 200-315 J

500

400

320

280

180

15-49

Refractory alloys A or B 180-320

Q

500

400

320

280

180

15-61

50200

2.7-23

9-75

414

60

A

Chromium plate A or D Hard

L

320

280

220

180

120

27-91

90300

4.933.5

16110

690

100

A

315-475

C or D

4.3-23

Source: Metcut Research Associates Inc.

(a) (b) (c)

Abrasive type: A, aluminum oxide; B, cubic boron nitride (metal bonded); C, silicon carbide; D, diamond (metal bonded). Abrasive grade: soft (F, G, H, and I); medium (J, K, L, M, and N); and hard (O, P, Q, R, and S). Honing fluid: A, 70% kerosene-30% sulfurized or chlorinated oil; B, kerosene.

Table 4 Selection of honing stone grit sizes for nonmetals using power or manual-stroke operations Material

Glass

Ceramic

Carbide

(a) (b)

Honing stone material Type(a) Grade(b) ANSI and ISO A S D P D L A S D P D L A S D P D L

Grit size for surface roughness (Ra), m ( in.) 0.025-0.125 0.15-0.25 0.30-0.50 0.53-0.75 (1-5) (6-10) (11-20) (21-30) 600 400 280 220 400 280 220 150 400 220 150 ... ... ... 600 400 ... 600 400 280 600 400 280 220 600 400 280 220 600 400 280 220 400 280 220 150

Working pressures >0.75 (>30) 150 ... ... 280 220 150 150 150 ...

Heavy, roughing Normal Light, feathering out Heavy, roughing Normal Light, feathering out Heavy, roughing Normal Light, feathering out

Abrasive type: A, aluminum oxide; D, diamond (bonded metal). Abrasive grade: medium (L); hard (P and S).

Table 5 lists guidelines for selecting aluminum oxide, silicon carbide, cubic boron nitride (CBN), and diamond and shows typical cutting speeds and stone pressures for these abrasives. Abrasive selection for honing in terms of workpiece material alone is:

• • • •

Aluminum oxide is widely used on steels Silicon carbide is generally used on cast iron and nonferrous materials Cubic boron nitride is used on all steels (both soft and hard), nickel- and cobalt-base superalloys, stainless steels, other alloys such as beryllium copper, and metals such as zirconium Diamonds are used on chromium plate, carbides, ceramics, glass, cast irons, brass, bronze, and surfaces nitrided to depths greater than 0.03 mm (0.001 in.)

Table 6 provides detailed information on the honing of specific materials with cubic boron nitride.

Table 5 Guidelines for selecting honing abrasives Abrasive type

Applications

Aluminum oxide (Al2O3)

Low cost, lowest hardness. Used for both soft or hard but rough, stringy materials; ordinary steels; and some nonferrous alloys. Wear ratio is 5:10. Good for deburring, interrupted cuts, poorly qualified bores, and scale and weld distortions Low cost. Used for either soft or hard but brittle materials or low shear-resistant materials. Results in more uniform surface finishes. Frequently used for finishing. Widely used on cast iron High cost. Rapid cutting and long-life stones. Wear ratio is usually 800:1200. Low noise levels. Used for hard steels, tool steels, stellites, nickel- and cobaltbase alloys, and superalloys. Minimum speed is 75m/min (250 sfm) when honing mild steels (AISI 1015-1026) High cost. Used for tungsten carbide, ceramics, glass, and cast iron. Very long life

Silicon carbide (SiC) Cubic boron nitride

Diamond

(a)

(b) (c)

Cutting speed(a) m/min sfm 5-30 16100

Honing stone pressure(b) kPa psi 10015-60 400

5-30

16100

100400

15-60

35-90

110180

4002000(c)

60300(c)

40-80

130260

6001200

85170

Cutting speed is the combination (vector sum) of axial and peripheral speed. The crosshatch angle generally runs between 45° and 80°--some evidence indicates the best removal rates come at 60° to 75°. A convenient way to arrive at a good rev/min for large-diameter bores ( 64 mm, or 2 in.) is: optimum rev/min = 30,000/diam in mm or 1200/diam in in. These values represent approximately 90 m/min (300 sfm). For small diameters, use safer lower speeds shown in Table 3. These values represent conservative starting values. A convenient starting point for any abrasive is the low end of the pressure range. To improve productivity, trials at successively higher pressures should be made until the stone wears too rapidly. When substituting CBN abrasives for conventional abrasives, a convenient starting point is at of the pressure used with the former abrasive, and then increase the pressure until the stone wears too rapidly.

Table 6 Operating parameters when honing with cubic boron nitride Material

Hardness, HRC

Stock removal rate

Diameter

Length

Honing oil

in./min 0.007

Stone wear rate per 0.025 mm (0.001 in.) of stock removal mm in. 0.0050 0.0002

1855 soderfors 16

Hard

mm/min 0.18

mm 12

in. 0.470

mm 54

MB-30

kPa 830

lb 120

1600

m 1.3

in. 50

M4

Hard

0.06

0.0025

0.0008

0.00003

30

1.190

44

MB-30

860

125

640

0.63

25

M4

Hard

0.14

0.0054

0.0030

0.00012

19

0.750

44

MB-30

1300

190

800

0.75

30

M4

Hard

0.04

0.0016

0.0015

0.00006

18

0.700

63

MB-30

1720

250

1270

0.38

15

D3

61-63

0.15

0.006

0.0013

0.00005

11

0.437

23

MAN-852

2070

300

1600

...

...

D7

Hard

0.11

0.0043

0.0036

0.00014

20

0.780

38

MB-30

2000

290

800

0.5

20

440C stainless M2 M2 Inconel 718 H13

59-62 60+ Hard Soft 46

0.05 0.06 0.08 0.04 0.006

0.002 0.0025 0.003 0.0016 0.00025

0.0018 0.0018 0.0036 0.0050 0.033

0.00007 0.00007 0.00014 0.0002 0.0013

57 31 31 7.5 63

2.247 1.250 1.250 0.300 2.500

50 75 100 125 455

MB-30 MB-30 MB-30 MB-30 MB-30

860 1380 1380 2480 Light

125 200 200 360 Light

400 310 230 1600 320

0.18 ... ... 0.38 0.58

7 ... ... 15 23

Source: Metcut Research Associates Inc.

in. 2 1 1 2

1 2 3 4 5 18

Pressure on stone surface

Spindle speed, rev/min

Surface roughness (Ra)

Aluminum oxide and silicon carbide stones are comparable in initial cost. However, one may be more economical than the other because of increased service life. Diamond stones initially cost up to 20 times as much as aluminum oxide or silicon carbide stones, depending on diamond concentration and thickness of abrasive layer. Diamond stones are almost mandatory, however, for honing extremely hard, wear-resistant materials such as tungsten carbide or ceramics. For honing nitrided cases or chromium plate, diamond stones have often been more economical on the basis of metal removed per unit of time. Table 7 lists parameters associated with single-stroke honing when using diamond-plated tools.

Table 7 Single-stroke honing with diamond-plated tools Workpiece diameter mm in. 10 1013 1316 1619 1925 2532 3241

-1 1-1 1

-

Spindle speed, rev/min 3000

Stroking feed per revolution at stroke rate, mm/min (in./min)(a) 1500 (60) 2300 (90) 3400 (134) mm/min in./min mm/min in./min mm/min in./min 0.51 0.020 0.76 0.030 1.14 0.045

5000 (200) mm/min in./min 1.70 0.067

Slow production

2400

0.64

0.025

0.91

0.036

1.42

0.056

2.11

0.083

Best operation

1900

0.81

0.032

1.19

0.047

1.80

0.071

2.67

0.105

Best operation

1500

1.02

0.040

1.52

0.060

2.26

0.089

3.38

0.133

1200

1.27

0.050

1.91

0.075

2.84

0.112

4.24

0.167

Fast production, but caution is advised

950

1.60

0.063

2.41

0.095

3.58

0.141

...

...

Do not use.

750

2.01

0.079

3.02

0.119

4.50

0.177

...

...

Do not use.

600

2.54

0.100

3.81

0.150

...

...

...

...

Do not use.

Results

1 4151

(a)

1

-2

One stroke into and out of a bore can achieve respectable removal rates (0.025 to 0.038 mm, or 0.001 to 0.0015 in.) on the diameter, close tolerances (±0.0025 mm, or 0.0001 in.) and reasonable finish (0.050 to 0.63 m, or 20 to 25 in., Ra) when a crowned, adjustable, 220-grit, plated-diamond hone is used on cast iron.

D e sign a t ion s. An abrasive marking system that applies to both grinding wheels and honing stones has been established

by the abrasives manufacturers. Details of this system are given in the article "Grinding Equipment and Processes" in this Volume. The four most significant characteristics are type of grit, grit size, hardness grade, and type of bond. These characteristics are identified by letters and numbers in the above order. For example, an aluminum oxide stone having a grit size of 180, a grade of R, and a vitrified bond is identified as A-180-R-V. This method of stone designation is used in this article. Tools A phenomenon of honing that makes it possible to develop a round, straight bore is the relationship of the cutting faces of the stones to the surface being honed, which is completely independent of the machine. The fact that either the tool or the workpiece floats (see Fig. 8) enables the tool to exert equal pressure on all sides of the bore, regardless of vibration in the machine or its environment.

Fig. 8 ( a) Tool or ( b) t he workholding fixt ure float s t o perm it t he bore and t he t ool t o align.

Except in special cases, such as with the square-axis requirement described next, honing will not change the axial location of a hole. The centerline of the tool follows the neutral axis of the bore as established by previous operations; the tool or workpiece must float so that the bore and the tool may align themselves. If practical, the workpiece should be mounted in a fixture that permits it to float. If the part is too large or unbalanced to float, however, the tool should be designed with universal ball joints. The fixtures must be designed to locate the workpiece approximately in line with the machine spindle and take the torque and axial thrust of the tool without distorting the workpiece. For bores that must be held square with an end surface, the part location method has been developed. In this technique, both the tool and fixturing are rigid. The workpiece is slipped over the tool and automatically aligned with it before clamping. Honing tools differ in design depending on whether they are to be used with manual or power stroking. M a n u a l St r ok in g. In manually stroked honing, the tool is rotating around a fixed axis, but the work has free lateral

motion and is effectively guided by the honing mandrel. A tool for manual stroking consists of one or more abrasive stones, a mandrel and a wedge, and guide shoes. The simplest type is illustrated in Fig. 9(a) and 9(b). The purpose of the unequal angular spacing of the stones with respect to the shoes (Fig. 9a) is to facilitate removal of stock from the high spots until roundness is attained. The wedge, which controls feed-out of the stone, can be actuated manually or by an automatic mechanism such as that shown in Fig. 6. The two parallel shoes (Fig. 9b) stabilize and guide the workpiece. Because the shoes wear very slowly, alignment with the bore is maintained. Shoes are made of materials that are wear resistant with respect to the workpiece material. Bronze, cast iron, steel, or even plastics have been used. Sintered shoes made by powder metallurgy can retain oil.

Fig. 9 Tools for m anual- st roke honing. ( a) End view showing how self- alignm ent is provided by t hree-point cont act of shoes and single honing st one wit h bore. ( b) Exploded view showing how axial displacem ent of wedge cont rols t he feed of t he honing st one. ( c) Manual- st roke honing t ool wit h t wo st ones; on t ools for honing longer bores ( up t o 460 m m , or 18 in.) , as m any as five st ones can be m ount ed in line

Honing tools of the type illustrated in Fig. 9 are available for honing bores with diameters from 1.6 to 100 mm ( to 4 in.). Tools for manual-stroke honing of long bores (up to 460 mm, or 18 in.) have two or more short stones mounted in line, as shown in Fig. 9(c). Tools and stones should be long enough to permit bridging of common irregularities (Fig. 10).

Fig. 1 0 Com m on irregularit ies t hat dict at e m inim um lengt h of honing t ool and st one

Tools incorporating modifications of the design shown in Fig. 9 are available for honing holes having surface irregularities such as keyways (Fig. 11a) and for honing blind holes with and without relief (Fig. 11b and c). When honing blind holes, stone and shoe should be of equal length and shorter than the depth of the hole. In Fig. 11(b) and 11(c), stone and shoes have been equally shortened to ensure uniform wear. Note the shoe clearance at the blind end.

Fig. 1 1 Tools for m anual- st roke honing of holes having keyways and blind holes wit h and wit hout relief

Pow e r St r ok in g. Heavier workpieces, which, because of their bulk and weight, are not adapted to free floating, are honed with power-stroked heads whose shafts, which provide the connection with the rotating and reciprocating machine spindle, are equipped with a set of universal joints. Tools for power stroking (fixtured honing) usually have stones spaced at equal distances around the circumference and may or may not include guide elements (Fig. 12). Expanding cones (Fig. 13) control feed-out. Although the cones may be actuated manually, they are usually actuated automatically by constantpressure or constant-feed methods. As in manual stroking, the tool may incorporate numerous modifications, depending on the application.

Fig. 1 2 Tool used for power- st roke ( fixt ured) honing of aut om ot ive cylinder bores

Fig. 1 3 Tool used for power- st roke ( fixt ured) honing of large bores. Expansion cones are used for feed-out of honing st ones. See t ext for discussion.

Figure 12 shows a typical tool used in fixtured honing. Tools of this type are used for multiple-spindle honing of parts such as motor blocks. For this application, tools must be self-aligning because it is impossible to align each cylinder bore with each spindle centerline. Therefore, each tool incorporates a double universal joint. The nonabrading guides that are placed between the stones (Fig. 12) are kept within 0.50 to 0.75 mm (0.020 to 0.030 in.) of the cylinder wall to prevent excessive eccentric wear of the stone, which will cause out-of-roundness or other deviations in bore shape. Figure 13 shows a typical tool assembly for honing large-diameter bores. In this assembly, which was used for honing a bore 1040 mm (41 in.) in diameter, the stones are mounted in a holder that will accept as many as four stones in line. Each

stone is 200 mm (8 in.) long, making the total effective length 810 mm (32 in.). Depending on bore size, up to 24 stone holders can be spaced around a tool of this type (although 12 holders are indicated in Fig. 12). Because this tool is used on a single-spindle machine, no universal joints are needed. Guide elements are seldom used with this type of tool, but in horizontal honing of long bores, outside support of the spindle (drive shaft) may be required. In power stroking, the honing stones must be encased to prevent wear on the tool body and to permit quick changing of worn stones. The material used for mounting the stones must not damage the bore surface by scratching or spalling. Two widely used types of holders are shown in Fig. 14: (a) a stone mounted in a plastic holder and (b) a stone cemented to a metal die casting.

Fig. 1 4 Mount ing for abrasive st ones used in fixt ured honing. ( a) St one encased in a plast ic holder. ( b) St one cem ent ed t o a m et al die cast ing

The body of the honing tool must have open areas that permit a copious flow of honing fluid around the abrasive stones during honing. These open spaces, both axial and radial, are evident in the tool shown in Fig. 12. Pa r t Loca t ion M e t h od. The principles of a proprietary system of work location for ensuring dependable alignment between the fixed honing tool axis and the hole, in a firmly clamped workpiece, are shown and described in the diagrams in Fig. 15. This system avoids reliance on a floating motion, which is accompanied by some friction; the workpiece is located by aligning its hole with the honing head and is clamped only after the necessary aligning adjustments have taken place. Connecting rods and a variety of other short bore parts are successfully honed individually in a two-position airfloat fixture, as illustrated. The part is inserted and approximately located in the fixture at the load position.

Fig. 1 5 Aut om at ic four- st ep process for aligning t he workpiece and t he honing t ool t o hone connect ing rods. ( a) Part is loaded in fixt ure, and fixt ure is m oved int o m achining posit ion on a cushion of air. ( b) The t ool, which is not rot at ing, is lowered int o t he bore wit h t he abrasives ret ract ed. ( c) St ones are expanded out ward against t he bore wall, while t he abrasives cent er t he bore and t he t ool. ( d) Wit h t he abrasives st ill expanded, t he clam p is lowered ont o t he part t o ready t he t ool for cycling.

The part location method allows the workpiece to be positioned concentric with the tool before it is clamped. The part is manually indexed into hone position on a thin cushion of air. The tool is then moved into the part bore, and the abrasives are expanded, moving the part into final hone position. The tool is stationary at this time, and the part is then clamped. This type of location, when used with a rigid tool having no universal action, creates a better geometry. The clamp is brought down on the part with a confirming surface to compensate for any out-of-parallel condition. This type of clamping eliminates torque influences. Therefore, there is no need for outside torque arms, and part distortion is eliminated. Another advantage in clamping the part in this fashion is that a support is put into the part allowing an increase in stone pressure and surface area without misalignment of the part and tool. Ga ge s Automatic size control in power stroking may be accomplished by adaptations of at least five types of gages--air, ring, expanding, plug, and bar gages. Use of these gages is described below, and methods are illustrated in Fig. 16 and 17.

Fig. 1 6 Air- t o- elect ronic- gage sizing for aut om at ic size cont rol in honing

Fig. 1 7 Gages for aut om at ic size cont rol of fixt ured workpieces honed in power-st roking m achines. ( a) Ring gage. ( b) Expanding gage. ( c) Plug gage. ( d) Bar gage. See t ext for discussion.

An a ir ga ge is an integral part of the honing tool. With each stroke of the tool, the gage portion enters the bore with sufficient clearance to permit pressurized air emitted from gage orifices to escape to the atmosphere. As the bore size increases, back pressure in the air system decreases. When the bore is to size, a previously calibrated pressure switch cuts off the cycle. This type of gage is particularly well adapted to controlling multispindle machines in honing cylinder bores. By adjusting the calibrated control, various diameters can be honed without interrupting the cycle. Air gaging can hold diametral tolerances of 0.075 mm (0.0003 in.) and can be adjusted within a range of 0.10 mm (0.004 in.). The device shown in Fig. 16, an air-to-electronic gaging system, converts the back pressure drop into an electronic signal that is fed to a programmable controller that automatically terminates the honing cycle when the bore diameter reaches its programmable size.

A r in g ga ge is mounted just above the workpiece (Fig. 17a). The ring, with an inside diameter equal to the specified bore diameter, is positioned so that only the plastic or metal tabs placed on the upper ends of the abrasive stones enter the ring at the top of each stroke. Because the abrasives and the tabs wear at the same rate, the diameter of the expanded tool is equal to the bore diameter. When the tool reaches gage ring size, friction from the tabs causes the ring to swing through a small arc and initiates the end of the cycle. This method of control is best adapted to bore diameters of 3.05 to 100 mm

(0.120 to 4.0 in.). It is capable of obtaining diametral accuracy of 0.0076 mm (0.0003 in.) or better, and adjustments within 0.013 mm (0.0005 in.) can be made to allow for wear or to vary the desired size. An e x pa n din g ga ge , which consists mainly of a split sleeve held together with a ring (Fig. 17b), reciprocates with the tool but is not attached to it. The gaging member, whose diameter is smaller than the bore diameter, enters the bore with each downstroke. At the bottom of the stroke, a lever on the side of the sleeve contacts a post, causing the split sleeve to expand until the gaging member touches the bore surface. Controls are preset so that two electrical contacts on the lever of the split sleeve meet and end the cycle when the diameter of the gaging member equals the desired bore diameter. This type of gage has been used successfully for size control of bores larger than 19 mm (0.750 in.) in diameter (there is no maximum size). A calibrated dial allows fine adjustment of the size within 0.25 mm (0.010 in.). A tolerance of 0.0075 mm (0.0003 in.) can be maintained at any dial setting. A plu g ga ge is independent of the tool and approaches the bore from the end opposite that at which the honing tool is

introduced (Fig. 17c). The plug, whose diameter equals the desired bore size, attempts to enter the bore with each stroke. When entry is made, controls are activated and the cycle is terminated. This method can be used to control accuracy within 0.005 mm (0.0002 in.). A ba r ga ge consists of two bars that float in the body of the tool (Fig. 17d). The bars are fastened to a split ring and held

against the bore surface by spring pressure. Bore diameter is measured by the distance between the contact faces of the bars. When the tool enters a bore, the bars are pressed inward, and two low-potential electrical contacts on the split ring are held open. As the bore becomes larger, the contacts move closer together; when they meet (as preadjusted), the cycle is stopped. The use of this type of gage is usually restricted to size control of bores larger than 50 mm (2 in.) in diameter. However, this gage accepts adjustments within 1 mm (0.040 in.). Also, control of honed diameters within 0.0075 mm (0.0003 in.) is feasible. Rot a t ion Spe e d Spindle speed depends mainly on the diameter of the bore being honed, because surface speed is usually the basic consideration in honing. The choice of an optimum surface speed is influenced by: • • • • •

Material being honed: Higher speed can be used for metals that shear easily, such as cast iron and some of the softer nonferrous metals Hardness: Harder workpiece surfaces require lower speed Surface roughness: Rougher surfaces that mechanically dress the abrasive stone permit higher speed Number and width of stones in the tool: Speed should be decreased as the area of abrasive per unit area of bore increases Finish requirements: Higher speed usually results in finer surface finish

Because of the variables listed above, rotation speed cannot be standardized. However, the speeds given in Table 3 serve as a starting point, and, in general, are close to those given in the examples in this article. Experience with a particular application may indicate advantages for higher or lower speed. Rotation speeds as high as 183 m/min (600 sfm) have been used successfully. Conversely, a reduction in spindle speed, and thus in surface speed, can reduce the number of rejects. Excessive speeds contribute to decreased dimensional accuracy, overheating of the workpiece, and glazing (dulling) of the abrasive. Overheating causes breakdown of honing fluid and distortion of the workpiece; the latter frequently affects final dimensions. Re cipr oca t ion Spe e d Speed of reciprocation, which depends largely on the length of the honing tool and the depth of the bore, is most usefully expressed in meters per minute (or surface feet per minute), the product of the number of stroke cycles per minute and twice the stroke length. Reciprocation speeds commonly used with a variety of work metals and alloys are listed in Table 3.

Because reciprocation speed, rotation speed, and crosshatch angle are related functions (see the next section), crosshatch angle can be controlled by varying the reciprocation speed when rotation is constant. Reciprocation speed also has some effect on the action of the abrasive; higher speed increases dressing action and thus usually produces a rougher finish. Con t r ol of Cr ossh a t ch An gle The crosshatch angle (Fig. 18a) obtained on a honed surface depends on the ratio of surface speed of reciprocation (stroking) to surface speed of rotation. When the rotation and reciprocation speeds are equal, the crosshatch angle is 90° (Fig. 18b). When rotation speed exceeds reciprocation speed, the crosshatch angle is less than 90° (Fig. 18c).

Fig. 1 8 Relat ion of crosshat ch angle ( a) t o rot at ion and reciprocat ion speeds. ( b) Rot at ion approxim at ely equal t o reciprocat ion. ( c) Rot at ion fast er t han reciprocat ion

The following formula can be used to determine the approximate angle that will result from given speeds (Fig. 18a):

(Eq 1) where Vr is the speed of reciprocation, Vp is the speed of rotation, and 2

is the crosshatch angle.

Although the above formula may be useful as a guide when determining the speeds necessary for obtaining the desired angle for a new job, it is often more practical to resort to trial and error. In power stroking, a common practice is to establish rotation speed and then to vary reciprocation speed to get the desired crosshatch angle. Reciprocation speed is less critical than rotation speed. However, an increase in the rate of reciprocation will sometimes improve the self-dressing characteristic of some abrasives and thus increase the stock removal rate; therefore, increased stock removal rates and fine finishes do not generally go together. If several thousandths of an inch of stock must be removed at a high production rate and less than a 0.50 m (20 in.) finish is required, two honing operations should be used.

A coarse abrasive (60 to 180 grit) is usually used first to generate geometry and size. Then a second, finer abrasive (240 to 600 grit) is used to refine the surface finish. The finish achieved with specific stones, however, depends on which material is honed. For some applications (engine cylinder bores are a notable example), crosshatch angle is important, and is noted in specifications. In the majority of applications, however, although an angle of 30° is commonly sought, any angle within the range of 20° to 45° usually is suitable. Angles within this range are practical with either manual or power stroking. In manual honing, an experienced operator can instantly alter the practice to suit conditions. For instance, the workpiece can first be stroked at a rate that has proved most efficient for stock removal. The technique can then be changed so that the desired crosshatch pattern is produced in the few final strokes. Conditions that will produce a crosshatch angle of approximately 30° in manual-stroke honing are given in Table 8 for a range of workpiece diameters and lengths.

Table 8 Conditions for producing a 30° crosshatch angle in manual-stroke honing Bore diameter mm in. 6.4

640

19

38

64

(a)

Rotation, rev/min 1600

1

2

320

200

Bore length mm in. 38 1 75 3 75

3

75

3

150

6

150

6

100

4

200

8

200

8

100

4

200

8

200

8

Stone length mm in. 32 1 32 1 57 2 64 2 64 2 114 4 83 3 83 3 165 6 83 3 83 3 165 6

Stroke length(a) mm in. 27.4 1.08

Stroke cycles/min(a) 155

65.5

2.58

65

57.1

2.25

75

55.1

2.17

94

131.3

5.17

39

114.3

4.50

45

74.2

2.92

70

175.8

6.92

29

148.1

5.83

35

74.2

2.92

72

175.8

6.92

30

148.1

5.83

36

In computing reciprocating surface speed: The distance traveled during one stroke cycle comprises a forward and a return stroke.

H on in g Pr e ssu r e A relatively wide range of pressures will yield acceptable efficiency and results. For example, in equipment using hydraulic force for feed-out, gage pressures have varied from 1040 to 3100 kPa (150 to 450 psi). However, honing is more often controlled by rate of feed-out than by gage pressure. In general, high feed-out rates are used for large diameters, and low feed-out rates are used for small diameters. As an example, a feed-out rate of 0.23 mm/min (0.009 in./min) could be used for roughing and a 0.18 mm/min (0.007 in./min) rate for finishing 215 mm (8 in.) diam cylinder bores in gray iron blocks for V-8 engines; for a 99.31 mm (3.910 in.) diam bore having a 150 mm (6 in.) length and made of the same material, the feed rate would be 0.075 mm/min (0.003 in./min) to 0.150 mm/min (0.006 in./min). Insufficient pressure will result in a subnormal rate of metal removal. When pressure is excessive, rougher finishes are obtained, because the abrasive is broken down too fast. This will result in increased stone cost, as well as decreased productivity caused by downtime required for replacing stones. Trial and error is the usual method of determining optimum pressure for a new application in production honing. A common procedure is to start with low pressure and then gradually build up, using workpiece finish as an indicator, until best conditions are found. Pressure must be kept up by automatic feed-out to compensate for stone breakdown and hole growth.

The data in Table 9, based on manual honing with a single stone, show why machines for honing various bores must apply a wide range of force. If a force of 220 N (50 lbf) is applied to a stone with a contact area of 7.74 mm2 (0.012 in.2), the pressure will be 29 MPa (4.2 ksi). Such a pressure would shatter the stone. But the same force applied to a stone with a contact area of 355 mm2 (0.550 in.2) creates a pressure of only 630 kPa (91 psi)--insufficient for cutting most materials.

Table 9 Relation of stone contact area and pressure with 220 N (50 lbf) Bore diameter mm in. 1.6

Bore length mm in. 16

Stone length mm in. 11

Stone width mm in. 0.71 0.028

Stone area mm2 in.2 7.75 0.012

Pressure(a) MPa psi 28.72 4166

3.2

19

14

0.89

0.035

12.9

0.020

17.24

2500

4.8

25

19

1.27

0.050

24.5

0.038

9.074

1316

6.4

38

1.78

0.070

56.8

0.088

3.915

568

13

75

3.68

0.145

210

0.326

1.062

154

16

90

5.59

0.220

354

0.550

0.627

91

16

230

5.59

0.220

1064

1.650

0.207

30

6.22

0.245

513

0.796

0.434

63

6.22

0.245

2568

3.981

0.089

13

38

1

38

1

115 510

1 1 3 3 9 4 20

32 57 64 190 83 413

1 2 2 7 3 16

Note: For manual honing with a single stone. Pressure under 220 N (50 lbf)

(a) H on in g Flu ids

Lubrication is more critical in honing than in most other metal-removing operations. No single honing fluid possesses a maximum of all properties needed for honing. Therefore, some compromise must be made, and mixtures of two or more liquids are commonly used. The oils used in honing serve two main purposes: • •

They promote cutting action by flushing workpiece metal and particles of abrasives from the honing stones, thus preventing the stones from loading and glazing They maintain an almost constant workpiece temperature, and thus minimize dimensional variation due to expansion and contraction

Honing fluid characteristics can directly influence quality or economy, or both. Viscous, gummy fluids, or fluids containing suspended solids, can cause a soft-bonded abrasive to lose efficiency and generate excessive heat from friction. Water-base solutions are superior as coolants, but they are poor lubricants and have insufficient viscosity to prevent chatter. Water also causes rust. Because of the latter two characteristics, water-base solutions are seldom used for honing fluids. Mineral seal oil is effective and is widely used for production honing. Mineral seal oil is a water-white product having a higher viscosity than kerosine (about 40 Saybolt Universal seconds, or SUS, compared to 31 SUS for kerosine). Its flash point also is higher than that of kerosine, and it is less likely to cause skin irritation. Mineral oils similar to those used for other machining operations have also proved satisfactory when one part of oil is diluted with four parts of kerosine. Buffers are often added to honing fluids, to minimize or prevent chatter of the honing stones and thus prevent their premature disintegration. Buffer materials absorb shock and recoil from fluctuations in force. Animal oils, including tallow, lanolin, and lard oils, are usually the most economical buffers, despite their relatively high cost, because of their long lasting qualities. These oils generally flow under pressure and cling to metal surfaces better than mineral oils.

Prepared proprietary oils that contain buffers are widely used. Before use, these oils (which may also contain rust inhibitors and deodorants) are diluted up to 95% with kerosine. The use of too much buffer may detract from its beneficial effects. An excessive amount: • • • • • •

Reduces the cutting action of the abrasive Produces smoother finishes Requires higher pressures or lower rotational speeds Lowers the ability of the fluid to dissipate heat Impairs fluid distribution Increases requirements for refrigeration and filtering

Regardless of the type of fluid used, it should be delivered to the honing stones in a constant and generous supply. The fluid also should be filtered through a system that removes particles coarser than 15 m (600 in.). The system should be kept free of water and stray oil (such as from the hydraulic system), which adversely affect the properties of honing fluids. In many plants, 17 to 20 °C (62 to 68 °F) is the preferred temperature range for honing fluids. Controlling the temperature becomes more important as tolerances become closer. If temperature is allowed to rise, dimensions may become inaccurate and the fluid may break down, causing excessive stone wear and changes in cutting characteristics. In production installations, heat exchangers are often used to maintain close control of honing fluid temperature. D im e n sion a l Accu r a cy Internal honing to tolerances of 0.025 to 0.0025 mm (0.001 to 0.0001 in.) is common. For some high-precision parts, tolerances as close as a few millionths of an inch are specified and achieved. Close tolerances can be produced and repeated in fixtured honing (power stroking) if sources of variation, such as machine and honing fluid condition, are closely controlled, as described in the example below. Ex a m ple 1 : Va r ia t ion s in D im e n sion s a n d Fin ish for 9 0 0 Cylin de r Block s. The data plotted in Fig. 19 represent results of a quality control check made on 99.31 mm (3.910 in.) diam cylinder bores in gray iron blocks for V-8 engines. Bores 2 and 7 were measured in 11 blocks from a production run of 900. The conditions employed in honing these bores are presented in Table 10.

Table 10 Processing details for honing cylinder bores in V-8 engine blocks Processing details(a) Machine production rate Spindle speed Spindle reciprocation Stock removal: Amount Time Honing fluid Stone life per set(c) Size control Dimensional tolerance Finish Crosshatch angle

(a) (b)

70 blocks/h 204 rev/min 78 strokes/min 0.051-0.102 mm (0.002-0.004 in.) 39 s Mineral seal oil(b) 450 blocks Spindle-mounted plug gage Max out-of-roundness and taper, 0.025 mm (0.001 in.) 0.50-0.89 m (20-35 in.) 22

°

Bores are classified in five sizes differing 0.013 mm (0.0005 in.) in diameter, for selective fitting of pistons. At 20 °C (68 °F), heat exchanger is required for maintaining this

(c)

temperature, and honing fluid must be free of water and tramp hydraulic oil. Silicon carbide stones

Fig. 1 9 ( a) Surface finish, ( b) t aper, and ( c) out - of- roundness variat ions obt ained in honing. Dat a r epresent m easurem ent s on cylinder bores 2 and 7 in 11 gray iron blocks for V- 8 engines, select ed from a run of 900. Measurem ent s were m ade on blocks 1, 50, 100, 200, 300, 400, 500, 600, 700, 800, and 900.

The honing fluid was maintained at 20 °C (68 °F) by the use of a heat exchanger, and was constantly filtered. Less than 2% of the 7200 bores honed required a repair operation because either taper or out-of-roundness exceeded the specified 0.025 mm (0.001 in.). Su r fa ce Fin ish Surface finish of 0.25 to 0.38 m (10 to 15 in.) can be obtained easily in production honing, and finish of less than 0.050 m (2 in.) can be achieved and reproduced. A range of roughness is sometimes specified. In other applications, a maximum surface roughness is specified. Under carefully controlled conditions, surface roughness can be maintained within a close range, as indicated in Fig. 19.

Size of grit in the honing stones is the main factor controlling surface finish. When grit is fine, the finish will be fine (other factors being equal); but as grit size is decreased, rate of stock removal is also decreased, as described in the following example. Ex a m ple 2 : H on in g Gr a y I r on t o a Fin ish of 0 .2 5 t o 0 .3 8

m ( 10 to 15

in .) .

In honing gray iron (hardness, 170 to 195 HB), a finish of 0.25 to 0.38 m (10 to 15 in.) was desired. Silicon carbide stones with a grit size of 180 produced a roughness of 0.63 to 0.75 m (25 to 30 in.). The required finish could be obtained with 320-grit stones, but the time required for honing made the use of this grit size impractical. The problem was solved by first rough honing with 180-grit stones and then finish honing, in another setup, with 320-grit stones. Rough finishes are sometimes improved by using a dwell time at the end of the honing cycle--that is, by continuing the rotation and reciprocation action for a few strokes after feed-out ceases and pressure drops off. In manual honing of a particular bore, use of this technique reduced surface roughness from the normal 0.50 to 0.25 m (20 to 10 in.). H on in g Pr a ct ice for I n t e r n a l D ia m e t e r s Honing is widely used for finishing bores in engine cylinders, cylinder liners, and bearing bores. Procedures for honing similar parts may vary from one plant to another, depending on quantity, available equipment, and established plant practice. As a rule, honing stones and techniques used for honing cast iron are different from those used for aluminum alloys. However, there are exceptions as in the case of an assembly of cast iron and aluminum in which the two metals were honed simultaneously, with the same abrasive, because it was the simplest way to achieve a proper fit. Sm a ll Bor e s. Conventional manual-stroking honing tools (Fig. 9) are available for use in bores as small as 1.6 mm (

in.) in diameter in parts such as fuel nozzles, miniature bearings, and heading dies, as in the following example. Ex a m ple 3 : H on in g Ve r y Sm a ll Bor e s. in.) in diameter. Bore length varied, Dies for cold heading tiny rivets and screw blanks had bores as small as 1.6 mm ( but was usually 25 to 50 mm (1 to 2 in.). Figure 20 shows one of the heading dies, which was made of tool steel, and a typical product of the die. Holes were drilled and reamed about 0.075 to 0.13 mm (0.003 to 0.005 in.) undersize before heat treatment. After hardening, they were honed, using manual stroking, to an accuracy of 0.0025 mm (0.0001 in.) for both roundness and straightness.

Fig. 2 0 Bore of die for cold heading t he rivet shown at t he left is t ypical of sm all bores finished by m anualst roke honing. Dim ensions given in inches

La r ge Bor e s. The maximum diameter and length of bore that can be honed is limited mainly by the size of the

equipment required for the workpiece and by the power required for the tools. Equipment with drive motors of up to 37 kW (50 hp) is available for honing steel shells of 1040 mm (41 in.) inside diameter (ID) and 19 m (63 ft) length. Cylinder shells for hydraulic hoists on regulating gates for dams are examples of large bores that are honed. In one honing operation, 0.75 mm (0.030 in.) of stock is removed from a 6.4 Mg (14,000 lb) shell of 760 mm (30 in.) ID and 7.9 m (26 ft) length to obtain a total envelope tolerance of 0.050 mm (0.002 in.). Before honing, the average out-of-roundness is 0.41 mm (0.016 in.). Sh or t Bor e s. Several different techniques are used for honing short bores. These are particularly applicable when bore diameter exceeds length. In the simplest method, several pieces are stacked with bores aligned, clamped tightly by any

suitable means, and honed as a unit. For example, 13 mm ( in.) long rings 38 mm (1 in.) in inside diameter can be honed in stacks of eight. In effect, this would be the same as honing a single piece 100 mm (4 in.) long. Stacked parts may be either manually or power stroked. However, for successful results from this technique, the parts must have parallel sides to permit building a straight stack that can be clamped tightly and provide a straight bore. Another technique that has proved successful for honing short bores is shown in the following example involving automotive-engine connecting rods (compare with newer method shown in Fig. 15). Ex a m ple 4 : Sh or t - Bor e H on in g Te ch n iqu e for Con n e ct in g Rods. A power-stroking horizontal machine was used in high production for honing 61.54 mm (2.423 in.) ID crankpin bores simultaneously in eight connecting rods. Figure 21 shows the fixture and the honing tool. Eight rods were stacked between 4.8 mm (

in.) wide parallel separator plates, resulting in an effective bore length of 260 mm (10

in.). The

tool had three banks of four honing stones. Each stone was 9.5 mm ( in.) square and 57 mm (2 in.) long. A twostation, rotary index table allowed the operator to unload eight completed rods and to load eight unfinished rods while eight other rods were being honed. A precheck plug probed the rods in the loading station to determine whether they had been bored to proper rough size. The operation completed one bank of rods in 45 s, floor-to-floor time, and a production rate of about 600 rods/h was obtained. In honing, 0.075 mm (0.003 in.) of stock was removed, a finish of 0.75 to 1.14 m (30 to 45 in.) was produced, and inside diameter was controlled within 0.13 mm (0.0005 in.).

Fig. 2 1 High- product ion honing of aut om ot ive part s. Fixt ure designed t o hone crankpin bores on eight aut om obile connect ing rods sim ult aneously, using a single honing t ool. Rot at ing fixt ure perm it t ed loading and unloading on one side while part s on t he opposit e side were honed.

Blin d h ole s are bores that have a bottom, shoulder, or other obstruction that prevents a tool from passing completely

through. The three most common types of blind holes are shown in Fig. 22. Most unrelieved blind holes can be honed satisfactorily, but there will always be some unfinished area at the bottom. The amount depends on length of bore, type of material, tolerance required, and amount of stock removed. Under the best conditions, dead-blind holes can be honed to within about 0.38 mm (0.015 in.) of the end. Any relief will improve results; as much relief as possible is preferred. Sometimes an unrelieved blind hole is in effect provided with a relief because specified tolerance and finish need not be met at the bottom of the hole.

Fig. 2 2 Three t ypes of blind holes

Special tools may be required, depending on whether or not relief (or on how much relief) is provided. For example, the unrelieved 13 mm ( in.) diam bore shown in Fig. 23(a) was manual-stroke honed to within about 0.38 mm (0.015 in.) of the end with a special tool having a hard-tipped honing stone (Fig. 11c). If adequate relief is provided, conventional tools are satisfactory. For example, cylinder heads in lawn mower engines (Fig. 23b) can be manual-stroke honed in high production with conventional tools, because of the generous relief (about 6.4 mm, or in., wide) at the blind end. Although both parts shown in Fig. 23 were manual-stroke honed, similar parts are frequently honed by power stroking.

Fig. 2 3 Blind holes honed by different m et hods. ( a) Unrelieved blind hole t hat required a special t ool ( see Fig. 11c) for honing. ( b) Relief t hat perm it t ed use of a convent ional honing t ool in t he bore of a cylinder head for a lawn m ower engine. Dim ensions given in inches

Delivering enough honing fluid to the work area is often a problem in honing blind holes. When a hole has a bottom opening (Fig. 23b), fluid can be pumped through a plastic tube inserted in the opening. When a hole has no bottom opening, the flow of fluid should be directed parallel to the mandrel, into the mouth of the bore. In manual honing, blind holes are more difficult to keep straight than open holes. A truing sleeve (dummy workpiece) is frequently used to keep the shoes and stones straight and parallel; also, the stone and shoe are made shorter than the blind hole. Experienced operators have found that using a series of short strokes with an occasional stroke all the way out of the mouth is the best practice, until the hole is close to final diameter. This keeps the bottom slightly larger than the mouth. Straight strokes are then used for finish honing.

Ta pe r e d Bor e s. Part size, angle of taper, and length-to-diameter ratio determine the method used in taper honing. Short

tapers are honed using a machine and tool such as that shown in Fig. 24. The machine has a head that can be positioned for any desired degree of taper, and the reciprocating tool holds a single stone. The workpiece is rigidly clamped in a fixture that rotates. This method is most commonly used for producing tapers on parts for which the length of honed area is less than the diameter. As the length of the taper increases in proportion to the diameter, however, the practicality of the method decreases, because the longer and more slender tools lack adequate rigidity.

Fig. 2 4 Machine and t ooling for honing short , t apered bores

Applications of this method of taper honing include special bearing rings and parts that use end tapers for sealing, and bores in gears that must fit tapered shafts. For example, drum-to-barrel seals in a 20 mm gun must have a taper of 0.050 mm/mm (0.050 in./in.) of length at each end and roughness less than 0.25 m (10 in.). To meet these requirements, 0.01 to 0.05 mm (0.0005 to 0.002 in.) of stock must be removed from the critical surfaces. Taper honing long bores in large parts is far more complex than honing short tapers. A major portion of the stock is removed by step honing. In this operation, a straight stroke is used, its length being progressively reduced to form a rough taper consisting of a series of small steps. The taper is then finished in a second operation in which a sine bar regulates the increase and decrease of the diameter on the return and forward stroke of the honing cone. Spe cia l Sh a pe s. Machines and tools have been developed for honing various special shapes. For female splines,

honing stones must be narrower than the spline width (preferably no wider than half the spline width) to allow for oscillation. Machines and tools for honing splines are designed to produce simultaneous reciprocation and oscillation, rather than reciprocation and rotation. Relief bores are commonly honed by contour boring. Spe cia l Applica t ion s of H on in g A few special uses of honing should be enumerated as a means of indicating the potential of the honing method beyond the field of its basic and most extensively accepted applications. These are related to the honing of internal cylindrical surfaces by using regular abrasives for obtaining specific dimensional conditions of the work surface. Among these related processes are: • • • • •

External honing Gear tooth honing Plateau honing Flat honing Electrochemical honing

•

Hone forming

Ex t e r n a l H on in g. Honing has been used to only a limited extent for finishing outside diameters, largely because required dimensions and finish can be produced at less expense by other processes, such as centerless grinding. However, advances in metrology and improved honing techniques have resulted in an increase in the number and scope of applications of external honing.

Special machines and special adaptations of conventional machines (such as lathes) have been tooled to hone outside surfaces of metal parts. With these machines, either power or manual stroking may be employed. Fix t u r e d e x t e r n a l h on in g (power stroking) is widely used for pieces that are not adaptable to competitive methods. A

notable example is the finishing of grooves in bearing races. Special machines that simultaneously rotate the workpiece and oscillate the stones (Fig. 25) produce the crosshatch lay pattern characteristic of a honed surface.

Fig. 2 5 Fixt ured honing of grooves on ext ernal surface of bearing rings wit h sim ult aneous oscillat ion of honing st one and rot at ion of workpiece

M a n u a l e x t e r n a l h on in g is applicable to the removal of small amounts of stock from external diameters of a wide

variety of sizes and shapes. The honing of lengths up to 3 m (10 ft) is common practice. Conventional honing machines are generally used for rotating workpieces up to 610 mm (24 in.) long. Lathes or drill presses are preferred for longer workpieces. Tools such as that illustrated in Fig. 26 are available for honing parts ranging in outside diameter from about 3.05 to 69.85 mm (0.120 to 2.750 in.). With this setup, the sides of the tool are gripped and stroked over the rotating workpiece. Feedout and cutting rate are controlled by applying pressure to the honing-control lever, which will move through a preset distance. Size is controlled automatically by setting the micrometer stone feed-out so that the honing-control lever will be against the stop pin when the correct size is attained. The only adjustment needed during the honing operation, even in production runs, is a slight additional stone feed-out to compensate for stone wear. A turn of the honing-control lever will instantly disengage the stone from the work for quick gaging or unloading, but will not change the setting on the micrometer stone feed-out.

Fig. 2 6 Assem bly used for m anual- st roke honing of out side diam et ers. See t ext for discussion.

With the setup shown in Fig. 26, a line of stones with opposing guide shoes, or opposing stones, can be used. For honing long parts (up to 610 mm, or 24 in.), multiple holders that contain as many as three stones (or shoes) in line may be used for correcting waviness. The torque arm can be used to offset the tendency of the tool to turn. A guide bar mounted on the machine acts as a stop for the torque arm. This type of tool can produce dimensional accuracy to 0.0025 mm (0.0001 in.) or better and surface roughness as low as 0.050 m (2 in.). Manual-stroke external honing has replaced lapping in some applications, because: • • •

Honing is usually faster Soft metals can be honed without being impregnated with abrasive The use, in honing, of multiple-length stones and shoes allows better control of bow and waviness

Long anodized aluminum tubes for in-flight refueling are honed externally in a lathe, the honing tool being moved by hand, and the nozzle for the honing fluid moving with the tool. Crankpins of some crankshaft are honed the same way at overhaul. Ge a r - t oot h h on in g is an abrasive process designed to improve geometric accuracy and surface conditions of a hardened gear. The teeth of hardened gears are honed to remove nicks and burrs, to improve finish, and to make minor corrections in tooth shape. Gear teeth are honed on high-speed machines specially designed for the process (Fig. 27). The honing tool is like a gear driving the workpiece at high speed (up to 30 m/min, or 100 sfm) while oscillating so that the teeth slide axially against the workpiece.

Fig. 2 7 Honing t eet h of helical gears

Spur gears and internal or external helical gears ranging in diametral pitch from 24 to 2.5, in outside diameter from 19 to 673 mm ( to 26 in.), and up to 75 mm (3 in.) in face width have been honed on these machines. Finishes of 0.75 m (30 in.) are easily achieved, and finishes of 0.075 to 0.10 m (3 to 4 in.) are possible. Both taper and crown honing can be done. Tools used in honing gear teeth are of two types, a helical gear shape tool made of abrasive impregnated plastic, and a

metal helical gear with a bonded abrasive coating that is renewable. The plastic tool, which is discarded at the end of its useful life, is widely used. The metal tool is used mainly for applications in which plastic tools would be likely to break; also, it is used primarily for fine-pitch gears. Plastic tools are supplied with abrasives of 60-grit to 500-grit size. Size of abrasive, gear pitch, and desired finish are usually related as:

Grit size 60 100 180 280 500

Gear pitch 16 16-20 >20 >20 >20

Finish m 0.75-0.89

in. 30-35

0.63-0.75 0.38-0.50 0.25-0.30 0.075-0.10

25-30 15-20 10-12 3-4

Honing tools do not load up, and a plastic honing tool can wear until its teeth break. Stock removal of 0.025 to 0.050 mm (0.001 to 0.002 in.) measured over pins is the recommended maximum. M e t h ods. The two methods used to hone gear teeth are the zero-backlash method and the constant-pressure method. In

the zero-backlash method, which is used for gears made to commercial tolerances, the tool head is locked so that the distance between the center of the work gear and the center of the honing tool is fixed throughout the honing cycle. In the constant-pressure method, which is used for gears produced to dimensions outside commercial tolerance ranges, the tool and the work gear are kept in pressure-controlled tight mesh. Applica bilit y. The use of honing for removing nicks and burrs from hardened gears can result in a considerable cost

saving in comparison to the usual method. In the usual method, the gears are tested against master specimens on sound test machines. Nicks indicated are searched for and removed using a hand grinder. The gear is then retested to make certain the nick has been removed. When honing is used, all of these various tests and procedures can be eliminated.

Some shape correction can be achieved in the removal of 0.050 mm (0.002 in.) of stock by honing. A helical gear 127 mm (5 in.) in diameter may show lead correction of 0.010 mm (0.0004 in.), involute profile correction of 0.0075 mm (0.0003 in.), and eccentricity correction of 0.010 mm (0.0004 in.). The advisability of using honing for salvaging hardened gears hinges on cost considerations. As the error in tooth shape increases, honing time increases and tool life decreases. On the other hand, if the gears represent a large investment in production time and material, honing may be the most economical method. Because honing is not designed for heavy stock removal or tooth correction, it cannot be substituted for grinding or shaving of gears. Rotary shaving usually leaves gear teeth smooth within 0.25 to 1.00 m (10 to 40 in.). Pla t e a u h on in g produces a special plateau finish, which removes the surface peaks but retains the deep valleys. Such a finish has been found desirable in engine performance because the valleys act as oil reservoirs for improved lubrication, especially during engine break-in.

A plateau finish is produced by first rough honing to final size. Then the surface is finished with a finer-grit stone for about 45 s, depending upon the amount of plateauing desired. The plateauing operation, with a 600-grit stone, removes so little stock that the bore diameter is not measurably increased. Fla t h on in g is a term designating a method and the equipment by which the flat surfaces of component parts produced

by other methods are improved with regard to both flatness and parallelism of opposite surfaces. One of these surfaces may be that on which the part is located during the honing of the opposite face, or both faces may be honed simultaneously on machines operating with two honing disks. The equipment used is similar in appearance to rotary face-grinding machines, but it is adapted to honing, a method which differs from grinding particularly in the low cutting speed of the abrasive disk, the applied speed being comparable to that used in conventional honing. The bonded abrasive disks used in flat honing are generally not intended for substantial rates of stock removal and thus can have very fine grains, promoting the development of a high-grade finish, even of the order of 0.025 m (1.0 in.) Ra when needed. The spindle of the honing disk used on flat honing machines can be raised and lowered by an air or hydraulic cylinder. Single- or double-surface flat honing machines are designed for high-production uses, finishing typically 1200 to 1800 parts per hour in a fully automated operation controlled by a timer. On two-wheel machines (see Fig. 28), the top wheel, lower wheel, and workholder each have separate drives and controls. Machines are available with automatic controls to gradually increase pressure on the top wheel during the honing cycle. Automatic size control is also available. The workpiece carrier is part of an epicyclic sprocket holder.

Fig. 2 8 Two- wheel flat honing m achine

Ele ct r och e m ica l H on in g. In this process, metal is removed by introducing an electrolyte into a gap between a

cathodic honing tool body and an anodic workpiece. Direct current from the power source is conducted to the tool through a brush assembly that acts as the cathode, while the workpiece becomes the anode. The tool mandrel is made of metal and has a series of small holes which provide channels for the electrolyte circulating under controlled pressure. The electrolyte has the additional role of being a coolant and flushes away the chips that have been sheared off by the honing stones. Abrasive honing stones are nonconductive and are limited to removing the electrochemically loosened metal particles, thereby controlling the geometric form and the size and texture of the produced surface. Thus, the electrochemically honed surface has the same characteristic crosshatch pattern and is essentially stress free, as is the surface produced in a regular honing process by purely mechanical action. The abrasive stones continuously remove from the work surface the oxides developed by the electrochemical action, leaving the surface clean. This makes use of an electrolyte much less corrosive than that needed in the electrochemical

machining process. Nor is it necessary for the electrolyte to be dispensed at high pressure, a process characteristic that calls for great structural strength of the equipment in the mechanically unassisted electrochemical machining. The electrochemical honing process is carried out in a manner quite similar to regular honing, combining the rotation and reciprocation of the tool at controlled speeds. However, the tool mandrel must have good conductivity and all elements of the equipment that are exposed to the corrosive environment must be made of corrosion-resistant materials. H on e for m in g ( H F) is a recent development that constitutes a marriage of two different processes, honing and electrodeposition. The process is used to simultaneously abrade the work surface and deposit metal. It produces work surfaces that combine the benefits of both processes, that is, a geometrically and dimensionally controlled surface with a functionally favorable texture developed on a cladding that is concurrently electrodeposited on the base metal. In some of its basic principles, the method is a reversal of electrochemical honing.

The gap between the anodic honing tool body and cathodic workpiece is kept small. Current densities used in HF operations are many times those employed in other electrodepositing methods. Rate of deposition varies from 0.018 to 0.05 mm/min (0.0007 to 0.002 in./min) and up to 0.64 mm (0.025 in.) of metal has been deposited effectively in the laboratory. In the following description of the equipment and process, a cylindrical internal surface or bore is considered; however, the method is applicable to flat or external round surfaces as well. Hone forming is carried out on special equipment, which includes the machine for supporting the fixture and for actuating the rotating and reciprocating tool with its controlled feed motion. Additional elements of the equipment are the rectifier to supply the direct current, the solution tank, and the circulating system. The insoluble anode is part of the honing tool and is connected to the positive side of the rectifier. The workpiece functions as the cathode and is connected to the negative side of the rectifier. The gap between the anode (tool) and the cathode (workpiece) is kept small, while through that gap at controlled velocity passes the solution, which, together with the workpiece, is contained in a sealed circuit. Figure 29 shows the components of a typical hone forming machine.

Fig. 2 9 Principal com ponent s of a t ypical hone form ing m achine

After the workpiece has been cleaned of oil and other potential contaminants, it is mounted in a special honing machine and, as the first phase of the process, the bore is honed in a conventional manner. This part of the operation produces a clean base metal surface with accurate cylindrical form. Subsequently, the metal-depositing process is started, keeping the anode at a controlled distance from the work surface, which acts as the cathode. The electric power supply is turned on, and the electrolyte is circulated between the anode and the cathode. During this process the honing action continues, but with reduced stone pressure, thus ensuring that the deposited metallic layer will develop into a work surface of the desired form (roundness), size (diameter), and surface texture (crosshatched). The process is terminated when the desired size of the bore has been reached, as determined either indirectly, by the elapsed time, or directly, by means of a gage. The entire process is carried out within a very short time, of the order of about one minute, because the cladding of the surface is many times faster than in conventional bath plating and has the added benefit of obtaining a precisely honed work surface with controlled characteristics. Other advantages of the new process are the elimination of most of the timeconsuming preparatory and postprocess operations, such as masking, washing, neutralizing, and so forth, which are necessary for attaining the same results with the application of conventional procedures. In hone forming applications, the workpiece must be conductive. To date, most materials used in hone forming are copper, bronze, tin, nickel, cobalt, or chromium plated and are hone formed on workpiece materials such as iron, steel, stainless steel, and bronze. Theoretically, any surface that can be honed can be processed with hone forming. Although the immediate uses of this new method are mostly in the area of work salvaging out-of-tolerance parts and reconditioning worn surfaces, its potential is much wider and includes production line operations.

M icr oh on in g The term microhoning is considered a descriptive designation for a method that uses hones (bonded abrasives) for tools and produces an accurately controlled surface whose parameters are measured in micro (very small) units. Essentially, the method is applied to round surfaces, both external and internal, of cylindrical, tapered, or spherical general form, and less frequently, to flat surfaces with limited areas. All of these contain surface elements that are either straight lines or circular arcs. The method operates with tools made of bonded abrasives, with the active surfaces having forms complying with the contacted area of the work surface. The abrasive tools, while being held against the surface of the rotating work with a controlled, light force, generally effect a rapid, short-stroke, reciprocating (oscillating) motion in a direction parallel with the surface elements of the processed work surface. Microhoning can be done on centers or with centerless through-feed machines. With centerless microhoning, 60 to 70° of the total workpiece circumference is in contact with one or more fine-grit stones. Com pa r ison of M icr oh on in g w it h Re gu la r H on in g Characteristics common to regular honing and microhoning methods are: • • • •

The use of bonded abrasives for tools Operation with low cutting speeds and producing light cutting force The combination of a rotational and a linear motion between the tool and the work; both methods produce a cross-hatch pattern, but in microhoning it is less distinct The objectives of the process, that is, improving the form, size, and surface texture of the work surface without causing metallurgical damage

Particular differences between regular honing and microhoning are described below.

Th e St r ok in g Le n gt h . Honing is essentially a long stroking method, capable of covering work surface lengths several

times that of the stones. During the process the stones travel a substantial distance (usually expressed as a fraction of the stone length) over the ends of the processed work surface. In microhoning, the stone reciprocation is over a very short length, typically of the order of about 1.0 to 4.1 mm (0.040 to 0.160 in.). For that reason the stone motion is often referred to as an oscillation. Fr e qu e n cy of t h e Re cipr oca t ion . In honing, the number of strokes per minute varies from very few, in the case of

very long strokes, to about 100 or 200 in general operations. The stroking, although usually hydraulically actuated, in the case of low-production or medium-production machines may also be manual. Microhoning operates with a very rapid reciprocating motion, its frequency varying, according to the applied system and stroke length, in the range of 300 to 2500 cycles per minute. The reciprocating motion has mechanical or pneumatic drive. Rot a t in g M e m be r . In honing, it is the tool that rotates (with the rare exception of the external honing of very long

parts); in microhoning, it is always the workpiece that rotates in front of the tool whose reciprocation is along a fixed path. Con t a ct Ar e a . In honing, usually several stones, at essentially uniform circumferential distances, make simultaneous contact with the work surface, each stone taking part in the stock removal.

In microhoning, a single stone, except for the rare multistone applications on large diameters, covers only one segment of the work surface, although the width of the stone produces a wraparound effect. W or k Su r fa ce Con figu r a t ion . Stroking in honing is always along a straight-line path, which limits the general form of the surface adapted to honing. The long stroking of honing also excludes surfaces that do not permit free stroking, although the honing of surfaces with one free end, for example, blind holes, is possible. Honing is used nearly exclusively for bores.

Microhoning, which operates with very short strokes, can be applied also to surfaces with circular cross-sectional contours, such as spherical or toroidal shapes. Equally adaptable to external and internal surfaces, the short stroking also permits the microhoning of surfaces bounded by shoulders on both ends, requiring a very narrow undercut, comparable to that provided in cylindrical or internal grinding. Am ou n t of St ock Re m ova l. Honing can be used for efficiently removing substantial amounts of stock. In rough

honing, of the order of 0.25 to 0.38 mm (0.010 to 0.015 in.) can be removed, occasionally, considerably more. In finish honing, about 0.05 to 0.15 mm (0.002 to 0.006 in.) can be removed from the part diameter. This stock-removing capacity of the honing method permits the correction of various types of bore form irregularities. Stock removal in microhoning is usually of the order of 0.0025 to 0.0075 mm (0.0001 to 0.0003 in.), consequently, the general form irregularities, for example, ovality, taper, and so on, can be corrected only within a very limited range. Ch a r a ct e r ist ics of t h e Pr odu ce d W or k Su r fa ce . Honing is generally used to produce surface finishes in the range

of 0.8 to 0.2 m (32 to 8 in.) Ra. Such surface texture values are adequate, or even desirable, for many types of bores finished by honing. Often even rougher surfaces are required, or, in rare cases, a finish of the order of 0.10 m (4 in.) Ra is specified. It is feasible to produce even finer finish values by honing, but generally at the price of rapidly decreasing productivity. Microhoning can produce, in efficient operation, surface finishes of the order of 0.10 to 0.05 m (4 to 2 in.) Ra, and, upon occasion, even finer finishes. This single parameter, the average surface roughness, does not convey, however, all the functionally important surface characteristics that can be developed in regular production by microhoning. These will be discussed in greater detail under a separate heading. Ba sics of t h e M icr oh on in g Pr oce ss The operating principles of the microhoning process were illustrated earlier in Fig. 3. The diagram illustrates that an abrasive stone, whose operating face complies with the general form of the work surface, is forced against the rotating workpiece to exert a specific pressure while carrying out a short-stroke reciprocating motion.

In operation, the abrasive stone will first make contact and act upon the protruding elements of the work surface. Protuberances can result from surface form irregularities such as lobbing, chatter marks, waviness, and so on, and can also constitute the peaks and crests of a rough surface. Generally, irregularities from both sources are present, with the roughness superimposed on the form irregularities. By abrading the protruding elements of the surface, the action of the microhoning stone will gradually equalize the work surface, correcting some of the form irregularities and reducing the roughness. The major protuberances of the surface will be abraded first by the concentrated action of the total stone pressure; consequently, the rate of surface improvement will be rapid in the starting phase of the operation. As the contacted work surface area continuously extends, the rate of stone penetration will become slower, partly because the specific pressure on the contacted areas decreases, and partly because the stone surface starts to glaze, due to the reduced dressing action of the work surface. The latter condition, although detrimental to the abrading ability of the stone, is beneficial to the development of a better finish on the microhoned surface, which is the prime objective of the process. The optimum balance between the effective abrading and the dependable smoothening ability of the stones can be controlled by the proper selection of the process variables, such as reciprocating and rotating speeds; applied pressure; and form, dimensions, and composition of the stones. The described progress from fast abrading to the slower smoothening action could be extended to produce very substantial improvements of the work surface by microhoning. However, it would take an excessive amount of time for a stone that is capable of rapid abrading action to lose its cutting ability to the degree that it could produce a very smooth surface. For that as well as for other reasons, that is, to retain the unimpeded abrading capacity of the stone, the microhoning process is usually operated in the range of its highest efficiency. The approximate extent of that range can be expressed conveniently by the ratio of surface roughness improvement, which usually varies from 4:1 to 8:1, although these values are not absolute limits. The lower ratios commonly apply to the effective range for coarse surfaces, such as improving the finish from 6.3 to 1.5 m (250 to 60 in.) Ra, while the roughness of a fine-ground surface can be reduced efficiently from 0.40 to 0.050 m (16 to 2 in.) Ra. Su r fa ce t e x t u r e is seldom measured, although it can affect certain functional characteristics, such as the sealing capacity and lubricant retainment of the surface. At a relatively higher roughness level, a cross-hatch pattern can be observed on the microhoned surface; that pattern tends to change, as the process progresses, into a nondirectional pattern, or to disappear entirely, resulting in a reflective surface.

These changes of the surface texture, brought about by microhoning, when applied to the distinctly directional pattern produced by grinding are generally beneficial to the functional properties and service life of the work surface. A rough surface will make contact with a mating part through its protruding crests, and the extensive lower-lying areas will have no functional role. Reducing the roughness of the surface without changing its alternating crests and valleys will only partially remedy the unfavorable conditions of high wear and friction in the contact area. The optimum contact condition is created by cutting down the crests and transforming them into wide plateaus, a process brought about by microhoning. The effect of microhoning on an originally rough surface is visualized by the schematic cross-sectional diagrams in Fig. 30, which illustrates in five steps (corresponding to the indicated processing times) the manner in which the rugged surface resulting from a preceding machining is improved by changing it into a surface containing wide plateaus.

Fig. 3 0 A rough surface, ( a) , being gradually im proved by m icrohoning, t hat is, by cut t ing down t he prot ruding crest s and finally developing a surface consist ing of wide plat eaus. The elapsed t im e and corresponding surface roughness obt ained at each int erval for t he five cross sect ions are: ( a) 0 s, 1.25 m ( 50 in.) , ( b) 10 s, 0.38 m ( 15 in.) , ( c) 20 s, 0.28 m ( 11 in.) , ( d) 30 s, 0.20 m ( 8 in.) , ( e) 0.050 m ( 2 in.) . The diagram s exaggerat e t he cross- sect ional cont ours of t he work surface for visualizing t he effect of m icrohoning on t he work surface t ext ure.

Pr oce ss Pa r a m e t e r s in M icr oh on in g Type a n d Gr it of Abr a sive St on e s. Both types of abrasive materials are used for microhoning stones:

• •

Aluminum oxide abrasives (which fracture less easily), for carbon and alloyed steel The more friable silicon carbide, for very soft or very tough types of steels, as well as for cast iron and most types of nonferrous metals

The hardness of the stones, controlled by the percentage of the bond, varies from J (very soft), to P (very hard). The former is used for extremely hard alloys, chromium plates, and so on, while the hardest bonds are needed for cast iron and nonferrous metals. The dimensions of the active stone face are determined by the size of the work surface. For external cylindrical surfaces, the width of the stone is about 60 to 80% of the part diameter, but generally not more than about 25 mm (1 in.). For work diameters larger than about 150 mm (6 in.), microhoning heads, with several stones arranged along an arc in compliance with the work surface, are frequently used. The length of the stone is usually somewhat less than the length of the work surface, but not more than about three times the width of the stone. For the microhoning of longer work surfaces, an additional traverse movement is needed. The thickness of the stones is controlled by the mounting dimensions, the access length (which is particularly essential for honing inside shoulders), and an adequate wear depth; this latter component is substantially more, in some applications, than the stone width.

The grit size of the stones is generally selected from a wide range to suit the condition of the machined surface and the objectives of the process. The type of microhoning work required can vary from roughing (applied either as a preparatory operation to a subsequent finishing or as a single process for general-purpose work) to fine, or even very fine finishing. For coarse microhoning applied to improve the finish and the functional properties of rough-machined surface, grit in the range of 60 to 220 is used. For general-purpose microhoning work to produce a finish of the order of 0.40 to 0.20 m (16 to 8 in.) Ra, grit sizes 320 to 500 are needed. Fine finishing is carried out with grit sizes in the 600 to 800 range, while for very fine finishing, stones with 1000 to 1200 grit may be selected. St on e Pr e ssu r e . One of the characteristics of the microhoning method is the use of a relatively low stone pressure to

avoid a deep penetration into the work surface, which could leave furrows and generate heat. The shearing effect of the grains is limited to the protruding elements of the surface, and that action can be achieved by causing the stones to bear against the workpiece with a light force. For average work, stone pressure in the range of 140 to 275 kPa (20 to 40 psi) is generally used, raising the pressure for roughing to about double these values. For very fine finishing, particularly for soft material, stone pressure as low as 14 to 34 kPa (2 to 5 psi) may be applied. W or k Spe e d. Most commonly, microhoning is applied to the surface of rotating workpieces that are of cylindrical and,

less frequently, of tapered shape, referred to as OD. However, flat surfaces, either uninterrupted, such as the end faces of bearing rollers, or of annular shape, such as flange areas, are also adapted to microhoning. Longitudinal flat surfaces, which require linear traverse movement, can also be microhoned, although that method is seldom applied because flat surfaces, unless they are located in a recessed position, are generally adaptable to lapping (see the article "Lapping" in this Volume). The work speed, therefore, is generally specified as the surface speed of the rotating workpiece, with its most commonly applied values falling within the following ranges: • •

For roughing, 12 to 15 m/min (40 to 50 sfm) For finishing, 30 to 60 m/min (100 to 200 sfm)

To produce a very fine finish and also to transverse workpieces at a high feed rate, substantially higher surface speeds, up to about 120 m/min (400 sfm) are also applied. At the lower work speed, the microhoning process generally develops a very fine but still distinguishable crosshatch pattern, which may be the desirable surface condition in many applications, although it reduces the reflectivity or shine of the work surface. At higher work speeds, that pattern disappears, and a brighter surface is developed. St r ok e Le n gt h a n d Spe e d of t h e St on e Re cipr oca t ion . The fast reciprocation of the stones in a short stroke is

one of the essential characteristics that sets microhoning apart from regular honing. The length of the stroke differs somewhat in various systems, some of which employ a single stroke length, for example, 4.76 mm (0.1875 in.), while others are designed to provide variable stroke length, adjustable over a range of about 2.00 to 5.10 mm (0.080 to 0.200 in.). The actual linear speed of the motion of the stone is the function of the stroke length and the rate of reciprocation. These two factors are also referred to as the amplitude and the frequency of stone reciprocation. Stone speed can be adjusted to the requirements of the operation by varying either of these factors. The resulting linear stone speeds differ over a rather wide range according to the system used and the condition of adjustment. Typical extreme values are 3.18 to 20.3 m/min (125 and 800 in./min). M icr oh on in g Applica t ion s Microhoning lends itself to cylindrical (internal and external), flat, tapered, toroidal, spherical, and barrel-shaped surfaces on automotive products such as ball-and-roller bearings, transmission shafts, crankshafts, shock absorber piston rods, valve shaft diameters, and universal-joint spiders. Figure 31 shows two typical setups used to machine two different types

of bearing races. Table 11 lists typical production rates and finishes obtained with microhoning on automotive components.

Table 11 Typical microhoning production rates for automotive components Automotive component

Tappet head Crankshaft Stem pinion bearing Distributor shaft Pressure plate(a) Brake drum(a) Tappet body Camshaft main bearing Gear thrust face Tapered bearing race

(a)

Finish after grinding m in. 0.7630-40 1.00 0.7630-40 1.00 0.3815-25 0.63 0.7630-40 1.00 2.50-5.0 100200 5.0-6.3 200250 0.2510-20 0.50 0.3815-25 0.63 0.2510-20 0.50 1.0040-50 1.25

Number of spindles on microhoning machine

Surface configuration Spherical or flat

Finish after microhoning m 0.125-0.20

1

Production rate, parts/h

in. 5-8

900

10

Cylindrical

0.125-0.20

5-8

80

2

Cylindrical

0.050-0.10

2-4

120

12

Cylindrical

0.075-0.125

3-5

720

2

Flat

0.18-0.30

7-12

100

1

Internal cylindrical

0.38-0.63

15-25

150

12

Cylindrical

0.050-0.10

2-4

800

10

Cylindrical

0.050-0.10

2-4

80

9

Flat

0.050-0.10

2-4

500

2

Cylindrical

0.050-0.10

2-4

450

Starting finishes for pressure plates and brake drums are for turned surfaces rather than ground. In some cases, pressure plates are ground to a finish of 0.50 to 0.63 m (20 to 25 in.).

Fig. 3 1 Set ups for m icrohoning ( a) ball bearing races and ( b) roller bearing races

Microhoning is very effective for finishing ball-shape components to a unique degree of sphericity. Such shapes, executed to an excellent finish, are required in engineering production for valve balls, pump pistons, and so on. An uncommon, but important application of spherical microhoning is in the manufacture of prostheses, specifically, artificial hip joints. The accurate sphericity and very high finish of the ball and socket surfaces, which is accomplished by microhoning these

surgically installed elements, ensures their dependable functioning and capability to withstand load of 1.18 kN (265 lbf) or more.

La ppin g Revised by Pel Lynah, P.R. Hoffm an Machine Product s

I n t r odu ct ion LAPPING is a low-speed low-pressure abrading operation that accomplishes one or more of the following results: • • • • •

Extreme dimensional accuracy of lapped surface (flat or spherical) Close parallelism of double-lapped faces Refinement of surface finish Extremely close fit between mating surfaces Removal of damaged surface and subsurface layers that degrade the electrical or optical properties

In general, the quality that can be obtained by lapping is not easily or economically obtained by other processes. Loose abrasive, carried in an appropriate vehicle, is used on cast iron laps in 99% of the lapping applications, but there are some isolated fixed-abrasive applications that are classified as lapping. It is difficult to make a clear distinction between lapping and honing. Lapping is the lower-pressure, lower-speed, and lower-power application of the use of fixed abrasives. Furthermore, the fixed abrasives of honing are usually limited to the conventional resinoid or vitrified face wheels, while the fixed-abrasive lapping operation often uses unconventional media, such as urethane-impregnated pads, polyvinyl alcohol and abrasive mixed, foamed, and cured to a hard, cellular block, plated or surface-bonded diamond, or thin abrasive-filled vinyl films. The usual definition of lapping is the random rubbing of a part against a lap, usually of cast iron composition, using an abrasive mixture in order to improve fit and finish. Lapping operations usually fall into one of two categories: individual-piece lapping and matched-piece lapping. I n in dividu a l- pie ce la ppin g, abrasive is rubbed against the workpiece with a special tool called a lap (usually of

material softer than the workpiece), rather than with a mating workpiece surface. When loose abrasive is used, the lap is usually made of soft cast iron (typically close-grain cast iron or meehanite metal) or a soft nonferrous metal. Laps made of bonded abrasive also can be used, as discussed above. Individual-piece lapping is most effective on hard metals or other hard materials. It is used to produce optically flat surfaces, to produce accurate planes from which other planes can be located (as for gage blocks), and to finish parallel faces. The machine lapping of individual pieces, either one per cycle or in multiple-piece loads, represents the bulk of the production lapping currently done in industry. Single-side flat lapping machines, double-side planetary machines, cup lapping machines for spherical surfaces, and specialized single- or double-plate machines (such as ball, roller, or pin laps) constitute the vast majority of lapping installations. I n m a t ch e d- pie ce la ppin g, sometimes called equalizing, two workpiece surfaces separated only by a layer of

abrasive mixed with a vehicle are rubbed against each other. Each workpiece drives the abrasive so that the grit particles act on the opposing surfaces. Irregularities that prevent the surfaces from fitting together precisely are thus eliminated, and the surfaces are mated. In many cases, a part is first lapped individually and is then mated with another part by this method, before the two are stocked as a pair of lapped-together parts. Matched-piece lapping enables mating parts (such as the heads and blocks of internal combustion engines) to form liquidtight or gas-tight seals without the need for gaskets. It also eliminates the need for piston rings in fitting some plungers to

cylinders. Other common uses of matched-piece lapping include fitting tapered valve components (Fig. 1) and mating two or more gears in a set.

Fig. 1 Tapered valve com ponent s finished by m at ched- piece lapping for precise fit of m at ing surfaces

Pr oce ss Ca pa bilit ie s Parts that are processed by lapping are constructed of a variety of materials, ranging from metal parts for tooling, gaging, or sealing to electronic crystals such as quartz piezoelectric frequency devices and silicon semiconductor material for integrated circuit manufacture. Tungsten carbide, ceramic, and glass components; aluminum computer disks; tool steel slitter blades; saw blanks; and jade decorative tiles are among the applications that demonstrate the diversity of the lapping process. Th e size or w e igh t of t h e w or k pie ce s that can be lapped is limited only by the available equipment. Parts finished by lapping range in weight from a fraction of an ounce to hundreds of pounds. W or k pie ce Sh a pe . Tools and methods have been devised for lapping virtually every shape of workpiece on which a

lapped surface is desired. Lapping is most widely used for finishing flat surfaces or outside and inside cylindrical surfaces. The process can also be applied to balls, rollers, cones, double-curved surfaces, assembled bearings, and shapes such as gear teeth. M a t e r ia l Re m ova l. Lapping is intended as a final finishing process that would be, in general, an impractical or

uneconomical means of removing stock. In most applications, less than 0.13 mm (0.005 in.) of material is removed from a surface by lapping. However, occasionally (usually in flat lapping), 0.38 mm (0.015 in.) or even more may be removed. In a few cases, it has proved more economical to remove stock by lapping than to add a preliminary grinding operation.

Se le ct ion of Abr a sive Silicon carbide and fused alumina are the abrasives most widely used for lapping. Silicon carbide is extremely hard (2500 HV). Its grit is sharp and brittle, making it nearly ideal as an abrasive for many lapping applications because it continually breaks down to expose new cutting edges. Silicon carbide is used for lapping hardened steel or cast iron, particularly when an appreciable amount of stock is to be removed. Fused alumina (2000 HV) is also sharp, but it is tougher than silicon carbide and breaks down less readily. Fused alumina is generally more suitable for lapping soft steels or nonferrous metals than silicon carbide. Boron carbide (2800 HV) is next to diamond in hardness and is an excellent abrasive for lapping. However, because it costs 10 to 25 times as much as silicon carbide or fused alumina, boron carbide is usually used only for lapping dies and gages, which is often done by hand and in small quantities using little abrasive. An example is synthetic sapphire for electronic applications. The raw material cost is expensive, justifying a high abrasive-processing cost. Relative costs for various quantities and grit sizes of silicon carbide, fused alumina, and boron carbide are compared in Table 1.

Table 1 Relative cost (as of 1988) of abrasives used in lapping Type of abrasive Aluminum oxide

Silicon carbide

Boron carbide

Average size m in 5.0 200 17.5 700 64.0 2520 5.0 200 22.5 900 55 2200 10 400 20 800 40 1600

Grade 1950 1600 1220 2950 2400 2240 3800 3400 3280

Cost $/lb 5 lb 10 lb 3.66 2.99 3.09 2.44 2.11 1.49 11.89 10.99 3.85 3.18 3.07 2.46 47.91 45.94 42.23 40.43 36.63 35.00

25 lb 2.80 2.27 1.34 10.57 3.01 2.26 44.45 39.10 33.84

Diamond (6500 HV), the hardest of all materials is also used as an abrasive for lapping metals. It is available as a paste or a slurry. Softer abrasives, such as emery, garnet, unfused alumina, and chromium oxide, also are used for lapping, but to a far lesser extent. As indicated in Table 2, these softer abrasives are used for lapping soft metals or for the final lapping of parts on which a highly reflective surface is required. In lapping to produce a reflective surface, no significant amount of stock is removed, and the finish is not necessarily finer than a matte finish.

Table 2 Types and grit sizes of abrasives for various applications of lapping Abrasive Relative hardness Grit size All-purpose compounds Hard and sharp 100, 220, 320, 400 Silicon carbide Medium soft 220, 240, 280 Corundum Compounds for roughing, finishing, or polishing Medium soft 400, 500, 600 Corundum 700, 800 Hard 500, 600, 900 Alumina 2-10 m (80-400 in.) Medium hard 900 Alumina 5, 10, 15 m (200, 400, 600 1-3 m (40-120 in.) Soft Alumina 1, 2 m (40-80 in.) Hard and sharp 600, 800, 1000 Silicon carbide Medium soft 600, 800 Garnet 10 m (400 in.) Medium soft 800 Emery Chromium oxide Medium soft 1 m (40 in.) Soft Ferric oxide 1 m (40 in.) Medium hard Cerium oxide 1, 2 m (40, 80 in.)

Typical applications Tool-room lapping Tool-room lapping

in.)

Roughing softer steels Finishing softer steels Roughing harder steels, stainless, chromium plate Finishing hard steels Polishing hard steels Polishing hard steels Polishing stainless, chromium plate Polishing Roughing hardened steels; cast iron Finishing brass, bronze Polishing brass, bronze Polishing softer steels Polishing stainless Polishing soft metals Polishing

Th e gr it size s most commonly employed in lapping range from 100 to 1000 (Table 2). However, abrasives are usually available in grit sizes from about 50 to 3800, and even finer.

For lapping hardened steel to remove about 0.0051 mm (0.0002 in.) of stock and to produce a finish of less than 0.050 m (2 in.), a grit size of 280 is appropriate. If finishing requirements are less stringent, 180-grit abrasive will be more economical because it removes metal faster than finer grit does. As the amount of stock to be removed increases, coarser grits are required. For the removal of considerable amounts of stock, it is more economical to employ a roughing operation, followed by a finishing operation.

When a substantial amount of stock is being removed, a fine finish can be produced without the use of a small grit size, because the originally coarse grit breaks down as lapping proceeds and progressively produces a finer finish. If this technique is used, there must be enough stock allowance for lapping so that the deeper scratches formed initially by the coarse grit will be removed by the time final dimensions are reached. Gr a din g. When an abrasive of a specified grit size is purchased, some of it will be finer and some coarser than the stated size. The degree of grading is an important consideration in the selection of any abrasive. Abrasives increase in cost as the grading becomes closer. However, the use of a low-cost, loosely graded abrasive is not always economical, as demonstrated in the following example.

Ex a m ple 1 : Close ly Gr a de d Abr a sive for Gr e a t e r Econ om y. A low-cost grade of silicon carbide that ranged in grit size from 100 to 800 was used for lapping piston rings. A change to an abrasive that was closely graded to a grit size of 600 reduced the overall cost of abrasive by 50%, even though the initial cost of the 600-grit abrasive was twice that of the low-cost grade. The savings was made possible because the 600grit abrasive contained more of the grit size that is most efficient for lapping; consequently, only one-fourth as much of it was required for removing the same amount of stock. The 600-grit abrasive also gave a smoother finish with less smudging.

Se le ct ion of Ve h icle Vehicles, or binders, for loose abrasives include a wide variety of compounds. Some shops prepare their own formulations or modify standard compositions. However, more consistent results can be obtained with standardized, commercial compounds. Two major factors in the selection of a vehicle are the material being lapped and the lapping method to be employed (inside or outside diameter, flat or spherical). Any vehicle should: • • • • • • •

Retain abrasives in uniform suspension and deagglomeration Serve as a cushion between surfaces being lapped (to minimize lap-to-part contact and yet avoid rolling action on abrasive particles) Adhere to laps and therefore minimize waste of compound Be noncorrosive to the material being lapped Be nontoxic to operators Be easily removable by cleaning Respond to temperature variations with the viscosity characteristics (stability or flexibility) desired in a given application. Although rapid changes in viscosity are usually undesirable, in some applications it is important that the vehicle be able to change quickly from a grease to an oil when under slight heat and pressure and then revert quickly to greaselike consistency when pressure is released

Most vehicles have an oil or grease base, although some are made of water-soluble compounds. The consistency of oilbase vehicles varies from that of mineral seal oil (a water-white product having a viscosity slightly higher than kerosene) to that of heavy grease. Common spindle oil is often used as a vehicle. Commercial compounds contain mixtures of animal fat, vegetable oils, and mineral oils. Vehicles with an oil or grease base are usually used for lapping ferrous metals. For specific applications in which grease or oil would be objectionable (such as copper-base alloys and other nonferrous metals), water-soluble vehicles are available. These vehicles, which are readily removed with water, are low-viscosity compositions of starches, bentonite, and soluble oils with rust inhibitors. Contamination is a potential problem in many nonmetal applications, especially in electronic components; therefore, clean-ability is the primary consideration. Most commercial and proprietary vehicles for these materials are glycerinebase formulations; this provides a good-quality suspension, good film-forming properties (and therefore good lubricity), and a water-soluble mixture that is readily cleanable. Clay or mica is occasionally mixed in to fill the voids between the abrasive particles, thus enhancing the suspension. Ionic, charged, and submicron particles are also used as suspension

agents. These particles affix themselves to the abrasive grain and produce electric repulsion forces to disperse and suspend the abrasive. Pure water is used in the lapping of glass and ceramic. The abrasive slurry is generously applied and recirculated. The resulting high fluid flow is used to keep the abrasive stirred and dispersed.

La ppin g Ou t e r Cylin dr ica l Su r fa ce s Outer cylindrical surfaces are usually lapped by one of the following methods: • • • •

Ring lapping (a manual operation) Machine lapping between plates Centerless roll lapping with loose abrasive Centerless lapping with bonded abrasives

In addition to these techniques, special methods are used for specific applications, such as the lapping of piston rings and crankshafts. Choice of method depends on part configuration, size of the production lot to be lapped, and cost. Many outer cylindrical surfaces can be lapped with equal success by two or more methods. Rin g La ppin g Ring lapping is the simplest method of lapping outer surfaces. Designs of ring laps vary, but the assembly illustrated in Fig. 2 is typical. The lapping ring (or ring lap), usually made of cast iron, is manually stroked back and forth over the workpiece, which is chucked in a lathe or polishing head and rotated. Lapping compound (usually of paste consistency) is often applied to the surface of the workpiece. A manually adjusted screw is tightened, as required, to maintain a slight drag on the lap.

Fig. 2 Typical ring lapping assem bly. Drilled holes and slot s perm it uniform adj ust m ent .

The ring lap should always be shorter than the workpiece, and if size permits, it should have adjustable slots. Finishing the bore of a ring lap is critical. It should be drilled, reamed, and honed (or lapped) to a size very close to the starting diameter of the workpiece; the screw adjustment should be used only to compensate for the slight decrease in workpiece diameter as lapping proceeds.

Applica bilit y. Ring lapping, when performed by a skilled operator, offers at least two advantages over machine

methods: • •

Parts can be produced to extremely close tolerances Out-of-roundness can be corrected to a degree not feasible by machine lapping

Aside from requiring operator skill, however, ring lapping is tedious and expensive, and it should be considered only when one or more of the following conditions prevail: • • • • • •

Equipment for any other method is not available Workpiece is out-of-round Weight of workpiece is unbalanced Workpiece has two or more different diameters that must be lapped Workpiece has flats, keyways, or other interruptions on its cylindrical surface Only a few pieces are to be finished

The following example describes a specific application in which some of the above conditions existed and in which ring lapping was therefore the most suitable method. Ex a m ple 2 : Pr e pr odu ct ion Ve r su s Pr odu ct ion La ppin g. In many cases, ring lapping is used for parts being developed, and a more economical method is used when the parts are in production. This procedure was followed for the valve needle shown in Fig. 3, which was ring lapped in small quantities during development but was machine lapped between plates (see Example 3) in production lots. These needles, in diameters of 6.4 to 9.5 mm (

to

in.), were made of alloy tool steel and hardened to 60 to 65 HRC.

Fig. 3 Valve needle and ring lap for finishing sm all, preproduct ion quant it ies. Dim ensions given in inches

For ring lapping, each needle was chucked by its stem and rotated in a lathe at 650 rev/min. The lap (Fig. 3), which was made of cast iron, was stroked back and forth over the needle until grinding marks were eliminated. The lapping medium with which the needle was coated consisted of chromium oxide mixed with spindle oil. Lapping produced a finish of 0.050 m (2 in.) and maintained tolerances of 0.0013 mm (0.000050 in.) for straightness and 0.00064 mm (0.000025 in.) for roundness. To ensure straightness, the laps used had to be at least three-fourths as long as the area to be lapped. The laps also had to be inspected frequently and had to be reconditioned by being lapped with internal laps of similar material. M a ch in e La ppin g Be t w e e n Pla t e s In the machine lapping of outer cylindrical surfaces between plates, the laps are two opposed cast iron or bonded-abrasive circular plates that are held on vertical spindles of the machine (Fig. 4). The plates are usually 200 to 710 mm (8 to 28 in.) in diameter, although larger sizes are available. For the most part, plain-face laps are used, and for the greatest accuracy,

the width of the lap face should not exceed the length of the surface being lapped. The workpieces are retained between these laps in slotted plates and are caused to rotate and slide. They are given an eccentric, or in-and-out, motion to break the pattern of motion and to ensure that they move over the inside and outside edges of the lap. This prevents grooving of the lap. For short runs, an eccentric motion is unnecessary if the laps are kept flat by reconditioning.

Fig. 4 Typical vert ical lapping m achine for finishing cylindrical surfaces in product ion quant it ies. Dim ensions given in inches

Ca st I r on La ps. When cast iron laps are used, the lower lap is usually rotated and drives the workpieces. The upper lap

is held stationary, but it is free floating so that it can adjust to the variations in workpiece size. The lower lap regulates the speed of rotation because the workholder is not driven. The abrasive is used with a paste-type vehicle and is swabbed on the laps before the cycle is started. Oil or kerosene is then added during the cycle to prevent drying of the vehicle, which may result in scratching. Because the upper lap floats, several parts must be lapped simultaneously. A quantity of three parts will support the upper lap, but when only three parts are lapped, the machine will not produce straightness or a common size. Therefore, it is advisable to lap a minimum of five parts; if this quantity is not available, the machine should be loaded with dummy parts. The best practice is to put as many parts as possible in a load. This reduces the pressure on each part and slows the operation. Thus, the operator has more control and can secure desired tolerances more easily. Finishes as fine as 0.025 m (1 in.), with stock removal of 0.0025 to 0.010 mm (0.0001 to 0.0004 in.), are feasible when cast iron laps are used. Diametral tolerances as low as 0.00050 mm (0.000020 in.), roundness within 0.00013 mm (0.000005 in.), and taper of less than 0.00025 mm (0.000010 in.) have been achieved. However, such accuracy depends greatly on the accuracy achieved in prior machining operations. Bon de d- Abr a sive La ps. When bonded-abrasive laps are used, both laps are rotated, with kerosene or a similar lubricant used as a coolant and to wash away chips or loose abrasive. Because both laps are driven at higher speeds than those used for cast iron laps, the lapping action is more severe. Consequently, the machine will not produce the extreme accuracy possible with machines using cast iron laps. In addition, because bonded-abrasive laps must be dressed with diamond tools, it is not possible to make them as flat as cast iron laps, on which the machines regenerate flatness.

The quantity of parts being lapped is less critical for machines using bonded-abrasive laps than for machines using cast iron laps, because both bonded-abrasive laps are rigidly supported on spindles and separately driven. As few as three parts can be successfully processed in this type of machine. Applica bilit y. Machine lapping between plates is an economical method of finishing outside cylindrical surfaces,

provided its use is warranted by production quantities and is permitted by part configuration. The process can be used for lapping parts a few hundredths of a millimeter to 75 or 100 mm (3 or 4 in.) in diameter and 6 to 230 mm ( to 9 in.) long. Parts commonly lapped by this method include plug gages, piston pins, hypodermic plungers, ceramic pins, small valve pistons, cylindrical valves, small engine pistons, roller bearings, diesel injector valves, plungers, small rolls, and miscellaneous cylindrical pins. Either hard or soft materials can be lapped, provided they are rigid enough to accept the pressure of the laps. Hard materials respond well to lapping and achieve luster. Hard materials are also easier to control for tolerance because the hardness slows the operation. Soft materials lap more rapidly and (especially when bonded-abrasive laps are used) often have a scratchy or dull appearance. This can be prevented by using a polishing abrasive, such as levigated alumina, which reduces the cutting ability of the bonded abrasive. Lim it a t ion s. A part with a diameter greater than its length is difficult or impossible to machine lap between plates. For

parts of this type, other methods of outer cylindrical-surface lapping are more practical. Parts with shoulders require special workholders that permit the shoulder section to be placed on the inside or outside of the lap face. Parts with keyways, flats, or interrupted surfaces are difficult to lap by machine, because the variations in pressure that occur are likely to cause out-of-roundness. If the relief extends over the entire length of the piece, this method of lapping cannot be used. Parts with raised hubs in the middle require special laps that are cut in such a way that they clear the hub. Clearance is necessary between the hub and the work surface to allow for oscillation of the workpiece. Thin-wall tubing can be lapped, but if the walls are so thin that deflection is significant, it will be difficult to maintain roundness. Parts that are hollow on one end but solid on the opposite end present problems in obtaining roundness and straightness, because the hollow end will deflect more under the weight of the upper lap. Plugging the hollow end of the part will sometimes solve these problems. Because it is impractical to keep more than one working surface on the face of cast iron laps flat, workpieces with work surfaces that have different diameters require a separate operation for each surface. It is usually impractical to machine lap workpieces with diameters that are greater than the diameter to be lapped. The outside edges of the plates lap at a faster rate than the inside edges; therefore, care must be taken to prevent the workpieces from becoming tapered. One method of overcoming the problem consists of using short lapping cycles and, at the end of each cycle, turning the workpieces end for end in the slots in the workholder. In addition, the workpieces should be removed from the slots after each short cycle, mixed, and then replaced at random in different slots. This prevents the inadvertent placement of all larger pieces at one side of the workholder and smaller pieces at the opposite side. Because the upper lap floats, this placement of the workpieces would make it difficult to produce accurate parts. Taper can be minimized by positioning the workholder so that the parts in the slots are at a 15° angle to a radius, as illustrated in Fig. 5.

Fig. 5 Set up for lapping product ion quant it ies of t he valve needle shown in Fig. 3

Because machine lapping between plates uses diametrically opposed laps, it cannot correct the out-of-roundness produced by centerless grinding. However, out-of-roundness of the type produced by grinding on centers can be corrected. The following example describes procedures for machine lapping between plates. Ex a m ple 3 : La ppin g Va lve N e e dle s t o Close Tole r a n ce . The valve needles described in Example 2 and illustrated in Fig. 3, although ring lapped in preproduction, were machine lapped between cast iron plates in production. Before being machine lapped, the parts were carefully ground for roundness and then (because the lap would ride on those parts that were largest in diameter) segregated into groups of 0.0025 to 0.005 mm (0.0001 to 0.0002 in.) diametral variation. Both upper and lower laps were grooved to prevent the breakdown of sharp edges during lapping. A laminated phenolic workholder designed to hold a maximum load of parts (Fig. 5) was eccentric to the laps to provide an oscillating motion. In this operation, the cycle was stopped so that the parts could be measured with an electrolimit gage, a visual shadow gage, or an air gage. If the desired size had not been attained, more finish lapping compound was added and lapping was continued. Lapping produced a finish of 0.050 m (2 in.), roundness within 0.00064 mm (0.000025 in.), and straightness of 0.0013 mm (0.000050 in.). To recondition the laps, finish lapping compound was applied to the bottom lap. The laps were then brought together and rotated until the edges of the grooves were sharp. The laps required occasional regrinding to maintain a minimum groove depth of 1.6 mm (

in.) and width of 0.76 mm (0.030 in.).

Ce n t e r le ss Roll La ppin g In centerless roll lapping, only a single piece is processed at a time. Therefore, this method is best suited to the lapping of small quantities of parts (usually, fewer than ten). A typical machine (Fig. 6) consists essentially of two 150 mm (6 in.) long cast iron rolls (one 150 mm, or 6 in., and one 75 mm, or 3 in., in diameter) and a reciprocating device for holding down the workpiece and controlling size. The end of the fiber stick that holds down the work is provided with a 120° V-groove. In operation, abrasive compound is applied to the rolls, and both rolls are rotated in the same direction-away from the operator and counter to the direction of workpiece rotation. The larger roll rotates at about 180 rev/min, the smaller roll at about 90 rev/min. The workpiece feeds across the rolls at about 50 mm/min (2 in./min) as the hold-down device is stroked back and forth to within 13 mm ( end of the workpiece.

in.) of each

Fig. 6 Typical cent erless roll lapping

The rate at which the workpiece feeds depends on the diameter of the piece. For example, if a 13 mm ( in.) diam workpiece feeds at 50 mm/min (2 in./min), a 25 mm (1 in.) diam piece in the same setup will feed at approximately 25 mm/min (1 in./min). Slow stroking is necessary to obtain the best surface finish and control of size. Stock removal in centerless roll lapping is usually 0.005 to 0.0075 mm (0.0002 to 0.0003 in.), depending on the finish obtained in the previous operation. The main advantage offered by centerless roll lapping is quick setup. Therefore, the process is readily adaptable to frequent size changes in short production runs. Limitations on the shape of parts for centerless roll lapping are similar to, but more stringent than, those that apply to machine lapping between plates. When both processes are equally suitable for a given application, the quantity of parts to be lapped determines which will be used. Ce n t e r le ss La ppin g W it h Bon de d Abr a sive s Centerless lapping is a variation of centerless grinding (see the Section "Grinding, Honing, and Lapping" in this Volume). The machines for the two processes are similar in appearance, but the lapping machine is constructed to produce finishes of 0.050 m (2 in.) or better, diametral accuracy of 0.0013 mm (0.000050 in.), and roundness within 0.00064 mm (0.000025 in.). The lapping and regulating wheels are 560 mm (22 in.) wide, which is much wider than those ordinarily used for centerless grinding. Therefore, the work remains in contact with the lapping wheel longer and receives a finer finish. The regulating and the lapping wheels (both are bonded abrasive) can be angled so that their axes are not parallel. Ordinarily, the regulating wheel is adjusted to a positive angle of 1 to 3° (depending on the production and finish requirements), and the lapping wheel is adjusted to a negative angle of about -4°. When trued, both wheels assume a slight hourglass shape, which then allows them to wrap around the workpiece as it passes between them. They also contact the workpiece at an angle to its axis, which is different from the axial-line contact of a grinding wheel. This eliminates lapping marks. The finest finish obtainable in a centerless lapping machine requires at least three operations, each with a progressively finer lapping wheel, and a full flow of clean fluid (such as kerosene) as a coolant. During the first operation, the workpiece is supported on a blade faced with hard steel or carbide. For correcting out-of-roundness, the center of the workpiece should be slightly above the center of the wheels. In the first operation, a maximum of 0.013 mm (0.0005 in.) of stock is removed, and a finish of 0.10 to 0.15 m (4 to 6 in.) is obtained. For the second and third operations, the workpiece is supported on a rubber blade and is centered on the wheels so that scratches are minimized. During the second operation, a maximum of 0.0025 mm (0.0001 in.) of stock is removed, and a finish of 0.050 to 0.075 m (2 to 3 in.) is obtained. During the third operation, practically no stock is removed, and a finish of about 0.050 m (2 in.) is obtained. Applica bilit y. Centerless lapping is a high-production operation that is particularly suited to centerless ground parts that

can be continuously fed, either manually or automatically. Parts 6 to 150 mm (

to 6 in.) in diameter by 380 mm (15 in.)

long can be centerless lapped, and when a long bar feed is used, it is possible to lap parts 13 to 75 mm ( to 3 in.) in diameter and 4.6 m (15 ft) long. Typical parts finished by centerless lapping are pistons, piston pins, shafts, and bearing races. Because little stock is removed in this process, only a small amount of correction can be made. Therefore, parts must be previously ground to the required straightness and roundness. Parts with shapes that have no irregularities are ideally suited to centerless lapping, but irregularities such as those on the part shown in Fig. 7 can be tolerated. Such parts may, however, present problems in holding tolerances because of the undercut and the keyway. Cross holes also add to the difficulty of holding extremely close dimensions in centerless lapping. The production rates attained in centerless lapping and in two other processes for the part illustrated in Fig. 7 are compared in the following example.

Fig. 7 Part lapped by t hree different m et hods. Dim ensions given in inches

Ex a m ple 4 : Pr odu ct ion Ra t e s for Ce n t e r le ss Ve r su s Ce n t e r le ss Roll Ve r su s Tw o- Pla t e M a ch in e La ppin g. The part shown in Fig. 7 (52100 steel hardened to 61 to 63 HRC) was lapped by three methods for the removal of 0.005 mm (0.0002 in.) of stock to produce a finish of 0.025 m (1 in.). Productivity was as follows:

Method of lapping Centerless Centerless roll Two-plate machine

Pieces per hour 700 10 100

Centerless lapping was done with bonded-abrasive wheels (355 mm, or 14 in., in diameter and 560 mm, or 22 in. long; grit size: 500) at a rotation speed of 52 rev/min. A blade fixture was used. Centerless roll lapping was carried out in a machine with two 150 mm (6 in.) long cast iron rolls--one 150 mm (6 in.) in diameter and one 75 mm (3 in.). The 1000-grit abrasive used was contained in a paste vehicle. Rotation of the smaller roll was at 100 rev/min. In machine lapping between plates, the two laps, which were of cast iron, were 400 mm (16 in.) in diameter and 75 mm (3 in.) thick. The spider fixture was used. Rotation speed was 100 rev/min, and an 800-grit abrasive was used in a water vehicle.

La ppin g of Ou t e r Su r fa ce s of Pist on Rin gs

Special procedures are required for lapping the outer cylindrical surfaces of parts that are considerably greater in diameter than in axial length. Piston rings are typical of such parts. Lapping is especially necessary for a chromium-plated piston ring because the ring will quickly ruin a cylinder unless the minute chromium nodules are removed. Because of slight variations in machining, rings may have areas that exert low pressure. During lapping, material will be removed faster from the high-pressure areas, thus causing a more even distribution of pressure around the ring, removal of chromium nodules, and smoothing of the surface. A specific procedure employed in the lapping of piston rings is described in the following example.

Ex a m ple 5 : La ppin g Eigh t Pist on Rin gs Sim u lt a n e ou sly. The stacking setup shown in Fig. 8 was used for simultaneously lapping eight chromium-plated steel piston rings (hardness: 775 HV). The lap consisted of an outer cylinder that was a solid casting and an inner sleeve that could be replaced when worn out. Replacing only the inner sleeve was more economical than replacing the entire cylinder. The piston rings were reciprocated in the sleeve at 150 cycles/min by a special machine. During each reciprocation, the stack of piston rings was rotated 45°. The abrasive, which was fed in through slots near the center of the sleeve, contained 10% of 600-grit aluminum oxide mixed with 90% (by weight) of a commercial lapping oil.

Fig. 8 Set up and fixt ure for lapping eight pist on rings sim ult aneously

This practice was used for rings 51 to 216 mm (2 to 8 in.) in diameter and 1.6 to 6.4 mm ( to in.) thick. Productivity was 2 to 48 rings/min, depending on ring size. Metal removal ranged from about 0.025 to 0.038 mm (0.0010 to 0.0015 in.). Size, which was controlled by the number of cycles, was checked by measuring the gap between the ends of the ring when installed in a gage.

La ppin g of Cr a n k sh a ft s Crankshaft journals and pins and a variety of similar cylindrical surfaces are often lapped when they require a finish better than that ordinarily produced by production grinding. The pin and journal surfaces on crankshafts, for example, are ground to finishes of 0.63 to 1.40 m (25 to 55 in.), but a finish of 0.10 to 0.20 m (4 to 8 in.) is required for some applications. This finer finish has been inexpensively achieved by the method described in the following example, in which all surfaces are lapped in one setup.

Ex a m ple 6 : M a ch in e La ppin g of Cr a n k sh a ft s.

Crankshafts were lapped in a machine with a 355 mm (14 in.) swing and 810 mm (32 in.) between centers. Work shoes automatically clamped coated-abrasive paper or cloth (fused alumina, 240 to 320 grit) around the crankshaft surfaces to be lapped (Fig. 9). Mineral seat oil was constantly applied to cool the surface being lapped. At a rotation speed of 125 rev/min and a reciprocation of 80 cycles/min, about 0.005 mm (0.0002 in.) of metal was removed to produce a finish of less than 0.25 m (10 in.). Production rate was 70 crankshafts per hour.

Fig. 9 Set up for lapping a cylindrical surface on a crankshaft

La ppin g of I n n e r Cylin dr ica l Su r fa ce s Holes or bores are lapped by using either solid or adjustable laps, usually made of cast iron. The laps can be rotated by any one of a variety of machines, including honing machines, but lathes or polishing heads are most commonly used. The lap, which carries the abrasive, is rotated while the workpiece is manually stroked over it, a procedure similar to manual honing. Machines that reciprocate either the workpiece or the lap (usually the lap) in addition to rotating the lap also are used. These machines resemble those used for power stroking in honing (see the article "Honing" in this Volume). In internal-surface lapping, virtually no stock is removed because stock can be removed at lower cost by honing. When lapping follows honing, it is usually just a touch-up operation. A variation of internal-surface lapping is employed for matching or mating male and female cylindrical components. For example, fuel injection plungers are often lapped into cylinders to produce matched pairs. In these applications, the male part, which rotates, becomes the lap. It is swabbed with abrasive, and the mating part is then manually stroked over it. Skilled operators are required for these operations. Adj u st a ble la ps (Fig. 10) are available in almost any diameter larger than 1.6 mm ( in.). These laps are expanded manually, as required, during operation. Although various means can be used for expanding the lap, adjusting screws (Fig. 10) are most often used. Ring gages are examples of parts that are lapped to extremely accurate dimensions with adjustable laps. Tolerances specified for various diameters and classes of gages are given in Table 3. To obtain the accuracy required by the specifications in Table 3, it is necessary to adhere strictly to the following sequence of operations in lapping:

•

Determine the initial condition of the workpiece (amount of stock to be removed, and taper and

• • • • • • • • •

roundness) Insert the shank of the lap into the polishing-head chuck Coat the lap with abrasive compound Loosen the screw in the lap Place the workpiece on the lap Tighten the screw in the lap until the workpiece turns with slight resistance Start the machine and reciprocate the workpiece evenly by hand over the length of the rotating lap, maintaining uniform resistance, until the piece runs smoothly from end to end Stop machine, remove the workpiece and allow it to cool Measure the workpiece to check for size, taper, barrel or hourglass shape, out-of-roundness, and bellmouth Repeat the process, except for the method of reciprocation, until the desired results are obtained. At this point, the method of reciprocation depends on the conditions found in measuring the ring. Over-size and out-of-roundness require slow, even reciprocation. Bellmouth requires reversal before the point of lowered resistance is reached. Barrel or hourglass shape and taper require faster reciprocation, localized at the high areas. Extremes of the above (0.005 mm, or 0.0002 in.) must be corrected by grinding

Table 3 Lapping tolerances for various sizes of ring gages Diameter mm 0.737-20.96 20.96-38.35 38.35-63.75 63.75-114.55 114.55-165.35 165.35-228.85

in. 0.029-0.825 0.825-1.510 1.510-2.510 2.510-4.510 4.510-6.510 6.510-9.010

Tolerance for gages of class: X Y mm in. mm 0.0010 0.000040 0.0018 0.0015 0.000060 0.0023 0.0020 0.000080 0.0030 0.0025 0.000100 0.0038 0.0033 0.000130 0.0048 0.0040 0.000160 0.0061

in. 0.000070 0.000090 0.000120 0.000150 0.000190 0.000240

Z mm 0.00254 0.00305 0.00406 0.00508 0.00635 0.00813

in. 0.000100 0.000120 0.000160 0.000200 0.000250 0.000320

Fig. 1 0 Com ponent s of an adj ust able lap

Solid la ps are low-cost round bars that are accurately finished to size. They are usually made of cast iron, but for

extremely small workpieces, they can be of copper or other nonferrous metal. Ordinarily, solid laps are used only when one or two odd-size pieces require finishing. However, solid laps must be used for finishing inside diameters of less than 1.6 mm ( in.) (because this is the minimum diameter of adjustable laps), and they can also be used by skilled operators to correct bow or snake. The chief disadvantage of solid laps is that they are useless once they become only slightly worn, unless they are refinished for lapping smaller bores. Blin d H ole s. Lapping blind or partly blind holes presents problems because of uneven distribution of abrasive. The

difficulty can sometimes be overcome by an improvement in lap design, as in the following example.

Ex a m ple 7 : La ppin g Blin d H ole s.

The part shown in section in the lower portion of Fig. 11 was made of 52100 steel and hardened to 60 to 62 HRC. Parts of this type had main bores 6 to 9 mm ( to in.) in diameter and 25 to 38 mm (1 to 1 in.) in length. The bores were originally lapped with expansion laps made of cast iron and slotted in a straight line (Fig. 11). This type of lap did not allow enough lapping compound to reach the bottom of the bore; consequently, lapping action at the outer end of the hole was more rapid, and excessive taper was produced.

Fig. 1 1 Subst it ut ion of spiral for st raight slot in lap for bet t er dist ribut ion of abrasive in lapping a blind hole. Dim ensions given in inches

The problem was solved by redesigning the lap, providing it with a right-hand spiral that formed a carrier for the lapping compound (Fig. 11). The redesigned lap was held in a speed lathe (horizontal chucking head) and rotated clockwise at about 650 rev/min; the part was rotated counterclockwise. A coarse abrasive (grit size: 600) was used. Lapping removed 0.005 mm (0.0002 in.) of stock at a rate of 40 pieces per hour. The bores were held to a diametral tolerance of 0.0013 mm (0.000050 in.) and to straightness within 0.00064 mm (0.000025 in.). Size and straightness were controlled by the use of an air gage and two master setting rings.

La ppin g Fla t Su r fa ce s Flat surfaces can be lapped by either manual or mechanical methods. In general, manual methods are used only when small quantities of parts are to be lapped or when special requirements must be met. Most flat lapping on a production basis is done with rotating two-plate machines similar to those used for lapping cylindrical surfaces. M a n u a l M e t h ods The hand rubbing of flat workpieces on a block or plate lap charged with loose abrasive is the simplest method of flat lapping. The lap, usually made of cast iron, has regularly spaced grooves about 1.6 mm ( in.) deep to retain the lapping medium (usually an abrasive-containing paste). The flat workpiece is rubbed on the lap in a figure eight or similar motion that covers almost the entire lap surface, so that the lap will remain flat for a considerable amount of work. This method of lapping is slow and tedious and requires a high degree of skill for optimum results. It is used only when a few parts must be lapped or when more efficient equipment is not available. Another and somewhat faster method of flat lapping single pieces makes use of a single-spindle vertical drill press. The lap, which is stationary, is mounted on the stand of the drill press. The workpiece is held by the spindle, which rotates it against the lap. Light pressures are applied by hand. This method is slow, and the lap is likely to wear unevenly. However, it may be preferable to other flat lapping methods for some applications. For example, certain round, flat

sealing parts require a lapped surface with the concentric-line pattern produced by this method because random scratch patterns are not leakproof. To achieve the concentric-line pattern, the rotating part is held against a rotating or nonrotating lap of the same size so that there is no motion except the rotation between the workpiece and the lap. When concentricity of scratch pattern is not important, the method described above is sometimes modified for faster lapping by oscillating the rotating workpiece back and forth over the lap. The workpiece should be mounted on the spindle of the drill press so that it is flexible and can float as it turns. A layer of rubber can be placed behind the workpiece to provide a suitably flexible mounting. Care should be taken when lapping for concentric patterns because this process violates one of the basic rules of lapping, namely, the random and nonrepeating motion between the lap and the work. Without this random motion, a small scratch passing over the same spot soon becomes a deep groove, with the removed material picked up and welded to the lap. M e ch a n ica l M e t h ods The two general types of machines for flat lapping are: • •

Single-face lapping machines having a single horizontal rotating lap Dual-face or four-way planetary lapping machines having two laps, one above or on top of the work and one below

Sin gle - Fa ce La ppin g M a ch in e s. The simplest single-face lapping machine is the horizontal rotating annular lap, with (usually) three or four conditioning rings held in place on the lap face but free to rotate (Fig. 12). The dynamic action of this system drives the rings in the same rotational direction as the lap (Fig. 13). When abrasive is applied to the lap, the rings and the lap wear into intimate contact and are of ball-and-socket shape. This is the fundamental operation of the single-face lap. If the rings are manipulated so that the lap is alternately the ball and the socket, then the lap flatness, and consequently the work flatness, can be held to within the required flatness tolerance.

Fig. 1 2 Typical horizont al single- face flat lap lapping m achine wit h t hree condit ioning rings. Ring posit ion flat t ening is achieved by using gravit y pressure. Court esy of P.R. Hoffm an Machine Product s

Fig. 1 3 Dynam ics of t he lapping m achine shown in Fig. 12. Work t o be lapped is placed wit hin t he condit ioning rings, which are held in place but are free t o rot at e. The work t ends t o abrade t he lap plat e, but t he rot at ing act ion of t he condit ioning rings causes t he lap plat e t o wear evenly, m aint aining a flat surface. St andard m achines handle part s from 3.2 t o 810 m m ( t o 32 in.) in cross sect ion. St eel, t ool st eel, bronze, cast iron, st ainless st eel, alum inum , m agnesium , brass, quart z, ceram ics, plast ics, and glass can be lapped on t he sam e lap plat e.

The manipulation of the rings usually follows one of two methods. If the ring is positioned with the center of the ring on the center of the plate track, the result will be a lap socket and ring ball. This is because the relative velocity between plate and ring, with both turning in the same direction, is higher at the lap inside diameter than at the outside diameter, thus wearing the center faster and creating a socket or bowl shape. If a tooth form is generated on the outside diameter of the conditioning ring and a corresponding driver is mounted at the lap center in mesh with the teeth, then the conditioning ring will be forced to rotate counter to the lap, the relative velocity ratio will be reversed, and the lap will wear in a ball shape. An alternative method for reversing the wear pattern is to shift the ring position toward the lap outside diameter. As the ring overhangs the plate, higher unit pressure is exerted on the lap outside diameter than on the inside diameter. This pressure increase causes increased wear, which exceeds the velocity wear and causes the lap to become ball shaped. With either type of machine, the rings act as tooling. The parts are placed in the ring or into carriers, which are then placed in the rings. For parts that are too large to fit in the rings, all but one can be removed, and fixtures can be made to hold the parts on the lap. The flatness is then controlled with both the ring and the parts, using the pressure flatting concept previously discussed. Various options are available to enhance the capability of this machine: • • • • • •

Pressure heads to apply down pressure to the parts through a pressure platen, which loosely fits the inside diameter of the rings Abrasive systems that automatically apply a predetermined rate of slurry flow Variable-speed drive to the lap Variable-speed driven rings Plate cooling Sizing devices

A variety of mechanical and electronic gaging systems permit the in-process monitoring of stock removal rates and workpiece thickness. These systems are frequency linked to the lapping machine so that a specific depth of material is removed and the machine stops when the target thickness is reached. These gaging systems include timers, hardened stops, electronic devices, and piezoelectric devices.

Tim e r s are extensively used when the process is rigidly controlled and repetitive parts are produced. Because the

removal rate is very constant, final size can be held closely, especially if the raw part sizes are in a narrow band. H a r de n e d st ops are used in conjunction with the timer to space the pressure plates, thus stopping the lapping action. Ele ct r on ic de vice s are used to measure the drop of the pressure plate. This is an indication of the material removed. Pie zoe le ct r ic D e vice s. The most unique sizing method is used on thin material and involves placing a piezoelectric wafer in with the parts. The wafer should be of the same thickness as the parts being lapped. As the wafer is lapped together with the parts being machined, the size is monitored by reading the frequency of the piezoelectric wafer. As an electrical signal is passed through quartz or similar piezoelectric materials, the oscillation frequency of the crystal changes as the lapped wafer decreases in thickness. This phenomenon is exploited by using in-process gaging to stop the lapping process at a fixed frequency/thickness set point. In a production environment, size control of ±0.0013 mm (±0.000050 in.) has been achieved using this method. D u a l- Fa ce La ppin g M a ch in e s. Most of the dual-face lapping machines being produced are of the planetary type, with

the workholders (carriers) nested between a center drive and a ring drive (Fig. 14). These drives can be either gear- or pintype configurations, but must have positive engagement.

Fig. 1 4 A 355 m m ( 14 in.) planet ary fixed- plat e dual- face lapping m achine. The t op plat e has been rem oved t o expose t he cent er drive gear, ring drive, and part s carriers. Court esy of P.R. Hoffm an Machine Product s

The parts being worked are propelled by the carrier in a serpentine path between two flat lap plates, on which abrasive has been placed (charged) or is continuously fed in the form of a slurry. In the simplest form, the bottom lap is fixed, and the top lap is restrained from rotating but is allowed to float so that it always bears on the largest pieces and laps all the pieces to the same size. In a machine of this type, the part is dragged between the plates by the carrier, and all the lapping power is directed into the system through the center (sun) and ring drives. Th e pla n e t a r y fix e d- pla t e m a ch in e has a decided disadvantage in that no successful method has been developed to

flat the plates in the machine while processing parts. An additional disadvantage is that the carrier forces are high, thus causing teeth failure of thin carriers and edge chipping on fragile parts. Th e t op pla t e dr ive m a ch in e was developed to eliminate the need to remove the plates to grind them flat (Fig. 15). In

a machine of this type, the top plate is driven twice as fast in the same direction as the carriers translate the bottom plate. As a result, the removal from the top and bottom of the work is equal, but the top and bottom plates wear as mirror images of each other. If the bottom plate is concave, the top plate is convex, and vice versa. Using variable speed on the

ring and center drive, the operator can work the flatness of the bottom lap plate in the same manner as the single-face lap and keep both top and bottom laps flat within a required tolerance.

Fig. 1 5 A 1220 m m ( 48 in.) planet ary t op plat e drive dual-face lapping m achine loaded wit h a variet y of com ponent s having various configurat ions and t hicknesses t o dem onst rat e t he processing capabilit ies of t he equipm ent . Shown clockwise, st art ing wit h t he bot t om - m ost carriers, are t he hardened high- carbon st eel shear blades; 102 by 102 m m ( 4 by 4 in.) phot om ask blanks used t o proj ect m icrocircuit s ont o elect ronic subst rat es; st ainless st eel sender bars, which will be et ched int o opt ical posit ion indicat ors; valve plat es m achined from cut off hot - rolled bar st ock for use in hydraulic m ot ors; and a hardened alloy st eel rot ary slit t er blade.

The machine shown in Fig. 15 is a 1220 mm (48 in.) double-face lap that has been loaded with sample parts appropriate for this type of processing. It should be noted that this machine is loaded in this manner for demonstration purposes only. W or k pie ce Size a n d Sh a pe Parts 3.2 mm ( in.) in diameter or smaller and as large as 810 mm (32 in.) in diameter can be flat lapped. Table 4 indicates the number of parts of given diameters that can be accommodated by single-face flat lapping machines with laps 305 to 2135 mm (12 to 84 in.) in diameter, in conjunction with various numbers and sizes of carrier rings. In addition, machines having plate diameters of 2440, 2690, 3050, and 3660 mm (96, 106, 120, and 144 in.) are available. Large quantities of small parts can be lapped on large machines. Large parts can be lapped only on large machines.

Table 4 Load capacities of planetary flat lapping machines Workpiece diameter(a)

915 mm (36 in.) diam laps; three

910 mm (36 in.) diam laps; four

6.4

Production capacity (quantity of parts) 305 mm (12 in.) 610 mm (24 610 mm (24 in.) diam laps; three in.) diam laps; diam laps; four three 248 mm 108 mm (4 in.) 210 mm (8 in.) in.) ID ID conditioning ID conditioning (9 rings conditioning rings rings Each Full Each Full Each Full ring load ring load ring load 235 705 1300 3900 920 3680

368 mm (14 in.) ID conditioning rings Each Full ring load ... ...

321 mm (12 in.) ID conditioning rings Each Full ring load ... ...

Each ring ...

13

55

165

310

930

220

880

700

2100

560

2240

19

23

69

130

390

90

360

300

900

245

13 8

39 24

70 45

210 135

55 32

220 128

165 105

495 315

5

15

31

93

22

88

70

2 1 1 ... ... ... ... ... ... ...

6 3 3 ... ... ... ... ... ... ...

17 7 4 1 1 ... ... ... ... ...

51 21 12 3 3 ... ... ... ... ...

12 5 2 1 1 ... ... ... ... ...

48 20 8 4 4 ... ... ... ... ...

39 16 9 4 1 1 ... ... ... ...

mm

25 32 38 50 75 100 150 200 250 360 380 400 360-660

(a)

in.

1 1 1 2 3 4 6 8 10 14 15 16 14-26

1220 mm (48 in.) diam laps; four 432 mm (17 in.) ID conditioning rings

1830 mm (72 in.) diam laps; four

2135 mm (84 in.) diam laps; four 813 mm (32 in.) ID conditioning rings

Full load ...

692 mm (27 in.) ID conditioning rings Each Full ring load ... ...

Each ring ...

Full load ...

...

...

...

...

...

...

980

...

...

...

...

...

...

130 80

520 320

235 145

940 580

620 390

2480 1560

... ...

... ...

210

55

220

100

400

270

1080

375

1500

117 48 27 12 3 3 ... ... ... ...

31 13 7 3 1 1 ... ... ... ...

124 52 28 12 4 4 ... ... ... ...

55 23 13 5 2 1 ... ... ... ...

220 92 52 20 8 4 ... ... ... ...

145 60 34 14 8 5 ... ... ... 1

580 240 136 56 32 20 ... ... ... 4

205 85 48 21 11 7 3 2 2 ...

820 340 192 84 44 28 12 8 8 ...

When using this table for determining capacity of machine with work separator, use the value for diameter next larger than that of the workpiece to be processed. Conditioning rings are carrier rings.

With the correct workholders and operating conditions, parts thinner than 0.05 mm (0.002 in.) have been lapped. Workpiece height limitation is determined by machine clearances and by the ratio of height to lapped surface area. This ratio must be kept low enough to prevent tipping or rocking of the part during lapping, unless a workholder or fixture is provided. The restraint or support required for typical shapes that can be flat lapped is indicated in Fig. 16.

Fig. 1 6 Part s requiring different am ount s of rest raint or support in flat lapping. Sym m et rical com ponent s sim ilar t o ( a) and ( b) do not require workholders for flat lapping. Asym m et rical com ponents sim ilar t o ( c) and ( d) require workholders t o keep t hem separat ed from each ot her. Part s sim ilar t o ( e) require workholders t o keep t hem from t ipping.

Fla t n e ss Flat lapping can produce flatness within one light band (0.29 m, or 11.6 in.) and smoothness better than 0.050 m (2 in.). In some optical applications, flatness is better than one-tenth light band. To obtain these results, the workpieces must be stable during the operation so that the possibility of stress relief that will result in distortion is eliminated. In addition, flatness of the lap or laps must be kept within the flatness tolerance for the workpiece. Size Tole r a n ce a n d Pa r a lle lism Parts having parallel faces, such as disks and seal rings, can be held to tolerances varying from ±0.0025 mm (±0.0001 in.) for small parts to ±0.025 mm (±0.001 in.) for large parts. The difficulty in maintaining accuracy increases for parts of uneven configuration; such parts may require fixtures, which then determine the accuracy attainable. The closer the tolerance requirements, the more difficult the operation is likely to be. Surfaces of disks, seal rings, and similar parts having parallel surfaces can be lapped on either the double-lap machines, which lap both sides in one operation, or the single-lap machines, which require two operations (one for each side). The single-face lap can be used to produce parallel parts, but is very skill sensitive and requires attention to such details as cleanliness, lap and pressure plate flatness, and the use of multiple transpositions. Parallelism requirements of less than 0.0002 mm/mm (0.0002 in./in.) of diameter or length in production lots dictate the use of a dual-face machine. These machines, with adherence to a process specification, can produce parts with opposing faces parallel within 0.00002 mm/mm (0.000020 in./in.). The operation becomes simpler as the number of parts to be lapped at one time increases and as the distribution of thick and thin parts improves. When only a few workpieces are lapped, or if the thickness of the part varies more than about 0.075 mm (0.003 in.), it may be necessary to distribute the thicker parts selectively or to remove and redistribute the parts

before lapping the second side. Allowance for stock removal in this type of operation should be 1 that the parts are out-of-parallel plus the amount of the variation in part sizes.

to 2 times the amount

Pr oce du r e Flat lapping is typically employed on piston rings when side flatness is extremely critical and on other ringlike parts. Although most ring-shaped parts can be lapped without elaborate procedures, special problems often arise. There are two main problems in lapping piston ring sides: • •

The rings are flexible and can be easily distorted under load Some piston rings have a very smooth, polished finish on the circumference that must not be marred during lapping

Some load must be applied during lapping because the rings are too light to lap without added load, but the applied load must not be great enough to distort the ring. An aluminum pressure plate is used because it is light but also rigid enough to stay flat. In the modern dual-face lap, parameters such as the applied pressure due to the weight of the top plate are computer controlled, and the machines can be easily set up to lap parts requiring specialized processing. In the two examples that follow, procedures successfully employed for the flat lapping of a ringlike part (Example 8) and a more complex configuration (Example 9) are described in detail. Ex a m ple 8 : La ppin g of Be a r in g Ra ce s. The setup shown in Fig. 17 was used for lapping the flat sides of inner and outer bearing races made of carburized or through-hardened steel in diameters from 30 to 215 mm (1 to 8 in.). The machine had two bonded-abrasive laps (400-grit silicon carbide) that rotated in opposite directions at 88 rev/min. The head of the machine, which was like that of a vertical drill press, was air actuated. It raised and lowered the top lap and could apply adjustable downward pressure (up to 414 kPa, or 60 psi) to the top lap. The bearing races were hand loaded in a horizontal circular fiber carrier that was eccentrically mounted over the bottom lap. The carrier rotated at 7

rev/min, and was adjusted so that eccentricity was

up to 25 mm (1 in.) and part clearance was about 6.4 mm ( in.). A viscous cutting oil was fed to the laps during the operation, and the laps were dressed two or three times during an 8-h shift. A maximum of 0.013 mm (0.0005 in.) of stock was removed from the races, leaving a finish of 0.075 m (3 in.). When checked with an indicator (0.0025 mm, or 0.0001 in., per division), no deviation from parallelism was detected.

Fig. 1 7 Set up for lapping inner or out er bearing races

Ex a m ple 9 : Fla t La ppin g of Sw in g Ch e ck Va lve W it h Spe cia l La ps a n d Fix t u r e s. The seats of swing check valves were lapped to the required flatness using the setup and relatively simple fixture shown in Fig. 18. This type of setup can be used for lapping similar offset flat surfaces. Cast iron laps were used, along with loose abrasive. The portion of the workplace that protruded beyond the lapped surface extended over the inside or outside diameter of the lap and was loosely held in position by a cutout nesting fixture. To apply additional weight on the uneven back surfaces of the workpieces, rubber buttons on the underside of a pressure plate were placed so that they rested on the high points of the workpieces when lowered into lapping position. The pressure plate (not shown in Fig. 18) was indexed so that the buttons always rested in the correct position on the workpieces.

Fig. 1 8 Set up for lapping seat s of swing check valves

La ppin g En d Su r fa ce s End lapping and conventional flat lapping are similar in principle, and the same machines are often used for both processes. As in flat lapping, fixtures are often necessary to prevent workpieces from tilting. In many cases, a workpiece whose height is several times the width of its lapped area is supported in a special fixture and processed in standard flat lapping machines. Engine connecting rods are notable examples. Faces that hold the shims for the connecting rod caps in assembly are supported by means of specially designed fixtures, and they are forced, by weight or air pressure, against a rotating lap. Their mating caps are often lapped at the same time. For straight cylindrical parts (not tapered), fixtures for end lapping are simple, and various methods can be used, as in the following example.

Ex a m ple 1 0 : En d La ppin g of St r a igh t Cylin dr ica l Rolle r s. A special machine was used for the high-production end lapping of rollers. Figure 19 shows the important components of this machine, which included an enclosed track that allowed the rollers to travel in a loop. The rollers were thus exposed to the bonded-abrasive wheels twice per trip around the track. The number of cycles required varied with the initial condition of the rollers. The laps, which were commercial vitrified wheels, rotated in the same direction at 1350 rev/min and were cooled with a low-viscosity honing oil.

Fig. 1 9 Set up for end lapping large quant it ies of hardened st eel rollers

Pr oble m s in Fla t a n d En d La ppin g The problems encountered in flat and end lapping (and often in other methods of lapping, as well) are usually related to surface roughness or scratches, drop-off at edges, low rate of stock removal, improper size control, and failure to attain desired flatness and parallelism. These problems can be solved through proper control of the following conditions: • • • • • • •

Preparation of the workpiece Selection and application of the abrasive and the vehicle Sweep path of the workpiece over the lap Condition of the lap surface Interchange or transposition of workpieces during the lapping cycle Duration of the lapping cycle Cleanness of the environment

Pr e pa r a t ion of t h e W or k pie ce . Parts to be lapped must have smooth, regular edges because irregularities at the edge

cause the abrasive grains in the lapping compound to burst free and form the foxtail pattern shown in Fig. 20(a). (This effect may result also from etching on the lapped surface.) To provide the smooth edges, it may be necessary to barrel finish workpieces in fine abrasive, or to buff them, prior to lapping. Glass parts are frequently beveled on automatic machines, and the edges of electronic materials are rounded off by grinding prior to lapping. This not only eliminates foxtailing but also presents a much stronger edge to the driving carrier, thus minimizing edge chips and scratches.

Fig. 2 0 Two undesirable consequences of im proper preparat ion of workpieces for lapping. See t ext for discussion.

Workpieces that have drop-off at edges, from previous operations, are subject to hydraulic cut in lapping. In hydraulic cut (Fig. 20b), the lapping abrasive builds up in the area of drop-off and perpetuates the condition. If drop-off cannot be eliminated, hydraulic cut can be minimized by the use of lesser amounts of abrasive in lapping. This remedy is usually uneconomical, however, because the stock removal rate decreases when a lesser amount of abrasive is used, and the likelihood of glazing is increased. Glazing can be prevented by the use of a diamond-base abrasive that contains aluminum oxide as the main cutting agent. The diamond abrasive should be of smaller grit size than the aluminum oxide; m (10 in.) diamond has been successfully used with 1200-grit aluminum oxide. Metal particles must be removed from the surfaces of workpieces to be lapped so that the laps do not become charged with these particles and scratch the workpiece. Workpieces that have been ground should be freed of metal particles by being barrel finished in fine abrasive before being lapped. Previously lapped workpieces also may have particles adhering to the surface. To remove these particles, the workpieces must be rubbed on a dead lapping block (a lightly charged, noncutting block). Abr a sive s a n d Ve h icle s. Coarse finish, scratches, loss of flatness or parallelism, and inadequate or excess stock removal rate often result from the use of improper abrasive or vehicle or from improper method of application of the lapping medium. In the sections "Selection of Abrasive" and "Selection of Vehicle" in this article, the basic considerations in the choice of these media for a given application are discussed in detail. Excessive amounts of lapping abrasive cause dull finishes, loss of size control, edge drop-off, and low rates of stock removal. Abrasive films that are too thin react similarly, except that they produce bright finishes and may result in glazing. Se le ct ion of Sw e e p Pa t h . The sweep path illustrated in Fig. 21(a) is advantageous for rectangular pieces with height-

to-area ratios greater than 2 to 1. One of the longer edges of the lapped area should be the leading edge. To keep the lap flat, the carrier must be positioned so that part of a workpiece sweeps over the outside and the inside of the lap face. The eccentric throw of the carrier must be reduced when the workpiece is large in relation to the width of the lap face; thus, practically straight-line lapping occurs. As a result, it is difficult to keep the lap flat, to control the size of the workpiece, and to obtain parallelism. The best lapping action is obtained when every new grain cut line crosses every previous grain cut line at an angle.

Fig. 2 1 Typical sweep pat hs of workpiece over lap face in flat lapping. See t ext for discussion.

The path shown in Fig. 21(b) is advantageous for small rectangular pieces over 1.0 mm (0.040 in.) thick and with a height-to-area ratio less than 2 to 1. The amount of sweep is determined by the throw of the carrier. Excessive throw of workpieces of small area is detrimental because of the high rate of side motion at reversal points. Insufficient throw of workpieces of large area is detrimental because the lapping pattern approaches a straight line. The sweep paths illustrated in Fig. 21(c) and 21(d) are derived from planetary action and are advantageous for cylindrical and square workpieces, as well as rectangular workpieces with height-to-area ratios of less than 2 to 1. These sweep paths help in reducing hydraulic cut because the leading edge constantly changes, especially on shorter sweep paths. This provides a nondirectional pattern. As workpiece size is increased in relation to carrier size, the center of the workpiece tends to travel in a circular path and may increase the difficulty of maintaining flatness and parallelism. Con dit ion of La p Su r fa ce . A dull gray surface on the lap is best for stock removal but not for producing highly

reflective finishes. Conversely, bright lap surfaces are excellent for producing highly reflective finishes, but poor for stock removal. When using laps having bright surfaces, the abrasive film must be thin. If not, scratching will result. However, a film that is too thin may cause glazing and loss of control of size and flatness. Sand and other impurities in the lap surface sometimes break free and cause scratches, especially when heavy, viscous films of lapping compound are used. A scleroscope hardness range of 27 to 32 (89 to 99 HRB) has proved optimum for cast iron lap surfaces. Harder laps often cause glazing and scratching. Softer lap surfaces cause loss of flatness and parallelism and also produce grayer finishes. I n t e r ch a n gin g or t r a n sposin g w or k pie ce s during the cycle is advantageous for producing consistent results. An

increase in handling time and cost is the only disadvantage of this practice. Either of two methods can be used: •

•

Workpieces in a load are interchanged between runs, and each load is processed independently until finished size is attained. Drop-off can be minimized by turning the workpieces end for end (reversing their leading edges) during the interchange Workpieces in a series of loads are intermixed at random outside the machine after each load is lapped for a given time. The process is repeated until size is reached. This method is usually the more

economical

D u r a t ion of La ppin g Cycle . Required lapping time is usually determined experimentally for a new job. Extremely

short cycles (in the range of 1 to 3 min) are not economical, because too great a portion of the total time is used for loading and unloading. However, short cycles are preferred for size control. When abrasive is not continually fed to the laps, there is an optimum time for any job that must be determined by trial. Cle a n lin e ss of t h e e n vir on m e n t is important to the success of any lapping operation. Scratches, edge drop-off, and

related difficulties can often be attributed to the dropping of air-borne particles into open machines, the use of contaminated abrasives, or inadequate cleaning of workpieces.

La ppin g Sph e r ica l Su r fa ce s Any of several methods can be used for lapping spherical surfaces. Size, quantity, and required accuracy of the workpieces are the main factors that determine choice of method. I n sin gle - pie ce la ppin g, concave or convex laps are individually contoured to the workpiece. The laps should be

made of fine-grain cast iron, which is suitable for lapping virtually any metal if the process is continued for enough time. Any of several types of machines that have one or two rotating spindles, such as drill presses or milling machines, can be used. When machines with one spindle are used, rotating the lap against the workpiece is preferred because preparing a lap with either a straight or tapered shank to be held in a chuck is simpler than improvising a method of holding a workpiece for rotation. A magnetic chuck can hold workpieces made of ferrous metals. Clamping devices must be used for parts made of nonferrous materials. A crank, which is held by the chuck of the turning machine, is constructed with a ball-end crankpin that fits a drilled hole in the back of the lap (Fig. 22) and causes the lap to rotate over the surface of the workpiece. The part is in line with the spindle of the turning machine; the crankpin is offset from center as required for the diameter of the workpiece. For best results, the lap should be heavy enough to provide the required lapping pressure (about 70 kPa, or 10 psi) on the workpiece. If necessary, added pressure can be applied to the lap by a hand feed lever, such as that on a drill press, but it should be applied carefully because excessive pressure can throw the lap off the work.

Fig. 2 2 Set up for single- piece spherical lapping wit h a one- spindle m achine using a rot at ing lap. See t ext for discussion.

When machines with two spindles are used, one spindle holds and rotates the workpiece while the other holds the lap in a floating position and oscillates it through an angle large enough to lap the required portion of the surface. A typical setup is illustrated in Fig. 23. One of the spindles must be designed so that it slides as the lap and the workpiece wear and thus keeps a constant pressure on the workpiece.

Fig. 2 3 Set up for single- piece spherical lapping wit h a t wo- spindle m achine

M u lt iple - Pie ce La ppin g. Quantities of spherical parts are lapped on a concave or convex lap (Fig. 24). The lap can

replace the lap on a conventional planetary lapping machine, or it can be fastened into the counterbore of the existing flat lap with holding brackets. Contoured laps should be made of fine-grain cast iron and should be machined to the required radii.

Fig. 2 4 Set up for m ult iple- piece spherical lapping

For continuous lapping operations, the conditioning (carrier) rings used should be made of the same material as the lap so that the rings and the lap wear at the same rate. When more than one part is lapped within each conditioning ring, workholders made of fiber, wood, or similar materials are required. Workholders are machined to the same curvature as the lap plate and cut out so that the workpieces can fit into them. Heavy workpieces can be lapped without additional weight applied, but pressure plates can be placed atop lighter pieces to provide sufficient weight for lapping. These plates can be of various weights to suit requirements, but they must be shaped to conform to the final workpiece shape.

Planetary lapping machines converted to use for spherical workpieces are used to lap parts having radii of 1525 to 3050 mm (60 to 120 in.), the applications for a specific machine being limited by the dimension of its conditioning rings. The capacity of the machine will be the same as for flat lapping (Table 4). In plain planetary lapping machines that are prepared for contour lapping but not equipped with conditioning rings, workpieces are held by hand or by improvised fixtures. In this case, the workpieces should be moved over the entire area of the lap with reasonable consistency so that the lap is worn evenly and its life is prolonged.

La ppin g of Ba lls Spherical parts such as balls for ball bearings can be economically lapped to close dimensional tolerances and smooth finishes. Two machines used in production are the multigroove lapper and the single-groove lapper. A multigroove lapper is shown in Fig. 25. The machine consists of two cast iron laps (130 to 150 HB); one is stationary and one rotates. Each lap has a series of concentric grooves that correspond with the grooves in the other lap. The radii of the grooves are the same as the radii of the balls being lapped. The stationary lap is cut out for the entrance and exit of the balls, which are fed in at spaced intervals.

Fig. 2 5 Mult igroove lapper used for lapping bearing balls

In the multigroove machine, the speed of the rotating lap is normally between 50 and 65 rev/min, and the load is usually 33 to 44 kN (7500 to 10,000 lbf). Stock of 0.010 to 0.015 mm (0.0004 to 0.0006 in.) is removed from the diameters of the balls; the time required is 8 to 24 h, depending on the dimensions and finish specified. Balls up to about 44.5 mm (1 in.) in diameter can be lapped in this type of machine to an accuracy better than 0.00064 mm (0.000025 in.) and a finish better than 0.050 m (2 in.). The single-groove machine is used for balls larger than about 44.5 mm (1 in.) in diameter and for balls of any size that require greater accuracy and better surface finish than are obtainable in the multigroove machine. In the single-groove machine, which also employs two opposing cast iron laps, the shaft may be either horizontal or vertical. There is no cutout in either lap. Balls to be lapped are alternated with spacer balls of slightly smaller diameter until the groove is filled. The speed of the rotating lap may vary from 30 to 150 rev/min, depending on ball size and material and on the accuracy and surface finish desired. The load may be as low as 135 N (30 lbf) when extremes of accuracy and finish are sought, and the lapping time may extend to three days. Usually, 0.005 to 0.0075 mm (0.0002 to 0.0003 in.) is removed from the ball diameters. Accuracy to less than 0.125 m (5 in.) and surface finish of 0.01 m (0.5 in.) are possible.

Assuming that the pitch diameter of the groove is 38 mm (15 in.) and that spacer balls are used, 188 balls 3.2 mm (

in.)

in diameter or 9 balls 67 mm (2 in.) in diameter could be finished in a single-groove machine in one run. However, less than the groove complement (even a single ball) can be lapped by using dummy balls of another material between the spacer balls to fill the groove. The material used to make the dummy balls must be different in some way (for example, color, density, magnetic properties, or reflectivity) from the material in the workpieces for easy identification. In both types of machines, there is a constant flow of fluid lapping compound, which usually consists of fused alumina of grit size 5 to 10 m (200 to 400 in.) suspended in 10% soluble oil and 90% water. Mineral oil can be used instead of soluble oil and water. Although costs are greater for single-groove lapping than for multigroove lapping, in either process costs are proportional to the time required for obtaining desired accuracy and finish. Most hard materials can be lapped to greater accuracy and better surface finish than soft materials. Rate of stock removal can be increased by using harder abrasives such as boron carbide or diamond. However, the use of harder or coarser abrasives results in some loss of surface smoothness.

La ppin g t o Acce le r a t e W e a r - I n Internal lapping is often employed as an accelerated wearing-in process for matching and aligning components of bearing assemblies. In most applications of this type, virtually no stock is removed, and sometimes the desired surface correction is so slight that no abrasive is needed. Typical tooling and techniques are described in the following example.

Ex a m ple 1 1 : La ppin g of Be a r in g Asse m blie s. The setup illustrated in Fig. 26 was used for the simultaneous lapping of two roller-bearing cage assemblies into mating outer races. All components of these assemblies except cages were made of steel carburized and hardened to 60 HRC. As shown in Fig. 26, the cage assemblies were mounted at each end of a horizontal driving spindle. The outer races were positioned by horizontally actuated pressure heads through which the lapping compound was supplied.

Fig. 2 6 Schem at ic view of set up for m at ched- piece lapping roller- bearing cage assem blies int o m at ing out er races

Lapping was performed by rotating the cage assemblies against the races at 1063 rev/min under pressure of 70 to 100 kPa (10 to 15 psi) for 5 min. The lapping compound was prepared by mixing 0.68 kg (24 oz) of medium-hard, 800-grit silica with 38 mL (10 gal.) of paraffin oil. The compound was recirculated and was changed every ten shifts (80 h). This machine could process cage assemblies of various outside diameters to a maximum of 190 mm (7

in.).

La ppin g of Spr in glik e Pa r t s Special equipment and techniques are sometimes required for lapping unusual parts. Springlike parts such as those used in numerous precision mechanisms are examples of parts requiring specially designed equipment and techniques.

La ppin g of Ge a r s Gear lapping corrects the minute errors in involute profile, helix angle, tooth spacing, and concentricity created in the forming or cutting or in the heat treatment of the gears. The lapping can be done by running a set of gears in mesh or by running one gear with a gear-shaped master lapping tool. Gear lapping is most often applied to sets of hardened gears that are required to run silently in service. Gear lapping is strictly a mating process and is not intended for stock removal. Two gears that have been matched by lapping should be operated as a set, and they should be replaced as a set, rather than singly. Gears are lapped in special machines, which can be arranged for manual, semiautomatic, or automatic operation. In semiautomatic operation, loading and unloading are manual; in automatic operation, loading and unloading are done automatically in accordance with a programmed cycle. Angular, spur, and helical gears can be lapped, but the process is mainly applied to spiral bevel gears and hypoid gears. A typical setup used for lapping hypoid gears is shown in Fig. 27. Gears of up to 915 mm (36 in.) pitch diameter can be lapped in semiautomatic or automatic operation. Manual lapping, in special equipment, is employed for gears of approximately 2540 mm (100 in.) pitch diameter down to the smallest gear that can be manufactured. Production lapping machines can be adjusted to lap gears with shaft angles of 0 to 180°.

Fig. 2 7 Set up for t he lapping of hypoid gears

La ppin g M e dia . Optimum grit size varies with different types and sizes of gears. A 280-grit abrasive is used for spiral

bevel gears; a finer abrasive (about 400-grit) is more suitable for hypoid gears because the sliding action is greater. As a rule, coarser grit is used for gears having a coarse pitch, and finer grits are used for gears having a fine pitch. When compound is brushed on, as in manual operation, a paste-type vehicle is used. However, in semiautomatic or automatic lapping, the abrasive should be mixed with a thin oil (such as mineral seal oil) so that it can be pumped to the workpieces (Fig. 27). Pr oce ssin g Te ch n iqu e s. It is important to roll all mating gears together before lapping to detect nicks and burrs, which can be removed by a small portable hand grinder before the gears are lapped. This preliminary rolling also inspects tooth contact, which should be in the same location for each set of gears and is especially important in automatic lapping.

During the lapping operation, the pinion (smaller gear) is used as the driver, and the larger gear is the driven member. The driven spindle is also used for applying the necessary tooth-contact load by adjusting to a slight drag. Running cycles as short as 15 s at about 76 m/min (250 sfm) can frequently produce desired results. However, longer time cycles may be necessary, depending on initial gear-tooth finish and service requirements. Low noise level is the criterion of successful gear lapping. Because gear lapping is strictly a mating process and no stock removal is intended, measurements are not made as in most other lapping processes. Minor corrections in tooth bearing shape and position can be obtained, however. Lapping does improve the finish of gear teeth, but improved finish is seldom the purpose of gear lapping.

H igh - Spe e d M a ch in in g I n t r odu ct ion HIGH-SPEED MACHINING is a relative term from a materials viewpoint because of the vastly different speeds at which different materials can be machined with acceptable tool life (Ref 1). For example, it is easier to machine aluminum at approximately 1800 m/min (6000 sfm) than titanium at 180 m/min (600 sfm). Because of this difference and because speed determines to a significant degree whether a material will form continuous chips or segmented shear-localized chips, one way of defining high-speed machining is to relate it to the chip formation process (see the section "Mechanics of Chip Formation" in this article). Localized shear occurs when the negative effect on strength of increasing temperature due to intense plastic deformation is equal to or greater than the positive effect of strain hardening. In this context, highspeed machining for a given material can be defined as that speed above which shear localization develops completely in the primary shear zone. Although appealing from a technical standpoint, the foregoing is not very useful as a practical definition. For this reason, it is generally preferable to define machining speeds quantitatively in terms of specific ranges. One suggestion is that 600 to 1800 m/min (2000 to 6000 sfm) should be termed high-speed machining, 1800 to 18,000 m/min (6000 to 60,000 sfm) very high-speed machining, and greater than 18,000 m/min (60,000 sfm) ultrahigh-speed machining (Ref 2). In the case of very-difficult-to-machine alloys such as titanium, it is preferable to use the term high-throughput machining rather than high-speed machining in order to maintain a proper focus on realistic machining conditions. The use of high-speed machining technology in industry has become more prevalent in recent years. The development of tougher, more refractory tool materials and of high-speed machining spindles has contributed to this growth in acceptance. High-speed machining can be used to machine parts that require the removal of significant amounts of material and to machine long, thin webs. The need throughout industry and the defense establishment to reduce costs and increase productivity has created new interest in high-speed machining. The rationale for continuing research and development in high-speed machining technology is evident when one considers the funds expended in metal removal annually in the United States. Of the more than $100 billion spent for metal removal, 75% can be attributed to the four conventional processes of turning, milling, drilling, and grinding (Ref 3). Machining costs can be significantly reduced only by determining ways to increase metal removal rates.

Ack n ow le dge m e n t s ASM International is grateful to D.G. Flom (retired: General Electric Corporate Research & Development) and R. Komanduri (National Science Foundation) for their valuable contributions to this article. Much of this article is based on the work of Flom and Komanduri conducted from 1980 to 1985.

H ist or ica l Ba ck gr ou n d ( Re f 4 ) The concept of high-speed machining was conceived by Dr. Carl J. Salomon during a series of experiments from 1924 to 1931 (German patent 523594, 1931). The patent was based on a series of curves of cutting speeds plotted against generated cutting temperatures. These experiments were performed on nonferrous metals such as aluminum, copper, and bronze. Salomon obtained speeds up to 16,500 m/min (54,200 sfm) using helical milling cutters on aluminum. His contention was that the cutting temperature reached a peak at a given cutting speed; however, as the cutting speed was further increased, the temperature decreased. Figure 1 shows a simplified presentation of this concept.

Fig. 1 I dealized cut t ing speed versus cut t ing t em perat ure plot as devised by Salom on. Source: Ref 4

As the cutting speed is increased from 0 in the normal mode, V1, the temperature will increase in a direct relationship until a peak value, Tcr, is achieved. The cutting speed at Tcr is commonly called the critical cutting speed, Vcr. If the cutting speed is further increased, it was predicted that the cutting temperature would decline. Near Vcr, Salomon suggested that there was an unworkable regime in which cutters were not able to withstand the severe process temperatures and forces. The shape of the curve was thought to be dependent on the exact nature of the base material being cut. When the cutting speed was sufficiently increased, the resulting temperatures, at V2, were reduced to those of the normal cutting temperatures, at V1, and the materials and cutters would once again permit practical cutting procedures. The same cutting temperature, Ta, found in the normal speed range, V1, could possibly be reproduced in the high-speed range, V2. The studies conducted by Salomon are now mainly of historical interest since current research is developing more definitive data using more sophisticated techniques. In addition, interpretations of Salomon's theory have been responsible for confusion and false expectations concerning high-speed machining. The first systematic investigation of high-speed machining in the United States was undertaken by R.L. Vaughn at Lockheed Aircraft Corporation (Ref 5, 6, 7). In 1958, Vaughn studied a series of variables involved in traditional machining that became very important in high-speed machining. According to Vaughn, the rate at which metal can be machined is affected by: • • • • •

Size and type of machine Power available Cutting tool used Material to be cut Speed, feed, and depth of cut

These five general variables can be broken down further into: • • • • • • •

Rigidity of machine, cutter, and workpiece Variations in speed from the slowest to the fastest, depending on machine used Variations in feed and depth of cut from light to heavy and whether cut dry or with the aid of lubricant and/or coolant Type and material of cutting tool Variations in cutter shape and geometry Type and physical characteristics of work material Specific requirements of desired cutting speed, tool life, surface finish, horsepower required, residual

stress, and heat effects

Recent advances in the development of computer control systems have provided the capability of accurately manipulating high-performance, automatic production machines. Progress in the field of bearing design, alternative spindle power sources, automatic tool changing, tool retention devices, and cutter materials has also made contributions toward proving Vaughn's experiments. A second series of studies was initiated in the 1970s. These studies were contracted by the U.S. Navy with Lockheed Missiles & Space Company (Ref 8, 9, 10). The objective of these studies was to determine the feasibility of using highspeed machining in a production mode, initially with aluminum alloys and later with nickel-aluminum-bronze. A team of researchers demonstrated that it was economically feasible to introduce high-speed machining procedures into the production environment to realize major improvements in productivity. This resulted in a significant increase in overall interest and a very active period of both experimental and applied research, as can be noted in a review of the literature. For example, important data concerning the effect of cutting speed and cutter geometry on cutting temperature when turning 2014-T652 aluminum were provided by F.J. McGee (Ref 11). These data indicated that cutting temperature curves tend to peak near the melting point of the aluminum alloys. In the late 1970s and early 1980s, a third series of contracts was awarded by the U.S. Air Force to the General Electric Company to provide a data base for the machining of aluminum alloys, titanium alloys, nickel-base superalloys, and steels. These studies were spearheaded by D.G. Flom (Ref 1, 12, 13, 14, 15, 16, 17, 18, 19) and R. Komanduri (Ref 20, 21, 22, 23, 24, 25, 26, 27), and the results will be the focus of this article.

M e ch a n ics of Ch ip For m a t ion ( Re f 2 6 ) Two types of chip formation have been observed in high-speed machining, depending on the type of work material to be machined and its metallurgical conditions. They are the continuous chip and the shear-localized (segmental) chip, a term arising from the intense deformation (or shear localization) between the segments. Continuous chips are likely to occur in the high-speed machining of a metal or alloy of body-centered cubic/face-centered cubic crystal structure, high thermal diffusivity, and low hardness, such as aluminum alloys or soft low-carbon steel. The shear-localized chip formation process has been studied and understood in some detail only recently. Shear localization occurs with such materials as titanium alloys, nickel-base superalloys, and hardened alloy steels, which are characterized by low thermal diffusivity, hexagonal close-packed crystal structures, and high hardness. Figure 2 shows the various surfaces involved in the formation of shear-localized chips. The process is divided into two stages. The first involves plastic instability and strain localization in a narrow band in the primary shear zone, leading to catastrophic shear failure along a shear surface. The surface originates from the tool tip almost parallel to the cutting velocity vector and gradually curves concavely upward until it meets the free surface. In the second stage, a gradual buildup of the segment (a low-deformation process) is caused by the flattening of the wedge-shaped work material ahead of the advancing tool. Initial contact between the segment being formed and the tool face takes place at the apex of the tool and is of extremely short duration. The contact increases as the flattening progresses. There is almost no relative motion between the bottom surface of the chip segment being formed and the tool face until near the end of the flattening stage. The chip segment being formed pushes gradually against the previously formed chip segment. The contact between the segment being formed and the one before it shifts gradually, beginning close to the work surface and gradually shifting toward the tool face as flattening progresses. As the upsetting of the segment being formed progresses, intense concentrated shear bands (white etched bands) appear between the segments in a longitudinal midsection of a chip (the regions are marked by arrows in Fig. 3b). These bands are caused by the buildup of stresses in the primary zone and are actually formed during this upsetting stage.

Fig. 2 Schem at ic of t he shear- localized chip form at ion process t hat occurs in t he high-speed m achining of cert ain m at erials. 1, undeform ed surfaces; 2, part of t he cat ast rophically shear- failed surface separat ed from t he following segm ent due t o int ense shear; 3, int ense shear band form ed due t o cat ast rophic shear during t he upset t ing st age of t he segm ent being form ed; 4, int ensely sheared surface of a segm ent in cont act wit h t he t ool and subsequent ly slid along t he t ool face; 5, int ense localized deform at ion in t he prim ary shear zone; 6, m achined surface. Source: Ref 26

Fig. 3 Exam ples of cont inuous ( a) and segm ent ed ( b) chip form at ion. Arrows indicat e areas of shear localizat ion. Source: Ref 1

Once shear-localized chips are formed above a certain speed, they persist with increases in speed. No further transition into different chip forms occurs, at least up to 30,000 m/min ( 100,000 sfm). Because they are easier to dispose of, shear-localized chips are preferable to continuous chips, especially at higher speeds where individual segments of a chip are completely isolated. Formation of the shear-localized chip, however, has not been accompanied by any rapid reduction of tool wear at high speeds. Whether or not formation of this chip reduces the forces and stresses on the tool has yet to be proved. With several metals and alloys, the degree of segmentation depends directly on cutting speed (Ref 1). An example is AISI 4340 steel, for which continuous chips are formed at 120 m/min (400 sfm), as shown in Fig. 4(a). At 975 m/min (3200 sfm), however, completely segmented and detached chips are formed (Fig. 4b). Similarly, Inconel 718, a nickel-base superalloy, forms relatively continuous chips below 60 m/min (200 sfm), but within the range of 60 to 120 m/min (200 to 400 sfm), segmentation begins. At higher speeds, severe detachment occurs (Fig. 5). Titanium alloys such as Ti-6A1-4V are unique in that they form segmented chips at all speeds regardless of their heat treatment conditions.

Fig. 4 Effect of cut t ing speed on chip form at ion of AI SI 4340 st eel. ( a) Cut t ing speed of 120 m / m in ( 400 sfm ) . ( b) Cut t ing speed of 975 m / m in ( 3200 sfm ) . Source: Ref 1

Fig. 5 I nconel 718 chips form ed at increasingly higher cut t ing speed. Unit s are given in surface feet per m inut e. Source: Ref 1

An a lyt ica l m ode lin g of the chip formation process has been conducted in high-speed machining within the framework

of continuum mechanics (Ref 28, 29, 30, 31, 32, 33). The theoretical and metallurgical features of chip formation have been organized into a detailed engineering analysis, the main features of which are:

• • • •

Chip geometry and morphology Kinematics Deformation zones Cutting forces

The theory applies both to continuous and segmental types of chips and is based on constitutive laws for the machined materials. The continuous, ribbonlike chip can be described in terms of shear lamellae joined by intense, thin shear bands. The thickness of a lamella depends on how far a crack can propagate before being arrested by the plastic state of the material at the crack tip. This distance is a material property and does not depend on cutting speed. Initial confirmation of this prediction is obtained in machining aluminum. Chips in the speed range of 150 to 4500 m/min (500 to 15,000 sfm) show virtually no change in lamellar thickness with speed. The metallurgy associated with shear fracture is important. The theoretical work strongly suggests that a miniature shear crack in front of the tool edge is essential for the observed chip formation.

Analysis of the transition from the continuous-type chip to the segmental-type chip as a function of cutting speed has led to the formulation of a stability principle, based on the properties of the stress-strain-strain rate surface under adiabatic deformation conditions. Such surfaces exhibit a locus of instability expressed by:

where is stress and is strain for various strain rates (Fig. 6). Because the strain and velocity conditions in metal cutting are a coupled set of variables, a speed can be identified at which the chip formation mechanism becomes unstable and chip segmentation begins. Within the range of practical cutting speeds, the onset of chip segmentation depends primarily on speed through influence of the latter on strain rate. In addition, thermal and mechanical properties, for example, hardness, play a significant role. The onset of segmentation changes the character of the cutting force and the stress distribution on the tool face.

Fig. 6 Variat ion of shear st ress, inst abilit y is given by

/

, versus shear st rain,

, at different st rain rat es

. The focus of shear

= 0. Source: Ref 1

In addition to the modeling just described, a two-dimensional finite-difference computer program has been investigated for modeling high-speed machining processes (Ref 28). The code solves the basic equations of continuum mechanics throughout a fixed mesh to obtain the deviatoric stress components, hydrostatic pressure, mass density, material velocity components, and internal energy as functions of space and time. Three calculations using the code have been performed-one at 60,000 m/min (200,000 sfm) and two at 15,000 m/min (50,000 sfm). In the latter calculations, one workpiece was twice as hard as the other. At the highest speed used, the workpiece behaves like a fluid; material strength effects become more important as cutting velocity is reduced. In addition, the material yields to a smaller depth beneath the cutting tool as velocity is decreased. The depth of yielding is further reduced by doubling the yield strength of the workpiece (simulating strain hardening). It is significant that the results of computer calculations are consistent with trends observed in practice. In addition to use of the finite-difference method, a partially successful finite-element analysis has been made of a simplified, continuous-chip, orthogonal machining process. Figure 7 shows the results of a finite-element analysis used to simulate segmented chipping.

Fig. 7 Result s of a finit e- elem ent analysis used t o sim ulat e chip segm ent at ion during high-speed m achining. The result s correspond t o a cut t ing speed of 1800 m / m in ( 6000 sfm ) and a rake angle of 5°. ( a) I nit ial geom et ry, t im e = 0.0 s. ( b) Geom et ry at 0.005 s. ( c) Geom et ry at 0.008 s. ( d) Geom et ry at 0.0085 s. Source: Ref 34

Simple constitutive equations can be used to describe the plastic-flow behavior of materials over wide ranges of strain, strain rate, and temperature (Ref 28). Of the three materials studied--6061-T6 aluminum, Ti-6Al-4V, and AISI 4340 steel-the steel exhibits the greatest strain-hardening rate. Orthogonal machining tests conducted on the same three materials at very low speeds have been used to determine strain distributions in and adjacent to the deforming chips. A modified thermal-mechanical criterion of instability predicts that the critical shear strains for aluminum and steel are very much higher than those for the titanium alloy at moderate cutting speed; the differences in these strains increase markedly with

increasing speed. The strain rate for transition from a simple mechanical to a thermally induced shear failure process cannot yet be determined from first principles.

H igh - Spe e d M a ch in in g Pa r a m e t e r s Cu t t in g For ce Ve r su s Spe e d. It has been confirmed that cutting force decreases with increasing speed until a

minimum is reached at a speed characteristic of the given workpiece material. Beyond this characteristic speed, the force tends to slowly increase. For example, force for AISI 4340 steel continues to decrease with increasing speed until about 1500 m/min (5000 sfm), at which point the force begins to increase with speed (Fig. 8). Most of these data were generated on two lathes--a high-speed 500 mm (20 in.), 110 kW (150 hp) engine lathe and an 800 mm (32 in.), heavy-duty, 95 kW (125 hp) engine lathe (Ref 35). The two curves at each feed in Fig. 8 represent data from each lathe. Additional confidence in these results is provided by force measurements for AISI 4340 steel machined at the Denver Research Institute (DRI) during ballistic experiments.

Fig. 8 Variat ion of cut t ing force wit h speed for AI SI 4340 st eel. Source: Ref 1

As shown in Fig. 8, the DRI data lie on an extrapolation of the curve generated at a feed of 0.23 mm/rev (0.009 in./rev). Similar to AISI 4340 steel, aluminum 6061-T6 exhibits an initial decrease in force with increasing speed up to about 3000 m/min (10,000 sfm), beyond which the force increases slightly. In contrast to the results for AISI 4340 steel and aluminum, the cutting force for titanium is relatively unaffected by speed. The decrease in force with speed observed for several materials does not mean a lowering of horsepower requirements; however, one can take advantage of the lower forces at high speed (especially with lighter chip loads) to machine accurately both thin webs and unsupported sections of a part. The practical top cutting speed for aluminum does not appear to be limited by cutting tool wear; spindle speed and power are the controlling factors. It should be obvious, however, that high spindle speeds alone do not ensure high metal removal rates. High feed rates and adequate depths of cut are also needed. Depending on the types of cut being made (straight or contoured), the speed of response of the machine tool and its control may be critical. Cu t t in g Te m pe r a t u r e Ve r su s Spe e d. The evidence indicates that the chip/tool interface temperature increases with

speed, approaching the melting point of the work material (Fig. 9), rather than falling off at very high speeds, as had been claimed by Salomon (Ref 36). Because the melting temperature of aluminum alloys is low ( 540 °C, or 1000 °F) and well below the temperature limitations of carbide and coated carbide tools, the top cutting speed for aluminum alloys appears unlimited from a cutting tool point of view. Table 1 lists the softening points of tool materials, along with the melting points of some common workpiece materials.

Table 1 Softening points of tool materials and melting points of workpiece materials Tool material High-speed steel Tungsten carbide Aluminum oxide Cubic boron nitride Diamond

Softening point °C °F 600 1110 1100 2010 1400 2550 1500 2730 1500 2730

Workpiece material Aluminum Superalloys Steel Titanium Zirconium

Melting point °C °F 600-660 1110-1220 1300-1400 2370-2550 1450-1500 2640-2730 1600-1650 2910-3000 1800-1850 3270-3360

Fig. 9 Variat ion of chip/ t ool int erface t em perat ure wit h speed for AI SI 4340 st eel. Adapt ed from Ref 36

Su r fa ce Fin ish Ve r su s Spe e d. There are indications that surface finish tends to improve with increasing speed, but

these results are not conclusive. Some of the reasons could be due to: • • •

Burnishing action due to tool wear Smaller chip load used Dynamic response of the cutter

Cu t t in g Tool M a t e r ia ls ( Re f 3 7 ) Contrary to predictions based on early work in high-speed machining (see the section "Historical Background" in this article), there does not appear to be an upper speed range at which cutting temperatures decrease with increasing speed. The general observation is that as speed is increased, the temperature approaches the melting point of the workpiece material as a maximum. Because tool wear is usually strongly dependent on temperature, tool wear is the major factor limiting cutting speed, the main exception to this being aluminum machining.

W e a r M e ch a n ism s. A 10-year research program on tool materials for high-speed machining concluded that the two

major wear mechanisms associated with high-speed machining are high-speed chemical dissolution wear and high-speed diffusion-limited wear (Ref 37). In the range of cutting speeds used in high-speed machining, the chemical dissolution of the tool material into the workpiece is the most important contributor to wear. In essence, the tool material dissolves into the flowing chip. The tool material that is the most resistant to dissolution exhibits the least wear. The second consideration is that of diffusion-limited wear. As the cutting speed is increased, the cutting temperature rises to a level at which seizure of the chip material occurs everywhere on the tool face. This layer of adherent material becomes saturated with tool constituents and serves as a diffusion-boundary layer, reducing the rate of transport of tool material into the chip and consequently the wear rate. The wear phenomenon becomes increasingly diffusion-limited, and the observed dramatic decrease in wear occurs with increasing speed. Because the diffusivity increases exponentially with temperature, a further increase in cutting speed beyond the speed for minimum wear produces a rapid increase in wear rate, as shown in Fig. 10.

Fig. 1 0 Effect of cut t ing speed on t he wear rat e of cubic boron nit ride t ooling. Workpiece: AI SI 4340 st eel ( 35 HRC) . Source: Ref 37

Cu t t in g Tool Se le ct ion . Consideration of the wear mechanisms of tool materials during high-speed machining

suggests three possibilities for tool development:

• • •

Pick a tool material that is so chemically stable with respect to the workpiece that chemical dissolution of the tool does not occur to a significant extent, even at the melting point of the workpiece Promote the transition to the diffusion limited wear regime Isolate the tool from the workpiece rather than a modification of the tool material itself. If a protective layer could be introduced between the tool and the chip, transport of the tool constituents into the chip could be prevented. The use of viscous lubricants in the profile milling of titanium has proved successful (Ref 38)

Tool Syst e m s for Alu m in u m Alloys. In aluminum alloys, the cutting temperature is limited by the low melting point

and high thermal conductivity of aluminum. Chemical effects are minimal, and wear is primarily a result of the abrasion of the tool material by hard second-phase particles. Abrasion resistance increases with the hardness of the tool material; high-speed steel tooling and cemented carbide tooling are suitable for machining most structural alloys, and polycrystalline diamond is preferred for the highly abrasive cast aluminum-silicon (10 to 20% Si) alloys (see the article "Machining of Aluminum and Aluminum Alloys" in this Volume for a discussion of machining high-silicon aluminum

alloys). Existing tool materials are adequate for machining aluminum alloys at any conceivable speed, with spindle speed and horsepower design limitations setting the upper limit on cutting speed. Tool Syst e m s for St e e l. The oxides are the only potential tool materials that are not limited by their chemical stability. Therefore, the most promising area for tool material development is in improving the toughness and flow strength of the oxides. A second possibility involves the development of tool materials with improved hot strength. The development of tool materials other than cubic boron nitride (CBN) with high hot strength in the range of 1300 to 1400 °C (2400 to 2600 °F) might allow a transition from dissolution-limited to diffusion-limited wear, with a corresponding increase in tool life. Finally, the wear of CBN in the machining of steels of moderate hardness (35 to 50 HRC) should be investigated to determine whether a transition to a low-wear regime (similar to that which has been observed in hard steels) occurs at sufficiently high cutting speeds. If so, machining in the range of 1200 m/min (4000 sfm) and above may be feasible. Tool Syst e m s for Su pe r a lloys. The recommendations are similar to those for steel. The oxides are quite chemically stable with respect to nickel and cobalt, making the development of tough oxide materials a priority. In addition, the transition to diffusion-limited wear is known to occur at high speeds; therefore, any new high hot strength compositions will likely find application in the machining of superalloys. A possible example is the class of tool materials based on silicon nitride (Si3N4) and alloys of Si3N4 and aluminum oxide (Al2O3), referred to as SiAlON (see the article "Ceramics" in this Volume). These tool materials are very effective in machining nickel-base alloys at high speeds. In light of the relatively poor chemical stability of these materials, it is suspected that they represent the second example (in addition to CBN) of a tool material that has sufficient hot strength to enable the very high speed wear transition. Tool Syst e m s for Tit a n iu m Alloys. Titanium has low thermal conductivity, low specific heat, and a high melting point. These properties ensure that cutting temperatures will be high at even moderate cutting speeds. In addition, titanium is highly chemically reactive with all known tool materials, causing rapid wear. It is quite likely that the most wear resistant materials--tungsten carbide and diamond--have already been identified. Therefore, the most promising area for investigation is the development of effective lubrication techniques to reduce the interaction between the tool and the chip. In addition, new tool geometries, such as the ledge tool described below, have also increased the productivity of machining titanium alloys.

Alt e r n a t ive Cu t t in g Tool Ge om e t r ie s ( Re f 2 7 , 2 8 ) Rapid tool wear remains a problem in the machining of titanium and other difficult-to-machine alloys, even though cutting speeds for titanium have been recently increased three- to fivefold through judicious choice of cutter grades and geometries, fluids, and machining parameters. Partial solutions to the tool life problem lie in new tool geometries and the use of rotating cutters. Th e Le dge Tool. The concept underlying the ledge tool (Fig. 11) is very simple: The tool contains a ledge that is allowed to wear away at a controlled rate with only minimal increase in force. Thus, tool wear has not been reduced, but tool life has been greatly increased. The size of the ledge (overhang) equals the depth of cut desired, and its thickness equals the ultimate flank wear width to be tolerated. In turning, the square tool with the ledged side is brought against the workpiece so that clearance is available between the edge of the ledge and the finished surface of the workpiece (that is,

the end cutting edge angle, 1°), as shown in Fig. 11. Because the depth of cut is the same or less than the width of the ledge, only the ledge portion of the tool does the cutting and wearing. The ledge wear back is due to a combination of flank wear and microchipping wear. With this tool, cutting speeds for titanium alloys can be increased five times over conventional speeds, with long tool life ( 30 min) and good finish (Fig. 12). The sparking that accompanies tool wear at high cutting speeds has been eliminated by surrounding the tool/workpiece interface with inert gas (N2) or by submerging the workpiece in a cutting fluid. A combination of a flood lubricant and an inert gas also serves to eliminate sparks and to keep the workpiece cool.

Fig. 1 1 Schem at ic of ledge t ool, which is designed t o increase product ivit y during t he high-t hroughput m achining of t it anium . ( a) Ledge t ool m ount ed on a convent ional t oolholder. ( b) A t urning operat ion using a ledge t ool. Source: Ref 27

Fig. 1 2 Variat ion of ledge wear back wit h cut t ing t im e for different carbide grades when t urning Ti- 6Al-4V at 180 m / m in ( 600 sfm ) wit h a dept h of cut of 0.75 m m ( 0.030 in.) and a feed of 0.023 m m / rev ( 0.009 in./ rev) . Source: Ref 27

Ledge tools have also been evaluated in the face milling of forged Ti-6Al-2Sn-4Zr-2Mo. The results have been comparable to those in turning except the rate of ledge wear back was about three times faster in milling; on a microscopic level, the mode of tool wear is the same. Figure 13 shows the variation of ledge wear with cutting time and volume of material removed.

Fig. 1 3 Variat ion of ledge wear back wit h cut t ing t im e ( a) and wit h t he volum e of m at erial rem oved ( b) for different carbide grades when face m illing a Ti- 6Al- 2Sn- 4Zr- 2Mo alloy ( 36 HRC) . Tool: 0.75 m m × 1.0 m m ( 0.030 in. × 0.040 in.) . Cut t ing speed: 155 m / m in ( 515 sfm ) . Chip load per t oot h: 0.23 m m ( 0.009 in.) . Axial dept h: 0.75 m m ( 0.03 in.) . Source: Ref 28

Another incremental approach for increasing tool life when machining titanium alloys is a new cutting geometry comprised of a high clearance angle (10 to 15°) together with a high negative-rake angle (-10 to -15°). This geometry will also allow the use of a conventional insert on a modified toolholder (Ref 39). Rot a r y t ool m a ch in in g is another technique that holds promise for extending tool life. In this method, a tool with a circular cutting edge is allowed to rotate about its own axis, either self-propelled by the cutting process or driven externally at the desired speed. As cutting proceeds, new portions of the cutting edge are continuously brought into contact with the workpiece at the cutting zone. Thus, increased tool life can be anticipated because of lower temperatures and reduced chemical reactions at the moving chip/tool interface, and reduced cutting forces can be anticipated because of increased options for modifying the chip formation.

Applica t ion s of H igh - Spe e d M a ch in in g ( Re f 2 6 ) High-speed machining is used in the defense and airframe industries to manufacture aircraft engine propulsion components and in the automobile industry. When high-speed machining is accompanied by higher feed rates and spindle power, the higher spindle speeds allow higher removal rates. Increased productivity, however, necessitates high-speed, high-power, compact spindle designs; low-inertia feed tables; fast feed drives; quick-response numerical control; and a totally integrated machining system. A typical integrated machining system consists of multiple machining cells that will fabricate large machine parts under hierarchical computer control in an effective, cost-efficient manner (Ref 40). Benefits of the integrated machining system include: • • • •

Reduced labor requirements Improved throughput with the application of high-speed and high-throughput machining Enhanced production flexibility and reduced work-in-process through establishment of a serial production environment Improved product quality and enhanced industrial base

Figure 14 shows an isometric view of an integrated machining system. Such a system can take up to more than 9000 m2 (100,000 ft2) of floor space.

Fig. 1 4 I nt egrat ed m achining cent er for t he high- speed and high- t hroughput m achining of alum inum and t it anium , respect ively. AGV, aut om at ed guide vehicle; AS, aut om at ed st orage; RS, ret rieval syst em . Source: Ref 40

Air fr a m e a n d D e fe n se . Most airframe manufacturers have implemented high-speed machining. Its primary

application is in end milling with small-size cutters. Aluminum alloys are the common work materials used, so tool wear is not a limitation, especially with carbide cutters. The ideal candidates for high-speed machining are parts whose machining time is a significant fraction of the floor-to-floor time. The ideal cuts are long, straight ones that enable the use of high feed rates and consequently high removal rates. Although some parts may fall under this category, there are others that are suitable for high-speed machining but require extensive pocketing and complex contouring. Both pocketing and contouring can involve frequent accelerations and decelerations in feed rates. Tool-changing time can be reduced by using as few cutters as possible, eventually using only one, smallest-diameter cutter capable of generating all the radii on the part. Modifying the design of the part may enable the use of high-speed machining and may be desirable if the modification can be achieved without compromising the part performance specifications. Further, although high-speed steel tools may be satisfactory for some applications, carbide cutters will not only extend tool life but also provide about three times the stiffness, which is essential for machining long, thin webs with slender end mills. The integrated machining system shown in Fig. 14 is used to machine both aluminum and titanium. For the high-speed machining of aluminum parts, metal removal rates of 3300 cm3/min (200 in.3/min) are possible using machines with up to 55 kW (75 hp) spindle drives and spindle speeds approaching 20,000 rev/min. For the high-throughput machining of titanium alloys, metal removal rates of the order of 165 cm 3/min (10 in.3/min) using 95 kW (125 hp) spindle drives and up to 833 rev/min spindle speeds are possible. Many high-speed machining applications, however, do not require such machine capabilities. For example, a 15 kW (20 hp), 20,000 rev/min spindle is used to machine A7 wing spars made of 7057-T6 aluminum (Ref 26). The feed rate is 15,000 mm/min (600 in./min) on long, external tapered flanges and 7500 mm/min (300 in./min) in pocket areas; the metal removal rate is 1300 cm3/min (80 in.3/min). Retrofitted machining centers with 20 kW (26 hp), 18,000 rev/min high-speed spindles are used for the high-speed machining of 7075-T6 aluminum parts for the Trident missile (Ref 26). The worktable has a feed capacity of 5000

mm/min (200 in./min). The high-speed system incorporates such safety features as a double lock for the toolholder and sensors that can shut down the machine quickly if tool breakage creates an imbalance. A major commercial airline manufacturer uses six-axis machining centers for the contour milling of aluminum honeycomb for engine nacelles (Ref 26). The system has a 3.3 m (11 ft) diam rotary table capable of a full 360° rotation. The spindle is mounted on the gantry and can tilt 240° and rotate 360°. A unique feature of this system is the use of a graphite-fiber reinforced plastic ram. This ram provides the required stiffness and reduced weight necessary for the high feed rates of low-inertia parts. The high-speed machining systems utilize high-speed (24,000 rev/min) 11 kW (15 hp) spindles. Air cr a ft En gin e Pr opu lsion . Nickel-base superalloys and titanium alloys are the work materials most often used in

aircraft engine propulsion components. These cause rapid tool wear at high speeds, which constitute a major limitation in the high-speed machining of propulsion parts. Until recently, the cutting speeds possible with nickel-base superalloys were about 30 m/min (100 sfm) with carbide cutters and about 9 m/min (30 sfm) with high-speed steel tools. The development of CBN and ceramics such as hot-pressed Al2O3 plus TiC, SiAlON, and silicon carbide (SiC) whiskerreinforced Al2O3 has made possible an increase in cutting speeds from 30 to 180 m/min (100 to 600 sfm) for machining nickel-base superalloys. These ceramic tool materials are much tougher and more consistent in performance than the ceramics introduced in the 1950s. SiAlON and SiC whisker-reinforced Al2O3 are recommended for roughing, and hotpressed Al2O3-TiC for finishing applications. Although new, tougher ceramic tool materials, perhaps based on ceramic composites, will undoubtedly be developed in the future, it is unlikely that cutting speeds will exceed 600 m/min (2000 sfm), because tool wear will continue to be a limitation. As described earlier in this article, innovative tool designs and geometries, such as the ledge tool, have improved productivity in titanium machining. Au t om obile I n du st r y. High-speed machining in the automobile industry is performed on gray cast iron and aluminum

alloys, especially the high-silicon type. Silicon nitride ceramic and polycrystalline diamond are the important tool materials that permit higher speed with longer tool life. Gray cast iron can be machined at a speed of about 1500 m/min (5000 sfm) with Si3N4 tools, and aluminum alloys with high silicon content (10 to 20%) can be machined at about 750 m/min (2500 sfm) with polycrystalline diamond tools. In working with these materials, the current high-speed machining systems need to provide more power and stiffness and improved chip-handling means, controls, and safety features. High-speed machining with these materials may require spindle power from 150 to 375 kW (200 to 500 hp). Chip removal rates can reach 16,000 cm3/min (1000 in.3/in.), necessitating efficient chip disposal systems. Because products are mass produced in this industry, the nonmachining time should be minimal.

I m ple m e n t in g H igh - Spe e d M a ch in in g ( Re f 2 6 ) There are several factors to be considered by companies planning to employ high-speed machining systems. Whether high-speed machining is economically appropriate to the application, the relative value of various systems, and how well the system can be integrated into existing operations must all be evaluated. Other areas to consider are the research and development support required, system reliability and maintenance, overall safety aspects, general acceptance on the manufacturing shop floor, capital and other investment costs, the skills required, corporate goals, and the financial health of the company. The overall goal should be to improve productivity, reduce costs, and produce parts of a given size, shape, finish, and accuracy at competitive cost. Productivity and overall costs depend on cutting time, noncutting time, labor, and overhead. High-speed machining can decrease cutting time by increasing cutting speed. Noncutting time can be decreased by the automatic loading and unloading of parts, automatic tool changing, in-process inspection, in-process sensing, and adaptive control. Labor costs can be decreased because with high-speed machining fewer operators are needed to work on fewer, more efficient machine tool systems. Similarly, overhead costs can be reduced by operating fewer, more efficient machine tools on two or more shifts and during holidays and at night. In order to evaluate the influence of cutting speed on productivity, floor-to-floor time can be regarded as the sum of the cutting time and the noncutting time. Figure 15 shows the percentage decrease in floor-to-floor time with an increase in cutting speed for different ratios of cutting time to floor-to-floor time, using conventional, currently used speeds of whatever the value may be for a given material and process as a base. When the cutting time is a significant fraction of the floor-to-floor time and when tool wear at high speed is not significant (solid lines, Fig. 15), the cutting speed can be increased considerably to effect a significant reduction in floor-to-floor time. If, however, this ratio is low (bottom curve,

Fig. 15), an increase in cutting speed by even an order of magnitude or more will result in only a marginal decrease in floor-to-floor time. In this case, unless the noncutting time could be decreased significantly, high-speed machining would not be advantageous.

Fig. 1 5 Variat ion of percent decrease in floor- t o- floor t im e wit h cut t ing speed. See t ext for det ails. Source: Ref 26

When an increase in cutting speed does not contribute to a significant reduction in floor-to-floor time and when tool wear is significant at high speeds, as in the machining of titanium alloys (dashed line, Fig. 15), high-throughput machining can be adopted to reduce noncutting time and to increase productivity. Similarly, where heavier cuts can be made on the part without affecting either the part finish and accuracy requirements or the cutting tool performance, greater depths of cut (as in high removal rate machining) are recommended instead of higher speed (see the article "High Removal Rate Machining" in this Volume). Fu t u r e N e e ds ( Re f 1 ) . High-speed machining can be cost effective only if other aspects of machining, including

reduction of noncutting times and labor costs, can also be improved. For this reason, automated machining, such as the integrated machining system discussed earlier, is receiving much attention. Dynamic in-process inspection is an integral part of automated machining. Sensors and diagnostics have been identified as critical to effective in-process inspection. Recently, proximity sensors for measuring tool wear, vibration sensors for measuring tool-touch (between cuts), and lasers for measuring workpiece dimensions have been studied. The integrated machining system shown in Fig. 14 utilizes a coordinate-measuring machine as its inspection module. Some of the specific needs for the increased use of high-speed machining are: • • • •

Integrated sensing techniques for monitoring machining processes Sensors and diagnostics for detecting tool breakage (including incipient tool breakage) Linkage of signal analysis to tool wear mechanisms and machining variables Self-teaching adaptive control systems

• • • • • •

Faster response in machine tool controls when machining aluminum Better cutting tools (composition and geometry) for machining titanium Water-base cutting fluids for machining titanium More reliability and predictability in cutting tools Greater machine tool rigidity for increased productivity More notch-resistant cutting tools for superalloys

In addition to the above-mentioned technical needs, significant reductions in noncutting time can be achieved by procedures that to a large extent involve improved management. These procedures include setup, part load/unload, queing, chip cleanup, tool change, maintenance, scheduling, and general operator efficiency. Until improvements in these procedures are achieved, cutting times will remain only a small fraction of the overall manufacturing sequence, and the full advantages of high-speed machining will not be realized.

Re fe r e n ce s 1. D.G. Flom, High-Speed Machining, in Innovations in Materials Processing, G. Bruggeman and V. Weiss, Ed., Plenum Press, 1985, p 417-439 2. B.F. von Turkovich, Influence of Very High Cutting Speed on Chip Formation Mechanics, in Proceedings of NAMRC-VII, 1979, p 241-247 3. The 12th American Machinist Inventory of Metalworking Equipment 1976-78, Am. Mach., Dec 1978, p 133-148 4. R.I. King and R.L. Vaughn, A Synoptic Review of High-Speed Machining from Salomon to the Present, in High-Speed Machining, American Society of Mechanical Engineers, 1984, p 1-13 5. R.L. Vaughn, Ultra-High Speed Machining, Am. Mach., Vol 107 (No. 4), 1960, p 111-126 6. R.L. Vaughn, "Recent Developments in Ultra-High Speed Machining," Technical paper 255, Vol 60, Book 1, Society of Manufacturing Engineers, 1960 7. R.L. Vaughn, "Ultra-High Speed Machining--Feasibility Study," Final Report, Contract AF 33 (600) 36232, Production Engineering Department, Lockheed Aircraft Corporation, June 1960 8. R.L. Vaughn, L.J. Quackenbush, and L.V. Colwell, "Shock Waves and Vibration in High-Speed Milling," Technical Paper 62-WA-282, American Society of Mechanical Engineers, Nov 1962 9. R.L. Vaughn and L.J. Quackenbush, "The High-Speed Milling of Titanium Alloys," Technical Paper MR 66-151, Society of Manufacturing Engineers, April 1966 10. R.I. King and J. McDonald, Product Design Implications of New High-Speed Milling Techniques, Trans. ASME, Nov 1976 11. F.J. McGee, "Final Technical Report for Manufacturing Methods for High-Speed Machining of Aluminum," Technical Registry No. 6089, Manufacturing Methods and Technology Branch (DRDMIEAT), U.S. Army Missile Research and Development Command, Feb 1978 12. D.G. Flom, Ed., "Advanced Machining Research Program (AMRP)," Semi-annual Technical Report, Air Force Contract No. F33615-79-C-5119, GE Report No. SRD-80-018, Feb 1980 13. D.G. Flom, Ed., "Advanced Machining Research Program (AMRP)," Annual Technical Report, Air Force Contract No. F33615-79-C-5119, GE Report No. SRD-80-118, Aug 1980 14. D.G. Flom, Ed., "Advanced Machining Research Program (AMRP)," Semi-annual Technical Report, Air Force Contract No. F33615-79-C-5119, GE Report No. SRD-81-018, Feb 1981 15. D.G. Flom, Ed., "Advanced Machining Research Program (AMRP)," Semi-annual Technical Report, Air Force Contract No. F33615-79-C-5119, GE Report No. SRD-81-062, Aug 1981 16. D.G. Flom, Ed., "Advanced Machining Research Program (AMRP)," Semi-annual Technical Report, Air Force Contract No. F33615-79-C-5119, GE Report No. SRD-82-027, Feb 1982 17. D.G. Flom, Ed., "Advanced Machining Research Program (AMRP)," Annual Technical Report, Air Force Control No. F33615-79-C-5119, GE Report No. SRD-82-070, Aug 1982

18. D.G. Flom, Ed., "Advanced Machining Research Program (AMRP)," Semi-annual Technical Report, Air Force Contract No. F33615-79-C-5119, GE Report No. 83-SRD-012, Feb 1983 19. D.G. Flom, R. Komanduri, and M. Lee, "Review of Past Work in High-Speed Machining," Paper presented at the TMS-AIME Meeting (Louisville, KY), The Metallurgical Society, Oct 1981 20. R. Komanduri and J. Hazra, A Metallurgical Investigation of Chip Morphology in Machining an AISI 1045 Steel at Various Speeds up to 10,100 SFPM, in Proceedings of NAMRC-IX, Society of Manufacturing Engineers, 1981 21. R. Komanduri and B.F. von Turkovich, New Observations on the Mechanism of Chip Formation When Machining Titanium Alloys, Wear, Vol 69, 1981, p 179-188 22. R. Komanduri, "Titanium--A Model Material for Studying the Mechanism of Chip Formation in HighSpeed Machining," Paper presented at the TMS-AIME Meeting (Louisville, KY), The Metallurgical Society, Oct 1981 23. R. Komanduri, Some Clarifications on the Mechanics of Chip Formation When Machining Titanium Alloys, Wear, Vol 76, 1982, p 15-34 24. R. Komanduri and R.H. Brown, The Mechanics of Chip Segmentation in Machining, J. Eng. Ind. (Trans. ASME), Vol 103, Feb 1981, p 33-51 25. R. Komanduri, T. Schroeder, J. Hazra, B.F. von Turkovich, and D.G. Flom, On the Catastrophic Shear Instability in High-Speed Machining of an AISI 4340 Steel, J. Eng. Ind. (Trans. ASME), Vol 104, May 1982, p 121-131 26. R. Komanduri, High-Speed Machining, Mech. Eng., Dec 1985, p 65-76 27. R. Komanduri, D.G. Flom, and M. Lee, Highlights of the DARPA Advanced Machining Research Program, J. Eng. Ind. (Trans. ASME), Vol 107, Nov 1985, p 325-335 28. D.G. Flom, "Advanced Machining Research Program (AMRP) Final Technical Report," Air Force Contract No. F33615-79-C-5119, GE Report No. 83-SRD-040, Oct 1983 29. D.R.C. Durham, "Physical Metallurgy of Deformation Localization," Topical report prepared for DARPA, UV-81-DD1, Aug 1981 30. B.F. von Turkovich, Influence of Very High Cutting Speed on Chip Formation Mechanics, in Proceedings of NAMRC-VII, 1979, p 291-297 31. D.R.C. Durham and B.F. von Turkovich, Material Deformation Characteristics at Moderate Strains and High Strain Rates, from Metal Cutting Data, in Proceedings of NAMRC-X, 1982, p 324-331 32. B.F. von Turkovich and D.R.C. Durham, Machining of Titanium and Its Alloys, in Proceedings of the Symposium on Advanced Processing Methods for Titanium (Louisville, KY), The Metallurgical Society, 1981, p 241-256 33. C.A. Brown, A Practical Method for Estimating Machining Forces from Tool-Chip Contact Length, Ann. CIRP, Vol 32 (No. 2), 1983, p 91-96 34. G. Meir, J. Hashemi, and P.C. Chou, "Finite-Element Simulation of Segmented Chipping in High-Speed Machining," Report MR88-120, Society of Manufacturing Engineers, 1988 35. T.A. Schroeder and J. Hazra, High Speed Machining Analysis of Difficult-to-Machine Materials, in Proceedings of NAMRC-IX, Society of Manufacturing Engineers, 1981 36. J.P. Kottenstette and R.F. Recht, Ultra-High-Speed Machining Experiments, in Proceedings of NAMRC-X, 1982, p 263-270 37. B.M. Kramer, On Tool Materials for High-Speed Machining, in High-Speed Machining, American Society of Mechanical Engineers, 1984, p 127-140 38. J. Jensen, "High-Speed Milling of Titanium," M.S. thesis, Massachusetts Institute of Technology, 1983 39. R. Komanduri and W. Reed, Jr., Evaluation of Carbide Grades and a New Cutting Geometry for Machining Titanium Alloys, Wear, Vol 92, 1983, p 113-123 40. Integrated Machining System--An Overview, LTV Aircrafts Products Group, Military Aircraft Division, 1988

H igh Re m ova l Ra t e M a ch in in g I n t r odu ct ion HIGH REMOVAL RATE (HRR) MACHINING involves the use of extremely rigid, high-power, high-precision machines, such as roll turning lathes, to achieve material removal rates far beyond the capacity of conventional machine tools. Material removal rates as high as 6000 cm3/min (370 in.3/min) have been reached using multiple ceramic cutters on high-carbon (0.8% C) heat treated cast steel rolls. This article will review the machine requirements, cutting parameters, and applications associated with HRR machining. Additional information can be found in Ref 1, 2, 3, 4, and 5.

Ack n ow le dge m e n t The editors would like to thank Jack Binns, Sr., inventor of the "Super-Lathe," for his valuable contributions to this article.

M a ch in e Re qu ir e m e n t s The key to success in HRR machining is extreme rigidity in machine, workpiece, and cutting tool setups. Figure 1 shows a "Super-Lathe" capable of making cuts as deep as 25 mm (1 in.) at feeds up to 1.3 mm/rev (0.050 in./rev) on hardened steel and chilled cast iron rolls. Metal removal rates as high as 4500 kg/h (10,000 lb/h) can be achieved with such machines, which have power ranges from 100 to 450 kW (150 to 600 hp) and can produce up to 400 kN (95,000 lbf) of output torque at the spindle. Workpieces of softer materials can also be turned, producing large, segmented chips, as shown in Fig. 2.

Fig. 1 Superlat he used for HRR m achining. This 300 kW ( 400 hp) m achine, which is com put er num erically cont rolled, can perform rough cut s on 1290 m m ( 50.75 in.) diam rolls and finishing cut s on workpieces up t o 1200 m m ( 47.25 in.) in diam et er at m axim um roll lengt hs of 7.8 m ( 25 ft , 7 in.) . Court esy of J. Binns, Sr., Binns Machinery Product s

Fig. 2 A large AI SI 1055 st eel ( 180 HB) chip cut on an HRR m achining lat he. Court esy of J. Binns, Sr., Binns Machinery Product s

Th e la t h e be d, which is the main structural member of the "Super-Lathe," is designed to withstand applied loads with

minimal deflection. The structure is an elongated steel box fabricated from hot-rolled ASTM A 36 steel that has a high elastic modulus. The six exterior plates of the box become the principally stressed members. When fabricated correctly, the unit needs no vibrational damping, and when subjected to maximum rated loads, maximum unit stresses of only approximately 3450 kPa (500 psi) will develop in the critically stressed members. Th e t a ilst ock of the machine features a rectangular quill that can withstand radial loads two or three times as great, for a

given amount of deflection, as an equivalent circular quill with an identical bore and wall thickness. The tailstock is also fabricated from A 36 structural steel plate. The method of clamping the tailstock to the bed again provides for maximum rigidity. There are four clamp bars at the bottom edge of the tailstock; when pulled up with 100 J (75 ft · lbf) of torque, the force required to separate the tailstock from the bed is approximately 5500 kN (625 tonf). Th e h e a dst ock is also made of heat-treated wrought steel. In the power shafting, no shaft between gears is smaller in

diameter than the length between the gears. The main spindle bearing, the bull gear, and the ring gear (when used) are all the same nominal diameter as the maximum size of the roll the lathe will turn. Additional information on superlathe gearing and electrical drive systems can be found in Ref 2 and 3. Workpieces are gripped by an eight-jaw chuck. Th e ca r r ia ge , which carries the cutter, must be capable of withstanding the maximum torque force and feed forces with

only a fraction of a thousandth of an inch of deflection. The weight and mass of the cross slide and saddle are important considerations because inertia can substantially reduce the changes in deflection of the carriage as the feed and cut forces vary at high frequency. For example, cutting forces suddenly drop from a maximum to almost zero when each segment of the chip ruptures from the workpiece. Antifriction ways (with linear ball and roller bearings) to the cross slide and bed are a major factor in steadying the carriage. Toolin g for HRR machining is also of considerable importance. Because the cutter shank is the link between the insert

cutting edge and the machine, the same considerations in terms of rigidity must be given it as the bed, headstock, and carriage. A number of criteria must be considered in shank design. The first is that the larger and more massive the shank, the better it will cut. A large monobloc shank is both more rigid and a better heat sink than either a small shank or one comprised of a small interchangeable toolholder placed in a large shank. In addition, a properly matched shank socket and insert will reduce insert breakage. Because aluminum oxide (Al2O3) and Al2O3 + TiC inserts are low in transverse rupture strength as compared to alternative cutting tool materials (Table 1), it is essential that the internal stresses from the cutting forces and thermal distortion be confined as much as possible to the compressive state.

Table 1 Transverse rupture strengths of common cutting tool materials Tool material Al2O3 Al2O3 + TiC SiAlON Tungsten carbide Coated tungsten carbide

Transverse rupture strength MPa ksi 690 100 860 125 780 113 1380-2760 200-400 1030-2070 150-300

Inserts should be mechanically clamped in place with only enough force to secure them while the cut is starting. After the cut has begun, the cut force will hold the inserts. When initially clamping the inserts, it is important to have them solidly against the stops in the socket. Eva lu a t in g M a ch in e Rigidit y. The manner in which chips flow from the work indicates the variation in deflections of

the cutting edge with respect to the workpiece and therefore provides a good evaluation of rigidity. Adequate rigidity has been achieved when the chip peels off as a smooth, continuous ribbon from the work (Fig. 3). After a short period of smooth flow, the ribbon will fall apart into short lengths from its own weight or movement, then drop into the chip chute. If the chips fly off like snowflakes, there is insufficient rigidity, and the insert will soon break.

Fig. 3 Cont inuous 115 m m ( 4 in.) wide ribbon of m achined m at erial peeling off an HRR cut t er. Court esy of J. Binns, Sr., Binns Machinery Product s

Cu t t in g Tools a n d Cu t t in g Pa r a m e t e r s Although cemented carbide and cubic boron nitride (CBN) tools can be used for HRR machining, Al2O3 + TiC tooling is normally used, particularly for finishing cuts on ferrous alloys (surface finishes of 0.25 to 0.75 m, or 10 to 30 in., can be produced). Alumina-base inserts with a density of 4.30 g/cm3 and a hardness of 95 HRA are commonly used. Carbide tools are used for some roughing cuts. Both round and rectangular insert configurations are used, with the latter being the most common (Fig. 4). For rectangular inserts, negative-rake angles of 5 to 25° with 5 to 10° clearance are used;

side cutting edge angles (lead angles) range from at least 30° up to 90° for lap cutting at fast speeds. Depths of cut to 75 mm (3 in.) are possible. Speeds range from 6 to 240 m/min (20 to 800 sfm), with feeds reaching as high as 6.4 mm/rev (0.250 in./rev) at low speeds. The effects of cutting parameters on tool life are reviewed in Ref 3 and 4. Detailed information on cemented carbide, CBN, and ceramics can be found in the Section "Cutting Tool Materials" in this Volume.

Fig. 4 Ceram ic insert ( bot t om ) used for HRR m achining, along wit h t he result ing chip ( t op) . The insert cut t ing edge is t hat adj acent t o t he scale. Court esy of J. Binns, Sr., Binns Machinery Product s

As stated earlier in this article, metal removal rates for ferrous materials can range as high as 6000 cm3/min (370 in.3/min) with multiple inserts. A 375 kW (500 hp) lathe can make cuts at more than 105 m/min (350 sfm) at a feed of 2 mm/rev (0.08 in./rev) and depths of cut of 75 mm (3 in.) in as-forged 52100 bearing steel rolls. This results in about 7700 kg (17,000 lb) of chips produced per hour, which equals an energy rate of metal removal of 44 cm3/min/kW (2 in.3/min/hp). Turning tests on Inconel 718 and Ti-6Al-4V have produced metal removal rates of more than 800 cm3/min (50 in.3/min). The workpiece finish was between 0.75 and 2.5 m (30 and 100 in.) With aluminum, it has been shown that removal rates of 3300 cm3/min (200 in.3/min) are attainable on each insert of a gang tool (Ref 5).

Re fe r e n ce s 1. Staff Report, Mech. Eng., March 1984, p 55-59 2. J. Binns, Sr., "Rough Turning and Hogging With Ceramic Cutters," Papers presented at the Creative Manufacturing Seminars, American Society of Tool and Manufacturing Engineers, 1963-1964 3. J. Binns, Sr., "Ceramic Cutter Performance on Rough Turning and Hogging Cuts," Technical Paper 633, American Society of Tool and Manufacturing Engineers, 1964 4. J. Binns, Sr., Super-Lathe for Roll Turning, Iron Steel Eng., Oct 1961 5. D.G. Flom, Ed., "Advanced Machining Research Program," Annual Technical Report, Air Force Contract No. F33615-79-C-5119, General Electric Co. Report No. SRD-82-070, Aug 1982

N u m e r ica l Con t r ol Yoram Koren, The Universit y of Michigan

I n t r odu ct ion THE METAL CUTTING INDUSTRY changed drastically during World War II. The ambitious aircraft and missile projects of the U.S. Air Force, which required the manufacturing of complicated and accurate parts, made it clear that conventional machine tools could not fill future needs. Under contract to the U.S. Air Force, the Parsons Corporation in Traverse City, MI, undertook the development of a flexible manufacturing system designed to maximize productivity and to achieve desired accuracies for small and medium-sized production runs. The Parsons Corporation subcontracted the development of the control system to the Servomechanism Laboratory of the Massachusetts Institute of Technology (MIT). The challenge was met successfully by MIT, and in 1952 a three-axis Cincinnati Hydrotel milling machine controlled with digital technology was developed. This digital technology was termed numerical control (NC). Thus, the primary motivations for the development of NC systems for machine tools were the demand for high accuracy in the manufacture of complicated parts and the desire to shorten production time.

Evolu t ion of N u m e r ica l Con t r ol The first NC controllers in the 1950s used vacuum tubes and were extremely large. Controllers in the early 1960s used transistors in their logic circuits and digital control loops. These units were more reliable. The third generation of NC controllers used integrated circuit chips and consequently became less expensive, more reliable, and smaller. Many of these third-generation controllers are still in operation. In all of these systems, the numerical data required for producing a part is maintained on punched tapes and inserted to the controller through a built-in tape reader. An important advance in the philosophy of machine tool numerical control that took place in the early 1970s was the shift toward the use of computers instead of controller units in NC systems. This produced both computer numerical control (CNC) and direct numerical control (DNC). Computer numerical control is a self-contained NC system for a single machine tool including a dedicated computer controlled by stored instructions to perform some or all of the basic NC functions. With DNC, several machine tools are directly controlled by a central computer. Of the two types of computer control, CNC has become much more widely used for manufacturing systems (for example, machine tools, welders, and laser beam cutters) mainly because of its flexibility and the lower investment required. The preference of CNC over DNC is increasing as a result of the availability and declining costs of minicomputers and microcomputers. One of the objectives of CNC systems is to replace as much of the conventional NC hardware as possible with software and to simplify the remaining hardware. There are many ways in which functions can be shared between software and hardware in such systems, but all involve some hardware in the controller dedicated to the individual machine (Ref 1). This hardware must contain at least the servoamplifiers, the transducer circuits, and the interface components, as shown in Fig. 1.

Fig. 1 Block diagram of a CNC m achine

The software of a CNC system consists of at least three major programs: • • •

A part program A service program A control program

The part program contains a description of the geometry of the part being produced and the cutting conditions, such as spindle speed and feed rate. The service program is used to check, edit, and correct the part program and to run system diagnostics. The control program accepts the part program as input data and produces signals to drive the axes of motion of the machine. The CNC controllers of the 1980s are more powerful and more user friendly than earlier units. They incorporate troubleshooting features such as on-board diagnostics, which allow self-testing of the controller, and simulation mode, which is used to test part programs without generating axes motions. Many controllers offer high-level programming facilities, three-dimensional tool path animation with graphics, tool data base, and preselection of cutting parameters. The modern machine controller is a workstation in a Local Area Network (Ref 2). It is capable of communicating with other controllers and of being integrated into a flexible manufacturing system in the future, thus permitting the gradual construction of a full flexible manufacturing system. Numerical control was introduced and developed in the metalworking industry, and the largest concentration of NC equipment remains in metalworking shops. Numerical control has been successfully implemented for turning, milling, drilling, grinding, boring, punching, and electrical discharge machines. It is interesting to note that numerical control has made possible the development of machines with basic capabilities that far surpass those of conventional machines. For example, sophisticated NC milling machines maintain control over five axes of motion and can literally sculpt complex surfaces (Ref 3). A new breed of NC machine tool is the machining center and the turning center, which incorporate the functions of many machines into a single device. A machining center can access multiple tools to perform such operations as milling, drilling, boring, and tapping (Fig. 2). A turning center is a powerful lathe equipped with an automatic tool changer. Other types of NC machines include welding machines, drafting machines, tube benders, inspection machines, and wiring machines in the electronics industry.

Fig. 2 CNC m achining cent er. Court esy of Cincinnat i Milicron

Fu n da m e n t a ls of N u m e r ica l Con t r ol Numerical control equipment has been defined by the Electronic Industries Association (EIA) as a system in which actions are controlled by the direct insertion of numerical data at some point. The system must automatically interpret at least some portion of these data. In a typical NC or CNC system, the numerical data required for producing a part are maintained on a disk or on a tape and called the part program. The part program is arranged in the form of blocks of information. Each block contains the numerical data required to produce one segment of the workpiece profile. The block contains, in coded form, all the information needed for processing a segment of the workpiece: the segment length, its cutting speed, feed rate, and so on. Dimensional information (length, width, and radii of circles) and the contour form (linear, circular, or other) are taken from an engineering drawing. Dimensions are given separately for each axis of motion (X, Y, and so on). Cutting speed, feed rate, and auxiliary functions (such as coolant on and off, spindle direction, clamp, and gear changes) are programmed according to surface finish, tolerance, and machining requirements (Ref 1, 2, 3, 4, 5, and 6). Compared to a conventional machine tool, the NC system replaces the manual actions of the operator. In conventional machining, a part is produced by moving a cutting tool along a workpiece by means of powered slides that are engaged and disengaged by an operator. Contour cuttings are performed by an expert operator by sight. On the other hand, the operators of NC machine tools need not be skilled machinists. They only have to monitor the operations of the machine, operate the tape reader, and load and remove the workpiece. Most intellectual operations that were formerly done by the operator are now included in the part program. However, because the operator works with a sophisticated and expensive system, intelligence, clear thinking, and good judgment are essential qualifications of a good NC operator. Preparing the part program of an NC machine tool requires a part programmer. The part programmer should possess knowledge and experience in mechanical engineering fields. Knowledge of tools, cutting fluids, fixture design techniques, machinability data, and process engineering are all of considerable importance. The part programmer must be familiar with the function of NC machine tools and machining processes and must decide on the optimum sequence of operations. The part program can be written manually, or a computer-assisted language, such as the automatically programmed tool language, can be used. In NC machines, the part dimensions are presented in part programs by integers. In CNC machines, the dimensions in part programs are sometimes expressed as numbers with a decimal point, but are always stored in the computer as integers. Each unit of these integers corresponds to the position resolution of the axes of motion and is referred to as the basic length unit (BLU). The BLU is also known as the increment size or bit weight, and in practice it corresponds approximately to the accuracy of the NC system.

In NC and CNC machine tools, each axis of motion is equipped with a separate driving device, which replaces the handwheel of the conventional machine. The driving device may be a dc motor, a hydraulic actuator, or a stepping motor. The type selected is determined by the power requirements and the machine. An axis of motion in numerical control means an axis in which the cutting tool moves relative to the workpiece. This movement is achieved by the motion of the machine tool slides. The main three axes of motion are referred to as the X-, Y-, and Z-axes. The Z-axis is perpendicular to both X and Y in order to create a right-hand coordinate system. For example, in a vertical drilling machine (as one faces the machine), a +X command moves the worktable from left to right, a +Y command moves it from front to back, and a +Z command moves the drill up, away from the workpiece. The X-, Y-, and Z-axes are always assigned to create a right-hand cartesian coordinate system. Each axis of motion also has a separate control loop. The control loops of NC or CNC systems use two types of feedback devices (Fig. 1): tachometers to monitor velocity and encoders or other position transducers (for example, resolvers) to measure position. The controller compares the actual position with the required one and generates an error. The control loop is designed in such a way as to reduce the error, that is, the loop is a negative-feedback type. A common requirement of continuous-path NC and CNC systems is the generation of coordinated movement of the separately driven axes in order to achieve the desired path of the tool. This coordination is accomplished by interpolators. Numerical control systems contain hardware interpolators, but in CNC systems interpolators are implemented by software.

Adva n t a ge s of N C Syst e m s It has been clearly shown that NC manufacturing reduces the number of direct employees because fewer multipurpose NC machines are required and, in some cases, one operator can operate more than one machine. However, the ratio of indirect to direct employees might increase, and in turn, overall employment in the industry might rise despite increased automation and mechanization (Ref 6). Output would of course also increase. Realistically, then, the advantage is not in lowering labor costs but in increasing the output per man-hour. Handling costs with NC technology have decreased--in some cases remarkably. Setup times have been substantially reduced, and actual productive time has been substantially increased. A furthur savings of time is achieved while passing from one operation to another during the machining of the workpiece. With a conventional machine tool, the work must be stopped at such points because the operator must go to the next step. The operator must stop the cutting process frequently and measure the part dimensions to ensure that the material is not overcut. It has been proved that the time wasted on measurements is frequently 70 to 80% of the total working time (Ref 1). The rate of production is also decreased because of operator fatigue. In NC systems, these problems do not exist. Because the accuracy is repeatable with numerical control, inspection time is also reduced. Numerical control produces higher-quality parts and makes possible the accurate manufacture of more complex designs without the usual loss in accuracy encountered in conventional manufacturing. Producing a part that must be cut with an accuracy of 0.01 mm (0.0004 in.) or better may take a considerable amount of time using conventional methods. In numerical control using single-axis motion, obtaining such accuracies is the state-of-the-art, and they are maintained throughout the entire range of cutting speeds and feed rates. Another intangible advantage of numerical control is the production of complex parts that are not feasible in conventional manufacturing. Complex-contour cutting in three dimensions cannot be performed by manual operation. Even when it is possible, the operator must manipulate the two handwheels of the table simultaneously while maintaining the required accuracy; thus, it becomes possible only when the part is simple and requires relatively low accuracies. It is obvious that in such work the NC machines save a considerable amount of time. Compared to conventional machining methods, the NC machine tool has the following advantages: • • • • •

Complete flexibility; a part program is needed only for producing a new part Accuracy is maintained through the full range of speeds and feeds The possibility of manufacturing a part of complicated contour A shorter production time Higher productivity achieved by saving indirect time, such as setting up and adjusting the machine and using one operator to monitor several machining operations, or by using completely automatic operation

in unmanned production

Pr ogr a m m in g Most NC and CNC machine tools use off-line programming methods, which can be either manual or computer assisted, such as programming with the aid of the automatically programmed tool (APT) language. During off-line programming, the machine remains in operation while a new part program is being written. Typically, when a part program is ready, it is stored on a punched tape or a floppy disk. The tape or disk is taken to the machine shop and loaded into the machine tool controller, and the part is subsequently produced. Manual part programs are written by programmers. First, the programmers must determine the machining parameters and the optimum sequence of operations to be performed. Based on this sequence, they calculate the tool path and write a manuscript. Each line of the manuscript, which is referred to as a block, contains the required data for transferring the cutting tool from one point to the next, including all machining instructions that should be executed either at the point or along the path between the points. The EIA standard RS-273-A ("Interchangeable Perforated Tape Variable Block Format for Positioning and Straight Cut Numerical Controlled Machines") provides a line format for point-to point (PTP) and straight-cut NC machines. A typical line according to this standard is as follows:

N102 G01 X-52000 Y9100 F315 S717 T65432 M03 (EB) The letter and the number that follows it are referred to as a word. For example, X-52000 and M03 are words. The first letter of the word is the word address. Word addresses are denoted as follows: N, sequential number, G, preparatory function; X and Y, dimensional words; F. feed rate code; S, spindle speed code; T, tool code; M, miscellaneous function; and EB, end-of-block character. The EB character is not printed but only punched or coded, and it is usually the carriagereturn code, thus permitting a new line to begin immediately afterward. The EB character indicates to the NC controller that the current reading is completed and that the axes of motion must start up. When this motion is accomplished, the next block is read. According to EIA standards RS-273-A and RS-274-D ("Interchangeable Variable Block Data Contouring Format for Positioning and Contouring\Positioning for Numerically Controlled Machines"), two block formats are available: the word address format and the tab sequential format. In the word address format, each word must be headed by the word address. The controller uses the address letter to identify the word that follows it. In this type of format, words need not be arranged in any specific order within the block because the letter identifies the corresponding word. However, because the standards prescribe a definite sequencing of words, the format recommended above is used for PTP programming. In the tab sequential format, a tab character is inserted between each two words in the block, and the address letter is omitted. In this format, words must be arranged in a specific order. When a word is not needed in a particular block, it can be omitted, but the corresponding tab character must be punched. Manual part programming can be applied to PTP systems, but it is too complex for contouring systems. Therefore, a great number of computer software systems have been developed to assist in NC part programming. The APT system is the most widespread and the most comprehensive one. It is available on many computers and is widely used by many manufacturers of NC equipment. The first prototype of the APT software system was developed in 1956 by the Electronic System Laboratory of MIT. The program was further developed by the cooperative efforts of 21 industrial companies sponsored by the Aerospace Industries Association with assistance from MIT. As a result of these efforts, a system called APT II was produced in 1958, and a more effective system, the APT III, was distributed in 1961. The Illinois Institute of Technology Research Institute was elected to direct the future expansion of the program, and the capabilities of APT are continually being expanded. The current APT language has a vocabulary of more than 600 words. The APT program is a series of instructions that specifies the path the tool must follow to produce a part. To communicate the tool path to the computer, one must provide the computer with geometric descriptions of the part

surfaces. The APT language enables the programmer to accomplish this and then to specify the manner in which the tool should move along these surfaces. The geometric expressions describing the part and the motion statements represent about 70% of the average program. A geometric expression defines a geometric shape or form. For each geometric form, there are many different methods of definition. Definitions of at least 16 different geometric forms are contained in APT: the most useful ones are POINT, LINE, PLANE, CIRCLE, CYLINDER, ELLIPS, HYPERB, CONE, and SPHERE. Several examples of definitions are presented below. Most APT statements are divided into two sections, major and minor, which are separated by a slash. The major section appears to the left of the slash and is generally one word containing one to six letters. The minor section, if required, appears to the right of the slash and contains modifiers to the major portion. For example, a point can be defined by:

POINT/X-coordinate, Y-coordinate, Z-coordinate An example of a corresponding geometric statement is:

PT2 = POINT/3,4 where PT2 is the symbolic designation of a point whose X-coordinate is 3 and whose Y-coordinate is 4. A line can be defined by a point and a tangent circle (Fig. 3):

LINE/symbol for a point, symbol for a circle

, TANTO,

The modifiers LEFT or RIGHT are applied looking from the point toward the circle. Examples are:

L1 = LINE/P1, LEFT, TANTO, CIRI L2 = LINE/P1, RIGHT, TANTO, CIRI

Fig. 3 Line definit ion by a point and a t angent circle

Once the required part has been defined with the geometric expressions, tool movements are specified using the motion statements. Each motion statement will move the tool either to a new location or along a surface specified by the statement. Examples of motion statements are:

GOTO/HOLE2 GOLFT/L1, PAST, L2 It should be noted that L1, L2, and HOLE2 are identifying names. Identifying names are given to geometric expressions and cannot be APT words. In addition to the geometric and motion statements, there are other kinds of statements and features. One of the most useful statements is the CLPRNT. The CLPRNT is an instruction to the APT system to produce a printed list of all the cutter location coordinates that have been computed. The computation results are those of the APT program before postprocessing. The output of the APT program is sent as input to another program called the postprocessor. The latter is a program written specifically for each CNC machine tool system, and it includes information about the particular machine (size, accelerations, coolant, and so on). The output of the postprocessor is the NC part program that is loaded onto the machine controller to produce the required part.

N u m e r ica l Con t r ol Numerically controlled machine tool systems are in some ways related to industrial robot systems. In both, the axes of a mechanical device are controlled to guide a tool that performs a manufacturing process task (Fig. 4). In both systems, each axis of motion is equipped with a separate control loop and a separate driving device. The difference is in the process and the mechanical device. The process in numerical control may be drilling, milling, grinding, or welding, and in robotics, painting, welding, assembly, or handling. In numerical control, the mechanical handling device is the machine tool, and in robotics, the manipulator. Machine tools, however, are more rigid and usually have perpendicular axes of motion, which result in simpler control strategies and better position accuracy than those of robot arms.

Fig. 4 Schem at ic of an NC or indust rial robot syst em

A common feature in numerical control and robotics is that the required path of the tool is generated by the combined motion of the individual axes. In numerical control, the tool is the cutting tool, such as a milling cutter or a drill; in robotics, the tool is the instrument at the far end of the manipulator, which might be a gripper, a welding gun, or a paintspraying gun. Robot systems, however, are more complex than machine tool CNC systems for the reason discussed below. Machine tools require control of the position of the tool cutting edge in space. In many cases, the control of three axes is adequate. Robots require the control of both the position of the tool center point and the orientation of the tool; this is achieved by controlling six axes of motion (or degrees of freedom). Some robot systems use more than six axes, and some CNC machines use more than three axes of motion. A typical example is the addition of a rotary table to a three-axis milling machine. On the other hand, many robot systems use fewer than six axes, and in such cases the wrist section contains fewer than three degrees of freedom (Ref 7). Similarly, there are CNC machines with only two numerically controlled axes of motion. For example, only two numerically controlled axes are required for turning parts on a lathe because the parts are symmetrical.

Syst e m St r u ct u r e

Computer numerical control systems can be divided into PTP and continuous-path or contouring systems (Ref 1, 2, 3, 4, 5, and 6). A typical PTP system is encountered in a CNC drilling machine. In a drilling operation, the machine table moves until the point to be drilled is exactly under the tool, and then the hole is drilled. The table then moves to a new point, and another hole is drilled. This process is repeated until all the required holes on the part are drilled. The machine table is then brought to the starting point, and the system is ready for the next part. In more general terms, the PTP operation is described as follows. The machine tool moves to a numerically defined position, and then the motion is stopped. The tool performs the required task with the machine table stationary. Upon completion of the task, the machine tool moves to the next point, and the cycle is repeated. In a PTP system, the path and the velocity while traveling from one point to the next are insignificant. Therefore, as shown in Fig. 5, a basic PTP system would require only position counters for controlling the final position of the machine table in order to bring it to the target point (Ref 8). The coordinate values for each desired position are loaded into the counters with a resolution that depends on the basic length unit of the system. During the motion of the table, the encoder at each axis transmits pulses, each representing the tool travel of 1 BLU along the axis. Each axis of motion is equipped with a counter to which the corresponding encoder pulses are transmitted. At the beginning of a motion from a point, each axial counter is loaded by the corresponding required axial incremental distance (in BLUs) to the next point. During the motion of the table, the contents of each counter are gradually decremented by the pulses arriving from the corresponding encoder. When all counters are at zero, the machine table is in its new desired position.

Fig. 5 Block diagram of a point - t o- point syst em

In continuous-path CNC machines, the tool performs the task while the axes of motion are moving, as in milling machines. The task of the controller in milling is to guide the tool along the programmed path. In continuous-path systems, all axes of motion may move simultaneously, each at a different velocity. These velocities, however, are coordinated under computer control in order to trace the required path or trajectory. In a continuous-path operation, the position of the tool at the end of each segment (motion step), together with the ratio of axes velocities, determines the generated contour, and at the same time the resultant velocity also affects the quality of the surface. Continuous-path systems require interpolators for determining the path between given end points (Fig. 6). Part programs supply only the end points of the segments along the contour, and the system computer interpolates the path between the points and generates in real time the commands to the individual axes of motion. Typically, NC and CNC machines are capable of linear and circular interpolation.

Fig. 6 Block diagram of a cont ouring syst em

The principle of operation for linear interpolators of machine tools is relatively simple (Ref 9). The axes of motion in machine tools are perpendicular to each other, thus creating a cartesian coordinate system. A motion along a straight path with a length L employs the following relationships:

(Eq 1) (Eq 2) and

(Eq 3) where

L = x2 + y2 + z2 The distances x, y, and z are the components of L in the X, Y, and Z directions, respectively, and V is the required velocity along the path (referred to as feed rate in machining). Equations 1, 2, and 3 provide the basis of linear interpolator algorithms for machine tools. Circular interpolator algorithms are more complicated and can be found in the literature (Ref 10, 11, 12). Acceleration and deceleration are achieved by varying V in Eq 1, 2, and 3 according to a predetermined formula. Machining feed rates (the tangential velocity of the tool along the contour) in machine tools are small; and therefore, paths that include acceleration and deceleration periods are not often encountered.

Ada pt ive Syst e m s Computer numerical control machines in production are programmable systems that can repeat a sequence of programmed operations as long as necessary. However, these systems are unable to sense and respond to any changes in their working environments. For example, assume that a CNC lathe is used to turn a batch of cylindrical workpieces to a 200 mm (8 in.) radius from raw material having radii from 204 to 207 mm (8.04 to 8.16 in.). The machining feed rate in the part program is calculated as the maximum allowable feed rate needed to remove a maximum load of 4 mm (0.16 in.) without cutter breakage, although with the smaller load the feed rate can be increased without risking the cutter. This and similar situations can be remedied only if the systems are able to adapt to the changing conditions in their environments; such changes would be sensed with suitable sensors. This is basically the motivation for introducing adaptive control systems in machining. The objectives of these adaptive controls are to increase productivity in rough machining and to improve part accuracy in fine cutting (Ref 13). The article "Adaptive Control" provides a more thorough discussion of adaptive control for machine tools.

Re fe r e n ce s 1. Y. Koren, Computer Control of Manufacturing Systems, McGraw-Hill, 1983 2. P. Ranky, Computer Integrated Manufacturing, Prentice-Hall, 1986 3. R.S. Pressman and J.E. Williams, Numerical Control and Computer-Aided Manufacturing, John Wiley & Sons, 1977 4. N.O. Olesten, Numerical Control, Wiley-Interscience, 1970 5. M.P. Groover, Automation, Production Systems, and Computer-Aided Manufacturing, Prentice-Hall, 1980 6. J. Pusztaai and M. Sava, Computer Numerical Control, Reston Publishing, 1983 7. Y. Koren, Robotics for Engineers, Mc-Graw-Hill, 1985

8. J.T. Beckett and H.W. Mergler, Analysis of an Incremental Digital Positioning Servosystem With Digital Rate Feedback, J. Dyn. Sys. Meas. Contr. (Trans. ASME), Vol 87, March 1965 9. G. Ertell, Numerical Control, Wiley-Interscience, 1969 10. B.W. Jordan, W.J. Lenon, and B.D. Holm, An Improved Algorithm for Generation of Nonparametric Curves, IEEE Trans. Comp., Conf. C-22, No. 12, Dec 1973, p 1052-1060 11. O. Masory and Y. Koren, Reference-Word Circular Interpolators for CNC Systems, J. Eng. Ind. (Trans. ASME), Vol 104, 1982 12. A.N. Poo and J.G. Bollinger, Dynamic Errors in Type I Contouring Systems, IEEE Trans. Ind. Appl., Vol 1A-8 (No. 4), July 1972, p 477-484 13. G.A. Ulsoy, Y. Koren, and F. Rasmussen, Principal Developments in Adaptive Control of Machine Tools, J. Dyn. Sys. Meas. Contr. (Trans. ASME), Vol 105 (No. 2), June 1983, p 107-112

Ada pt ive Con t r ol Yoram Koren and A. Galip Ulsoy, Universit y of Michigan

I n t r odu ct ion ADAPTIVE CONTROL (AC) systems for metal cutting processes are a logical extension of computer numerical control (CNC), which was described in the article "Numerical Control" in this Volume. The term adaptive control in metal cutting is in fact a misnomer because these are actually process control systems and may or may not be adaptive in the sense commonly used in the control theory literature (Ref 1, 2). In this article, AC systems for metal cutting are described first, and then classified according to the AC strategies used. Finally, the economic benefits of AC systems are enumerated. In CNC systems, the relative position between the tool and the workpiece is controlled; however, the part programmer must specify the cutting speed and feed rates. The determination of these cutting parameters requires experience and knowledge regarding workpiece and tool materials, machine characteristics, coolant effects, and so on. Their selection directly affects such important economic factors as product dimensional accuracy, surface finish, metal removal rates, tool wear rates, and tool breakage. The main focus of adaptive control is the improvement of these production and product quality related factors in the actual machining process. This is accomplished using the measurement and control of certain process variables in real time as well as several alternative strategies for the AC algorithm. A schematic of a typical AC configuration for a machine tool is shown in Fig. 1. It is clear that the adaptive control represents a process control system that operates in addition to the CNC position, or servo, control system. In this discussion, the different types of AC systems that have been proposed are considered first and then the economic benefits of AC systems.

Fig. 1 Schem at ic of an adapt ive cont rol syst em incorporat ed int o a CNC m achine t ool syst em . Source: Ref 3

Adaptive control systems use a variety of sensors and control strategies. Depending on these factors, as well as the particular process being controlled, the AC systems can conveniently be classified as (Ref 3, 4): • • •

Adaptive control with optimization (ACO), in which an economic index of performance is used to optimize the cutting process using on-line measurements Adaptive control with constraints (ACC), in which the process is controlled using on-line measurements to maintain a particular process constraint (for example, a desired force or power level) Geometric adaptive control (GAC), in which the process is controlled using on-line measurements to maintain desired product geometry (for example, dimensional accuracy or surface roughness)

These three types of AC systems are discussed in some detail in the sections that follow. It is worth mentioning here that the ACO system is the most general. However, because it is difficult to implement an ACO system, the suboptimal ACC and GAC strategies are often used. These can typically provide much of the benefit to be expected from ACO systems and are easier to implement. The ACC-type systems are well suited for rough cutting, and the GAC-type systems are typically used in finishing operations. The economic benefits of AC systems can be significant, particularly under varying cutting conditions (Ref 5). However, the main objective is improvement in productivity, for example, by increasing the metal removal rate (MRR). This is illustrated in Fig. 2 for a milling operation with a variable depth and/or width of cut. With adaptive control, the feed rate can be increased when the depth and/or width of cut is small, and reduced if either becomes larger. By contrast, in conventional milling, the smallest feed rate would be selected based on the worst-case conditions for that particular part.

Fig. 2 Com parison of feed rat es in adapt ive and convent ional ( nonadapt ive) m illing when cut varies. ( a) Variable dept h. ( b) Variable widt h. Source: Ref 3

Some typical results demonstrating the economic benefits of adaptive control are given in Fig. 3. These compare machining costs in adaptive and conventional (non-adaptive) machining and show that improvements in productivity of 20% to 80% can be expected with the use of AC systems. The actual improvement depends on the material being machined and the complexity of the part being produced. Improvements or consistency in product quality can also be an objective of AC systems (usually GAC systems). Another benefit is the reduced part programming time because feeds and speeds are adjusted on-line. The economic impact of adaptive control also depends on the percentage of total production time allocated to actual cutting, and currently this is often only 5 to 20%. As automation reduces the percentage of production time required for part setup and tool changing, the percentage of time the machine spends in actual cutting increases, and the advantages of adaptive control become even more significant.

Fig. 3 Machining cost com parison for adapt ive and convent ional ( nonadapt ive) m achining. Bar graphs ( a) and ( b) com pare a Cincinnat i Milacron syst em wit h nonadapt ive m et hods under t he following condit ions: ( a) t he m achining of m ild st eel wit h a 0.05 m m ( 0.002 in.) t olerance and a cut 0.75 m m ( 0.003 in.) deep and 25 m m

( 1 in.) wide, and ( b) t he m achining of st ainless st eel wit h a 0.25 m m ( 0.01 in.) t olerance and a cut 2.5 m m ( 0.1 in.) deep and 25 m m ( 1 in.) wide. Bar graphs ( c) , ( d) , and ( e) com pare a Bendix syst em wit h nonadapt ive m et hods when variable- dept h cut s ( t op graph) and a 1 m m ( 0.05 in.) dept h of cut ( bot t om graph) are used. Machining condit ions are as follows: ( c) m achining of 4140 st eel wit h a high- speed st eel cut t er, ( d) m achining of 4140 st eel wit h a carbide cut t er, and ( e) m achining of st ainless st eel wit h a high- speed st eel cut t er. Source: Ref 3

Ada pt ive Con t r ol W it h Opt im iza t ion This section describes two systems, a milling and a grinding system, based on an ACO strategy. While ACO systems represent the most general class of AC systems, they have not found application in industry. The two systems described here indicate the ACO strategy and also help to illustrate the limitations that have prevented their use in industry. The first system described is a laboratory system developed in the mid-1960s for milling and used industrially in a few installations. The ACO system described for grinding is a more recent development that has found off-line application in industry. However, its on-line implementation has so far been limited to the laboratory. ACO Syst e m for M illin g. The best-known ACO system is probably the system developed by The Bendix Corporation

for the U.S. Air Force in the early 1960s (Ref 5, 6). The structure of this ACO system for milling is illustrated in Fig. 4. This system will not be described in detail here, but requires some mention because of its historical significance and because it illustrates many aspects of ACO systems. It uses several measurements of process variables (cutting torque, tool temperature, and machine vibration) and an economic index of performance for on-line adjustment of the cutting feed and speed. The performance index (J) can be expressed as:

(Eq 1) where MRR = waV is the metal removal rate (in.3/min), w is the milling width (in.), a is the depth-of-cut (in.), V is the milling feed rate (in./min), TWR = K1(MRR) + K2 + K3(dT/dt) is the tool wear rate (in./min), is the tool temperature (°F), T is the cutting torque (lbf · ft), W0 is the allowable width of flank wear (in.), f is the feed (in./rev), N is the spindle speed (rev/min), C1 is the cost of machine and operator per unit time ($/min), C2 is the cost of tool and regrind per tool change ($/change), and t1 is the tool changing time (min).

Fig. 4 Schem at ic of t he Bendix ACO syst em for m illing. Source: Ref 3

The constants K1, K2, and K3 depend on tool and workpiece materials, and is an adjustable parameter in the range of 0 1 that determines the type of performance index, J. When = 1, the J represents the reciprocal of the cost per unit produced; when = 0, the J represents the production rate; and for intermediate values of , the J represents a weighted combination of these two objectives, which can be adjusted to represent the profit (that is, maximum production with minimum cost). The objective of the ACO system, then, is to maximize this economic index of performance, J, subject to constraints on maximum spindle speed, minimum spindle speed, maximum torque, maximum feed, maximum temperature, and maximum vibration amplitude. This constrained optimization problem is solved on-line using gradient methods, with increments in feed of 0.003 mm/rev (0.0001 in./rev) and increments in spindle speed of 10 rev/min. Although this ACO system was successfully demonstrated in the laboratory, it was not widely accepted by industry. The main problem was the need for a tool wear sensor (for tool wear rate, or TWR, in Eq 1) that can operate in an industrial environment. The calculation of the tool wear rate from the temperature and torque measurements as used in the Bendix system was not satisfactory for the entire range of feeds and speeds, and reliable temperature measurements were difficult to obtain in a production environment. Accurate on-line tool wear estimation remains an important topic of current research, as discussed in the section "Trends in AC Systems" in this article. ACO Syst e m for Gr in din g. A successful ACO system for grinding is based on the measurement of the power at the

grinding wheel, as illustrated in Fig. 5. This measured power, P, is compared to the maximum allowable power for burning, Pb, and the power error generated, e, is used to adjust the workpiece spindle speed, nw, and the radial infeed velocity, vf. The adjustment strategy uses a function Gb to calculate the allowable power for burning and the optimal locus algorithm, based on a model of the grinding process mechanics, to maximize vf and calculate the corresponding value of nw such that P = Pb. Thus, the system is designed to operate at (or near) the power constraint for burning. The sampling period is denoted by T, and is 0.5 s. A surface roughness constraint can also be included. Experimental evaluation of the system was performed in the laboratory, and excellent results were obtained, as shown in Fig. 6. These results illustrate the successful operation of the system at the specified power level through ten cycles, despite the progression of grinding wheel wear; this was achieved by increasing the radial infeed velocity, vf. This optimal locus approach may also be

applicable to the development of ACO systems for metal removal operations other than grinding, and as such is a topic of current research (Ref 8).

Fig. 5 Schem at ic of an ACO syst em for grinding. Source: Ref 7

Fig. 6 Grinding of repet it ive cycles wit h power lim it const raint wheel, 32A46K8VBE; workpiece, SAE 4340 st eel, 240 HB; wheel diam et er, 440 m m ( 17.3 in.) ; workpiece diam et er, 165 m m ( 6.50 in.) ; grinding widt h, 32 m m ( 1.26 in.) ; m at erial rem oved from workpiece radius per cycle, 0.3 m m ( 0.012 in.) ; single- point dressing only before cycle 1 wit h a dressing dept h of 10 m ( 400 in.) , and a dressing lead of 5 m ( 200 in.) . Source: Ref 7

Ada pt ive Con t r ol W it h Con st r a in t s In general, commercial AC systems used in production today for rough milling and turning are of the ACC type. This is because ACO systems are rather complex and will require further research before they can be fully implemented. The ACC systems can provide much of the benefit of ACO systems and are relatively easy to implement. In this section, a general discussion of ACC systems is followed by a description of a particular ACC system for a turning operation. Fu n da m e n t a ls of ACC Syst e m s. Typically, ACC systems are based on the measurement of a single process variable,

such as force, torque, or motor current, and try to maintain that variable at some predetermined reference value. If this reference value is determined to ensure a relatively high production rate without excessive wear rates or breakage, it provides good, although suboptimal, performance. Most ACC systems attempt to maximize the metal removal rate, MRR, by maximizing one of the machining variables, for example, operating at a maximum feed rate compatible with maintaining a constant load on the cutter, as was illustrated in Fig. 2 for a slab milling operation. In this case the average

feed rate with the ACC system is larger than the programmed feed rate, particularly when there are significant variations in the depth and/or width of cut in the particular milling operation. The most commonly used constraints in ACC systems are the cutting force, cutting torque, and the machining power. The machining variables commonly manipulated to meet these constraints are the feed rate (V) and the spindle speed (N). The machining feed is defined by the ratio:

(Eq 2) where p = 1 in turning and drilling and is the number of teeth on the cutter in milling. The main cutting force component is typically modeled as being proportional to the depth of cut and the feed:

F = KSafu

(Eq 3)

where KS is the specific cutting force coefficient, and the exponent, u, is typically in the range 0.6 < u < 1.0. Both KS and u depend on workpiece and tool material properties and must be empirically determined by means of machining tests. The cutting torque T is proportional to the force times the workpiece radius in turning and to the force times the tool radius in milling and drilling. The machining power P is proportional to the force times the spindle speed. An ACC syst e m for t u r n in g is shown in Fig. 7 and described in Ref 9. The main cutting force component is measured and sampled (typically every 0.1 s). The sampled force value F is then compared to a desired, or reference, force value Fr. The force error is provided as input to the ACC controller, which in turn produces a feed rate command signal. A positive error increases the feed rate and consequently increases the actual force F (see Eq 3). Some results of laboratory experiments in which the depth of cut varies in a stepwise fashion are shown in Fig. 8. To completely eliminate the force error, an integral control strategy has been used; that is, the feed rate command signal is proportional to the integral of the force error.

Fig. 7 An ACC syst em for t urning. Source: Ref 9

Fig. 8 Turning experim ent s wit h an ACC syst em wit h a m edium cont roller gain and st ep changes in dept h of cut . The plot s show t he effect of changes in ( a) dept h of cut on ( b) feed and ( c) cut t ing force wit h increasing t im e. Source: Ref 9

However, the ACC system can still have poor performance problems, and even tool breakage or controller instability (Ref 10, 11). The experimental results in Fig. 9 illustrate the instability that occurs because of changes in the cutting process parameters (for example, depth of cut, spindle speed, and even feed rate). The cutting process, as shown in Fig. 7, is part of the control loop, and unless the controller gain is varied to compensate for the process parameter variations, consistent closed-loop performance cannot be obtained. Thus, the process parameter variations require the use of techniques from adaptive control theory (Ref 1). In other words, the controller gains must be adapted on-line to the changing process parameters. Such variable-gain ACC systems have been developed and are described in the section "Trends in AC Systems" in this article.

Fig. 9 Turning experim ent s wit h an ACC syst em wit h a high cont roller gain and st ep changes in dept h of cut . Plot s show t he effect of changes in ( a) dept h of cut on ( b) feed and ( c) cut t ing force wit h increasing t im e. Source: Ref 9

Ge om e t r ic Ada pt ive Con t r ol Geometric adaptive control (GAC) systems combine inspection of the finished product with on-line adjustment of feeds and speeds. In this section, some basic characteristics of GAC systems are described and then a particular GAC system for the turning of cylindrical parts is presented. Fu n da m e n t a ls of GAC Syst e m s. Another AC strategy is to base the selection of feed rates and speeds on the

dimensional accuracy or surface roughness of the part being machined. Such an approach puts the emphasis on product quality rather than metal removal rate. Thus, such AC systems are most suited for finishing operations. These GAC

systems require a direct or indirect measurement of the product simultaneously as it is being machined or possibly after it has been machined. For example, as illustrated in Fig. 10, indirect measurement can be obtained by means of the sensing of spindle deflections in a milling operation (Ref 12, 13). Direct measurements typically require laser gaging systems or other optical methods (Ref 14), and are usually performed immediately after machining.

Fig. 1 0 Two views of t he deflect ion of an end m ill during m achining showing t he various com ponent s of force, F; m om ent , M; and deflect ion, . ( a) Front view. ( b) Side view. Source: Ref 12

GAC Syst e m for Tu r n in g. An example of a GAC system is illustrated in Fig. 11. This system, developed at the

research laboratories of a major automobile manufacturer, measures the dimensional accuracy and surface roughness of the part after machining to manipulate the feed rate in a turning operation. Near real time measurement and control are used in the turning of cylindrical parts. For a given cutting speed, the principal factors affecting surface roughness are the feed and the tool wear. Surface finish and dimensional accuracy are measured for each part at the completion of the cutting operation, and corrections then affect subsequent parts. The goal is to adjust the feed rate to reduce variability in the surface finish that arises because of tool wear. The results obtained from some laboratory experiments are shown in Fig. 12, 13, and 14. Figure 12 shows the feed being adjusted to maintain the desired surface finish. Roughly speaking, there are four distinguishable regions: • • • •

Break-in region, where the initial feed is adjusted (pieces 1 to 10) Steady-state region, where the desired feed is maintained (pieces 11 to 32) Wear out region, where roughness increases because of wear (pieces 33 to 46) Failure region, where the tool must be replaced (pieces 47 to 53)

A comparison of the diameter and surface finish histograms in Fig. 13 and 14 shows the significant improvement in dimensional accuracy and surface finish achieved with the GAC system.

Fig. 1 1 A geom et ric adapt ive cont rol syst em for t urning cylindrical part s. Source: Ref 14

Fig. 1 2 ( a) Surface finish in a GAC syst em as det erm ined by ( b) feed rat e. Source: Ref 14

Fig. 1 3 Hist ogram plot s of diam et er dim ensional accuracy wit h ( a) convent ional ( nonadapt ive) cont rol and ( b) geom et ric adapt ive cont rol. Source: Ref 14

Fig. 1 4 Hist ogram plot s of surface roughness wit h ( a) convent ional ( nonadapt ive) cont rol and ( b) geom et ric adapt ive cont rol. Source: Ref 14

Tr e n ds in AC Syst e m s In this section, current research issues and future trends in AC systems are briefly summarized. First, the tool wear and breakage problem is discussed. This is followed by a discussion of variable-gain AC systems, which can adapt controller gains to changing process parameters. Finally, some brief comments are made regarding the role of AC systems in computer-integrated manufacturing (CIM). ACO Syst e m s a n d Tool W e a r / Br e a k a ge . As discussed in the previous sections, the most general AC strategy is adaptive control with optimization. However, ACO systems are typically very difficult to implement, the major difficulty being the need for a practical on-line tool wear sensor. Current research on the use of force sensing together with modelbased estimation techniques for on-line tool wear sensing appears promising (Ref 15). These methods have already been shown to work successfully when flank wear, Wf, is dominant (see Fig. 15) and have also been applied to flank wear estimation in turning with varying depths of cut and feed rates (see Fig. 16). Acoustic emissions from the metal cutting process have also been extensively investigated for tool wear and breakage detection, as well as for chip entanglement detection (Ref 18). Acoustic emission signals, together with frequency domain signal processing and pattern recognition techniques, will need further development before they can be used effectively for tool wear sensing in an industrial environment.

Fig. 1 5 Est im at ed versus m easured flank wear in t urning. Est im at ed values were obt ained from force m easurem ent . Source: Ref 16

Fig. 1 6 Act ual flank wear versus est im at ed flank wear in four separat e t est s wit h varying dept hs- of- cut .

St raight line indicat es t he expect ed result s for perfect est im at ion. Source: Ref 17

Va r ia ble - Ga in ACC Syst e m s. Another area of recent research is related to variable-gain ACC systems (Ref 2). These are AC systems for metal cutting that are in fact adaptive in the sense commonly used in the control literature. In these process controllers, the controller gains are varied to maintain consistent closed-loop performance despite process parameter variations (Fig. 17). A process model is used to estimate on-line the changing process parameters, and the controller gain is then varied according to the adjustment policy. The turning system described previously in the section "Adaptive Control With Constraints," which incorporates a variable gain controller, has been shown to yield excellent results (Fig. 18). There has been considerable interest in designing variable-gain ACC systems for turning (Ref 19, 20) and for milling (Ref 11, 21).

Fig. 1 7 A variable- gain ACC syst em for t urning. The est im at ed force, Fe, is com pared t o t he m easured force, F, t o obt ain t he m odel error, Em . This is used t o adj ust t he m odel param et ers. The error, E, bet ween t he reference force, Fr , and t he m easured force, F, is used for t he cont rol. The cont roller gains are dependent on t he m odel param et ers. Source: Ref 19

Fig. 1 8 A variable- gain ACC syst em for t urning wit h est im at ion of t he process gain, Km , and adj ust m ent of t he cont roller gain, Kc, for st ep changes in t he dept h of cut , a, wit h increasing t im e. Also shown are t he variat ions of cut t ing force, F, and feed rat e, f. Source: Ref 19

I n t e gr a t ion of Ada pt ive Con t r ol I n t o CAD / CAM / CI M Syst e m s. An important issue in the future development

of AC systems is their complete integration into CNC systems, as well as their role in a CIM system hierarchy. Such issues are extremely important for unmanned manufacturing and will require additional research to extend our current understanding of AC systems. Important issues include the interface between computer-aided design (CAD) and computer-aided manufacturing (CAM), and the application of expert systems and other methods, from artificial intelligence to AC systems, as well as process monitoring and diagnostics. Detailed information on interfacing adaptive control techniques with CAD/CAM systems is available in the article "CAD/CAM Applications in Machining" in this Volume.

Re fe r e n ce s 1. G.C. Goodwin and K. Sin, Adaptive Filtering, Prediction, and Control, Prentice-Hall, 1984 2. A.G. Ulsoy, Y. Koren, and F. Rasmussen, Principal Developments in the Adaptive Control of Machine Tools, J. Dyn. Syst., Meas., Cont. (Trans. ASME), Vol 105 (No. 2), 1983, p 107-112 3. Y. Koren, Computer Control of Manufacturing Systems, McGraw-Hill, 1983 4. R.S. Pressman and J.E. Williams, Numerical Control and Computer-Aided Manufacturing, John Wiley & Sons, 1977 5. J. Huber and R. Centner, "Test Results With an Adaptive Controlled Milling Machine," Paper MS68-638, American Society of Tool and Manufacturing Engineers, 1968 6. R. Centner, Final Report on the Development of an Adaptive Control Technique for a Numerically Controlled Milling Machine, "Technical Documentary Report ML-TDR-64-279," U.S. Air Force, Aug 1964 7. G. Amitay, S. Malkin, and Y. Koren, Adaptive Control Optimization of Grinding, J. Eng. Ind. (Trans. ASME), Vol 103 (No. 1), 1981, p 102-111 8. Y. Koren, The Optimal Locus Method for Control of Manufacturing Processes, J. Dyn. Syst., Meas., Cont. (Trans. ASME), 1988 9. O. Masory and Y. Koren, Adaptive Control System for Turning, Ann. CIRP, Vol 29 (No. 1), 1980 10. O. Masory and Y. Koren, Stability Analysis of a Constant Force AC System for Turning, J. Eng. Ind. (Trans. ASME), Vol 107 (No. 4), 1985, p 295-300 11. L.K. Lauderbaugh and A.G. Ulsoy, Model Reference Adaptive Force Control in Milling, J. Eng. Ind. (Trans. ASME), 1988 12. G. Stute, Adaptive Control, in Proceedings of the Machine Tool Task Force Conference, Vol 4, Sect 7.14, Wright Aeronautical Laboratories, U.S. Air Force, 1980 13. T. Watanabe and S. Iwai, A Control System to Improve the Accuracy of Finished Surfaces in Milling, J. Dyn. Syst., Meas., Cont. (Trans. ASME), Vol 105, 1983, p 192-199 14. C.L. Wu, R.K. Haboush, D.R. Lymburner, and G.H. Smith, Closed Loop Machining Control for Cylindrical Turning, Modeling, Sensing and Control of Manufacturing Systems, American Society of Mechanical Engineers, Nov 1986, p 189-204 15. Y. Koren, Flank Wear Model of Cutting Tools Using Control Theory, J. Eng. Ind. (Trans. ASME), Vol 100 (No. 1), Feb 1978 16. K. Danai and A.G. Ulsoy, An Adaptive Observer for On-Line Tool Wear Estimation in Turning, Part I, Theory; Part II, Results, Mech. Syst. Signal Proc., Vol 1 (No. 2), 1987, p 211-240 17. Y. Koren, A.G. Ulsoy, and T.R. Ko, Estimation of Tool Wear Under Varying Cutting Conditions, in Proceedings of the 14th NSF Conference on Production Research and Technology, National Science Foundation, 1987, p 73-77 18. E. Kannatey-Asibu, Jr. and D. Dornfeld, Quantitative Relationships for Acoustic Emission From Orthogonal Metal Cutting, J. Eng. Ind. (Trans. ASME), Vol 103 (No. 3), 1981, p 330-340 19. O. Masory and Y. Koren, Variable-Gain Adaptive Control Systems for Machine Tools, J. Manuf. Syst., Vol 2 (No. 2), 1983, p 165-174 20. L.K. Daneshmend and H.A. Pak, Model Reference Adaptive Control of Feed Force in Turning, J. Dyn. Syst., Meas., Cont. (Trans. ASME), Vol 108 (No. 3), 1986, p 215-222

21. M. Tomizuka, J. Oh, and D. Dornfeld, Model Reference Adaptive Control of the Milling Process, Control of Manufacturing Processes and Robotic Systems, American Society of Mechanical Engineers, p 55-64, Nov 1983

CAD / CAM Applica t ion s in M a ch in in g Em ory W. Zim m ers, Jr., Lehigh Universit y

I n t r odu ct ion CAD/CAM is a term that refers to computer-aided design and computer-aided manufacturing. This technology is moving in the direction of greater integration of design and manufacturing, which have traditionally been treated as distinct and separate functions in a production firm. When the machining function is included as part of a plantwide CAD/CAM system, many benefits accrue. For example, more accurate information is provided to and received from the machining operation, and the total cycle time from concept to finished machined part is reduced. Com pu t e r - a ide d de sign (CAD) can be defined as the use of computer systems to assist in the creation, modification,

analysis, or optimization of a design. The computer systems consist of the hardware and software used to perform the specialized design functions required by a particular customer. The CAD hardware typically includes the computer, one or more graphics display terminals, keyboards, and other peripheral equipment. The CAD software consists of the computer programs necessary to implement computer graphics on the system and application programs to facilitate the engineering functions of the customer. Examples of these application programs include stress-strain analysis of components, dynamic response of mechanisms, heat-transfer calculations, and numerical control part programming. Many of these techniques, which were historically available only to the component designer, are becoming available to personnel responsible for planning and performing machining operations. This method replaces intuitive decision making with a more scientific approach leading to higher product quality and improved production rates. Com pu t e r - a ide d m a n u fa ct u r in g (CAM) can be defined as the use of computer systems to plan, manage, and control

the operations of a manufacturing plant through either direct or indirect computer interface with the production resources of the plant. As indicated by the definition, the applications of CAM fall into two broad categories, computer monitoring and control and manufacturing support. Computer process monitoring involves a direct computer interface with the manufacturing process for the purpose of observing the process and associated equipment, for example, collecting data from the machining process. The computer is not used to control the operation directly. The control of the process remains in the hands of machinists, who may be guided by the information compiled by the computer. Computer process control goes one step further than monitoring by not only observing the process but also controlling it, based on the observations. With computer monitoring, the flow of data between the machining process and the computer is in one direction only, from the process to the computer. In process control, the computer interface allows a two-way flow of data. Signals are transmitted from the machining process to the computer, just as in the case of computer monitoring, but, in addition, the computer issues command signals directly to the machining process, based on control algorithms contained in its software. An example of this is adaptive control machining (see the article "Adaptive Control" in this Volume). In addition to the applications involving a direct computer process interface for the purpose of process monitoring and control, CAM also includes indirect applications in which the computer serves a support role in the manufacturing operations of the plant. In these applications, the computer is not directly linked to the manufacturing process. Instead, the computer is used off-line to provide plans, schedules, forecasts, instructions, and information by which production resources can be managed more effectively. An example of CAM for manufacturing support is numerical control (NC)

part programming by computers in which control programs are prepared for automated machine tools (see the article "Numerical Control" in this Volume). In the presentation to follow, elements of a CAD/CAM system relevant to machining are presented. In addition, the hardware, interfaces, and benefits of integrating machining with the CAD/CAM system of the manufacturing plant are discussed.

Fu n ct ion s of a CAD / CAM Syst e m The basic functions that should be included when considering the relationship of CAD/CAM and machining are: • • • • • • • •

Design Analysis Drafting Process planning Part programming Program verification Part machining Inspection

When electronically linked, these provide the foundation for efficient machining operations. Each of these functional areas will be examined to gain a better understanding of the activity involved in each and the interrelationships among the various activities. D e sign . Starting from a rough sketch or possibly just an idea, the designer, manufacturing engineer, and other qualified

personnel use the graphic display as if it were a drafting board to construct the geometry of a part, tool, or fixture design. Several parts and/or fixtures can be simultaneously displayed so that the designer can ensure that they will fit together properly and that there will be no interference upon assembly. Computer-aided design systems are at their best when designing families of parts or machining fixtures (parametric design). A component can be designed so that the variable parameters are input data. The component can then be called up as a library function, and the values of the required parameters can be specified. In this way, the design of the entire family of parts, and ultimately the machining instruction set, can be completed in little more time than is required to produce the design for one part or fixture. Most CAD/CAM systems provide some form of layering. A layer is simply one way of describing a part or machining fixture. This can be a particular geometric view, for example, a rough cast shape or a finished shape, or a nonpictorial way of describing the part, such as a bill of material or Automatically Programmed Tools (APT) part program. Layering is used heavily throughout the design process, in drafting, in planning for machining, and in NC part programming. An a lysis Ca pa bilit ie s. Although traditionally thought of as being applicable only in the initial stages of the

manufacturing cycle, analysis capabilities are being more fully utilized in support of the machining process. M a ss Pr ope r t ie s. The analysis capability of a CAD/CAM system that has perhaps the widest application is that of mass properties. This feature automatically gives the perimeter, area, center of gravity, and moments of inertia of any cross section. The properties of the cross section when rotated about any axis to form a solid are also given. These include exposed surface area, volume, weight, center of gravity, and moments of inertia. Applications range from moments of inertia (for nonsymmetric parts to be machined) to contour roll die design. Many problems in machining relating to chatter or poor surface finish can be analyzed and solved using these capabilities. St a ck u ps. The ability to display an assembly of various components is used to good advantage in axial stack-up

calculations. In this application, the engineer or machinist has all the components of interest displayed on the screen. In the time it would normally take to describe his requirements to a draftsman, he is able to identify the areas for which he wants a nominal clearance and to obtain their values.

Fin it e - Ele m e n t An a lysis. Many parts or sections of parts are subjected to severe heat transfer and stress conditions

during the machining process, for example, in pocket milling or machining thin-wall sections. Typically, few calculations are made because of the time or skill level involved. One mathematical tool used to calculate the effects of these on a given part is finite-element model analysis. In this technique, the three-dimensional model is divided into a number of small cube- or wedge-shape elements, which are analyzed separately (Fig. 1). The total effect on the part is the sum total of all of the subdivisions.

Fig. 1 A finit e- elem ent m odel m esh used for st ress and t herm al analysis of m achined part s. Court esy of L. Niggem ann, Schlum berger Technologies, CAD/ CAM Division

The finite-element grid is generated automatically through software developed for finite elements. The user has the ability to change any portion of this grid pattern if so desired. The elements are numbered, and the precise coordinates of the nodal points are obtained automatically by the graphics system. This information is output from the system in a format that requires no human intervention before being input to a computer for analysis. Should the analysis indicate that the geometry requires modification, this is done rapidly at the cathode ray tube (CRT) terminal, and the entire process is repeated. With the availability of more local computing capability and with improvements in networking, use of these techniques has increased at the factory level. When this occurs, more precise calculations are substituted for intuition and trial and error approximations by the machinist, setup person, or NC part programmer. D r a ft in g. To perform this function, a draftsman accesses the data base and retrieves the part design. A series of steps is

followed: • • • • • • • •

The detail work is performed Lines and fillets are trimmed Layout lines are removed Areas are crosshatched Drill holes are identified Dimensions are added (generally available in English or metric units, or both) Tolerances are noted All necessary text is placed on the drawing

Company standards for drawings can be easily enforced by programming the standard into the system. Aids to the draftsman include transformations to assist in developing oblique, isometric, and perspective views, and the capability to

replicate details that appear at various places on the design. Information critical to the machining process is prepared at this stage. Common applications include surface contour and styling drawings, layouts for preliminary design, exploded views, perspective drawings, dimensioned drawings, isometric drawings, nondimensioned master drawings for production design, fixture layouts, interference problem layouts, assembly drawings, and subassembly drawings. Pr oce ss Pla n n in g. The engineering drawing of the workpart must be interpreted in terms of the manufacturing

processes to be used. This step, referred to as process planning, is concerned with the preparation of a route sheet. The route sheet is a listing of the sequence of operations that must be performed on the workpart. It is called a route sheet because it also lists the machine tools through which the part must be routed in order to accomplish the sequence of operations. The foundation of the most commonly used retrieval-type computer-aided process planning (CAPP) systems is parts classification and coding and group technology. In this approach, the parts produced and their associated machining fixtures are grouped into families according to their manufacturing characteristics. For each part family, a standard process plan is established. The standard process plan is stored in computer files and retrieved for new workparts that belong to that family. Some form of parts classification and coding system is required to organize the computer files and to permit efficient retrieval of the appropriate process plan for a new workpart. Figure 2 illustrates the flow of information in a CAPP system.

Fig. 2 I nform at ion flow in a ret rieval- t ype com put er- aided process planning syst em

For some new parts, editing the existing process plan may be required. This is done when the manufacturing requirements of the new part are slightly different from the standard. The machine routing may be the same for the new part, but the specific operations required at each machine may be different. The complete process plan must document the operations

as well as the sequence of machines through which the part must be routed. Because of the alterations that are made in the retrieved process plan, a CAPP system such as this is sometimes also called a variant system. It is important that accurate, updated machining information be included as part of the overall process. This requires feedback and updating as the inevitable machining improvements are made on the factory floor. In many CAD/CAM systems, this occurs as electronic data transfer from a shop floor terminal. If this is not done, it is likely that the system will produce suboptimal machine information on such things as feeds and speeds. Pa r t Pr ogr a m m in g. In producing an NC part program, the accurate geometric definition of the part that is present in

the data base is used by the part programmer. This data base geometric definition eliminates the step of defining the part geometry from drawings, which is the first step in traditional (APT-like) NC programming. Auxiliary statements concerning cutting speed, feed rate, coolants, and tool changes must be entered. The definition of the cutter path is then normally specified interactively, rather than by the traditional APT process of mentally envisioning the cutter path. The part program can often be done through a graphic presentation of the machining operation (if the CAD system is of the refresh type, this is often accomplished through animation). Figure 3 shows such a graphic presentation. The resulting NC part program is usually in APT/CL (cutter location) file format.

Fig. 3 Graphics display of a t ooling fixt ure and com ponent part wit h t he associat ed t ool pat h clearly defined t o generat e t he necessary CNC m achine code out put . Court esy of L. Niggem ann, Schlum berger Technologies, CAD/ CAM Division

Typically, NC programming can be done for two, two and one-half, three, four, and five axis machining and for machines including milling machines, lathes, punch presses, drilling machines, and most of the other digitally controlled machining operations. Pr ogr a m Ve r ifica t ion . With many CAD/CAM systems, verification using computer simulation software is available at

varying levels of sophistication. The basic objective is to reveal what can be anticipated during the actual machining process, primarily by visual verification of the images created during the simulation. The image may appear on a graphics display terminal, a plotter, or other hard copy device. Often the hard copy output can be included in the operator setup and operations sheets used by the machinist. The advantages include making the verification process more productive and predictable, providing an efficient mechanism for allowing NC data to be optimized, reducing scrap, reducing inspection of first-run parts, and freeing production machines for production work. It should be noted that the level of acceptance of the simulation model is dependent upon the accuracy and completeness with which it depicts the real world. The graphic representation may be a wire frame model, a solid model, or a shaded image. There may be the capability of displaying the raw stock from which the part is to be cut. The software may or may

not be able to display the full geometry of cutting tools, as well as toolholders, and/or the clamps and fixtures for locating and holding the workpiece. Desirable features for manipulating the image include the capability to orient the part for optimum viewing advantage, to display more than one view on the screen, and to modify a view by sectioning, reorienting, zooming, or other operations. Color can be effectively used to enhance comprehension, for example, by indicating rapid moves (in red), a gouge cut in a fixture, or the path cut by each tool (using different colors). Of great use is the ability to interrogate the model to identify the command in the tool path that caused a detected error. CAD/CAM and NC programming systems usually offer wire frame models (Fig. 4) for tool path simulation. There are usually several options for tool path display. These include the display of the centerline of the cutter (which may include the tool axis vector), the display of a two-dimensional representation of the cutter (a disk for mills and the insert, and perhaps a toolholder outline for lathes), a display of a three-dimensional representation of the cutter, and a display of textual information about the cut, for example, identification of point-to-point cycles (drill, bore, and so on), programmed feed rates, and cutting sequence number identification.

Fig. 4 Prelim inary wire fram e geom et ry of a connect ing bracket shown in t hree different window orient at ions t o facilit at e visualizat ion and enhance t he design process. Court esy of L. Niggem ann, Schlum berger Technologies, CAD/ CAM Division

The display of the tool path is usually fairly rapid; therefore, the ability to slow down the display, step through each cutter location position in individual steps, or zoom in on specific details may be required to discover whether the workpiece is being cut correctly (or whether the fixture is inadvertently being cut incorrectly). Also, actions or cycles that are initiated by Machine Control Data (MCD) codes, such as special machining cycles, touch probe or robotic functions, or repetitive programming codes for loops or macros that compress the length of MCD to fit Machine Control Unit (MCU) memory, are usually not understood by the NC programming software. Therefore, the actions or movements resulting from those commands would not be available to the tool path simulation display. Any hidden errors could only be discovered during a tryout on the machine tool. In many cases, however, the NC program is uncomplicated enough to instill a high level of confidence by means of the wire frame simulation model alone. A machine tryout is still required on complex parts, however, to reduce the potential for catastrophic error. Pa r t M a ch in in g. In the machining process itself, the CAD/CAM system provides information (often to the machine

tool site) in the form of instructions for the operator and/or instructions to the machine tool. To fully utilize the CAD/CAM technology, this linkage should be electronic and timely. Information provided can include setup and operation instructions, along with detailed machining settings such as feed and speeds. On advanced systems, some of this information is provided in a graphic format. When a direct interface to a numerically controlled machine tool is installed, an additional benefit is often derived. Information can be provided to the CAD/CAM system regarding operational status, which is valuable for optimizing overall plant performance. This type of information often includes uptime status of the machine tool and the causes of delay, such as a broken tool or a waiting interval for tooling.

I n spe ct ion . The CAD/CAM system can facilitate the inspection of machined parts with complex shapes. The use of

various cut planes allows the part to be measured by an inspector, after which the CAD/CAM system generates a view of the part. In some systems, provision is made for rapid electronic feedback to specific machining operations by means of the CAD/CAM system. When this is done, scrap is reduced and "drift" resulting from tool wear or other causes can be detected in time to take corrective action.

Be n e fit s of Usin g CAD / CAM for M a ch in in g Tool D e sign . The ability to obtain a view of a part from any desired angle (provided by CAD/CAM) is a great aid in

designing tooling. For example, in the design of holding fixtures for machining compound angles, information that takes days to calculate by traditional methods can be easily obtained by automatic measurement from the CAD/CAM system. An a lysis. The analysis routines available in a CAD/CAM system help to consolidate the analysis process into a more

logical work pattern. Rather than having a back and forth exchange between manufacturing and analysis groups, a single person can perform the analysis while remaining at an engineering work station. Generally, the engineer can retain his train of thought more easily. There is a time savings to be derived from analysis routines, both in engineer time and in absolute time. This savings results both from the rapid performance of the analysis and from the time no longer lost when the design is routed by traditional means from the shop to the analyst and back again. Un de r st a n da ble D r a w in gs. Interactive graphics are just as adept at creating and maintaining isometric and oblique

drawings as they are at producing the simpler orthographics. All drawings can be generated and updated with equal ease. Thus, an up-to-date version of any drawing type is always available. In general, understanding a drawing relates directly to the projection used. Orthographic views are less comprehensible than isometrics. Both are less understandable than a perspective view. Most conventional construction drawings are line drawings. The addition of shading increases comprehension (Fig. 5). Different colors further enhance understanding. Thus, the machinist or machine tool setup person can reference an up-to-date part description in the format most useful for the machining task to be performed.

Fig. 5 Com put er- generat ed shaded m odel of a connect ing bracket . Court esy of L. Niggem an, Schlum berger Technologies, CAD/ CAM division

N u m e r ica l Con t r ol Pa r t Pr ogr a m m in g. One typical NC part programming language is Automatically Programmed

Tools, described in the article "Numerical Control" in this Volume. A powerful and versatile language, APT has three potential disadvantages. The first is that the user must learn a language with its own syntax and grammar. Unless he already has computer programming experience, he is exposed to some concepts that are entirely alien to him. A second disadvantage is that the part programmer must interpret the engineering drawings (with the possibility of error) and define the geometry of the part for APT. The third disadvantage is that the programmer must mentally visualize the tool path as he is programming. Experience has shown that these disadvantages either prevent capable machine shop personnel from becoming part programmers or else make the learning time unacceptably long. On the other hand, CAD/CAM uses

language easily understood by machine shop personnel. This eases the problem of computer shock and eliminates the need to learn a completely new language. Because the geometry of the part is defined in the CAD/CAM data base, there is no need to go through the process of extracting the part geometry from the drawings. The geometry is already given, precisely as the designer specified it. The graphic display and interactive nature of the system eliminate the need to envision the cutter path, because the user is provided with visual verification of every step in the process. These advantages have a significant impact on the time necessary to produce an NC tape and to train an NC part programmer. The experience that the General Electric Company had with the design and manufacture of the T-700 engine indicated that NC tapes were produced in half the time taken by conventional APT programming and that the time required to train a methods person to make NC tapes was reduced by a factor of six. Additional information on NC part programming can be found later in this article. I m pr ove d Accu r a cy. When CAD/CAM is used, there is a high level of dimensional control, far beyond the levels of

accuracy attainable manually. Mathematical accuracy is often to 14 significant decimal places. The accuracy delivered by interactive CAD/CAM systems for machining three-dimensional curved-space designs is much greater than that offered by manual methods. Computer-based accuracy is beneficial in many ways. Parts are labeled by the same, recognizable nomenclature and number throughout all drawings. In some CAD/CAM systems, a change entered on a single item can appear throughout the entire documentation package, effecting the change on all drawings that use that part. Such accuracy also shows up in the form of fewer engineering change orders, more accurate material and cost estimates, and tighter procurement scheduling. These last two points are especially important, for example, for long lead time material purchases.

CAD / CAM H a r dw a r e Com pon e n t s As described below, the hardware for a CAD/CAM system includes a graphics terminal(s), digitizer (optional), alphanumeric terminal, hard copy unit, output plotter(s), secondary storage device, machine control unit, and auxiliary hardware, all of which interface with the central processing unit. Figure 6 illustrates the general hardware setup for a CAD/CAM system.

Fig. 6 General hardware configurat ion of a CAD/ CAM syst em

Th e Gr a ph ics Te r m in a l. The most visible part of a CAD/CAM system from the point of view of the user is the graphics terminal. Virtually all state-of-the-art CAD/CAM systems are provided with some form of CRT similar to those used in televisions, oscilloscopes, and radars. In addition, CRTs are available in many sizes and configurations and with various capabilities. The majority of CAD/CAM systems provide either a "storage" CRT or a "refresh" CRT.

Cu r sor Con t r ol. The cursor is generally a bright spot on the screen that indicates where lettering or drawing will be

placed. There is normally a method by which the position of the cursor can be read by the computer. Therefore, controlling the cursor position enables one to enter locational information into the computer. This information can be used in a geometric sense, such as to define a point. A more sophisticated use interprets the cursor position as the selection of an item from a menu (the screen is divided into sections, one of which corresponds to each menu section). There are three basic divisions in cursor control devices. The first type includes those that control the cursor without a direct physical relationship to the screen (from the view-point of a user). For example, there are two thumbwheels used to control the cursor; one controls the horizontal position of the cursor, and the other controls the vertical position. The thumbwheels are turned to change the position of the cursor. Another type of device is the light pen. The cursor follows the light pen, with the computer seeking the position of the light by trying to place the bright cursor under the pen. The third type of device is a digitizing tablet and pen. This type of tablet detects the position of the pen and moves the cursor to a corresponding position on the screen. D igit ize r s. A manual digitizer has a cursor that can be moved about a large, smooth table. It is oriented by means of a

standard x-y coordinate system through which the computer can detect its location. The cursor generally has a cross hair target. Drafters find the digitizer not unlike their drafting boards. It tilts, raises, and lowers. Some digitizers are available with backlighting for easy tracing or with free-floating cursors. Automatic raster scanners can be used to digitize continuoustone photographs and line drawings. Automatic line followers reduce a drawing to a series of lines rather than a stream of points. By recording only the information needed to reproduce the lines, the output is condensed. However, a line follower operator must intervene in the entry process to select the direction to be followed at line intersections. Alph a - N u m e r ic Te r m in a l. Nearly all CAD/CAM systems employ an alpha-numeric terminal along with the graphics

terminal. Often terminals are accessible at the machine tool itself. Such a terminal is used to enter supplemental information, commands (if there is no cursor-type menu), alpha-numeric input, and system commands by means of a keyboard. The terminal receives messages from the system, design analysis output, program listings, and error messages. The advantage of a separate terminal is that these messages can be displayed and entered without erasing or over-writing the image on the graphics screen. The terminal is either a CRT or a hard copy terminal. Each type has its advantages. The CRT provides much faster output and does not produce masses of scrap paper. The hard copy terminal provides a permanent listing, often desirable in debugging or for permanent records. H a r d Copy Un it . There is normally a hard copy unit provided to make copies of the screen in a few seconds. These

copies can be used as permanent records of intermediate steps or on occasions when a rough hard copy of the screen is desired. These units are not suitable for final drawings because they are not as accurate as plotters. Ou t pu t Plot t e r s. There are many types of plotters available for use on CAD/CAM systems, including high-speed drum plotters, flatbed plotters, and computer output microfilm plotters. Th e Ce n t r a l Pr oce ssin g Un it . Functioning as the heart of a CAD/CAM system, the central processing unit (CPU) has traditionally been a mainframe for large installations or a minicomputer for smaller turnkey systems. The CPU does the necessary mathematical calculations and directs all activity within the system. It acts as the master controller and manager of workstation activity such as storage and plotting.

It is important to note that the traditional view of CAD/CAM systems as a stand-alone entity is evolving. The CAD/CAM functions and support hardware/software should be considered in the context of an integrated factorywide machining system. As this occurs, computational activity will be distributed in a cost-effective manner, and cross-departmental communication of technical/operational information will be more efficient. The host CPU in most configurations directs plotters in precisely what to draw, copies disk data onto magnetic tapes for semiactive storage, reads magnetic tapes of drawings and/or documentation for revision or other reuse, and transmits data to and from other computers in the CAD/CAM system.

Se con da r y St or a ge . The CAD/CAM data base is normally maintained on magnetic disks along with much of the

CAD/CAM system software. Disks or similar direct-access devices provide rapid access to the information stored on them because of their random access configuration. CAD/CAM systems generally have the capability of handling multiple disk drives, totaling many hundreds of megabytes of storage. Magnetic tape drives are normally available. They are used as disk backup and for permanent archival file storage. Information access is not nearly as rapid from a magnetic tape as it is from a disk because the tape is a sequential storage device. Therefore, to access any piece of information, all the information preceding it must first be read. For archival files and backup, however, the infrequent use makes the slow access time perfectly acceptable. Information transfer rates from magnetic tape are typically half as fast as for disks. However, this is still extremely rapid; magnetic tape is an extremely compact and inexpensive method of storing large amounts of data. M a ch in e Con t r ol Un it . This hardware device is located where the machining operation occurs. It consists of the electronics and hardware that read and interpret the program of instructions and convert it into mechanical actions of the machine tool. The typical elements of a conventional NC controller unit include the interface to the CAD/CAM factory system or stand-alone reader, a data buffer, signal output channels to the machine tool, feedback channels from the machine tool, and the sequence controls that coordinate the overall operation of the foregoing elements. It should be noted that nearly all modern NC systems have a microcomputer as the controller unit.

A reader is an electromechanical device for reading the tape or other magnetic medium containing the program of instructions. The data contained on the tape are read into the data buffer. The signal output channels are connected to the servomotors and to other controls in the machine tool. Through these channels, the instructions are sent to the machine tool from the controller unit. To make certain that the instructions have been properly executed by the machine, feedback data are sent to the controller by means of the feedback channels. The most important function of this return loop is to ensure that the table and the workpart have been properly located with respect to the tool. M a ch in e Con t r ol Pa n e l. The control panel, or control console, contains the dials and switches by which the machine

operator runs the NC system. It may also contain data displays to provide information to the operator. Although the NC system is an automatic system, an operator is still needed to turn the machine on and off, to change tools (some NC systems have automatic tool changers), to load and unload the machine, and to perform various other duties. To be able to discharge these duties, the operator must be able to control the system, and this is done through the control panel. Addit ion a l H a r dw a r e . The hardware described in the above sections is mainly based on mainframe use. In addition to this, an interface module is required to transfer the data from the mainframe to the machine tool. Two types of interface modules are widely used. A direct-interface module is used when a postprocessing program is available on the mainframe. A postprocessor converts the code generated by the NC program into machine-readable instructions. A direct interface, for example, the NC machine controller, transfers the machine-readable instructions from the main-frame to the machine. An indirect-interface module is used when a postprocessing program is not available on the mainframe. Examples of this type of controller are modified personal computers (PCs) configured as part of the machine tool controller and advanced NC machine controllers. In this case, code generated by the NC program on the mainframe is transferred to the PC or the machine controller. A postprocessing program on a PC or controller converts the code into machine-readable instructions and is then transferred to the machine.

As personal computers have become more powerful and less expensive, they are becoming extremely popular in NC machining. Many of the functions (design, analysis, NC programming, and so on) are now available on PCs. Minimum hardware requirements to perform these functions on PCs are an 80286 or equivalent processor, a 640 KB (kilobyte) random access memory, a 10 MB (megabyte) hard drive, an RS232 port, a math coprocessor, and a high-quality plotter for hard copy.

CAD / CAM D a t a Ba se All of the functions of a CAD/CAM system revolve around its data base. Depending on the overall structure of the machining equipment, the physical hardware location or locations may vary. The discussion below will cover logical, not physical, organization. The use of a common data base allows the simplification of many tasks. Once an item is entered into the data base, it can be accessed and used by any subsequent operation. For example, in creating an NC part program prior to machining, there is no need to reenter geometry of the part because it can be recalled from the data base.

Essentially, everything contained on the disks of the system is part of the data base. Thus, the data base includes much of the CAD/CAM system software (such as system commands, multiple menus, macros, plotter output routines, NC postprocessors, and compilers). The data base also contains programming routines, part models, designs, drawings, and assemblies, as well as the alphanumeric information associated with the machining task, such as the results of optimized metal cutting routines and text for machinist or technician reference. The data base can be set up in various ways, depending on the requirements of the vendor. Much of this activity centers around the part drawing. Some vendors store a drawing as a copy of the design geometry, while others store it as a pointer to the appropriate model along with the appropriate transformation information. This compacts the data base and simplifies the process of updating all the drawings of a part when the design is changed, because the only change necessary would be one to the model. The value of these transformation techniques can easily be seen. If one wished to rotate an entire drawing by a certain amount and also make it larger, one could simply multiply each line in the drawing by the proper transformation matrix. Three-dimensional work is much more complex (and involves about three times as much calculation), but it is easy to imagine the extension of the matrix transformation concept to points located in three-dimensional space. Use of the proper transformation matrix can produce combinations of shearing, local scaling, rotation, reflection, translation, perspective, and overall scaling. The matrix transformation technique allows the construction of views from any desired viewpoint. Stored in the data base is a full, three-dimensional, geometric description of the part. To obtain a view of the part, one multiplies each component by appropriate transformation matrices. Thus, it is possible to store a drawing of a part essentially as a transformation matrix and a pointer to the specific mathematical model of the part. For example, many systems are configured to store the mathematical description of a part and generate drawings or views based on matrix mathematics transformations. The pointer references the specific part description, and the transformation matrix generates the required view. This allows a tremendous space savings if more than one drawing of a part is stored. Of course, the entire process is invisible to the machinist or other user: The vendor-provided CAD/CAM software determines the transformation matrix, and the computer performs the numerical calculations. A general understanding of this concept is useful, however, in that it helps to remove the mystique from a CAD/CAM system and reinforces the fact that no individual item in a CAD/CAM system is incomprehensible when viewed in isolation without the large number of components surrounding and camouflaging it.

CAD t o CAM I n t e r fa ce s Experience has shown that the machining benefits obtained from an integrated CAD/CAM system are far greater than those realized by applying CAD and CAM as separate technologies. Hence, bridging the gap between CAD and CAM is of great importance. Ideas start out as concepts of the three-dimensional (3D) part that must be machined. Unfortunately, traditional documentation usually consists of transferring those ideas into two-dimensional (2D) representations. To automate production, the 2D drawing must be reinterpreted and, in a certain sense, recreated as a 3D part for the purposes of creating an NC part program prior to machining. With an integrated CAD/CAM system, the common data base allows the use of the 3D part design already created while the model was being built during the design and drafting process. All dimensions are stored to an accuracy typically far greater than that of the most precise machining tool; hence, the preparation for NC machining essentially consists of generating the tool path and creating the necessary files. Nearly all CAD/CAM systems that are capable of producing NC tapes do so with output in the form of a cutter location (CL) file. This is suitable for input to a postprocessor. Typically, postprocessing is accomplished on another computer, but many vendors also provide postprocessors that can be run on the same computer as the CAD/CAM system. There are many advantages associated with using the CAD/CAM data as a basis for NC programming. Among the advantages are: •

Tooling is standardized

• • • • • • •

Most simple, calculational errors are eliminated Documentation is kept up to date in the common data base Programmers are freed from many tedious operations, leaving more time for truly creative work and eliminating many errors caused by boredom Personnel do not have to be sophisticated programmers Frequently a user-defined library of macros can be set up to handle difficult tasks, allowing semitechnical personnel to operate the system Productivity and task turnaround times improve immensely. Figures typically quoted range from a 2:1 to a 10:1 increase in productivity Visual verification of tool paths prior to machining provides for less tooling (tool collisions are essentially eliminated), less wasted motion and discarded material, and fewer dry runs

In systems that take full advantage of the CAD/CAM data base, a tool path is often generated from a geometric boundary. This is in contrast to the conventional method of describing a path by means of a series of commands, such as "go right," "go forward," or "tangent to." The surface to be followed can normally be specified with the cursor control by pointing the cursor to the proper surface. For example, APT postprocessor statements are inserted at appropriate points in the specification of the cutter path. Once a tool path has been created, it contains all the necessary information to produce the complete machine motion. Tool paths range from a two and one-half axis motion to full 3D contouring. Automatic generation of pocketing around open and closed boundaries and point-to-point operation proceed directly from an initial setup specification from the user. Full 3D contouring includes profiling and pocketing to a surface and/or surface intersection. Three and five axis motions may also be programmed. The tool paths may be displayed as they are created as an interactive means for establishing optimum machine tool flow or in a determined sequence showing intermediate tool points, in other words, in the actual machining sequence. If a refresh-type terminal is used, animation of the cutter path may be available. This means that the operator has all the necessary capabilities to create full simulation of the NC process before actually attempting a machining operation; he can "see" the part being run on the graphics display with geometric verification of output. Typically, it is possible for a library of user-created routines to be set up that allow the performance of complicated tasks by merely inserting a few parameters. An example is preset tool paths for a family of parts that are recognizable as such by associated descriptive text and other parameters. Routines for automatic calculation of the number and depth of cuts (rough and final finish cuts) based on tables contained in common storage can also be set up. Another example is separate user routines that are called (invoked) by standard NC routines. These special routines are essentially nontechnical in nature and contain extensive prompting. This opens up the possibility of having fewer technically qualified people operate the system by providing closely tailored programming using simple, conversational questions in a menu-type format.

N C Pa r t Pr ogr a m m in g Part programming involves the planning and specification of the sequence of processing steps to be performed on the NC machine. It also involves, although less directly, the preparation of the input medium by which the processing instructions are communicated to the machine tool or cell. With computer-assisted part programming, the machining instructions are written in Englishlike statements in the NC programming language, which are processed by a computer either to prepare the physical medium or for electronic filing. When using one of the NC programming languages, part programming essentially consists of two tasks: • •

Defining the geometry of the workpart Specifying the tool path and/or operation sequence

No matter how complicated the workpart may appear, it is composed of basic geometric elements. Nearly any component that can be conceived of by a designer can be described by points, straight lines, planes, circles, cylinders, and other

mathematically defined surfaces. It is the part programmer's task to enumerate the component elements out of which the workpart can be formed. Each geometric element must be identified, and the dimensions and location of the element must be explicitly defined. After defining the workpart geometry, the programmer must construct the path that the cutter will follow to machine the part. This tool path specification consists of a detailed step-by-step sequence of cutter moves. The moves are made along the geometry elements that have previously been defined. By using NC programming language statements, the tool can be directed to machine along the workpart surfaces, proceed to point locations, drill holes at those locations, and so on. The job of the computer in computer-assisted part programming consists of: • • • •

Input translation Arithmetic calculations Cutter offset computation Postprocessing

The part programmer submits his NC part program to the computer in the NC programming language he is using. The input translation component converts the coded instructions contained in the program into computer-usable form in preparation for further processing. The arithmetic calculations unit of the system consists of a comprehensive set of subroutines for solving the mathematics required to generate the part surface. These subroutines are called (invoked) by the various part programming language statements. The arithmetic calculations unit is in fact the fundamental element in the part programming package. This unit frees the programmer from time-consuming geometry and trigonometry calculations so he can concentrate on the workpart processing. The purpose of the cutter offset computation is to offset the tool path from the desired part surface by the radius of the cutter. This allows the part programmer to define the exact part outline in his geometry statements. Numerically controlled machine tools have different features and capabilities. They use different NC tape formats. Therefore, the final task of computer-assisted part programming is to take the general instructions and make them specific to a particular machine tool system. The unit that performs this task is called a postprocessor. The postprocessor is really a separate computer program that has been written to prepare the punched tape for a specific machine tool. The input to the postprocessor is the output from the other three components. The output is the NC tape written in the correct format for the machine on which it is to be used.

D ir e ct N u m e r ica l Con t r ol There are a number of problems inherent in conventional NC that have motivated machine tool builders to seek improvements in the basic NC system. Among the difficulties encountered in using conventional NC systems are problems with the tape or the tape reader and the fact that the controller unit is hard wired; that is, its control features cannot be easily altered to incorporate improvements. It was with these problems in mind that the machine tool builders developed the concept of using the general-purpose computer to control NC machines. Their concept has come to be called direct numerical control (DNC). Direct numerical control can be defined as a manufacturing system in which a number of machines are controlled by a computer through a direct connection and in real time. The tape reader, which was the least reliable component of the NC system, is omitted in DNC. Instead of using the tape reader, the part program is transmitted to the machine tool directly from the computer memory. The system consists of four components: • • •

A central computer Bulk memory, which stores the NC part programs Communication links

•

Machine tools

The computer calls the part program instructions from bulk storage and sends them to the individual machines as the need arises. It also receives feedback data from the machines. This bidirectional information flow occurs in real time, which means that requests from each machine for instructions must be satisfied almost instantaneously. A diagram of a typical DNC system is shown in Fig. 7.

Fig. 7 Diagram of a direct num erical cont rol syst em

Com pu t e r N u m e r ica l Con t r ol Over the years, the cost and physical size of the digital computer have been greatly reduced. At the same time, its computational capabilities have increased. The result of these improvements has been the maturing of computer numerical control (CNC). Computer numerical control is an NC system that uses a dedicated, stored-program computer to perform the basic NC functions. Because a digital computer is used in both CNC and DNC, there is often confusion surrounding the two system types. The principal differences between the two are: •

• •

Direct numerical control computers distribute instructional data to, and collect data from, a number of machines or machining cells. With CNC, only one machine or a small number of machines in a cell can be controlled Direct numerical control computers occupy a location that is typically remote from the machines under their control. Computer numerical control computers are located close to their machine tools Direct numerical control software is developed not only to control individual pieces of production equipment but also to serve as part of a management information system in the manufacturing sector of a firm. Computer numerical control software is developed to augment the capabilities of a particular machine tool

Almost all CNC machine tools provide tape editing on the shop floor. This means that the NC tape can be optimized during tape tryout at the site of the machine tool. Frequently, CNC machines have local CAD capabilities, allowing the operator to access the workpiece or other graphic information needed on the shop floor. The general configuration of a CNC system is shown in Fig. 8.

Fig. 8 General configurat ion of a CNC syst em

D ist r ibu t e d N u m e r ica l Con t r ol A logical extension of direct numerical control is referred to as distributed numerical control. This is complementary to and consistent with the concept of an integrated CAD/CAM system for machining. The new distributed numerical control combines the centralized data base feature of DNC with the distributed computer power available in CNC machine tools. This creates a communications network for the shop floor that eliminates paper tape and provides an automated part program library, status reporting on machine tool operation, and the capability to run part programs of practically infinite length. The configuration of a DNC system is such that the CL file is generated by compiling the APT part program or by generating a tool path from geometry created with a CAD software package. This CL file is sent to the machine-specific postprocessor, which generates the instruction set to operate the NC machine tool. The postprocessor is a software package running on either the CAD/CAM computer system or a satellite computer. A statement in the CL file specifies which postprocessor is to be used to generate the instruction set, which is then sent to the proper NC machine tool over a communications link. This same link is used to gather production data from the NC machine tools. A block diagram of a typical distributed numerical control system is shown in Fig. 9.

Fig. 9 Diagram of a dist ribut ed num erical cont rol syst em . EDM, elect rical discharge m achining

Some of the reasons for acceptance of DNC as an effective approach to manufacturing include: • • • • • • • •

The availability of low-cost microprocessor adapters that can interface to any vendor's machine tool The price reduction of microcomputers and magnetic disk drives, which makes these computers costeffective management tools for the distribution of part programs The development of data communications facilities such as fiber optics that can operate reliably in the often hostile environment of the shop floor The growth in the population of CNC machine tools, which makes their integration into a shop network a logical productivity improvement The establishment of standards for part programs and data communications The familiarity with computers that shop owners and managers have acquired The success of CAD/CAM and Materials Requirement Planning (MRP), which is prompting managers to look for additional ways of increasing productivity through computers The realization by management that the success of the factory of the future is dependent upon the integration of the production process with the engineering and planning operations

This integration process is the basis of distributed numerical control. The integration of production tools and CAD systems will provide the means and flexibility whereby the real benefits of flexible manufacturing can be realized. In order to optimize the benefits of DNC, certain elements of a state-of-the-art system should be recognized. Editing, printing, and plotting part programs are all necessary functions of a distributed numerical control system. Editing provides a means of changing part programs on the shop floor or even manually writing a part program without tying up an expensive machine tool for this purpose. Distributed numerical control terminals should be provided for editing and printing part programs. Plotting or other graphic verification of part programs should also be a by-product of a distributed numerical control system.

File management provides an automatic library for part programs. It should include automatic storage, purging, deleting, copying, scheduling, and renaming of the revision of a part program. The distributed numerical control file manager must provide a technique for tracking, storing, and updating revisions of part programs. The DNC system should also offer rugged terminals that can exist in the manufacturing environment. The terminals should have a flexible security structure that allows management to assign responsibility and functionality where it is needed, not where the distributed numerical control system supplies it. Many operators on a factory or shop floor do not have computer training or experience. Most CNC operators have keyboard experience and a sense for computer operation, but they are neither computer operators nor NC programmers. A DNC system should provide a menu technique for accessing the functions of the system. A well-designed system will also allow NC programmers to access additional functions where needed. A distributed numerical control system must be designed to connect economically to any machine. It is an acknowledged industry fact that NC controllers have not been standardized. Different shops have machine tools of varying NC vintage and capability. A distributed numerical control system must interface with all NC machine tools as well as robots and other numerically driven machines. A state-of-the-art DNC system will provide interfaces and integrations with a CAD/CAM or part programming system. Computer networks are also a future productivity tool. Many computer manufacturers have advanced computer networks, and many offer a standard network. It is important that the distributed numerical control system be easily integrated into a widely accepted computer network by standard, well-maintained products. The integration into a standard computer network will simplify future extension of the distributed numerical control system of the plant to many of the shared vendor data bases that major manufacturers will be establishing. An industry-accepted operating system will permit the integration of other manufacturing tasks to the distributed numerical control system of the plant.

Se le ct e d Re fe r e n ce s • • • • • • • • • • • • • • • • • • •

Autofact '88 Conference Proceedings, Society of Manufacturing Engineers, 1988 K. Blache, Success Factors for Implementing Change: A Manufacturing Viewpoint, Society of Manufacturing Engineers, 1988 CAD/CAM, 2nd ed., Society of Manufacturing Engineers, 1985 CAD/CAM, Videotape, Society of Manufacturing Engineers CAD/CAM Networking, Videotape, Society of Manufacturing Engineers N. Chiantella, Ed., Management Guide for CIM, CASA/SME, 1986 S. Cousins, Ed., Integrating the Automated Factory, Society of Manufacturing Engineers, 1988 D. Goetsch, Fundamentals of CIM Technology, Delmar Publications, 1988 G. Graham, Ed., Automation Encyclopedia: A to Z in Advanced Manufacturing, Society of Manufacturing Engineers, 1988 M.P. Groover, Automation, Production Systems, and Computer Integrated Manufacturing, PrenticeHall, 1987 T.G. Gunn, Computer Applications in Manufacturing, Industrial Press, 1981 J. Harrington, Computer Integrated Manufacturing, Industrial Press, 1973 L.G. Lamit and V. Paige, Computer-Aided Design and Drafting (CADD), Merrill Publishing, 1987 Y.C. Pao, Elements of Computer-Aided Design and Manufacturing, John Wiley & Sons, 1984 P.G. Ranky, Computer Integrated Manufacturing, Prentice-Hall, 1986 U. Rembold, C. Blume, and R. Dillman, Computer Integrated Manufacturing Technology and Systems, Marcel Dekker, 1985 K. Taraman, Ed., CAD/CAM Integration and Innovation, Society of Manufacturing Engineers, 1984 E. Teicholz, Ed., CAD/CAM Handbook, Society of Manufacturing Engineers, 1985 E. Teicholz and J.N. Orr, Computer-Integrated Manufacturing Handbook, McGraw-Hill, 1987

•

R.W. Yeomans, A. Choudry, and P.J.W. Ten Hagen, Design Rules for A CIM System, North-Holland, 1985

Abbr e via t ion s a n d Sym bols o •

Abbreviations and Symbols

a

•

A

•

Ac

•

Af

•

As

•

A

• •

•

wheel depth of cut in grinding; crystal lattice length along the a-axis

•

area

•

area of cut

•

area of sliding contact on the rake face

•

shear plane area

•

ampere

•

angstrom

•

alternating current

•

adaptive control

ac

•

AC

•

ACC •

•

ACI

•

ACO

•

AFM

• • •

adaptive control with constraints Alloy Casting Institute adaptive control with optimization abrasive flow machining

•

AGMA

•

AGV

• •

•

AIA

•

AISI

•

AJM

• • •

•

ANSI

•

APT

•

• •

American Gear Manufacturers Association automated guided vehicle Aerospace Industries Association American Iron and Steel Institute abrasive jet machining American National Standards Institute automatically programmed tool

ASP •

Anti-Segregation Process

•

ASTM

•

at. %

• •

•

atm

•

AWJ

• •

American Society for Testing and Materials atomic percent atmosphere (pressure) abrasive waterjet

•

AWM

•

AWS

•

b

•

•

B

•

bal

•

American Welding Society

•

crystal lattice length along the b-axis

•

grinding width

•

balance or remainder

•

body-centered cubic

•

Baumé (specific-gravity scale)

•

bcc

•

Bé

•

BLU •

•

BUE

•

c

built-up edge

•

specific heat of workpiece material; cost; crystal lattice length along the c-axis

CAD

•

CAM

•

CAPP

• • •

•

CBN

•

CCPA

•

computer-aided design computer-aided manufacturing computer-aided process planning cubic boron nitride

•

Cemented Carbides Producers Association

•

carbon equivalent

•

CE

•

CFR

•

CHM

• •

•

CIM

•

CL

•

cm

Code of Federal Regulations chemical milling

•

computer integrated manufacturing

•

cutter location

•

centimeter

•

CM

•

CMM

• •

•

CNC

•

cpm

•

CPM

• •

chemical machining (milling) coordinate measuring machine computer numerical control cycles per minute

•

Crucible Particle Metallurgy

•

cycles per second

•

cps

•

CPU •

CR

basic length unit

•

•

•

abrasive waterjet machining

central processing unit

• •

CRT

•

CSD

•

CVD

• • •

•

CVN

•

CW

•

d

•

•

D

•

DE

•

DS

•

diameter; distance

•

equivalent diameter

•

wheel diameter

•

workpiece diameter

•

direct current

•

DIN

•

DNC

•

DOC

• •

Deutsche Industrie-Normen (German Industrial Standards) direct numerical control

•

depth of cut

•

natural log base, 2.71828

•

modulus of elasticity (Young's modulus)

•

end-of-block (character); electron beam

•

EB

•

EBM

•

ECEA

• •

•

ECM

•

ECG

• •

•

EDM

•

EIA

•

ELI

• • •

•

ELP

•

EP

•

Eq

•

Charpy V-notch (impact test or specimen)

depth of cut; used in mathematical expressions involving a derivation (denotes rate of change); diameter

dc

E

chemical vapor deposition

•

•

•

controlled spray deposition

continuous wave

DW

e

cathode ray terminal

•

•

•

cold rolled

electron beam machining end cutting edge angle electrochemical machining electrochemical grinding electrical discharge machining Electronic Industries Association extralow interstitial

•

electropolishing

•

extreme pressure

•

equation

et al.

• •

ETP

•

f

•

ft

•

F

•

Fc

•

Ff

•

Fm

•

Fn

•

FN

•

Fr

•

Fs

•

Ft

•

fcc

and others

•

electrolytic tough pitch

•

feed rate

•

feed per tooth

•

force

•

primary (horizontal) cutting force

•

feed force

•

momentum force

•

normal force on shear plane

•

normal force in grinding

•

radial or thrust force

•

shear force

•

tangential force

•

face-centered cubic

•

Fig.

•

FMS

•

ft

•

figure

•

flexible manufacturing system

•

foot

•

GAC

•

gal.

• •

•

GPa

•

h

•

h

geometric adaptive control gallon

•

gigapascal

•

hour

•

undeformed chip thickness; height

•

HAG

•

HAZ

•

HB

•

higher-accuracy grinding

•

heat-affected zone

•

Brinell hardness

•

HBN

•

hcp

• •

•

HIP

•

HK

• •

hexagonal boron nitride hexagonal close-packed hot isostatic pressing Knoop hardness

•

hp

•

HPG

•

HPSN

• • •

•

HR

•

HSLA

• •

•

HSS

•

HV

•

Hz

•

horsepower higher-productivity grinding hot-pressed silicon nitride Rockwell hardness (requires scale designation, such as HRC for Rockwell C hardness) high-strength low-alloy (steel) high-speed steel

•

Vickers hardness

•

hertz

•

IBR

•

IC

•

integrally bladed rotator

•

inscribed circle

•

inside diameter

•

inch

•

ID

•

in.

•

INCRA •

International Copper Research Association

•

inch per tooth

•

ipt

•

ISO •

International Organization for Standardization

•

joule

•

J

•

JIT

•

k

•

K

•

K

•

KIc

•

kg

•

kgf

•

just-in-time

•

thermal conductivity

•

Kelvin

•

stress intensity factor

•

plane-strain fracture toughness

•

kilogram

•

kilogram force

•

kilometer

•

kilobyte

•

km

•

KB

•

kPa

•

ksi

•

kV

•

•

kilopascal

•

kips (1000 lbf) per square inch

•

kilovolt

kW

kilowatt

•

length

•

liter

•

length

•

pound

•

pound force

•

• •

L

•

L

•

lb

•

lbf

•

LBM

•

ln

•

laser beam machining

•

natural logarithm (base e)

•

common logarithm (base 10)

•

log

•

LSG •

low-stress grinding

•

meter

•

m

•

MB

•

mg

•

megabyte

•

milligram

•

megagram (metric tonne)

•

Mg

•

min •

•

MHS

•

mL

•

mm

• • •

•

MPa

•

MPIF

•

• •

material handling system milliliter millimeter megapascal Metal Powder Industries Federation

MRP •

•

MRR

•

MRS

•

materials requirement planning material removal rate

•

minimum residual stress

•

millisecond

•

ms

•

MSDS

•

N

•

material safety data sheet

•

newton

•

speed of rotation (rev/min); number of cycles to failure

•

N

•

NASA

•

minimum; minute

•

National Aeronautics and Space Administration

•

numerical control

NC

•

NFPA

•

nm

•

No.

•

National Fire Protection Association

•

manometer

•

number

•

NPS

•

NPSC

• •

•

NPT

•

NPTF

•

ns

•

American National Standard Straight Pipe Thread American National Standard Straight Pipe Thread for Couplings American National Standard Taper Pipe Thread

•

Dryseal USA (American) Standard Taper Pipe Thread

•

nanoseconds

•

OD

•

ODS

•

outside diameter

•

oxide dispersion-strengthened

•

oxygen-free

•

OF

•

OFHC

•

OSHA

•

•

oz

•

p

•

p

•

P

•

Pg

•

Ps

•

Pa

•

PC

•

Occupational Safety and Health Administration

•

ounce

•

page

•

pitch

•

cutting power

•

gross power

•

specific power

•

pascal

•

personal computer

•

PCBN

•

PCD

• •

•

PCM

•

pH

•

PH

polycrystalline CBN (cubic boron nitride) polycrystalline diamond

•

photochemical machining

•

negative logarithm of hydrogen-ion activity

•

precipitation hardenable

•

PLC

•

P/M

•

oxygen-free high-conductivity (copper)

• •

ppm

programmable logic controller powder metallurgy

• •

PSA

•

psi

•

PTP

•

pressure-sensitive adhesive

•

pounds per square inch

•

•

PVD

•

r

•

R

physical vapor deposition

•

radius; ratio of uncut chip thickness to chip thickness

•

stress (load) ratio; radius; gas constant; resultant force

•

roughness average

Ra

•

Rmax Rq

•

Ry

•

Rz

point-to-point

•

•

•

parts per million

•

maximum peak-to-valley roughness height

•

root-mean-square roughness average

•

maximum peak-to-valley roughness height

•

ten-point height (roughness average)

•

reference

•

Ref

•

rem •

remainder

•

revolution

•

rev

•

RHR

•

rms

•

roughness height rating

•

root mean square

•

second

•

s

•

SAE •

•

SCC

•

SCEA

•

SEM

• • •

Society of Automotive Engineers stress-corrosion cracking side cutting edge angle scanning electron microscopy

•

SFFER

•

sfm

•

SFR

• •

•

SI

•

SP

silica-flour-filled epoxy resin surface feet per minute

•

smooth face rubber

•

Systàme International d'Unités

•

specific power

•

serrated rubber

•

SR

•

SUS •

Saybolt Universal seconds

•

t

•

tc

•

T

•

depth of cut; feed or uncut chip thickness; time; tool lifetime; thickness

•

chip thickness

•

temperature

•

TEA

•

TEM

• •

•

TIR

•

TRS

•

tsi

•

•

tons per square inch

•

TWR

•

•

uf

•

um

tool wear rate

•

specific energy

•

friction energy per unit volume

•

kinetic (momentum) energy per unit volume

•

shear energy per unit volume

us

•

UNS •

•

USM

•

UTS

•

V

•

V

•

Vc

• •

F

Vs

Unified Numbering System ultrasonic machining

•

ultimate tensile strength

•

volt

•

cutting velocity

•

chip velocity

•

slide infeed removal rate (grinding)

•

shear velocity

•

radial wear rate of grinding wheel

•

S

•

W

•

thermal shock resistance

•

•

•

total indicator reading transverse rupture strength

TSR

u

thermal energy method

•

•

•

triethanolamine

•

radial removal rate of work being ground

•

volume

vol

•

vol%

•

W

•

w

•

volume percent

•

watt

•

width; weight

•

wc

•

W

•

WJM

•

width of chip

•

chip weight

•

•

WRP

•

wt%

• •

•

YAG

•

yr

•

ZW

•

year

•

workpiece total volumetric removal rate (grinding)

•

°

•

°F

• • •

workpiece unit-width volumetric removal rate (grinding)

•

angular measure; degree

•

degree Celsius (centigrade)

•

degree Fahrenheit

•

direction of reaction

•

divided by

•

equals

•

approximately equals

•

not equal to

•

identical with

•

greater than

•

much greater than

•

greater than or equal to

•

infinity

•

is proportional to; varies as

•

integral of

•

less than

•

much less than

•

less than or equal to

•

maximum deviation

•

minus; negative ion charge

=

• •

>

• • • • • •

•

÷

•

•

weight percent yttrium-aluminum-garnet

Z'W

°C

work removal parameter

•

•

•

waterjet machining