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EDITORIAL REVIEW COMMITTEE P.W. Taubenblat, Chairman I.E. Anderson, FAPMI T. Ando S.G. Caldwell S.C. Deevi D. Dombrowski J.J. Dunkley Z. Fang B.L. Ferguson W. Frazier K. Kulkarni, FAPMI K.S. Kumar T.F. Murphy J.W. Newkirk P.D. Nurthen J.H. Perepezko P.K. Samal H.I. Sanderow D.W. Smith, FAPMI R. Tandon T.A. Tomlin D.T. Whychell, Sr., FAPMI M. Wright, PMT A. Zavaliangos INTERNATIONAL LIAISON COMMITTEE D. Whittaker (UK) Chairman V. Arnhold (Germany) E.C. Barba (Mexico) P. Beiss (Germany) C. Blais (Canada) P. Blanchard (France) G.F. Bocchini (Italy) F. Chagnon (Canada) C-L Chu (Taiwan) H. Danninger (Austria) U. Engström (Sweden) N.O. Grinder (Sweden) S. Guo (China) F-L Han (China) K.S. Hwang (Taiwan) Y.D. Kim (Korea) G. Kneringer (Austria) G. L’Espérance, FAPMI (Canada) H. Miura (Japan) C.B. Molins (Spain) R.L. Orban (Romania) T.L. Pecanha (Brazil) F. Petzoldt (Germany) S. Saritas (Turkey) G.B. Schaffer (Australia) Y. Takeda (Japan) G.S. Upadhyaya (India) Publisher C. James Trombino, CAE
[email protected] Editor-in-Chief Alan Lawley, FAPMI
[email protected] Managing Editor James P. Adams
[email protected] Contributing Editor Peter K. Johnson
[email protected] Advertising Manager Jessica S. Tamasi
[email protected] Copy Editor Donni Magid
[email protected] Production Assistant Dora Schember
[email protected] President of APMI International Nicholas T. Mares
[email protected] Executive Director/CEO, APMI International C. James Trombino, CAE
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international journal of
powder metallurgy Contents 2 5 7 9
44/2 March/April 2008
Editor's Note PM Industry News in Review PMT Spotlight On … Rajendra Kelkar Consultants’ Corner B. Pittenger
FOCUS: PM Machinability 13 Machining of PM Materials: A Secondary Shaping Operation of Primary Concern C. Blais
15 Characterization of PM Machinability: Practical Approach and Analysis D. Christopherson
21 Machining of PM Steels: Effect of Additives and Sinter Hardening B. Lindsley
33 Effect of Prealloyed MnS Content and Sintered Density on Machinability and Mechanical Properties P. Boilard, G. L’Espérance and C. Blais
41 Green Machining: Parameters, Applications, and Sintered Properties É. Robert-Perron and C. Blais
49 Face Turning of PM Steels: Effect of Porosity and Carbon Level A˘. Salak, M. Selecká, K. Vasilko and H. Danninger
DEPARTMENTS 62 Meetings and Conferences 63 PM Bookshelf 64 Advertisers’ Index Cover: Crater wear on carbide tooling. Photo courtesy Bruce Lindsley, Hoeganaes Corporation, Cinnaminson, New Jersey. The International Journal of Powder Metallurgy (ISSN No. 0888-7462) is a professional publication serving the scientific and technological needs and interests of the powder metallurgist and the metal powder producing and consuming industries. Advertising carried in the Journal is selected so as to meet these needs and interests. Unrelated advertising cannot be accepted. Published bimonthly by APMI International, 105 College Road East, Princeton, N.J. 08540-6692 USA. Telephone (609) 4527700. Periodical postage paid at Princeton, New Jersey, and at additional mailing offices. Copyright © 2008 by APMI International. Subscription rates to non-members; USA, Canada and Mexico: $95.00 individuals, $220.00 institutions; overseas: additional $40.00 postage; single issues $50.00. Printed in USA by Cadmus Communications Corporation, P.O. Box 27367, Richmond, Virginia 23261-7367. Postmaster send address changes to the International Journal of Powder Metallurgy, 105 College Road East, Princeton, New Jersey 08540 USA USPS#267-120 ADVERTISING INFORMATION Jessica Tamasi, APMI International INTERNATIONAL 105 College Road East, Princeton, New Jersey 08540-6692 USA Tel: (609) 452-7700 • Fax: (609) 987-8523 • E-mail:
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EDITOR’S NOTE
“
A paradox, a paradox, A most ingenious paradox! We’ve quips and quibbles heard in flocks, But none to beat this paradox.” So goes the lyric from The Pirates of Penzance, the Gilbert and Sullivan operetta which premiered in 1879. Over a century later, the PM industry is challenged by a tenet contrary to logic, and yet true. Paradoxically, PM is a net-shape, or near-net-shape, technology, yet roughly half of all PM components require a machining operation. Hence, the machining of PM materials is a secondary operation of prime concern. This timely Focus Issue, coordinated by Carl Blais, addresses the current state of the art of the science and practice of PM machinability. In particular, it addresses machinability characterization, the role of admixed and prealloyed additives in improving the machinability of ferrous components, including sinter-hardenable compositions, green machining, and operative mechanisms dictated by porosity. Contributing again to the “Consultants’ Corner,” Brian Pittenger addresses readers’ questions in the domain of powder transport. In particular, approaches to minimize or eliminate powder agglomeration in the feedshoe, and to enhance the consistency of flow of fine iron particles, are discussed. Beginning with the January/February issue, APMI International members may now access the Journal online. The e-version is identical in content to the printed form, plus it offers easy, convenient navigation features and powerful search capabilities. All that is needed to login is your personal user ID and password. The convenient features allow the user to navigate from page to page with a simple click, and the table of contents is always only one click away. The search capability allows the user to enter a word or phrase and the software will display and highlight the location results: no more thumbing through pages to remember where the information was located. E-versions of past issues will be archived in the APMI Members Only area.
Alan Lawley Editor-in-Chief
Based on processing limitations, and cost normalizing of properties, the prognosis for extensive commercialization of metal nanopowders in bulk press-and-sinter PM parts is, at best, questionable. Attendant incremental improvements in performance, compared with those of conventional metal and ceramic powders, do no compensate for their higher cost (IJPM, 2007, vol. 43, no. 6, pp. 28–30, and, 2008, vol. 44, no. 1, pp. 44–54). Applications are limited to energetic formulations based on aluminum nanopowders for propellants, explosives, and pyrotechnic compositions. In marked contrast, the market for organic electronic materials is predicted to top $15.8 billion by 2015 (www.nanomarkets.net)! This market embraces semiconductors, conductors, dielectrics, and substrates that will be used in the growing organic electronics industry. New kinds of organic semiconductors, based on hybrid materials, including formulations with carbon nanotubes, are expected to improve electron mobilities, switching speeds, and enhance environmental stability.
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Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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SCM's products include: • • • • • • •
North Carolina USA
Manufacturing Sites • Research Triangle Park, North Carolina USA • Suzhou, China Tel: 919-544-8090 • www.SCMmetals.com
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SHIFT UP TO T110 PM. Today’s part manufacturers require powders with the highest compressibility to achieve the near full densities needed for new automotive gears and sprockets. Inco T110 PM nickel powder offers a performance boost to sintered steels, without the loss in compressibility associated with prealloyed iron powders. Increased diffusion of T110 PM nickel during sintering can double hardenability and significantly improve mechanical properties when compared to standard nickel powder. And with over 100 times as many particles, superior distribution of nickel leads to better part uniformity and greater dimensional precision. At Inco Special Products, we provide nickel solutions for your materials challenges.
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PM INDUSTRY NEWS IN REVIEW The following items have appeared in PM Newsbytes since the previous issue of the Journal. To read a fuller treatment of any of these items, go to www.apmiinternational.org, login to the “Members Only” section, and click on “Expanded Stories from PM Newsbytes.”
Press Maker Reorganizes Process Equipment Cincinnati Incorporated, Cincinnati, Ohio, has integrated its Powder Metal Products press line into the Molded Materials Process Equipment (MMPE) Group as part of an expansion program into new growth markets. PM presses were previously sold and serviced through a separate department. Copper Infiltrating Alloy Wins Patent Ultra Infiltrant, Carmel, Ind., reports that the U.S. Patent office has approved its patent application for a wrought copper alloy system for infiltrating PM parts. Expected to be issued within six weeks, the patent contains 13 claims covering alloy composition, infiltrating temperature, and the material in wire form as well as wafer, disk, and washer configurations. Injection Molding Machine Builder Insolvent Battenfeld Kunststoffmaschinen GmbH, Kottingbrunn, Austria, has filed for insolvency in Vienna, reports Modern Plastics Magazine. The filing came on the heels of the company’s sale by its owners, the Adcuram Group AG, Munich, Germany, to another private equity firm on December 21, 2007. New Technology Center Technogenia Inc. has opened a new Lasercarb Technology Center near Houston in Conroe, Tex. The facility uses a diode laser cladding machine that deposits abrasion-resistant materials onto industrial parts for the oil and gas industries.
Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
Hoeganaes Names President David Kasputis was appointed president of Hoeganaes Corporation, Cinnaminson, N.J., on January 1. He succeeds Robert Fulton who held the position since July 1994. New High-Performance Permanent Magnet Material Researchers at the U.S. Department of Energy’s (DOE) Ames Laboratory, Ames, Iowa, have designed a PM alloy for permanent magnets made by metal injection molding (MIM). The new material, which provides good magnetic strength at 392°F, is aimed at making electric drive motors more efficient and less costly for electric cars, fuel-cell automobiles and plug-in hybrids. New Alloy-Powder Atomizing Unit Onstream Castolin Eutectic Ireland Ltd. (CEIL), Dublin, Ireland, is operating a new anti-satellite gas-atomizing facility supplied by Atomising Systems Limited (ASL), Sheffield, England. CEIL makes alloy powders based on nickel, cobalt, and iron for welding, thermal spray, and brazing applications. CMW Receives Safety Award CMW Inc., Indianapolis, Ind., received a safety award from its insurance carrier, AmCOMP Inc., for operating almost two years without a lost-time accident. The company’s safety committee meets monthly to address potential safety hazards and educates employees about best practices. PM2008 World Congress Program Now Online MPIF has just launched a dedicated
Web site for the 2008 World Congress on Powder Metallurgy & Particulate Materials at www.mpif.org/meetings/2008/ 08_intro.htm. Sponsored by the Metal Powder Industries Federation and APMI International and held only once every six years in North America, the congress will take place in Washington, D.C., June 8–12, at the new Gaylord National Resort & Convention Center at National Harbor, Maryland. The 2008 International Conference on Tungsten, Refractory & Hardmaterials VII, also organized by MPIF and APMI, will run concurrently with the congress at the same location. Rapid Prototyping via Laser Sintering Vaupell Rapid Solutions, a division of Vaupell Molding and Tooling, Inc., Hudson, N.H., has purchased a direct metal laser sintering (DMLS) system from EOS of North America Inc., Novi, Mich. Vaupell will produce tooling inserts, prototype parts, and other products. Metaldyne Supplies Chinese Automaker Metaldyne, Plymouth, Mich., an Asahi Tec unit, will supply powertrain and chassis parts to Chery Automotive Co., Ltd., in China. Chery will develop vehicle applications with Metaldyne including suspension parts, oil pumps, and powder-forged connecting rods. New PM Consulting Service Prima Problem Solving, Pittsburgh, Pa., has opened a new Operational Specialist Services division offering assistance in manufacturing ijpmand
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PM INDUSTRY NEWS IN REVIEW
operational improvement for the powder metallurgy industry. It will also provide help to other pressing, processing, and machining companies. Turkish PM Conference Issues Call for Papers Papers are requested for the Fifth International Powder Metallurgy Conference, October 8 to 12, 2008, in Ankara, Turkey. The abstractsubmission deadline is March 31. Mahindra Aiming at U.S. Market Mahindra & Mahindra Ltd., Mumbai, India, plans to sell a compact SUV and two pickup truck models in the U.S. next year, according to Automotive News. The three models will feature four-cylinder diesel engines and six-speed transmissions.
MPIF Launches E-learning Initiative The Metal Powder Industries Federation has just announced the launch of its long-awaited e-learning program. The online courses on various aspects of powder metallurgy, available to anyone with a computer, are designed to help companies involved in any phase of the PM industry with their workforce-training requirements. Swedish Powder Maker Reports 14 Percent Gain Höganäs AB, Sweden, reports 2007 sales of metal powders rose 14 percent to about $930 million (5,838 MSEK). Although powder shipments grew five percent for the year, the company attributes the sales increase mainly to higher metals prices.
GKN Reports on 2007 PM Results GKN plc, London, reports that 2007 PM sales, which includes GKN Sinter Metals and Hoeganaes Corporation increased modestly to £602 million (about $1.2 billion) compared with £582 million (about $1.16 billion) in 2006. Sinter Metals makes PM parts and Hoeganaes produces principally ferrous-based metal powders. New High-Performance Furnace Pusher Plate Sunrock Ceramics Company, LLC, Chicago, Ill., offers a new highperformance alumina pusher plate for hydrogen atmosphere hightemperature sintering furnaces. The HPA-CG material has been tested and qualified for production by leading furnace manufacturers and PM parts makers, the company reports. ijpm
PURCHASER & PROCESSOR
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1403 Fourth St. • Kalamazoo, MI 49048 • Tel: 269-342-0183 • Fax: 269-342-0185 Robert Lando E-mail:
[email protected]
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SPOTLIGHT ON ...
RAJENDRA M. KELKAR, PMT Education: BE Metallurgical Engineering, M.S. University of Vadodara, India, 1995 MS Materials Engineering, Illinois Institute of Technology (IIT), 2003 Why did you study powder metallurgy/particulate materials? One of my undergraduate professors encouraged me to consider PM before coming to the U.S. I was aware of this field, but possessed little knowledge about it. I took a chance and asked my advisor at IIT, professor Philip Nash, about working on his PM project; it was an immediate fit. This gave me an exposure to PM and I equated the technology to cooking—you mix powders and process them in different ways to see what the resulting alloy offers. My initial work at IIT resulted in the decision to make PM my career path. When did your interest in engineering/science begin? At a young age I liked to break/make things. As an adult, my interest continued and engineering offered a logical career path. What was your first job in PM? What did you do? In my current job at SSI Technologies, I work as a materials engineer in the R&D department. Job responsibilities include: material development; failure analysis; metallurgical laboratory analysis/testing; continuous-improvement initiatives; and plant-related issues. Materials development includes conventional PM, metal injection molding, full-density products, and porous parts. My manager, Peter DePoutiloff, has been a constant mentor and source of encouragement. Describe your career path, companies worked for, and responsibilities. I started my career with ISPL Industries Ltd., India, a manufacturer of high-tensile fasteners, as a trainee
Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
engineer in 1995. Later, in 1996, I joined Birla Copper in their smelting department as a process engineer/shift supervisor. Here I was exposed to the construction and commissioning of plants. Subsequently, I gained experience in converting, refining and anode casting of copper. After working for almost five years in India, my passion brought me to IIT, Chicago, to pursue a master’s program in Metallurgical and Materials Engineering. During my graduate tenure, I was able to work on projects involving the cryogenic treatment of M2 tool steels, and the sintering of Ti-6Al-4V. After working briefly for heat-exchanger and circuitboard manufacturers, I accepted a position with SSI Technologies in 2004. What gives you the most satisfaction in your career? I prefer to get involved with new and innovative projects. My current job fulfills my passion for R&D and innovation. When I see successful production parts developed, based on R&D, it gives me a deep sense of satisfaction. List your MPIF/APMI activities. I am a member of the Chicago chapter of APMI International and a regular attendee at their meetings. I successfully completed the PMT Level I examination. I participated in the 2007 PM Metallographic Competition and will present a paper on iron–silicon soft magnetic materials at the PM2008 World Congress in Washington, D.C.
Materials Engineer SSI Technologies, Inc. 3330 Palmer Drive Janesville, Wisconsin 53546 Phone: 608-373-2814 Fax: 608-755-1004 E-mail:
[email protected]
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SPOTLIGHT ON ...RAJENDRA M. KELKAR, PMT
What major changes/trend(s) in the PM industry have you seen? To enter new and viable businesses, the PM industry is moving towards full-density materials with a variety of complex shapes. This mandates the development of new materials and new processing techniques involving compaction, sintering, and post-sintering. Why did you choose to pursue PMT certification? As soon as I joined SSI Technologies, I thought it would be helpful to have the PMT certification in order to increase my knowledge base, and to benchmark my career.
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How have you benefited from PMT certification in your career? Certification has helped me to increase my understanding of PM. It was an important milestone in my career. What are your current interests, hobbies, and activities outside of work? Outside of work, my life centers around family (wife Madhavi and daughter Vishva) and parents. I like to travel, relax with friends, and pursue socio-cultural activities. ijpm
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CONSULTANTS’ CORNER
BRIAN H. PITTENGER* Q
We have found that powder agglomerates can form in the feedshoe due to shaking and vibration. This prevents complete and uniform filling of the die cavity. It also can create problems with the part. What can be done to eliminate this effect? Before getting into detailed options, there are several approaches that can be considered from a high-level perspective. These approaches include a change in the material, a change in the process, a change to the equipment, or a combination of one or more of these entities. Which approach is the best depends on many factors. It is also important to stress that there is no one solution that fits all cases, and that the details of a given application will dictate which approach is the most practical and reliable. Often, if not in almost all cases, the material being pressed cannot be modified without a substantial effort to confirm that there will be no adverse effects on the final part quality and performance. However, if this is a possibility, or there is some room to make changes, there may be options with respect to changes in the particle-size distribution of the powder blend components (fewer fines), a change (reduction) in the percentage of the component that is likely causing agglomerate formation (requires some investigation), or a reduction in the moisture content in the blend. A high moisture level may be the result of too high a humidity level in powder storage, or pressing rooms, in which case humidity control may be necessary (although the implications of electrostatics must then be considered). A change in the process may be an acceptable alternative that has the benefit of maintaining the current material. In some cases, shaking and vibration have been found to contribute little to the quality of fill and could be shut off altogether, but this is not the usual situation. We have also
A
found in some cases that ensuring more routine replacement of the wear plate on the feed shoe can eliminate agglomerates and minimize powder waste, since there is no longer the same leakage of powder. Similarly, changes to the edge profiling on the wear plate or the opening size and shape can eliminate the problem. Beyond this, one may want to consider a change to an alternative shoe concept. There are other options rather than a translating reservoir of powder sitting on the platens. You may want to consider one of many approaches of dosing or feeding the powder into the die cavity. There are air-assisted feed systems as well as a number of rotary methods to feed the powder directly into the die cavity. Whether one of these approaches will work in the given application requires that some trials be performed. These feed methods may have additional benefits beyond the elimination of agglomerates. You may find the fill more consistent and with little to no powder loss. It may also eliminate a source of mechanical problems and wear. Finally, success has been achieved in some pressing applications by pre-dosing with an exact weight out and then loading it entirely into the cavity. This may need careful consideration with respect to current tooling.
Q A
I am working to enhance the consistency of flow of fine iron particles and seek information on how to predict flow behavior and flow mechanisms. I have found that using fluid mechanics principles does not work. You are correct. Computational fluid dynamics models and fluid mechanics engineering equa-
*Senior Consultant, Jenike & Johanson, Inc., 400 Business Park Drive, Tyngsboro, Massachusetts 01879; Phone: 978-6493300; Fax: 978-649-3399; E-mail:
[email protected]
Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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CONSULTANTS’ CORNER
tions are poor tools for modeling most granular mechanics problems (excepting those in which the powder is fluidized). Instead, one needs to focus on the tools available that enable one to better understand the flow behavior of a granular bed. Andrew Jenike’s technical Bulletin 123, available through he University of Utah’s Experimental Station, is a reliable source for understanding how powders flow through bins, hoppers, and, to some extent, feeders. Another recent source of information on the flow of powders and bulk solids is a book by Dietmar Schulze published by Springer and titled “Powders and Bulk Solids—Behavior, Characterization, Storage and Flow.” Ultimately, it is important to consider what must be known about the flow behavior and flow mechanisms of particles. For example, is it a specific unit operation that is to be better understood, or a transfer from one process step to the next? Are we interested in dilation of a bed at incipient flow or packing behavior as the consolidation forces increase? In each case, a number of basic flow properties of the powder are at work. These include permeability, compressibility, internal friction, wall friction, and cohesive strength. All of these properties are functions of the applied consolidation pressure. They will also vary, based on particle-size distribution, surface morphology, storage history, and environmental conditions. Therefore, when examining flow behavior, one of the first steps, especially when the goal is to engineer a specific flow behavior into processing equipment, is to understand these flow properties under representative conditions for the powder being handled. I hope these comments provide some direction in powder flowability. For further information, as it pertains to typical powder metallurgy processing, you may want to consider one of the MPIF courses such as their Basic PM Short Course or the PM Parts Compacting Tooling Seminar. In addition, there are a number of publications targeted at PM packing and processing available through MPIF (www.mpif.org). You may find something specific to your needs there. Finally, if you are searching for the latest research and cutting-edge work by leaders in the field, I recommend that you attend some of the PM Modeling sessions at the PM2008 World Congress in Washington, D.C., June 8–12, 2008. I look forward to seeing you there! ijpm Readers are invited to send in questions for future issues. Submit your questions to: Consultants’ Corner, APMI International, 105 College Road East, Princeton, NJ 08540-6692; Fax (609) 987-8523; E-mail:
[email protected]
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International: powder injection molding. If you wish to produce complex ceramic and metal products using the PIM process, then come to the leading international specialists in this field: ARBURG. For you, we have the appropriate ALLROUNDER machine technology and the required know-how from our PIM laboratory. With our expertise, you will be able to manufacture efficiently and to the highest quality, prepare material, injection-mold components, debind and sinter - finished! You want to find out more about PIM processing? Simply talk to us!
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[email protected]
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gress rld Con PM Wo , 2008 2 June 8-1 30 4 Booth # n D.C. gto Washin
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PM MACHINABILITY
MACHINING OF PM MATERIALS: A SECONDARY SHAPING OPERATION OF PRIMARY CONCERN Carl Blais*
What a paradox! A focus issue on the machining of components produced by a shaping process that prides itself on being able to minimize, and even eliminate, the need for machining. Although powder metallurgy (PM) qualifies as a “near-net-shape process”, approximately 50% of all components produced by PM are submitted to one of many machining operations. Indeed, machining is generally required to generate geometrical features that cannot be incorporated in compaction. Moreover, PM components in high-performance applications often mandate tight tolerances, hence the necessity for machining. Although the machining of metals is not a new topic, and there are numerous handbooks on the subject, little has been published on the optimized cutting conditions to machine PM materials. This is not surprising since the vast majority of machining charts available have been published by cutting-tool manufacturers. Why help PM, a competitor? PM parts manufacturers attempted to apply the knowledge available on the machining of wrought products to the machining of PM materials. The success of this approach was marginal, at best, due to the significant differences in microstructure between wrought and PM materials of similar chemistry. Nevertheless, powder manufacturers began to develop their own machining guidelines. Developments in the field started with the production of new powders with improved machinability and were quickly followed by the optimization of machining parameters. Over the last decade, there have been significant advances in understanding the phenomena that govern tool wear while machining PM materials. This knowledge will certainly help increase the competitiveness of the PM industry and open up new high-performance applications for PM. This focus issue on PM machining embraces advances in machinability characterization, admixed and prealloyed additives used for improved machinability of unalloyed and sinter-hardenable ferrous components, green machining, and mechanisms. In the first article, Denis Christopherson provides a practical and insightful analysis of steps to follow to characterize and optimize the machinability of PM components. His quantitative approach is based *Professor, Department of Mining, Metallurgical and Materials Engineering, Université Laval, Québec City, Canada; E-mail:
[email protected]
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MACHINING OF PM MATERIALS: A SECONDARY SHAPING OPERATION OF PRIMARY CONCERN
on a “Define–Test–Evaluate” process in which machinability criteria are first determined while taking into account the limitations of a given machining process. The second phase is aimed at emulating the “real life” machining process to determine, in the third phase, the best solution for the system studied (material, cutting tool, and machining operation). The analysis is supported by a case study that emphasizes the efficiency of his approach. The second contribution, from Bruce Lindsley, gives a detailed summary of the mechanics of chip formation and tool wear. The paper also discusses the benefits and disadvantages of different additives on the machinability of unalloyed and sinter-hardenable PM steels. This review highlights the complex interactions that exist between additives for improved machinability and cutting-tool chemistry. The results presented constitute a solid base on which to plan a machining project involving PM materials. The paper by Patrick Boilard et al. continues the overall theme. It presents the different quantitative methods that can be used to characterize machinability in turning and drilling. It also characterizes in detail the effect of the concentration of prealloyed manganese sulfide (MnS) particles on the machinability of FC-0208 mixes. Their results illustrate the differences that exist in terms of machining behavior between two different machining processes. Moreover, their work shows that 1.0 w/o additive compared with a 0.35 w/o additive has a significant effect on machinability improvement, with minimal effect on the degradation of mechanical properties. Étienne Robert-Perron et al. review green machining, used primarily to circumvent the difficulties of machining heat-treated alloy PM steels. They highlight the need to use binder/lubricants
to not only increase green density but also to lubricate the chip–tool interface during green machining. The results are aimed at optimizing green machinability in terms of tool selection, lubrication methods, green strength, and tool wear, which is not negligible. A comparison of static mechanical properties of PM components green machined, and machined after sintering, is also presented. ˘ alak et al. quantify the effect of To conclude, S porosity and carbon level on the machinability in face turning of sintered plain carbon and alloy steels. Deep tur ning marks develop on the machined surface, accompanied by a built-up edge and are the mechanisms responsible for impairment of machinability by porosity. Resistance to plastic deformation increases with increasing carbon content and enhances machinability. Although it is never easy (nor wise) to make predictions, it can be expected that the machinability of PM components will continue to be a priority in ter ms of R&D. While significant improvements have been achieved, work still remains to be done especially for applications in which fatigue resistance is important. Indeed, the quality of machinability of PM materials still relies heavily on additions of second-phase particles. Therefore, work needs to be carried out to optimize machinability and dynamic properties. In order to reach such a goal, several parameters will have to be taken into account: sintered density, volume fraction, as well as chemistry and morphology of machinability-improving additives, tool selection, cutting parameters, etc., and all those at the lowest cost possible. Finally, one can only hope that a publication regrouping all this knowledge will be published to offer those involved with machining PM materials a single source of reliable information. ijpm
ERRATA: The following listings were printed incorrectly in the Web Site Directory (IJPM, vol. 44, no. 1, p. 65 and p. 76). Equipment Manufacturers Section ABTEX Corporation www.abtex.com Abtex Corporation is a single-source manufacturer for abrasive deburring brushes and the machines needed to apply them. Brush line additions include 6” and 8” O.D. composite hub radial wheels in 1/2”, 3/4” and 1” widths. Supplementing its end deburring machines, the deburring systems group now offers both wet and dry process planetary head flow-through systems for flat parts, and rotary indexing systems for addressing more complex part geometries PM Products or Parts Producers Section Mi-Tech Metals, Inc. www.mi-techmetals.com Mi-Tech Metals, Inc., located in Indianapolis, Indiana, produces tungsten heavy alloy and copper and silver tungsten composite materials. Additional materials include tungsten carbide and pure molybdenum and tungsten. Mi-Tech maintains inventory to meet immediate requirements and our extensive machine shop manufactures parts to print.
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PM MACHINABILITY
CHARACTERIZATION OF PM MACHINABILITY: PRACTICAL APPROACH AND ANALYSIS Denis Christopherson, Jr., PMT*
A practical approach is outlined to address powder metallurgy (PM) machinability issues and challenges, in particular cutting-tool wear, process characterization, and measurement techniques. Coupled with an actual production case study of tool life in a single-point turning operation, the approach serves as an example of PM machining analysis to improve performance.
INTRODUCTION Understanding and troubleshooting machining issues can be challenging. The added dimension of PM parameters and variables can concern seasoned problem solvers. This article offers a practical approach and steps to better understand the problem and work towards a solution. Discussion of cutting-tool wear, process characterization, and measurement techniques, coupled with a case study, provides an example of PM machining analysis that can be undertaken at any facility to improve machining performance. Figure 1 shows a limited fishbone diagram of possible PM machining variables. Clearly the magnitude of the variables cannot be studied in exhaustive fashion during one’s lifetime, and certainly not within a typical manufacturing deadline. Therefore, a methodical approach and prioritized effort must be practiced to obtain optimum output in a reasonable timeframe.
Figure 1. Fishbone diagram showing some of the parameters involved in PM machining characterization and troubleshooting *Manager, Research and Development, Federal-Mogul Sintered Products, 401 Industrial Avenue, Waupun, Wisconsin 53963, USA; E-mail:
[email protected]
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CHARACTERIZATION OF PM MACHINABILITY: PRACTICAL APPROACH AND ANALYSIS
Figure 2. Process-flow steps for PM machining characterization
Approach to Machining Characterization The process flow necessary for a successful approach to characterization of PM machinability is shown in Figure 2. 1 The first step, often bypassed but critical, is simply to Define the machining criteria. Success at one facility may be considered failure at another, and the limitations of equipment, parameters, and measurement capabilities must be defined and understood prior to the subsequent steps of Testing and Evaluation. Define: Before testing, the goals and expectations of the stakeholder (customer), as well as the constraints and limitations of the production line, must be understood. It serves little purpose to find optimum parameters, or develop a process, if the equipment or facility cannot adopt the solution. Testing should focus on the reasonable capabilities of the facility in question. Success criteria and the measurement method must also be defined. If the customer uses surface finish as the determining criterion for tool life, the test exercise must include surface finish among other potential measurement criteria. Definition of success should be discussed with the customer prior to testing. A mutual understanding of the goal must be part of the test plan and must be kept in view throughout the work. It is an uncomfortable situation to present the customer with a solution that cannot be achieved, cannot be measured, and cannot be implemented. Testing: Once the goals and criteria are defined, the testing effort can begin. Often, the test work is done off-line to limit manufacturing disruption. In this case, proper simulation must be achieved to
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successfully correlate with production performance. The closer the conditions are to the production process, the more likely the test results will translate into production practice. If the production process in question is a turning operation, drilling tests offer little applicability, and correlation becomes speculative. For the most meaningful testing effort, a turning simulation should be performed, using production-similar workholding, toolholding, and test components. The test simulation will seldom match production practice exactly, but the operator should strive to make the test as close to production practice as is reasonably possible. Once the testing simulation is designed, benchmarking the current production state to develop the measurement standard is necessary. The test standard from simulation should match the general performance in current production. While the tool life value or other results may not match exactly, the tool-wear characteristics and workpiece measurables should approach production experience. For example, the test simulation may provide 750 cuts per tool, while production experience is 500. The primary requirement is that the tool-wear mechanism exhibited by the simulation mimic production, and that workpiece measureables (such as surface finish and diameter) fall within a reasonable variation of production values. If the cutting tool fails in normal production experience by cratering wear, the simulation should also deliver cratering wear on the tool. If the simulation results in a built-up-edge (BUE) condition, correlation fails, and alteration of the test setup is necessary. Evaluate: Evaluation of the results from the test simulation constitutes the final step toward machining Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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CHARACTERIZATION OF PM MACHINABILITY: PRACTICAL APPROACH AND ANALYSIS
Figure 3. Metallurgically prepared cross section of workpiece (FC-0205) placed against image of cutting tool (TiN coated carbide). Note mirror image of workpiece and tool, including features of tool wear reflected on cut surface of workpiece
characterization. Proper evaluation is key to understanding reasons for production problems and the path to improvement. If the work definition and test simulation are done correctly, evaluation of the output is the simplest step of the characterization method. Evaluation should rely on more than one parameter, but the most important single component of study is the cutting tool. The cutting tool is a mirror image of the workpiece surface (Figure 3) and is a potent source of information about the machining operation and the mechanisms in action during the process. The observations and measurements of the cutting tool allow the investigator to “Read the Process,1” providing clues to per for mance limitations and paths toward improvement. Figures 4(a) through 4(d) show examples of cutting-tool wear characteristics in the machining of FC-0208 (modified) from which tool-wear mechanisms can be identified and directions toward improvement determined.2 It is important to note that the cutting-tool investigation should include both the tool flank and the tool face to obtain a full picture of the wear and chip flow interaction Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
with the tool. Figure 4(a) shows classic cratering wear, indicated by the appearance of a crater or “scooping out” of tool material when viewing the chip interaction region of the tool face. The crater is formed through a chemical reaction between the cutting tool and workpiece, initiated by excessive heat and chemical dissolution of the tool surface into the chip. As the crater forms behind the leading edge of the tool, the edge becomes weaker and eventually chips and fractures, further accelerating tool wear. Common actions to reduce or eliminate cratering wear include reducing cutting speed, improving lubrication/cooling, and changing the cutting-tool grade/coating to more stable options. Figure 4(b) shows fracture of the cutting tool. This example is relatively extreme, but is a tool removed from a production line in which this type of failure was common and was tolerated. Fracture failures are usually less dramatic, including notching of the cutting edge or chipping of the face and edge. Tool fracture can also be a secondary failure mode following a primary wear condition such as described in the discussion of cratering. Fracture occurs when the cutting parameters are too aggressive (excessive speed or feed), the tool edge is too weak, or if the machining system (toolholding, workholding, machine rigidity, etc.) is not robust. Any single factor or combination of these factors can result in “chatter,” vibration, or catastrophic fracture during the machining operation. Fracture is the most deleterious of tool-wear mechanisms, causing premature failure and inconsistent performance. Often, the inconsistency and unpredictability of the fracture event is more frustrating and disruptive to manufacturing than the failure itself. In most cases, fracture should be viewed as an unacceptable condition that requires immediate attention. Figure 4(c) shows the BUE condition. This occurs when machining soft or “sticky” materials such as aluminum, magnesium, or ferritic irons and steels. BUE is localized welding or adhering of the workpiece material to the cutting tool. The BUE progressively increases to a critical size and eventually breaks off, often removing some of the tool material in the process. As noted previously, BUE can be the primary failure mode resulting in secondary chipping and fracture of the tool. BUE also causes higher cutting pressures as the
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Cratering, fracture and BUE reflect premature tool failure and less than optimal performance. Steps can be taken to improve the machining process if these three conditions exist; however, the big picture must be considered. Often, nonoptimum conditions are accepted in the interest of higher production throughput or other overriding factors.
Figure 4. Examples of cutting tool characteristics. (a) cratering wear: Al2O3 coated carbide, (b) catastrophic fracture: TiN coated carbide, (c) BUE condition: TiN coated carbide, and (d) abrasive wear: TiN coated carbide
adhered material acts as a dull cutting edge. In turn, higher pressures often cause higher temperatures and accelerated tool breakdown. BUE can be controlled or eliminated by improving lubrication, increasing cutting speed and/or feed rate, or changing the tool grade, coating, and/or edge preparation. Figure 4(d) shows abrasive wear of the cutting tool. Abrasive wear is the common and natural mechanism of machining as one material slides across another. Abrasive wear is the desired wear mechanism and indicates that the machining process is within the favorable region or “Safe Zone” of parameters. The important argument offered from this review of tool-wear mechanisms is that abrasive wear is generally the desired tool-wear condition.
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Machining in the “Safe Zone” and “Sweet Spot” Reading the process to determine the failure mode provides necessary clues to machining improvement. Further understanding of the machining process is provided through an analogy offered in Figure 5 in which the machining process is compared with a baseball bat. Note the location of the machining “Safe Zone” and “Sweet Spot” on the bat.1 When hitting a baseball, there are preferred contact areas for the ball on the bat and other contact areas that should be avoided. Hitting the ball in the wrong area of the bat will likely result in failure to meet the goal of a hit or home run. A similar situation holds true with machining parameters. On the “machining baseball bat,” there is a region of parameters (cutting tool, speed, feed, etc.) resulting in the desired non-catastrophic, consistent performance. Approaching the outer edges of the safe zone reduces the robustness of the process. A variation of the input parameters may push the machining process outside the “Safe Zone,” resulting in premature tool failure. A process that delivers inconsistent performance and/or switches tool-failure modes is likely to occur near or outside the “Safe Zone.”
Figure 5. The machining process baseball bat analogy. The machining process performs best in the “Sweet Spot” of the “Safe Zone”
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CHARACTERIZATION OF PM MACHINABILITY: PRACTICAL APPROACH AND ANALYSIS
The most robust process is found in the “Sweet Spot” where significant variations are absorbed and little change in machining performance occurs over time. From the discussion of toolwear characterization, abrasive wear indicates “Safe Zone” operation and the best tool life resides within this zone corresponds to the “Sweet Spot.” Fracture wear is considered outside of the “Safe Zone.” To some extent, cratering wear and BUE may be considered within the “Safe Zone,” but near the edges. Dramatic cratering and BUE would be outside the “Safe Zone.” To bring the analogy full circle, a baseball hit in the “Sweet Spot” travels the longest distance. A ball hit outside the “Safe Zone” can result in sore hands! Case Study: Production Turning An actual production case study is outlined and offers a simple example of the effectiveness of “Reading the Process” and following the path to a solution. Here the customer reported poor tool life in a single-point turning operation on a multimaterial workpiece application, namely PM steel and wrought aluminum cut with the same tool in the same operation. Analysis of spent production tools showed a BUE condition limiting tool life. After reviewing customer criteria and parameter constraints, a simulation was designed to copy the production experience. Production PM steel components and aluminum were fixtured together to simulate the bimetal cutting seen by the tool in the production process. Test simulation in the laboratory on a CNC turning center using production parameters (tool grade, cutting speed, feed rate, depth of cut, etc.) showed approximately 100-piece tool life and comparable BUE to production experience, Figure 6(a). This test confirmed correlation of the simulation to the production line and provided confidence that solutions identified by the simulation testing would be relevant to production. Knowing that BUE is an undesirable condition (outside the “Sweet Spot”), the cutting speed was increased to combat the BUE tool-wear mechanism. The next simulation showed a 200% improvement in tool life, a significant improvement compared with the baseline. Tool analysis from this test showed that increasing the cutting speed changed the failure mode from BUE to cratering wear, Figure 6(b). While a significant improvement was realized, cratering wear indiVolume 44, Issue 2, 2008 International Journal of Powder Metallurgy
Figure 6. Test simulation cutting FC-0205 (modified) with a TiN coated carbide tool: (a) baseline showing BUE mechanism at 100-piece tool life, and (b) crater wear at 300-piece tool life, resulting from increased cutting speed
cates that further improvement is possible by seeking the “Sweet Spot.” Because cratering wear reflects a chemical reaction between the workpiece and cutting tool, possibly due to excessive heat, options to address this issue were considered. BUE at slower speeds and cratering at higher speeds suggested that a tool change may offer a bridge between the two failure mechanisms. A review of the literature showed that a TiB2 coating is used on carbide tools to machine “sticky” workpiece materials such as magnesium and aluminum.3 This coating has a higher hardness and lower coefficient of friction than the TiN coating used in the production process. It further improved the test simulation tool life to 540 passes, a 440% increase compared with the baseline, Figure 7. The customer implemented the new parameters (tool grade and cutting speed) and experienced better than three times tool life on the production line. Machining Characterization—Child’s Play Basic machining characterization does not require advanced and expensive equipment and technology. Following the approach and practice cited here can provide simple direction toward improvement of the machining process in a timely fashion. Optimization of complex machining processes may require a more in-depth analysis and matrix testing, but relief can be achieved relatively easily, as illustrated in the case study cited. Figure 8 shows images of typical cratering wear
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Figure 7. Increased resistance to BUE and cratering wear in machining FC-0205 (modified) by changing tool grade from TiN coated carbide (Figure 6) to TiB2 coated carbide resulting in 540 piece tool life (440% improvement. (a) face view of tool, (b) flank view of tool Figure 9. Toy microscope with image analysis software used to record tool-wear images shown in Figure 8
Figure 8. Additional examples of cratering wear and BUE. Bimetal cut of FC-0205 (modified) and wrought aluminum: (a) TiN coated carbide tool, (b) Al2O3 coated carbide tool
and BUE. While these images are not perfect in clarity and resolution, they are adequate for the characterization of tool wear. The reader will forgive the image quality when told that the pictures were taken using a child’s microscope and image analysis software package purchased at a local toy store for approximately $50, Figure 10.4 CONCLUSION Although troubleshooting machining problems can be challenging, a simple approach can be practiced to identify the causes and direct a path to a solution. Defining the tool failure mechanism is critical to “Reading the Process” and to developing a plan to improve the machining process. The approach to Define, Test, and Evaluate
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must show a relationship to production floor experience. Defining production-relevant goals and criteria from which to develop a Test simulation that correlates to production experience through proper Evaluation provides confidence in successful implementation. The “Machining Baseball Bat” analogy can aid in an understanding of the machining parameter relationship to performance. A parameter set within the “Safe Zone” offers desirable tool performance, and “Sweet Spot” conditions provide the most reliable and robust machining process. Away from the “Sweet Spot” and outside the “Safe Zone” typically results in catastrophic and unpredictable machining performance. Practical machining characterization can be performed with limited resources, as shown by optical microscopy performed with a $50 plastic toy. A basic understanding of failure mechanisms and their causes provides important clues to solving many PM machining problems. REFERENCES 1. D. Christopherson, “A Practical Approach to Analyzing PM Machinability”, Special Interest Program II; Machining of PM Components in the 21st Century: More than Just Chip Making, PM2TE2004, Chicago, IL, June 13–17, 2004. 2. D. Christopherson, “An Analytical Approach to Evaluating Machinability”, Powder Metallurgy Machinability Seminar, Metal Powder Industries Federation, Indianapolis, IN, Oct 13–14, 1999. 3. “Kennametal Grade System for Cutting Materials—Lathe Tooling Catalog”, Kennametal, Inc., 2006, p. A28. 4. http://www.toygroove.com/educational/qx5 ijpm
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PM MACHINABILITY
MACHINING OF PM STEELS: EFFECT OF ADDITIVES AND SINTER HARDENING Bruce Lindsley*
INTRODUCTION Ferrous PM is generally considered a net- or near-net-shape process. By the judicious use of compaction tooling and core rods, the amount of starting material that ends up in the final part is much higher for PM than that in wrought-steel counterparts. This high materials utilization is, in part, responsible for the growth of PM components. Nevertheless, many pressed-and-sintered parts are machined prior to final assembly, as certain features must be introduced by secondary machining operations.1,2 It is estimated that over 50% of all PM components require machining.3 These machining operations include drilling (transverse holes) and hard turning (undercuts). PM parts are generally considered more difficult to machine than their wrought counterparts. The machinability of PM steels differs from that of wrought steel due to the presence of porosity and, often, a heterogeneous microstructure.4 The porosity in PM steels makes lubrication more difficult and reduces heat transfer from the cutting surface. Additionally, porosity produces an interrupted cut with the tool, resulting in micro impacts and a fatigue condition at the edge of the tool. To achieve similar strengths as those of wrought products, PM alloys typically contain higher levels of carbon. As the apparent hardness of a PM material approaches that of a wrought product, the microindentation hardness must be significantly higher to offset the effect of porosity. The higher microindentation hardness, in combination with the porosity, is primarily responsible for the different machining response of PM steels compared with that of wrought steels. Improvement of ferrous PM component machinability will enhance the adoption of PM in applications. In order to accomplish this goal, it is necessary to understand the fundamentals of, and unique characteristics exhibited by, PM machining.
The machining of ferrous powder metallurgy (PM) alloys differs considerably from that of wrought materials. The role of porosity and heterogeneous microstructures complicates the machining process; in addition, the presence of martensite in the microstructure of highly alloyed and/or sinter-hardened PM components increases tool wear. One advantage of PM is that machinability additives can be admixed into the powder and therefore into the final part. The choice of additive will depend upon the cutting conditions and properties required of the final part. Manganese sulfide is the most widely used additive for improving machinability. Alternative additives can be chosen if corrosion resistance is an issue. Sinter-hardened parts are particularly difficult to machine as the martensitic microstructure increases the apparent hardness of the part, in which case advanced tooling materials such as cubic boron nitride (cBN) are recommended. Proprietary additives, in combination with these tool materials, can increase tool life significantly. This paper discusses the improvement in machinability with various additives and the effect of sinter hardening on the machinability of PM steels.
Fundamentals Machining of metals is accomplished by the removal of a thin layer from the surface of the workpiece. The one-time removal of a large volume of material is impractical as the stresses would be too high, too much heat would be generated, and both the workpiece and the cut*Manager of Product Development, Hoeganaes Corporation, 1001Taylors Lane, Cinnaminson, New Jersey 08077-2017, USA; E-mail:
[email protected]
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MACHINING OF PM STEELS: EFFECT OF ADDITIVES AND SINTER HARDENING
ting tool would be damaged. The removal of a thin layer utilizing a wedge-shaped tool results in the formation of chips and occurs in most machining operations, including drilling, turning, boring, and milling. Turning and boring are similar in that a fixed tool remains in contact with a turning workpiece during the machining operation. Cutting speed, feed, and depth of cut are critical parameters in these procedures, and are illustrated in Figure 1 for turning. Milling is different in that it involves an interrupted or intermittent cut as the cutting surfaces are rotated in and out of the workpiece. Drilling has unique attributes, the most important of which is that the cutting speed changes along the radius of the drill bit. The tool geometry and its interaction with the workpiece are different for each method, resulting in different cutting conditions and machining responses. In addition, tool materials (high-speed steel, tool steel, carbide, cubic boron nitride, cermet) often vary between methods.
Figure 1. Turning operation—schematic
(a)
Without going into detail for each operation, the important point is that machinability testing using one method may not be a good predictor of machinability in another method. Drill tests, which are often used for rapid evaluations, will not reliably predict machinability during turning, for example. Further, changes to the tools and cutting conditions within an operation may change the ranking of materials. Machinability testing should be performed under conditions closest to those that exist in the actual application. During machining, the layer removed is deformed over a large angle. The energy required for the plastic deformation of this volume of metal accounts for the majority of the total energy of machining. 5 The metal in the chip is heavily sheared and the original microstructure is heavily deformed;6 Figure 2 is an optical micrograph (OM) illustrating shearing in a machining chip from a PM steel. This heavy deformation during the formation of the chip increases the temperature by way of adiabatic heating. The resulting high temperature can be transferred to the tool, where temperatures measured by embedded thermocouples in turning tools have reached over 1,000°C.5,7 More heat is generated when higher-strength materials are machined, as more energy is required to deform the material. These high temperatures are often responsible for limitations in machining speed/throughput and for accelerated tool wear. Chip/Tool Interface Ideally, the deformed layer (chips) moves freely over the tool with little interaction. However, this is frequently not the case. As noted previously, the chips impart heat to the tool due to adiabatic
(b)
Figure 2. Cross section of chip formed during turning of an FC-0208 PM part. Courtesy B. Tougas6 (OM)
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MACHINING OF PM STEELS: EFFECT OF ADDITIVES AND SINTER HARDENING
heating during metal deformation and frictional heating. In addition, the chips adhere to the surface of the tool. This is not simply a function of friction between the tool and the chips; rather, the chips instantaneously weld to the tool and are then broken off by the subsequent chips.5 This type of friction welding requires two conditions: deformation of the metal, and clean surfaces. Both of these conditions are met during machining after the first layers of metal are removed and the surface of the tool is scrubbed clean by the chips. At lower speeds and/or with softer metals, this is manifested as the built-up edge (BUE) condition. Material from the workpiece is deposited onto the tool and is severely work hardened (up to 600 HV in steel) such that it is not easily removed and becomes the cutting surface. 5 The BUE changes the shape of the tool resulting in some
Figure 3. Tool/workpiece interaction—schematic. Locations of BUE and tool wear are shown7
(a)
plowing as the cutting mode. The surface finish is inferior and the surface layer of the workpiece may be heavily deformed. As the speed is increased, feed is reduced, or materials with a higher hardness are cut, less material builds up on the tool, and a point is reached at which little to no material builds up on the tool. This represents the lower speed and feed limit for good machining behavior. A processing window generally exists above this point where material can be removed with limited tool wear and good surface finish. As speed and feed continue to increase, tool wear accelerates, especially in the form of crater wear. Machinability charts have been developed for different materials with which to identify the best speed and feed conditions for good machining response. Two types of tool wear are dominant in machining operations that lead to tool failure or unacceptable surface finish, namely, crater wear and flank wear. The flank is that part of the tool contacting the work piece, whereas the crater is formed on the rake surface by the chips, Figure 3. Flank wear, also known as abrasive wear, is the slow removal of the tool cutting edge. Harder workpieces or alloys containing abrasive constituents (such as oxides) generally increase flank wear. Crater wear occurs as the chips interact with the tool on the rake surface. As cited previously, chips weld to the surface of the tool and are then broken off by subsequent chips. Some of the fractures occur in the tool material, and the near-surface layer of the tool is carried away with the chip. Examples of tool flank wear and crater wear are shown in Figure 4. The trajectory of the chips will
(b)
Figure 4. Representative examples of: (a) flank wear, and (b) crater wear on carbide tooling (OM)
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MACHINING OF PM STEELS: EFFECT OF ADDITIVES AND SINTER HARDENING
change as the tool wears and a crater forms; this may predict when a tool is ready to fail.8 Heat generated in the chips is transferred to the tool and degrades its properties. Regions of the tool that experience the highest temperatures are associated with the locations of greatest wear. In addition to the reduction in mechanical properties with heat, chemical interactions can occur between the tool material and the workpiece at high temperature. Diffusion of alloying elements, in particular nickel, from the workpiece to the tool can degrade the properties of the metallic-binder phase in carbide tools. The machinability of wrought steels can be improved via the incorporation of free-machining components. Free-machining steels typically contain MnS inclusions and may also contain lead, bismuth, selenium, and tellurium. Oxide control in the melt can also improve the machinability of steels by producing calcium–silicon rich oxides. These constituents improve machinability in several ways. They act as crack initiators, which facilitates chip formation. These free-machining additions are also soft and malleable and undergo extensive deformation ahead of the tool–workpiece interface which significantly reduces the shear stresses in the workpiece, both ahead of the tool–workpiece interface and at the chip–tool interface, thereby reducing the cutting forces. Elongated sulfides in PM machining chips can be seen in Figure 2. Finally, at the high temperatures encountered during machining, inclusions can form a coating on the tool, providing both protection and improved lubrication. The propensity of chips to microweld to the surface of the tool is reduced. Machining of PM Materials The machining of PM materials is different from that of their wrought counterparts. The presence of porosity and an often higher microindentation hardness change the machining response. The optimum cutting conditions change, requiring a reevaluation of the machining process. Conditions developed for wrought components are often not suitable for their PM counterparts. Simple replacement of wrought parts with PM parts into a machining operation promotes and enhances the belief that PM parts are harder to machine. Cutting speed, feed, and/or tooling need to change when PM parts are machined. Coolants are widely used for machining of wrought compo-
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nents, whereas many PM parts are machined dry, without coolant. Coolants can be absorbed into the pores and away from the cutting surface, reducing their benefit. In some cases, the use of a coolant will actually reduce tool life.9 The removal of coolant from the pores after machining also creates an additional operation. Given these differences, it is not surprising that PM materials often give machinists trouble when initially machining a PM part. PM, however, has the advantage of permitting the admixing of materials into ferrous alloys. Both powder producers and part makers have taken advantage of this approach by incorporating machinability additives, such as MnS, into the steel. These free-machining additives improve machinability by assisting in chip formation, by lubrication of the tool face, and by a reduction of crater wear 4 in a similar manner to that in wrought steels. These machinability enhancers are generally introduced as fine powders, although some iron powder grades contain sulfides in the prealloyed state. Some of the first machining additives mimicked those used in wrought metallurgy. Metallic additives such as lead, bismuth,10 and tellurium11 were added to PM steels with mixed results. Machinability improved with the addition of these elements, but the additives were found to vaporize during the sintering cycle and significant loss occurred, especially near the surface of the sintered compact, where machinability enhancement is critical. The vaporization of lead has obvious health concerns and was abandoned. Tellurium also vaporized, so it was added in the form of manganese telluride, or combined with copper additions to form Cu2Te precipitates that exhibited a lower susceptibility to tellurium loss.11 These additives were present in relatively low-density parts by today’s standards, so it is possible that either tellurium or bismuth could be revisited in higher-density parts with reduced loss to the sintering atmosphere. Sulfides are the most frequently used freemachining agents in both wrought and PM steels.12 Sulfur, molybdenum disulfide, and manganese sulfide have all been found to improve the machinability of PM steels.1–4,9,12–24,36–38 Figure 5 shows the improvement in drill life with the addition of 0.5 w/o of these additives.4 Drill life is often superior with sulfur and MoS 2 compared with MnS. However, several factors limit the use of sulVolume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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MACHINING OF PM STEELS: EFFECT OF ADDITIVES AND SINTER HARDENING
Figure 5. Effect of sulfur-containing additives on the drillability of FC-02084
Figure 6. Effect of sulfur-containing additives on the dimensional change of FC-02084
fur and MoS2 in PM. For example, both sulfur and MoS2 greatly increase the dimensional change of PM steels. A dimensional change increase of more than 0.6% was found with sulfur additions to FC0208, as seen in Figure 6.4 For this reason, in general, sulfur and MoS2 cannot be added to existing premixes to improve machinability. Mechanical properties may also decrease with the use of sulfur additions. In addition to dimensional change and mechanical property effects, both sulfur and MoS2 are not stable during the
sintering process. In one study, approximately 25% of the sulfur was lost with these additives.12 Interestingly, in the case of MoS2 additions, the mechanical properties increased with sulfur loss as molybdenum alloyed with the base iron. Studies have shown that the sulfur loss is related to the sintering atmosphere.13 As the hydrogen content increases, more sulfur reacts with the hydrogen to form H2S gas and is removed from the sintered compact. This is problematic for several reasons: the removal of sulfur will reduce the machinability of the part and alter the part dimensions; the H2S gas that evolves can have detrimental effects on furnace hardware, namely, reduced belt life and corrosion of the muffles with the evolution of H2S gas; and MoS2 is expensive with the current price of molybdenum at a historically high level. The use of MnS generally eliminates many of the problems associated with MoS2 and sulfur. The addition of MnS at low to moderate levels has a minor ef fect on dimensional change and mechanical properties, Table I. Mechanical properties of the mixes containing MnS were generally within 5% of those of the mix without additives. MnS is also stable during the sintering cycle and several studies have shown good sulfur retention. 12–17 However, it has been suggested that sulfur evolution from MnS causes some damage to the sintering furnace. 18 MnS has been found to be an effective machinability aid for most machining operations, including drilling and turning. Several studies have confirmed the benefits of MnS additions to the machinability of PM steels, 1–4,9,12–17,19–23 an example of which is shown in Figure 7. During the cutting process, MnS deforms along the shear plane, reduces tool contact time, and forms a lubricating layer on the tool analogous to free-machining wrought steels.5 Many part families that use a large fraction of the atomized steel produced in North America, including forged connecting rods, main bearing caps, and transmission carriers, contain admixed MnS
TABLE I. EFFECT OF MNS ADDITIONS ON MECHANICAL PROPERTIES OF FC-0205 Mix
Additive
Sintered Density (g/cm3)
Dimensional Change (%)
TRS (MPa)
Apparent Hardness (HRA)
0.2% Yield Strength (MPa)
Ultimate Tensile Strength (MPa)
Elongation (%)
1 2 3
None 0.35 w/o MnS 0.3 w/o MA
7.05 7.01 7.01
0.58 0.60 0.57
1,136 1,078 1,086
48 46 47
393 364 371
554 502 512
3.5 3.2 3.2
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(a)
(b)
Figure 7. Reduction in flank tool wear during turning with addition of 0.35 w/o MnS to: (a) FC-0208, and (b) FD-0405 21
for improved machinability. For the reasons cited, admixed MnS has been the most successful machinability additive used for several decades. While manganese sulfide has many beneficial attributes, it does have some limitations and potentially negative effects. Aside from possible evolution of sulfur-containing gases in the sintering furnace, MnS is not stable in the presence of water. Manganese sulfide is extremely hygroscopic, and will absorb water from the environment. High-humidity environments can rapidly oxidize the MnS to form manganese oxide or complex oxysulfides. When this occurs, the “pea soup” green MnS will turn brown and the absorbed moisture greatly increases the tendency for rust formation. This will occur in both improperly stored premixed powders and in sintered parts.
(a)
An example of a PM sprocket (FLN2-4405) exposed to ambient conditions for an extended period of time confirms the increased corrosion in the MnS-containing part, Figure 8. Oxidation of the MnS will also deteriorate its machinability enhancing properties.23 If MnS converts to MnO or complex oxysulfides, the inclusion will no longer shear and deform at the cutting temperatures, will not coat the tool with a protective layer, and will generally be more abrasive. Parts containing MnS should be machined dry as rapid oxidation will occur in the presence of a water-based coolant. Also, parts that are not machined immediately after sintering should be stored in a low humidity environment to avoid MnS degradation. Some efforts have been made to reduce MnS degredation. The addition of a small
(b)
Figure 8. Sprockets (FLN2-4005) with: (a) no additive, and (b) 0.35 w/o MnS after exposure to ambient conditions 22
26
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amount (6 w/o) of iron or molybdenum into the MnS has shown improvements in stability and machinability.23,24 One final limitation of MnS is that it tends to be less effective as the amount of alloying increases.13 This will be discussed further in relation to sinter-hardening. Some of the limitations cited can by minimized or avoided by prealloying the sulfur. Resulfurized steels are made by adding sulfur to an iron bath that contains manganese. The manganese reacts with the sulfur to form MnS upon atomization. Control of the process results in small, well-distributed sulfides throughout the powder. This technique has been used by several iron powder producers for several decades.10,25,26 Since the sulfides are contained within the iron powder, much of the concern with MnS oxidation is removed. Excellent mechanical properties and machinability performance have been achieved by this route.10,25–29 Some studies suggest that prealloyed MnS improves the fatigue response of PM parts.27 One drawback is reduced compressibility of the prealloyed MnS powder. A second limitation is contamination with sulfide-free powder grades. In addition to sulfides and metallic additives, several other machining aids have been used in PM steels. Powder producers and parts makers have taken advantage of premixing to introduce free-machining agents that are incompatible with the processing of wrought steels. Unlike sulfide machining agents, these products are proprietary to specific producers and are patent protected. A review of related patents identifies the following machining additives: hexagonal boron nitride,30 calcium fluoride and barium fluoride, 31,32 and complex oxides, including calcium magnesium silicon oxide.33 Hexagonal boron nitride is often used as a lubricant for high-temperature metalworking operations. It has a planar crystal structure, similar to that of graphite, and provides good lubricity. Small amounts of hexagonal boron nitride (hBN) have been shown to improve machinability with little or no effect on mechanical properties.30 Drill life was reported to increase by 2.5 to 37 times with the addition of hBN to iron–carbon and iron–copper–carbon alloys 34 and thrust forces also decreased.35 Causton36 observed mixed results with hBN in turning. 0.1 w/o hBN reduced flank wear more than MnS (admixed or prealloyed), yet the sulfur-containing steels resulted in a larger reduction in cutting forces. MnX is a sulfur-free proprietary additive that Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
has been shown to be effective in certain applications. It was originally marketed to facilitate the machining of diffusion alloyed materials, where a three-fold increase in the number of drilled holes per bit was found compared with a MnS mix.37 It can be used by itself or in combination with MnS. It is often the case that a combination of different machining additives improves the machinability more than each additive alone. In a study comparing the machinability (drilling) of FN-0205, a combination of MnS and MnX produced superior results compared with either additive alone.38 All additive permutations in the study were much better than the no-additive condition. Another proprietary additive, KSX, has been introduced in the last few years.39 It is a complex calcium oxide that is added in the form of fine particles. Machining results on powder-forged samples have shown good wear properties and improved fatigue performance with a lower additive content than a comparable MnS mix. Use of KSX is reported to leave a thin, protective silicon-rich film on the surface of the tool in a manner similar to that with calcium-treated wrought steels.5 A final proprietary additive, MA, has also been recently introduced.21 This additive is targeted for sinter -hardening applications, although good results have been found in traditional alloys, such as FC-0208. Like the previous sulfur-free additives, it has little effect on rusting. The benefits of this additive will be discussed in the context of sinter hardening. Other methods have been used to improve machinability in PM steels. Iron with a thin layer of sulfur on the surface inhibits carbon diffusion from the graphite into the iron. Using this concept, residual graphite after sintering can be obtained from samples made from a premix of sulfurized iron and graphite. The free graphite has been reported to greatly enhance the machinability of PM steels.40 The replacement of a small percentage of iron with prealloyed molybdenum steel in an FC-0208 composition was found to greatly improve drill life.4 The addition of the prealloyed steel modified the chip formation and may have reduced a BUE condition. This had the unique effect of increasing the apparent hardness of the steel while improving machinability at the same time. Post-sinter modifications to affect the pore network can also be made to improve machining response. Copper infiltration and resin impregnation of PM steels are generally used to
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(a)
(b)
Figure 9. Improvement in machinability of FLN2-4405 containing various additives: (a) dry, and (b) after oil dip 22
seal the porosity, but will also have a beneficial effect on machinability. Dipping parts in oil will also improve the machining response, as the oil will act as a lubricant during the machining process, Figure 9. The machinability of stainless steels presents an interesting challenge. While MnS additions may result in the best machining response, the corrosion resistance of stainless steel is degraded significantly with the addition of sulfur-bearing additives. Therefore, non-sulfide additives are generally used. MnX has been reported to improve the machinability of 316 stainless steel while not impairing its corrosion properties.41 The addition of a small amount of copper and/or tin to improve mechanical properties was also found to improve machinability.42 Some examples exist in which MnS has been used in stainless steel, but it is not generally recommended. Machining of Sinter-Hardened PM Steels With the advent of sinter-hardening alloys and furnaces, the machinability of these materials in the as-sintered condition is significantly more challenging. With hardened wrought components, parts are generally machined in a softened condition prior to quenching and tempering, or surface treatment. With sinter-hardened components, the cost of this heat treatment step is eliminated. However, the hard, martensitic microstructure is difficult to machine, and requires advanced tooling that can withstand higher forces and increased wear resistance. Both tool-wear rates and cutting
28
forces increase significantly when machining martensitic microstructures. One method to overcome this issue is to resort to “green” or “brown” machining. Parts can be readily machined (with respect to tool wear and machining forces) prior to sintering and hardening. Green machining offers the lowest cost, but requires high green strength and optimized cutting parameters. Brown machining is performed after a presinter that burns out the lubricant and initiates the sintering process so that the strength of the component is high enough to withstand the cutting forces. Temperatures during the presinter are kept low enough to avoid significant levels of carbon from going into solution. Thus, the matrix is relatively soft and the graphite acts as a machining lubricant. Distortion during sintering often requires green-machined holes to be remachined after sintering. The selection of dimensionally stable alloys is critical if green machining is employed. Some machining must be performed in the sintered state in order to meet final tolerances. This can be difficult as martensite is hard and abrasive. During the sinter -hardening process, increased alloying (with molybdenum, nickel, chromium, manganese, copper, and carbon) and/or an increased cooling rate cause the hightemperature austenite to transform to martensite. A representative martensitic microstructure is shown in Figure 10. Martensite is the hardest metallic constituent in steels (microhardness) due to carbon remaining in solution and to distortion of the crystal structure to accommodate the carVolume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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bon. Martensite hardness is greater than 60 HRC in fully dense steels for carbon contents >0.4 w/o. Higher carbon contents increase the distortion and hardness in martensite and change the morphology from lath martensite (≤0.6 w/o C) to plate or lenticular martensite (≥0.8 w/o C) with a mix of the two in between. This distortion results in a highly stressed state within the hardened component. These stresses should be relieved through a tempering operation before any machining is performed. The reason for this is twofold: untempered martensite is extremely hard and difficult to machine, and the dimensions of the part change after tempering, potentially requiring another machining operation to achieve size control. Tempered martensite remains relatively hard. A 1 h temper at 200°C will only reduce the hardness by 5–10 HRC from the as-quenched or sinterhardened state. This high hardness of martensite results in accelerated tool wear under most cutting conditions and for most tool materials. As seen in Figure 11, increasing amounts of martensite increase the rate of tool wear.8 In addition, as increasing carbon content increases the hardness of martensite, it increases the tool wear of fully sinter-hardened components. The effect of another admixed alloying element, copper, is more complex. The addition of copper to a sinter hardenable steel increases its hardenability, or its ability to form martensite. If the addition of copper results in increased martensite formation, tool wear will increase. However, if the steel is fully martensitic regardless of the copper content, then the addition of copper can reduce tool wear.8 One of the primary reasons for tool wear is the elevated temperature experienced by the tool during machining. When machining hardened materials, the tool temperature can rise to a level that is higher than that developed when conventional materials are cut. The combination of higher cutting stresses and elevated temperature result in rapid tool wear. Tool feeds and speeds often need to be reduced to maintain acceptable tool life.43 Advanced tool materials can be used that withstand these conditions. It has been found that carbide tools coated with TiCN exhibit the best wear resistance with respect to other carbide grades.44 However, cBN tools are superior to carbide tools for sinter-hardened materials.4,9,21 The mechanical properties and temperature tolerance of cBN are responsible for the improved performance. While this material is higher in cost than Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
Figure 10. Representative microstructure of sinter-hardened FLC-4608. (OM)
Figure 11. Tool-flank wear in turning of FLC2-4808 with a carbide tool, as a function of martensite content
carbide, the price has declined over the years to make it more cost competitive. The grade of cBN is important, as found by Andersson, et al.9 for a sinter-hardened chromium steel. Grades with increased toughness are preferable to grades with higher hardness. Lack of toughness is the primary reason why ceramic tools do not perform well under most machining conditions.4,9 These tools often fail catastrophically after a short period of time. An example of the improvement cBN can provide over coated carbide tools is illustrated in Figure 12. After 50 cuts, the carbide tool experienced accelerated wear and was near failure, whereas the cBN tool showed little wear. It is interesting to note that cBN does not
29
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Figure 12. Tool-life improvement in machining sinter-hardened FLC2-4808: ● baseline carbide tool, ◆ cBN tool, ■ carbide tool/FLC2-4808 with 0.3 w/o MA
always provide the lowest tool wear. In non-sinter-hardened FLN4C-4005, a coated carbide tool outperformed cBN.21 Machining additives continue to play an important role in the machining of hardened microstructures. It can be seen in Figure 12 that the proprietary MA additive reduced tool wear of the carbide tool to the levels of the cBN tool. This additive is particularly effective in sinter-hardening applications. MnS can also be used and has been found to be effective in lower-carbon martensites. 9 It should again be noted that sweeping conclusions about a tool material or additive cannot be made. The tool, substrate, and additive type need to be optimized for each machining system. Another approach to improve tool wear is to modify the temperature of the tool so that it does not overheat. In a study by Zurecki et al.,45 liquid nitrogen was used to cool the tooling when a hardened PM material was turned. Cryogenically cooled alumina tools outperformed comparable cBN tools cooled with a traditional coolant. In work to be presented at the PM2008 World Congress, 46 sinter -hardened FLC-4808 was turned using a cBN tool. Tool life increased by more than 100% with the use of cryogenic cooling. An additional tenfold increase in tool life was found with the addition of MA, even at a higher cutting speed. One of the benefits of cooling the cutting tool is that cutting speeds can be increased significantly without overheating the tool, thereby increasing throughput.
30
SUMMARY The machining of PM steel components is a complex problem. Machinability is not a material property, and cannot be measured in an absolute sense. Rather, machinability is a process and is dependent upon the many factors that contribute to how a part cuts. The use of machinability agents facilitates the machining process by reducing cutting forces and reducing tool wear. The best additive depends upon the cutting conditions and the properties required of the final part. MnS is the most popular additive and provides good machinability in many instances. Alternative additives can be chosen if corrosion resistance is an issue. Sinter-hardened parts present a challenge in machining as the martensitic microstructure increases the hardness of the part. Advanced tooling materials such as cBN are recommended. Certain proprietary additives, such as MA, in combination with these tool materials, can increase tool life significanly. The benefits to machinability of the various additives, base irons, and tool materials found in the literature must be weighed carefully as machinability changes under different cutting conditions. A tool material or additive that worked for the conditions used in R&D may not work for a particular industrial setup. Given the inherent flexibility that PM offers, it is best to try many combinations of tooling and additives. While this is time consuming and may not be possible if parts are not machined in-house, it is the best approach to improve machinability. REFERENCES 1. H. Sanderow, J. Spirko and R. Corrente, “The Machinability of P/M Materials as Determined by Drilling Tests”, Advances in Powder Metallurgy & Particulate Materials, compiled by R.A. McKotch and R. Webb, Metal Powder Industries Federation, Princeton, NJ, 1997, part 15, pp. 125–143. 2. D.S. Madan and A. Fitzgibbon, “Shelf Life of MnS Powder and MnS Containing Mixes”, Advances in Powder Metallurgy & Particulate Materials, compiled by M. Phillips and J. Porter, Metal Powder Industries Federation, Princeton, NJ, 1995, part 8, pp. 177–182. 3. S. Berg and O. Mars, “Investigating the Relationship Between Machinability Additives and Machining Parameters”, Advances in Powder Metallurgy & Particulate Materials—2001, compiled by W. B. Eisen and S. Kassam, Metal Powder Industries Federation, Princeton, NJ, 2001, part 6, pp. 50–55. 4. R. Causton and C. Schade, “Machinability: a Material Property or Process Response?”, Advances in Powder Metallurgy & Particulate Materials—2003, compiled by R. Lawcock and M. Wright, Metal Powder Industries Federation, Princeton, NJ, 2003, part 7, pp. 154–169. 5. E.M. Trent, Metal Cutting, 2nd edition, Butterworths, London, UK, 1984. 6. B. Tougas, “Contributions to the Improvement of Machinability of PM Components using Second Phase Particles”, MSc Thesis, Department of Mining, Metallurgical and Materials Engineering,
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Université Laval, Canada, 2008. 7. L.A. Kendall, “Tool Wear and Tool Life”, Metals Handbook Ninth Edition: Volume 16 Machining, ASM International, Metals Park, OH, 1989, pp. 37–48. 8. D. Raiser, W. Misiolek and B. Lindsley, “The Effect of Post Sintering Cooling Rate on Microstructure and Machinability of a PM Sinter Hardened Steel”, Advances in Powder Metallurgy & Particulate Materials—2006, compiled by W. Gasbarre and J. von Arx, Metal Powder Industries Federation, Princeton, NJ, 2006, part 6, pp. 27–39. 9. O. Andersson and S. Berg, “Machining of Chromium Alloyed PM Steels”, Advances in Powder Metallurgy & Particulate Materials— 2005, compiled by C. Ruas and T.A. Tomlin, Metal Powder Industries Federation, Princeton, NJ, 2005, part 6, pp. 1–8. 10. P.J. Andersen and J.S. Hirschhorn, “Use of Additives to Improve the Machinability of Sintered Steels”, Modern Developments in Powder Metallurgy, compiled by H.H. Hausner and P.W. Taubenblat, Metal Powder Industries Federation, Princeton, NJ, 1976, vol. 10, pp. 477–489. 11. P.W. Taubenblat, W.E. Smith and F.A. Bladt, “Iron-CopperTellurium—New Machinable Fe-Base PM Alloy”, ibid. reference no. 10, pp. 467–475. 12. R.J. Causton and T. Cimino, “Machinability of P/M Steels”, ASM Handbook Volume 7: Powder Metal Technologies and Applications, ASM International, Materials Park, OH, 1998, pp. 671–680. 13. B. Hu and S. Berg, “Optimizing the use of Manganese Sulfide in P/M Applications”, Advances in Powder Metallurgy & Particulate Materials—2000, compiled by H. Ferguson and D. Wychell, Metal Powder Industries Federation, Princeton, NJ, 2000, part 5, pp. 191–197. 14. A. Liersch, H. Dannenger and R. Raimund, “The Influence of Admixed Inorganic Additives on Properties and Machinability of Sintered Plain Iron and Steels”, ibid. reference no. 4, part 4, pp. 192–203. 15. H. Danniger, A. Lersch and R. Ratzi, “Influence of Additives in PM Steels on the Quality of Machined Surfaces”, Advances in Powder Metallurgy & Particulate Materials, compiled by J.J. Oakes and J.H. Reinshagen, Metal Powder Industries Federation, Princeton, NJ, 1998, part 2, pp. 474–479. 16. K.S. Chopra, “Manganese Sulfide in Machining Grade Ferrous P/M Alloys”, Modern Developments in Powder Metallurgy, Metal Powder Industries Federation, Princeton, NJ, 1988, vol. 21, pp. 361–379. 17. U. Engstrom, “Machinability of Sintered Steels”, Powder Metallurgy, vol. 26, no. 3, 1983, pp. 137–144. 18. H. Suzuki, M. Yoshida, H. Tanaka and Y. Ikai, “250MSA Resulfurized High Green Strength Steel Powder”, ibid. reference no. 1, part 15, pp. 17–27. 19. Y. Trudel, C. Ciloglu and S. Tremblay, “Selecting Additives to Improve Machinability of Ferrous P/M Parts”, Moder n Developments in Powder Metallurgy, edited by E.N. Aqua and C.I. Whitman, Metal Powder Industries Federation, Princeton, NJ, 1984, vol. 15, pp. 775–784. 20. D.S. Madan, “Effect of Manganese Sulfide (MnS) on Properties of High Performance P/M Alloys and Applications”, Advances in Powder Metallurgy & Particulate Materials, compiled by J.M. Capus and R.M. German, Metal Powder Industries Federation, Princeton, NJ, 1992, part 4, pp. 245–267. 21. B. Lindsley and C. Schade, “Machinability Additives for Improved Hard Turning of PM Steel Alloys”, ibid. reference no. 8, part 6, pp. 16–26. 22. B. Lindsley, S. Patel, S. Shah, G. Falleur and C. Chambers, “The use of Machinability Additives in Ferrous PM Parts to Improve Machinability”, Advances in Powder Metallurgy & Particulate Materials—2007, compiled by J. Engquist and T. Murphy, Metal Powder Industries Federation, Princeton, NJ, 2007, part 6, pp. 14–24. 23. S. Claeys and K. Chopra, “Enhanced Machinability and Oxidation Resistance with MnS Containing Additions to Iron”, ibid. reference no. 1, part 15, pp. 111–123. 24. K.Y. Lee, D.K. Park, H.B. Kim and I.S. Ahn, “Sintered Stability and Machinability of P/M Steel using Modified MnS Additions”, Advances in Powder Metallurgy & Particulate Materials—2002,
Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
25. 26. 27.
28.
29. 30.
31. 32. 33. 34.
35. 36. 37. 38. 39.
40. 41. 42. 43. 44. 45. 46.
compiled by V. Arnhold, C-L. Chu, W.F. Jandeska Jr. and H.I. Sanderow, Metal Powder Industries Federation, Princeton, NJ, 2002, part 12, pp. 72–85. P. Plamondon, G. L’Espérance and C. Blais, “Optimization of Prealloyed MnS Steel Powders for Improved Machinability “, ibid. reference no. 3, part 6, pp. 29–39. N. Akagi, M. Haas, M. Sato and Y. Seki, “Tensile, Machinability and Fatigue Properties of Pre-Alloyed, Free Cutting Steel Powder”, ibid. reference no. 13, part 6, pp. 17–23. J. Campbell-Tremblay, C. Blais, G. L’Espérance and P. Boilard, “Characterization of the Fatigue Per formance of P/M Components Produced with Powders Developed for Improved Machinability”, ibid. reference no. 9, part 10, pp. 150–159. S. St.-Laurent, M. Gagne and F. Chagnon, “Effect of Sulfur Content on Machinability of Forged Specimens made of Water Atomized Steel Powders Pre-Alloyed with MnS”, ibid. reference no. 1, part 15, pp. 95–109. F. Bernier, P. Boilard, J-P. Bailon and G. L’Espérance, “Machinability and Dynamic Properties of Sinter-Hardened Steel Parts”, ibid. reference no. 9, part 10, pp. 201–210. C. Ciloglu, M. Gagne, E. Laraque, J. Poirier, S. Tremblay and Y. Trudel, “Machinable-Grade, Ferrous Powder Blend Containing Boron Nitride and Method Thereof”, U.S. Patent No. 4927461, May 22, 1990. O. Andersson, “Particulate CaF2 and BaF2 agent for Improving the Machinability of Sintered Iron-Based Powder”, U.S. Patent No. 5545247, August 13, 1996. O. Andersson, “Particulate CaF 2 Agent for Improving the Machinability of Sintered Iron-Based Powder”, U.S. Patent No. 5631431, May 20, 1997. T. Kaneko and T. Esumi, “Sintered Material Having Good Machinability and Process for Producing the Same”, U.S. Patent No. 5679909, October 21, 1997. M. Gagne, “Sulfur Free Iron Powder Machinable Grade”, Advances in Powder Metallurgy & Particulate Materials, compiled by T.G. Gasbarre and W.F. Jandeska, Metal Powder Industries Federation, Princeton, NJ, 1989, vol. 1, pp. 365–375. E. Ilia, M. O’Neill, J. Lee, J. Poirier and S. St-Lauent, “Development of a Main Bearing Cap for an Inline 6 Cylinder Engine”, ibid. reference no. 4, part 9, pp. 22–35. R.J. Causton, “Role of Additives in P/M Machining”, ibid. reference no. 24, part 12, pp. 47–65. “Additive Improves Machinability Threefold”, Metal Powder Report, 1993, vol. 48, no. 9, p. 34 R.J. Causton, “Machinability of P/M Steels”, ibid. reference no. 2, part 8, pp. 149–170. H. Suzuki and T. Sawayama, “Comparison of Machining and Mechanical Properties Between Manganese Sulfide and KSX in Powder Forged (P/F) Materials”, ibid. reference no. 9, part 10, pp. 272–280. S. Uenosono, S. Unami and K. Ogura, “A New Improvement in Machinability of P/M Steel due to Retained Graphite Particles”, ibid. reference no. 2, part 8, pp. 171–176. P. K. Samal, B. Hu, I. Hauer and O. Mars, “Optimization of Corrosion Resistance and Machinability of PM 316L Stainless Steel”, ibid. reference no. 22, part 7, pp. 40–50. P.K. Samal, O. Mars and I. Hauer, “Means to Improve Machinability of Sintered Stainless Steel”, ibid. reference no. 9, part 7, pp. 66–77. C. Blais, B. Young and G. L’Espérance, “Hard Turning of Parts from Sinter Hardenable Powders”, ibid. reference no. 24, part 12, pp. 1–15. T. Wada, K. Hiro, J. Fujiwara and S. Hanasaki, “Machinability of Hardened Sintered Steel”, Proc. Seventh International Conference on Progress of Machining Technology, 2004, pp. 74–79. Z. Zurecki, R. Ghosh and J.H. Frey, “Finish-Turning of Hardened Powder Metallurgy Steel using Cryogenic Cooling”, Int. J. of Powder Metall., 2004, vol. 40, no. 1, pp. 19–31. R. Ghosh and B. Lindsley, “Role of Machining Additives and Cryogenic Cooling in Machining of Sinterhardened Materials”, To be published, Advances in Powder Metallurgy & Particulate Materials—2008, Metal Powder Industries Federation, Princeton, NJ, 2008. ijpm
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PM MACHINABILITY
EFFECT OF PREALLOYED MnS CONTENT AND SINTERED DENSITY ON MACHINABILITY AND MECHANICAL PROPERTIES Patrick Boilard*, Gilles L’Espérance, FAPMI** and Carl Blais***
INTRODUCTION The machinability of sintered PM parts depends on many variables ranging from chemical composition to the machining process to the microstructure.1 It is generally recognized that parts produced by PM exhibit inferior machinability compared with wrought parts with equivalent microstructures.2,3 This is attributed primarily to the presence of porosity within the parts which leads to interrupted cutting and lower thermal conductivity, thereby resulting in greater tool wear.4,5 Although the beneficial effect of density on mechanical properties (tensile, impact, and fatigue) is well documented,6 studies investigating the effect of density on machinability have generated a divergence of conclusions. Most studies conclude that there is no significant correlation between machinability and density (6.7 to 7.3 g/cm3 range) in the absence of MnS particles6,7 or with the addition of MnS particles.8,9 In contrast, results obtained on PM stainless steels have shown that an increase in density leads to longer drill life and hence improved machinability.1 However, drilling trials performed on various sintered steels showed an increase in cutting forces in the 5.3 to 7.1 g/cm3 density range, indicating a decrease in machinability with density.10 The effect of MnS particles on the machinability and mechanical properties of sintered PM parts has been extensively studied.4,8,11–14 It is found that prealloyed MnS particles are more beneficial in terms of machinability than admixed MnS particles. With prealloying, MnS is located within the individual powder particles and in the vicinity of sintered necks.15–17 MnS particles located within the individual powder particles can therefore promote crack initiation where it enhances machinability (i.e., in the bulk of the material). Additions of 0.2 to 0.8 w/o MnS are found to significantly increase machinability in drilling.1 No major adverse effect on mechanical proper-
Blends were prepared using base powders with different prealloyed MnS contents (up to 1.0 w/o) and pressed-and-sintered to two densities. Machinability was characterized in drilling and turning and some mechanical properties were evaluated. Results show that the machinability of samples with 1.0 w/o MnS were superior to those of the other samples, although the mechanical properties evaluated were lower for a given density. In drilling, machinability improved with an increasing level of prealloyed MnS particles and with density. The presence of a high level of prealloyed MnS particles was more important than density in the improvement of machinability. In turning, machinability increases with increasing levels of prealloyed MnS particles in the range 0.35 w/o to 1.0 w/o and tool wear is marginally lower for the higher density samples. Machinability in turning is influenced by the strength of the powder metallurgy (PM) material.
**Professor, *PhD Student, Center for Characterization and Microscopy of Materials – (CM)2, École Polytechnique de Montréal, Montréal, Canada; E-mail:
[email protected], ***Professor, Department of Mining, Metallurgical and Materials Engineering, Université Laval, Québec City, Canada
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EFFECT OF PREALLOYED MnS CONTENT AND SINTERED DENSITY ON MACHINABILITY AND MECHANICAL PROPERTIES
ties is found for low MnS contents (0.5 w/o), at densities <7.2 g/cm3.8,12 In a previous study, samples containing 1.0 w/o prealloyed MnS particles exhibited excellent machinability in turning, although a decrease in fatigue properties was reported.18 It is generally accepted that different machining processes lead to different machinability ratings when comparing similar materials. 19 Thus, an alloy which exhibits good machinability in drilling may exhibit poor machinability in turning. In the present study, machinability trials in drilling and turning were preformed on powder blends containing varying levels of prealloyed MnS particles (up to 1.0 w/o) pressed-and-sintered to two different densities. Mechanical properties (transverse rupture strength (TRS), impact toughness and tensile strength) were also characterized and correlated with machinability. MATERIALS AND EXPERIMENTAL PROCEDURES The base powders were produced by a hybrid process. 20 Powder mixes were prepared using base powders containing different amounts of prealloyed MnS particles in the 0.0 to 1.0 w/o range. The chemical composition of all blends was that of the FC-0208 designation.6 The copper content was kept constant at 2.0 w/o, and the graphite content was adjusted to take into account the oxygen content of the different base powders. This was done to obtain similar apparent hardness levels for all the materials at a given density. Comparing the machinability of blends of similar chemical composition with different apparent hardness levels would yield results that are difficult to interpret. In all cases, the carbon level after sintering was in the range 0.70–0.75 w/o. Samples Samples prepared for machinability trials were cylinders 50.8 mm (2 in.) high × 28.6 mm (1.5 in.) OD. The compacts were pressed to two different green densities, 6.7 and 7.0 g/cm3. Samples for mechanical property evaluation were prepared following standard MPIF procedures. 21–23 All samples were pressed to green densities of 6.7 and 7.0 g/cm 3, except for the impact toughness samples which were pressed to a green density of 6.7 g/cm3. All the samples were sintered at a temperature of 1,120°C (2,050°F) for 30 min in an atmosphere of 10 v/o H2 and 90 v/o N2. There was no detectible change in density of the compacts during sintering.
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Machinability Trials For drilling, trials were performed using a standard drilling testbench. 24 A load cell located under the motor allowed for measurement of the thrust force while the holes were drilled. The tools used to characterize the machinability were HSS tools, 3.18 mm (0.125 in.) dia. with a 118° point angle. Drilling trials were performed at 1,500 rpm and a drill feed of 0.06 mm/rev (0.0024 in./rev). The blind holes were 12.7 mm (0.5 in.) deep. A total of 200 holes were drilled for all materials at the two densities studied. All the drilling tests were performed dry. For turning, trials were performed on a Mazak Nexus 100 lathe. The following parameters were used throughout the trials: feed rate: 0.15 mm/rev (0.006 in./rev); cutting speed: 121 m/min (400 ft./min); and depth of cut (DOC): 0.25 mm (0.010 in.). A predetermined number of parts was chosen in the trials, for which the total amount of material removed was ~1,100 cm3. The cutting inserts were used until the maximum width of the flank wear, VBmax, was equal to 0.37 mm (0.015 in.). Flank-wear measurements were taken successively after machining sets of five samples. The inserts used were CNMG432FP (Kennametal cermets with a TiCN coating—grade KT 315). All the turning tests were performed dry. Different criteria are used to characterize the machinability of PM parts. These include average thrust force, tool wear, and the slope of the drillability curve.24,25 For drilling trials, all these criteria were used to characterize the machinability of the samples. For the turning trials, only tool wear was characterized utilizing optical microscopy (OM). For drilling, tool wear was characterized using scanning electron microscope (SEM) and image analysis.16 In order to quantitatively evaluate tool wear, the following steps were followed: (1) a SEM image of the original tool was taken, (2) machinability trials were performed, (3) a second image of the tool was taken after machining, and (4) the two images were superimposed and the images analyzed to obtain a quantitative measurement of tool wear. The presence of built-up edges (BUEs) has to be carefully examined; failing to correctly identify BUEs induces errors in the quantification of tool wear. RESULTS AND DISCUSSION The mechanical properties obtained for the different blends pressed to the two density levels are Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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EFFECT OF PREALLOYED MnS CONTENT AND SINTERED DENSITY ON MACHINABILITY AND MECHANICAL PROPERTIES
summarized in Table I; typical values reported in MPIF Standard 35 are also included. It should be noted that the higher density values reported in MPIF Standard 35 are for a sintered density of 7.2 g/cm3 whereas our samples were pressed to a density of 7.0 g/cm3. As expected, similar apparent hardness values were obtained for all samples at a given density. For TRS at a density of 6.7 g/cm3, only samples containing 1.0 w/o MnS exhibit property values significantly lower than those reported in MPIF Standard 35 (7% lower). For a density of 7.0 g/cm3, the TRS of the 1.0 w/o MnS material is again lower than that reported in MPIF Standard 35 (8% lower), but the TRS values of samples with up to 0.65 w/o MnS are similar to MPIF Standard 35 values, even though the density of these samples is lower (7.0 g/cm3 vs 7.2 g/cm3). Similar observations can be made for ultimate tensile strength (UTS); values for samples containing 1.0 w/o MnS are lower. Thus, UTS values are 17% and 20% lower for samples with 1.0 w/o MnS compared with the higher values at densities of 6.7 g/cm3 and 7.0 g/cm3, respectively (samples with 0.35 w/o MnS). UTS values for samples with MnS contents up to 0.65 w/o are similar and the values are again slightly below typical values cited MPIF Standard 35. The impact toughness values for the 6.7 g/cm3 density samples decrease slightly with an increasing amount of prealloyed MnS and are similar to the values cited in MPIF Standard 35. In summary, there is little or no effect of the amount of MnS on mechanical properties in the 0.0 to 0.65 w/o range, as reported in other studies.8,12 However, samples containing 1.0 w/o MnS exhibit lower mechanical properties, in particular TABLE I. MECHANICAL PROPERTIES EVALUATED Blend
Sintered Apparent Hardness TRS UTS Impact Strength Density (HRB) (MPa) (MPa) J (ft.·lb)
0.0 w/o 0.35 w/o 6.7 g/cm3 0.65 w/o 1.0 w/o MPIF 35 6.7 g/cm3
80.4 ± 1.6 82.0 ± 2.1 82.0 ± 1.7 77.8 ± 2.4 73
833.9 893.9 832.5 799.4 862
346.5 390.1 386.7 322.5 414
9.1 (6.7) 8.8 (6.5) 7.5 (5.5) 7.5 (5.5) 6.8 (5)
0.0 w/o 0.35 w/o 7.0 g/cm3 0.65 w/o 1.0 w/o MPIF 35 7.2 g/cm3
87.2 ± 2.1 88.9 ± 1.4 89.8 ± 2.0 88.4 ± 0.6 84
1,066.5 1,043.3 1,065.1 979.0 1,069
455.2 475.7 463.0 387.7 517
--------9.5 (7)
Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
for TRS and UTS. Machinability results obtained for drilling trials are shown for all four different blends in Figure 1. Values of the criteria used to characterize machinability in these drilling trials are presented in Table II. For all the blends, the average thrust force decreases with an increasing level of prealloyed MnS particles. As shown in Figure 2, this is the case for the two densities studied. The beneficial effect of MnS particles on machinability in drilling is clearly visible as thrust forces for samples containing 1.0 w/o MnS are significantly reduced compared with the thrust forces for samples devoid of MnS. Average thrust forces are reduced by ~40% and 20%, respectively, for the 6.7 g/cm3 and 7.0 g/cm3 density samples. In terms of the effect of density on machinability in drilling, the average thrust forces are lower for higher-density samples with MnS contents in the range 0.0 to 0.65 w/o. It is interesting to note that enhanced machinability of the higher-density samples is observed, even if the samples are harder. Indeed, larger apparent hardness values have often been associated with lower machinability at a given density.16 In the present study, however, the results show that the level of porosity dominates apparent hardness for samples containing 0.0 to 0.65 w/o MnS. As reported by Agapiou,1 this could be attributed to the larger amount of work hardening in the lower density samples, thereby increasing thrust forces and tool wear. For samples containing 1.0 w/o MnS, the average thrust forces in drilling are considerably lower than those for all the other samples, but density has little effect on performance in drilling. This suggests that a high level of prealloyed MnS particles will have a marked beneficial effect on machinability and will outweigh the effect of density. As shown in Figure 3, tool wear is also reduced for blends containing high levels of prealloyed MnS particles. Although the relation is not as clear as it is for the average thrust force, it must be noted that tool wear is more difficult to measure than thrust force. The most significant decrease in tool wear occurs when the amount of MnS is increased from 0.0 to 0.35 w/o. This demonstrates the beneficial effect that a small amount of MnS particles can provide in drilling operations. Furthermore, we find that tool wear for samples containing 1.0 w/o MnS is the lowest amongst the samples evaluated.
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EFFECT OF PREALLOYED MnS CONTENT AND SINTERED DENSITY ON MACHINABILITY AND MECHANICAL PROPERTIES
Figure 1. Results from drilling trials
TABLE II. MACHINABILITY RESULTS FOR DRILLING MnS Content 0.0 w/o 0.35 w/o 0.65 w/o 1.0 w/o 0.0 w/o 0.35 w/o 0.65 w/o 1.0 w/o
Average Thrust Force (N)
Tool Wear (%)
Slope of Drillability Curve
6.7 g/cm3
443.4 385.0 370.7 271.4
14.7 ± 0.6 7.0 ± 0.5 7.8 ± 0.3 6.4 ± 0.1
0.49 0.50 0.38 0.09
7.0 g/cm3
366.9 347.7 328.7 286.3
10.7 ± 1.0 7.3 ± 0.4 7.2 ± 0.7 5.8 ± 0.2
0.29 0.29 0.11 0.24
Sintered Density
In terms of density and tool wear, a large reduction is observed at the higher density for samples containing 0.0 w/o MnS. This effect, however, is not as important in samples with larger amounts of prealloyed MnS particles. The slope of the drillability curve is generally a good indicator of the wear rate of the tool.16,24 Thus, higher slopes imply that the thrust forces
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Figure 2. Average thrust force as a function of prealloyed MnS content
are increasing more rapidly, leading to higher rates of tool wear. As in the case for average thrust forces and tool wear, the slopes of the drillability curves are generally lower with increasing prealloyed MnS content, although the relation is again not as clear as it is for the average thrust force. Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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Figure 4. Tool wear as a function of material removed; density 6.7 g/cm3 Figure 3. Tool wear in drilling as a function of prealloyed MnS content
The results also show that the slopes are generally higher for the 6.7 g/cm3 density samples, consistent with the fact that a higher level of porosity leads to higher wear rates. The exception is for the 1.0 w/o MnS samples for which the curves are similar. In terms of the thrust forces, this is considered to be the result of the dominant effect of MnS compared with density for the higher levels of prealloyed MnS particles (1.0 w/o). In summary, prealloyed MnS particles can enhance machinability in drilling, even for relatively low levels of prealloyed MnS particles. However, samples containing 1.0 w/o MnS show excellent machinability in drilling when compared with other FC-0208 blends containing lower levels of prealloyed MnS particles. This leads to significant reductions in the machinability criteria studied (average thrust force, tool wear and slope of the drillability curve). Moreover, an increase in density generally leads to an improvement in machinability, as indicated by the different machinability criteria studied. Finally, of the machinability criteria evaluated in drilling, the average thrust force is most sensitive to the effects of prealloyed MnS and density, possibly because tool wear and the slope of the drillability curve are more difficult to measure. Machinability results obtained in turning are shown for all four different blends in Figures 4 and 5. Table III gives tool wear, measured at the end of the trials. For the two densities studied, machinability increases with increasing levels of prealloyed MnS particles in the range 0.35 to 1.0 w/o. Moreover, the results show that samples containing 1.0 w/o MnS exhibit superior machinability when comVolume 44, Issue 2, 2008 International Journal of Powder Metallurgy
Figure 5. Tool wear as a function of material removed; density 7.0 g/cm3
TABLE III. TOOL WEAR IN TURNING TRIALS MnS Content
Sintered Density
Tool Wear—VBmax (mm)
0.0 w/o 0.35 w/o 0.65 w/o 1.0 w/o
6.7 g/cm3
0.28 0.37 0.34 0.18
7.0 g/cm3
0.25 0.36 0.31 0.16
0.0 w/o 0.35 w/o 0.65 w/o 1.0 w/o
pared with lower prealloyed MnS contents, as reported in a previous study.18 Indeed, for a similar amount of material removed, tool wear is 37% and 35% lower than that of the second best material for the 6.7 g/cm3 and 7.0 g/cm3 density samples, respectively. For all materials, tool wear is slightly lower at the end of the trials for the 7.0 g/cm3 samples. Although the differences in tool wear are small, this indicates that the higher density samples exhibit superior machinability, in agreement with
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the other blends, and still benefit from excellent machinability, both in drilling and in turning.
Figure 6. Tool wear after turning trials as a function of UTS
the results in drilling. A surprising result, however, is observed in samples devoid of MnS which exhibit lower tool wear than samples containing 0.35 w/o and 0.65 w/o MnS. A possible explanation could be related to the UTS values of the different materials. Since these materials exhibit a low elongation-to-fracture during tensile testing, the UTS can be related to the elastic limit, which is the stress at which permanent deformation occurs. Comparing UTS values and tool wear in turning, we find a good correlation, as shown in Figure 6. Tool wear is lower in materials which exhibit a lower UTS, notably for samples devoid of MnS, which exhibit lower UTS values than samples with 0.35 w/o and 0.65 w/o MnS. It appears that in turning operations, machinability is influenced by the strength of the material. The differences in UTS between the MnS-free material and the other material containing various amounts of MnS (0.35 w/o to 1.0 w/o) probably reflects differences in processing. Therefore, a direct comparison of the performance in turning of the MnS-free material with the MnS-containing materials cannot be made. Furthermore, machinability in turning does improve for MnS-containing materials in the range 0.35 to 1.0 w/o. A final observation is made for samples containing 1.0 w/o MnS. These samples exhibit lower mechanical properties than the other samples containing smaller amounts of prealloyed MnS particles at a given density. However, as shown, it is possible for the material containing 1.0 w/o MnS at a density of 7.0 g/cm3 to achieve mechanical properties that are comparable to, or better than, those exhibited at a density of 6.7 g/cm3 for
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CONCLUSIONS Lower prealloyed MnS levels (in the range 0.0 to 0.65 w/o) do not significantly influence TRS, UTS, and impact toughness. However, samples containing 1.0 w/o prealloyed MnS particles exhibit consistently lower TRS and UTS values at a given density. In drilling, the machinability criteria evaluated show an improvement with increasing levels of prealloyed MnS particles. Machinability in drilling also improves with density, especially for lower levels of prealloyed MnS particles, even though apparent hardness is higher. For samples containing 1.0 w/o MnS, a high level of prealloyed MnS particles outweighs the effect of density in determining machinability. Of the machinability criteria evaluated in drilling, the average thrust force appears to be most sensitive to prealloyed MnS and density, possibly because tool wear and the slope of the drillability curve are difficult to measure. In turning, machinability improves with an increase in the level of prealloyed MnS particles (0.35 to 1.0 w/o). Samples containing 1.0 w/o MnS exhibit the best machinability. Tool wear is slightly lower for the 7.0 g/cm3 density samples. Samples containing 0.0 w/o MnS exhibit lower tool wear than those of samples containing 0.35 w/o and 0.65 w/o MnS. This is related to the UTS of the different materials. It appears that in turning operations, machinability is influenced by the strength of the material. Samples containing 1.0 w/o MnS exhibit low mechanical properties at a given density. However, increasing the density improves properties while achieving excellent machinability both in drilling and in turning. REFERENCES 1. J.S. Agapiou and M.F. Devries, “Machinability of Powder Metallurgy Materials”, Int. J. Powder Metall., 1988, vol. 24, no. 1, pp. 47–57. ˘ alak, M. Selecka and H. Danninger, Machinability of 2. A. S Powder Metallurgy Steels, 2005, Cambridge International Science, Cambridge, UK. 3. O.W. Reen, “The Machinability of P/M Materials”, Modern Developments in Powder Metallurgy, edited by H.H. Hausner and P.W. Taubenblat, Metal Powder Industries Federation, Princeton, NJ, 1977, vol. 10, pp. 431–452. 4. D. S. Madan, “An update on the use of Manganese Sulfide (MnS) Powder in Powder Metallurgy Applications”,
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5. 6.
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Advances in Powder Metallurgy and Particulate Materials, compiled by L.F. Pease and R.J. Sansoucy, Metal Powder Industries Federation, Princeton, NJ, 1991, vol. 3, pp. 101–115. R.M. German, Powder Metallurgy Science, Second Edition, 1994, Metal Powder Industries Federation, Princeton, NJ. “MPIF Standard 35, Materials Standards for PM Structural Parts”, 2007, Metal Powder Industries Federation, Princeton, NJ. S.A. Kvist, “Turning and Drilling of Some Typical Sintered Steels”, Powder Metallurgy, 1969, vol. 12, no. 24, pp. 538–565. D.S. Madan, “Effect of Manganese Sulfide (MnS) on Properties of High Pr for mance P/M Alloys and Applications”, Advances in Powder Metallurgy and Particulate Materials, compiled by J.M. Capus and R.M. German, Metal Powder Industries Federation, Princeton, NJ, 1992, vol. 4, pp. 245–267. H.I. Sanderow, J.R. Spirko and R. Corrente, “The Machinability of P/M Materials as Determined by Drilling Tests”, Advances in Powder Metallurgy and Particulate Materials, compiled by R.A. McKotch and R. Webb, Metal Powder Industries Federation, Princeton, NJ, 1997, vol. 2, part 15, pp. 125–143. P.J. Andersen and J.S. Hirschhorn, “The Influence of Some Metallurgical Variables on the Machinability of Sintered Steels”, Influence of Metallurgy on Machinability, compiled by V. Tipnis, ASM, Metals Park, OH, 1975, pp. 143–158. B. Hu and S. Berg, “Optimizing the Use of Manganese Sulfide in P/M Applications”, Advances in Powder Metallurgy and Particulate Materials—2000, compiled by H. Ferguson and D.T. Whychell, Sr., Metal Powder Industries Federation, Princeton, NJ, 2000, part 5, pp. 191–197. H. I. Sanderow and T. Prucher, “The Effect of Manganese Sulfide on the Mechanical Properties of P/M Steels”, Advances in Powder Metallurgy and Particulate Materials, compiled by A. Lawley and A. Swanson, Metal Powder Industries Federation, Princeton, NJ, 1993, vol. 4, pp. 97–108. K.S. Chopra, “Improvement of Machinability in PM Parts Using Manganese Sulphide”, Progress in Powder Metallurgy, compiled by C.L. Freeby and H. Hjort, Metal Powder Industries Federation, Princeton, NJ, 1987, vol. 43, pp. 501–510. R. J. Causton, “Machinability of P/M Steels”, Advances in Powder Metallurgy and Particulate Materials, compiled by M. Phillips and J. Porter, Metal Powder Industries Federation, Princeton, NJ, 1995, vol. 2, part 8, pp. 149–170. C. Blais and G. L’Espérance, “Turning and Drilling of Parts Made From Sinter Hardenable Steel Powders”, Powder Metallurgy, 2002, vol. 45, no. 1, pp. 39–47. C. Blais, G. L’Espérance and I. Bourgeois, “Characterisation of Machinability of Sintered Steels During Drilling Operations”, Powder Metallurgy, 2001, vol. 33, no. 1, pp. 67–75. N. Akagi, S. Kawai, M. Satoh and Y. Seki, “Machinability and Fatigue Properties of a Pre-alloyed, Free Cutting Steel Powder”, ibid. reference no. 11, part 6, pp. 17–23. J. Campbell-Tremblay, C. Blais, G. L’Espérance and P.
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Boilard, “Characterization of the Fatigue Performances of P/M Components Produced with Powders Developed for Improved Machinability”, Advances in Powder Metallurgy and Particulate Materials—2005, compiled by C. Ruas and T.A. Tomlin, Metal Powder Industries Federation, Princeton, NJ, 2005, vol. III, part 10, pp. 150–159. 19. T. Cadle, F. Bopp and C. Landgraf, “Why Sintered Powder Metal Standard Machinability Tests Don’t Work and an Alternative Approach That Does”, Advances in Powder Metallurgy and Particulate Materials, compiled by J.J. Oakes and J.H. Reinshagen, Metal Powder Industries Federation, Princeton, NJ, 1998, vol. 1, part 2, pp. 3–15. 20. L.G. Roy, “Production of Iron Powder—The Domfer Process”, ASM Handbook Volume 07: Powder Metallurgy Technologies and Applications, 1984, ASM International, Metals Park, OH, pp. 89–92. 21. “Standard 10, Method for Preparing and Evaluating
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Tensile Specimens of Powder Metallurgy Materials”, Standard Test Methods for Metal Powders and Powder Metallurgy Products, Metal Powder Industries Federation, Princeton, NJ, 2007. “Standard 40, Method For Determination of Impact Energy of Unnotched Powder Metallurgy Test Specimens”, ibid reference no. 21. “Standard 41, Method for Determination of Transverse Rupture Strength of Powder Metallurgy Materials”, ibid reference no. 21. A.F. De Rege, G. L’Espérance, J. Gingras and S. Martel, “Test Bench for the Routine Evaluation of Machinability of Powder Metallurgy Materials”, ibid. reference no. 9, vol. 2, part 15, pp. 43–58. D.S. Madan, “The Importance of Machinability in the Processing of P/M Parts”, ibid. reference no. 14, vol. 2, part 8, pp. 55–68. ijpm
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PM MACHINABILITY
GREEN MACHINING: PARAMETERS, APPLICATIONS, AND SINTERED PROPERTIES Étienne Robert-Perron* and Carl Blais**
INTRODUCTION The PM process allows the manufacture of components with relatively complex geometries requiring minimal machining operations. Thus, a discussion of the machining behavior of PM components may appear paradoxical since the PM process is touted to be net or near-net shape. However, due in part to the limitations of the process (for example, it is not possible to generate undercuts in axial pressing), as well as strict dimensional conformance, machining operations are necessary for approximately one-third of the ferrous PM parts produced in North America.1 In Europe, it is estimated that about 40% to 50% of all ferrous and steel PM parts undergo some machining.2 Unfortunately, the machining performance of PM components differs from that of wrought steel with a similar microstructure. Reasons for the inferior machinability of PM components derive from the porosity (10–15 v/o) that affects the interaction between the cutting tool and the work material.3 This behavior leads to a significant increase in temperature at the tool–chip interface which decreases the cutting-tool life. The machining of sinterhardenable components is even more difficult since these parts exhibit hard microstructures (martensite/bainite). An approach to circumvent this limitation of PM steel components is green machining, i.e., machining prior to sintering. At this stage of the manufacturing process, strong atomic bonds between particles as well as hard phases (martensite/bainite) are not yet formed. Therefore, the mechanical properties of green compacts arise only from the cold welding and mechanical interlocking of neighboring powder particles. Green machining is sometimes described as being mandatory for components that are sinter-hardened since machining after sintering is often seen as being “impossible” to perform. 2 However, green machining is not a straightforward procedure and several issues must be addressed. This paper highlights key parameters for successful machining of green PM components. Different machining operations performed on green compacts are presented as well as the attendant sintered properties. Direct reclamation of chips generated during green
The machining of green powder metallurgy (PM) components is an attractive approach to reduce manufacturing costs and to compete with other shaping processes. During the last decade, different strategies have been developed to increase the strength of green PM components, thereby making it possible to consider green machining. However, before utilizing green machining in an industrial environment, material and process issues must be studied and optimized. This paper highlights key parameters in green machining. Different machining operations performed on green PM components are discussed as well as the attendant sintered properties.
*Process Metallurgist, Quebec Iron and Titanium, Inc., 1625 Route Marie-Victorin, Sorel-Tracy, Québec, Canada; E-mail:
[email protected], **Professor, Department of Mining, Metallurgical and Materials Engineering, Université Laval, Québec City, QC, Canada.
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machining for subsequent application in highstrength PM components is also discussed. MACHINABILITY CRITERIA The performance of components machined in their green state is usually characterized in terms of two criteria: surface finish and the average width of breakouts. The surface finish is generally measured using a profilometer or by means of scanning electron microscopy (SEM). The average width of breakouts is the mean width of particles that are pulled out, measured in the area near the outlet edge (defined as the last edge seen by the cutting tool). Measuring the average width of breakouts in the area surrounding the inlet edge (defined as the edge machined when the cutting tool starts cutting the component) is not necessary since the quality of the latter edge shows a much lower dependence on the green machining process. The average width of breakouts is measured from SEM or optical micrographs (OM) using an image-analysis routine. Several measurements are performed on each micrograph and averaged to obtain the average width of breakouts. Figure 1 summarizes this technique for measuring breakouts near the outlet edge after the facing of a green PM ring. In this case, the outlet edge is the internal diameter since the cutting tool enters the component by the external diameter and leaves the part by the internal diameter. Usually, the depth of breakouts is approximately one third of the average width. KEY PARAMETERS IN GREEN MACHINING Green Strength Representative steel parts pressed to a green density of 6.80 g/cm3 have green strength values
in the range 12 MPa to 17 MPa (~1,750 psi to ~2,500 psi). These values are generally too low to allow for holding of the parts in the chuck of a lathe or a machining center and lead to catastrophic failure during machining. 2 To secure clamping during machining, and to ensure highquality components in terms of surface finish and average width of breakouts, green-strength values >35 MPa (5,000 psi) are typically necessary. 4 Such values are obtained either with warm compaction or binder/lubricant technologies. Warm compaction involves pressing a preheated powder at a temperature typically ranging from 90°C to 150°C to obtain green strengths about four times higher than those provided by room-temperature compaction.5 The binder/lubricant technologies consist of substituting for conventional lubricants (e.g., ethylene bis-stearamide (EBS)) with polymeric-based compounds that promote mechanical interlocking and cold welding between powder particles to enhance strength. Such polymeric binders/lubricants have the ability to form a strong network that strengthens the green specimens during compaction and/or by performing a curing treatment at approximately 190°C in air for 1 h.6,7 Binders/Lubricants High green strength is usually associated with improved green machinability. However, this assumption is not always true and may lead to the development of high-green-strength components with low green machinability. Other aspects, such as lubricating properties, as well as the level of admixed binder/lubricant, affect the quality of green machined PM components.8,9 For example, Table I shows that Lube A and the binder/lubricant B show similar green strengths (27 MPa and 28 MPa, respectively) but that the cutting forces TABLE I. MEASURES FOR COMPONENTS STUDIED (B/L = Binder/Lubricant, Cured = Heat Treatment at 190°C for 1 h in Air) FLN4-4405, Green Density 7.2 g/cm3 Mix
Figure 1. Representative micrograph of outlet edge after green machining FC-0208. Dark surface corresponds to areas where particles were pulled out by passage of the cutting tool. Average width of breakout is 166 µm. (OM)
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Green Strength Cutting Force Breakouts Width (MPa) (N) (µm)
0.75 w/o EBS@60°C 0.58 w/o lube A@60°C 0.55 /o B/L B@120°C 0.65 w/o B/L C@60°C 0.65 /o B/L C@60°C Cured
14 27 28 22 38
157 144 131 128 116
839 435 291 431 241
Source: Reference No. 9 Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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measured are significantly different (144 N and 133 N, respectively). This difference in cutting forces translates into an important divergence in machinability, as characterized by the average width of breakouts (435 µm vs. 291 µm). This difference is clearly a function of the lubricating properties in machining of the lubricating systems studied. Lubricant properties of binders/lubricants are frequently studied for compaction and ejection performance. Usually, polymeric-based compounds exhibit superior lubricating properties compared with conventional EBS wax. It is believed that the ability of some polymers to impart good lubricating properties is related to the regular arrangement of the macromolecular chains and their ability to slide over one another when submitted to shear stress.7 As shown, these characteristics are important during green machining for reducing the friction and, therefore, the cutting forces. Binders/lubricants that lubricate the interface between the cutting tool and the chip lower the cutting forces during green machining, thus minimizing particle pullout on the freshly machined surface, as well as near the outlet edge.8,9 Therefore, high-green-strength components would exhibit mediocre machinability if manufactured with binders/lubricants that result in low lubrication properties in machining. Similarly, components made with a low weight fraction of binder/lubricant (e.g., 0.20 w/o) would not exhibit a green machinability performance as good as that of parts made with 0.70 w/o binder/lubricant.8 Density Gradients It is well established that green density influences green strength. A high green density, reflecting a high compaction stress, promotes increased particle movement and deformation, which is the basis for cold welding, mechanical interlocking, and increased green strength. 3 However, due to non-uniform stress distributions during compaction, density gradients may be induced in green compacts. Die displacement during compaction, as well as friction between particles and die walls, are the main causes of pressure gradients that lead to the occurrence of density gradients. In double-action pressing, the area with the lowest density is usually located near the mid-height of the component.3 Since green density influences green strength, there are green-strength gradients in green compacts. In addition, the machinability of green comVolume 44, Issue 2, 2008 International Journal of Powder Metallurgy
ponents, in terms of surface finish and edge integrity, is a function of green strength. Other things being equal, when green strength varies within a component, the machinability of the green compact varies accordingly.10 Figure 2(a) illustrates the porosity gradients (interpreted as density gradients) measured by image analysis from cross sections of the ring shown in Figure 2(b). Figure 3 shows representative micrographs of machined surfaces (walls of grooves machined along the radial axis) taken at different locations along the height of the component (corresponding to the arrows 1, 2, and 3 in Figure 2(a)). These micrographs show that variations in green density significantly influence surface finish and the average width of breakouts of components machined in their green state. The micrograph in Figure 3(a) shows that when a groove is machined in an area with a relatively high level of porosity (arrow 1 in Figure 2(a)), edge integrity is severely affected. When the machining operation is performed in areas of lower porosity (arrows 2 and 3 in Figure 2(a)), edge integrity improves, as shown in Figures 3(b) and 3(c). Cutting Parameters and Tool Wear For sintered components, the cutting conditions are of crucial importance when machining green PM parts. However, the cutting parameters
Figure 2. (a) typical porosity gradient in a green component (b) ring used for characterization
Figure 3. Representative micrographs of machined surfaces of FLC-4608 from areas arrowed in Figure 2. (a) area corresponding to arrow 1, (b) area corresponding to arrow 2, and (c) area corresponding to arrow 3. (OM/unpolished and unetched)
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used for producing high-quality components in green machining differ significantly from those used in the machining of sintered products. It has been shown that for drilling or turning operations, the surface speed can be maximized to increase productivity since it has a minimal effect on quality, based on surface finish and edge integrity.4,9,11,12 Also, it is usually believed that the feed rate must be kept to ≤0.0254 mm/r in order to prevent pullout of particles near the edges, and for optimizing surface finish.4,9,10,12 Tool Wear Usually, tool wear is considered negligible when characterizing the machinability of green PM specimens. This assumption is inaccurate and may lead to the selection of suboptimum cutting conditions. Tool wear is of importance when studying the machinability of green PM components. It has a significant effect on the quality of the outcome, especially when the cutting conditions are not optimal. When machining green PM components with a
Figure 4. Comparison of cutting-tool wear in machining of green and sintered FC-0208 (FLOMET HGS™)
worn cutting tool, the particles near the outlet edge are not cut but are primarily pushed out, increasing the width of the breakouts.13,14 Figure 4 compares tool wear when facing PM components in the green and sintered states. It is seen that, even if the rate of tool wear is significantly lower for the machining of green compacts than for the machining of sintered parts, tool wear does occur in green machining. Figure 5 illustrates the outlet edges of rings machined (by facing) in the green state (surface speed = 366 m/min, feed rate = 0.0635 mm/r, depth of cut = 0.254 mm). Figure 5(a) was taken after the machining of 35 cm3, while Figures 5(b) and 5(c) were taken after the machining of 1,100 cm3 and 3,500 cm3, respectively. It is clear that the average width of breakouts increases with the rate of tool wear, and that the surface finish follows the same pattern.13,14 As seen in Figures 4 and 5, tool wear occurs in green machining and it affects green machinability in terms of surface finish and edge integrity. Moreover, when considering tool wear, the optimum cutting conditions change compared with those characterized by negligible tool wear. Indeed, up to now, published research on green machining assumed that tool wear was negligible in green machining. On the other hand, it has recently been shown that the generally accepted feed rate of 0.0254 mm/r reported in the literature induces an excessive dwell time of the tool during the green machining of large quantities of parts. This leads to premature wear of the tip of the cutting tool and to accelerated deterioration of the surface finish. It is now recommended that the feed rate be increased to 0.0635 mm/r. This latter feed rate prevents accelerated tool wear, thus minimizing its detrimental effect on surface finish and the average width of breakouts.13,14 For similar reasons,
Figure 5. Average width of breakouts near outlet edge (ID) of rings after machining of green FC-0208 (FLOMET HGS™) components: (a) 35 cm3, average width of breakouts: 137 µm, (b) 1,100 cm3, average width of breakouts: 204 µm, and (c) 3,500 cm3, average width of breakouts: 433 µm. (SEM/Secondary Electron Images)
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surface speed must be reduced to approximately 366 m/min when facing green PM components. Such a surface speed decreases tool wear during green machining and minimizes its effect on component quality. Under these optimized cutting conditions (surface speed 366 m/min and feed rate 0.0635 mm/r) it has been shown that the productivity of green machining of PM components is approximately twice that for the machining of sintered components.15,16 MACHINING OPERATIONS Drilling Different PM components can be successfully manufactured via green machining processes such as drilling. Figure 6(a) illustrates the outlet edge of a hole machined in a green PM part.In Figure 6(b) it is seen that, even if the outlet edge is sharp, some particles have been pushed out during green drilling, leading to an average width of breakouts of 115 µm.12 This represents the size of approximately one or two powder particles since D50 of the
base powder was 77 µm. This level of edge quality was obtained after drilling a component made from a FLC-4608 powder blend (FLOMET HGS™) with a green strength of 52.2 MPa (7,575 psi). The drilling parameters were: surface speed 140 m/min (7,000 rpm), feed rate 0.0254 mm/r, and KWCD00461 drill type (6.35 mm dia.), produced by Kennametal. The acceptable result is explained in part by the fact that a high helix angle was selected in drilling the hole which favors the extraction of chips from the cutting area (drill tip). Turning Turning is another operation that can be performed on green PM components. Figure 7(a) shows how to perform external turning on green PM rings. Rings are fixed to a wrought steel bar by a tightening screw which is held in the jaws of the chuck during machining. For internal turning and facing, green components can be held directly in the jaws of the lathe, as shown in Figure 7(b).14 As noted previously, clamping of green parts is
Figure 6. Representative micrographs of outlet edge of hole machined in FLC-4608 under optimum cutting conditions: (a) typical mosaic used to characterize the average width of breakouts (OM unpolished-unetched), and (b) typical edge showing breakouts. (SEM/Secondary Electron Images)
Figure 7. (a) fixture used for external turning: #1—steel bar fixed in chuck to hold sample; #2—sample to be machined; #3—tightening screw used to fix sample on steel bar, and (b) components held in jaws of chuck for internal turning and facing
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GREEN MACHINING: PARAMETERS, APPLICATIONS, AND SINTERED PROPERTIES
feasible as long as the green strength is sufficient (typically > 35 MPa (5,000 psi)). In the present case, a green strength of 45 MPa (6,525 psi) was obtained on a FC-0208 (FLOMET HGS™) powder. The average width of the breakouts measured after facing this green PM ring was 137 µm, while surface finish (Ra) was 3.00 µm. However, as cited previously, these values increased as the wear on the cutting tool increased. 14 Optimum cutting conditions for turning/facing such components were found to be: surface speed 366 m/min and a feed rate of 0.0635 mm/r. Tool holder SCLCR102 and cutting tool CPMT060204FW (Kennametal) were used since they minimize quality degradation related to tool wear in green machining.13 Grooving Another successful application of green machining is timing sprockets, in which a groove is machined along the radial axis in sprocket-shaped components, as illustrated in Figure 8(a). Such PM components, especially those made from sinter-hardenable steel powders, are often judged to be impossible to machine. The cyclic impact on the cutting tool from tooth to tooth generally leads to catastrophic failure of the cutting tool after the machining of a few sintered parts (e.g., five components were machined by the authors prior to tool failure (powder type: FLC4608)—cutting tool: GIP 6.00E-0.80 IC-9015 manufactured by ISCAR, surface speed: 122 m/min, and feed rate: 0.0127 mm/r). The green machining of such PM parts is feasible and prevents catastrophic tool failure. In this case, an FLC-4608 (FLOMET HGS™) sinter-hardenable powder was used with a green strength of 37 MPa (5,375 psi). The optimum cutting conditions were: surface speed 457 m/min and feed
rate 0.0254 mm/r, which ensures a level of productivity three times higher than that in the machining of sintered components.10 Following these cutting conditions, the average width of breakout near the outlet edge was 220 µm, as seen in Figure 8(b). It is important to note that no teeth broke during these green machining tests with this powder blend. SINTERED PROPERTIES Radial-Crushing Strength Even if PM components with satisfactory geometrical quality have been produced by green machining, there are, to our knowledge, only two studies reported in the literature14,17 concerning the sintered properties of such components. The first study focused on the radial-crushing strength of sintered PM rings manufactured via green machining. In this study, green PM rings (FC-0208) have been fully green machined (OD, ID, and the two faces), as illustrated in Figure 7. Radial-crushing strength after sintering has been compared with that of rings machined after sintering. The radial-crushing strength of samples machined in their green state was 735 MPa (107,000 psi) compared with 761 MPa (110,000 psi) when components were machined after sintering. The small difference (3%) between the two sets of results is not significant since the standard deviations were 78 MPa (11,300 psi) and 30 MPa (4,350 psi), respectively, for the parts machined in their green state and parts machined post-sintering. Components exhibited similar microstructures after sintering, namely, a mix of pearlite and bainite, with an apparent hardness of 78 HRB. Moreover, the dimensional change of the rings machined in their green state was identical
Figure 8. (a) simulation of machining of timing sprocket, (b) representative edge integrity in FLC-4608 characterized after machining of groove; average width of breakouts 220 µm. (OM)
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to that of the unmachined rings (+0.33%), which is characteristic of this type of powder system. Tensile Properties The second study compared the sintered tensile properties of cylindrical specimens machined in their green state with those machined after sintering.17 An FLC-4608 powder blend was pressed into bars and green machined to obtain tensile specimens. The specimens were sintered in a belt furnace in an atmosphere of 90 v/o N2-10 v/o H2 at 1,120°C for 25 min and rapidly cooled (global cooling rate 1.5°C/s from 650°C to 400°C) to transform the austenite into martensite. The second series of specimens was green machined to a diameter 0.05 cm larger than that of the final size of the tensile specimens. After sintering this second series of samples, under the same conditions cited, were remachined to final diameter.9 This procedure was necessary to obtain the same microstructure in the final samples and, therefore, to compare only the influence of surface finish on the sintered tensile properties between parts machined in the green state and parts machined post-sintering. Table II cites the tensile properties measured on the two series of specimens, each exhibiting a microstructure of tempered martensite. The apparent hardness of these two series is identical while the tensile properties (yield strength, ultimate strength, and ductility) of the greenmachined components were within 5% lower than those of the components machined after sintering. These two examples confirm that, for the same microstructure, the sintered properties of the components green machined are essentially identical to those of parts machined conventionally (post-sintering). Moreover, based on the sintered properties, no cracks or microcrack formation was initiated during green machining or handling of the green PM specimens. Indeed, if this had TABLE II. COMPARISON OF TENSILE PROPERTIES OF GREEN-MACHINED & SINTERED COMPONENTS AND COMPONENTS MACHINED AFTER SINTERING Property
Machined in Green State
Machined after Sintering (Pre-machined in Green State)
Apparent Hardness (HRC) Yield strength (MPa) Ultimate strength (MPa) Elongation (%)
33 ± 1 940 ± 12 980 ± 36 1.08 ± 0.04
33 ± 1 950 ± 20 1,028 ± 22 1.14 ± 0.04
Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
occurred, the sintered properties would have been significantly lower than those of the components machined after sintering. The marked increase in tool life and productivity, combined with the possibility of machining PM materials that are often considered nearly impossible to machine, make green machining an attractive approach to increase the performance of the PM process. CHIP RECLAMATION Up to now, the fabrication of PM components from chips generated during the machining of wrought materials or sintered products has not been an economical method for producing highquality components.18 The use of coolants/lubricants during machining and in subsequent treatments to obtain the desired size distribution of chips constitutes a severe obstacle. However, chips formed during the machining of green PM components are clean and similar in their size distribution to that of the base powder. It has been shown recently that these chips can be reused directly without subsequent treatment in the production of high-strength PM components.19 Up to 20 w/o of chips can be added directly to a compatible powder mix without significantly affecting compaction behavior and sintered properties. 19 When 20 w/o of reclaimed chips was added to a similar base powder, the apparent density, green strength, and the compressibility varied within 5% of the values of the reference blends (0 w/o chips). With 20 w/o chips, the flow rate decreased by 20% due to the plate-like morphology of the chips. After sintering, the transverserupture strength and the apparent hardness of components made using 20 w/o reclaimed chips were equal to those of the base system (0 w/o chips). Table III summarizes the characteristics of TABLE III. CHARACTERISTICS OF COMPONENTS MADE WITH 20 w/o RECLAIMED CHIPS (FLC-4608 FLOMET HGS™, CURED) Property
Apparent Density (g/cm3) Flow Rate (s/50 g) Green Strength (MPa) Compressibility at 6.80 g/cm3 (MPa) TRS after Sintering (MPa) Apparent Hardness (HRC) Dimensional Change (%)
FLC-4608 (20 w/o) Reclaimed Chips
FLC-4608 (0 w/o) Reclaimed Chips
3.16 33.5 44 495 1,390 25 0.25
3.16 27.9 46 470 1,350 25 0.31
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components made with 20 w/o reclaimed chips compared with conventional parts (identical chemical composition and sintering conditions). CONCLUSIONS The machining of green PM components is not a straightforward procedure since green strength, the type of binder/lubricant, density gradients, cutting parameters, and cutting-tool wear affect green machinability. These parameters must be optimized in green machining in order to produce high-quality components in terms of surface finish and average width of breakouts. Green machining offers several advantages compared with the machining of sintered components: (1) Provides a solution to the machining of sinter-hardened PM parts (2) Increases productivity (3) Reduces the wear rate of the cutting tool (4) Offers viability in the machining of interrupted components (e.g., gears and/or timing sprockets) (5) Allows for the recycling of chips for subsequent use in high-quality PM components Sintered properties (radial-crushing strength and tensile properties) of components machined in their green state are essentially identical to those of PM components machined conventionally (post-sintering). REFERENCES 1. D.S. Madan, “The Importance of Machinabilty in the Processing of P/M Parts”, Advances in Powder Metallurgy & Particulate Materials, compiled by M. Philips and J. Porter, Metal Powder Industries Federation, Princeton, NJ, 1995, part 8, pp. 55–67. 2. A. Salak, M. Selecka and H. Danninger, Machinability of Powder Metallurgy Steels, 2005, Cambridge International Science Publishing, Cambridge, UK. 3. R.M. German, Powder Metallurgy & Particulate Processing, 2005, Metal Powder Industries Federation, Princeton, NJ. 4. E. Robert-Perron, C. Blais, S. Pelletier and Y. Thomas, “Special Interest Program—Machining”, PowderMet2006, Metal Powder Industries Federation, Princeton, NJ, San Diego, CA, June 18–21, 2006. 5. S. St-Laurent and F. Chagnon, “Designing Robust Powder Mixes for Warm Compaction”, Advances in Powder Metallurgy & Particulate Materials, compiled by R.A. McKotch and R. Webb, Metal Powder Industries Federation, Princeton, NJ, 1997, part 3, pp. 3–18. 6. L. Tremblay and Y. Thomas, “Enhanced Green Strength Lubricating Systems for Machining Ferrous Materials”, Advances in Powder Metallurgy & Particulate Materials, compiled by C. Rose and M. Thibodeau, Metal Powder Industries
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Federation, Princeton, NJ, 1999, part 2, pp. 141–156. 7. L. Tremblay, J.E. Danaher, F. Chagnon and Y. Thomas, “Development of Enhanced Green Strength Lubricating Systems for Green Machining”, Advances in Powder Metallurgy & Particulate Materials—2002, compiled by V. Arnold, C-L Chu, W.F. Jandeska, Jr. and H.I. Sanderow, Metal Powder Industries Federation, Princeton, NJ, 2002, part 12, pp. 123–137. 8. E. Robert-Perron, C. Blais, S. Pelletier and Y. Thomas, “Binder/Lubricants in Green Machining”, Advances in Powder Metallurgy & Particulate Materials—2007, compiled by J. Engquist and T. Murphy, Metal Powder Industries Federation, Princeton, NJ, 2007, part 6, pp. 1–13. 9. P. Lapointe, C. Blais, S. Pelletier, Y. Thomas and S. StLaurent, “Characterization of Different Lubrication Approaches to Improve Green Machinability”, Proc. Euro PM2007, European Powder Metallurgy Association, Shrewsbury, UK, 2007, vol. 3, paper #783. 10. E. Robert-Perron, C. Blais, Y. Thomas, S. Pelletier and M. Dionne, “An Integrated Approach to the Characterization of Powder Metallurgy Components Performances During Green Machining”, Mater. Sci. Eng. A-Struct., 2005, vol. 402, pp. 325–334. 11. A. Benner and P. Beiss, “Green Tur ning of War m Compacted P/M Steels”, Advances in Powder Metallurgy and Particulate Materials—2001, compiled by B. Eisen and S. Kassam, Metal Powder Industries Federation, Princeton, NJ, 2001, part 6, pp. 1–15. 12. E. Robert-Perron, C. Blais, S. Pelletier and Y. Thomas, “Drilling of High Quality Features in Green Powder Metallurgy Components”, Mater. Sci. Eng. A-Struct., 2007, vol. 458, pp. 195–201. 13. E. Robert-Perron, C. Blais, S. Pelletier and Y. Thomas, “Machinability of Green Powder Metallurgy Components, Part I: Characterization of the Influence of Tool Wear”, Metal. Mater. T rans-A., 2007, vol. 38, no. 6, pp. 1,330–1,336. 14. E. Robert-Perron, C. Blais, S. Pelletier and Y. Thomas, “Machinability of Green Powder Metallurgy Components, Part II: Sintered Properties on Components Machined in Green State”, Metal. Mater. Trans-A., 2007, vol. 38, no. 6, pp. 1,337–1,342. 15. C. Blais and G. L’Espérance, “Turning and Drilling of parts Made from Sinter Hardenable Powders”, Powder Metall., 2002, vol. 45, pp. 39–47. 16. J. Campbell-Tremblay, C. Blais, G. L’Espérance and P. Boilard, “Characterization of the Fatigue Performances of P/M Components Produced with Powders Developed for Improved Machinability”, Advances in Powder Metallurgy and Particulate Materials—2005, compiled by C. Ruas and T. Tomlin, Metal Powder Industries Federation, Princeton, NJ, 2005, part 10, pp. 150–159. 17. E. Robert-Perron, C. Blais, S. Pelletier and Y. Thomas, “Tensile Properties of Sinter Hardened PM Components Machined in the Green State”, Powder Metallurgy, in press. 18. J. T. Strauss, “Consultants’ Corner”, Int. J. Powder Metall., 2005, vol. 41, no. 6, pp. 29–32. 19. E. Robert-Perron, C. Blais, S. Pelletier and Y. Thomas, “Chips Reclamation in Green Machining for Producing High-Performance PM Components”, Int. J. Powder Metall., 2007, vol. 43, no. 3, pp. 49–55. ijpm
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PM MACHINABILITY
FACE TURNING OF PM STEELS: EFFECT OF POROSITY AND CARBON LEVEL Andrej S˘alak*, Marcela Selecká**, Karol Vasilko*** and Herbert Danninger**** INTRODUCTION For an increasing number of structural components, which are also becoming more and more complex, the near-net-shape capability of the powder metallurgy (PM) process results in a decrease of scrap and cost compared with machined wrought steel parts. However, there are a significant number of shapes that are difficult (or impossible) to produce without machining. Examples include undercuts, slots, bevels, blind holes, holes perpendicular to the pressing direction, and threads. These require a machining operation (turning, drilling, or milling) to achieve tight tolerances and a high-quality surface finish.1 At present, about 60% of all PM components need some form of machining operation.2 The machinability of PM steels is affected by more factors than that for wrought steels. Many of these factors are not germane to the characteristics of conventional steels, such as alloying with nickel, copper, and molybdenum, with the consequence of a higher hardness of sintered steels compared with wrought steels, at a given tensile strength.3–5 Of the various characteristics of PM steels affecting machinability, porosity is cited as the chief detrimental factor.3,6 Many studies have focused on the effect of porosity on the machinability of sintered steels, primarily in turning, assuming that the detrimental effect of porosity was due to a constantly interrupted cut as the tool edge moved out of the metallic phase to the pores. This means that the material is cut by the tool passing over one pore after another.6–8 This description of the cutting process has been termed the “interrupted cutting theory,”3 but it did not consider the true sequence of events in the cutting process. Experimental proof of this cutting mode in relation to porosity is lacking. In contrast, analysis of machined surfaces of sintered parts in turning,9 and in drilling of structural steels10 and stainless steels11 showed that they are formed by deformed and work-hardened layers.6 This means that the cutting edge separates the chip from a pore-free workhardened area of the workpiece. This explanation has been called the
The machinability of sintered steels is inferior to that of the wrought steels they frequently replace; porosity is regarded as the primary detrimental factor. This study examined the effect of porosity on the machinability in the face turning of six sintered steels at three levels of porosity (11, 15 and 20 v/o) and two carbon levels (0.3 and 0.5 w/o). The machinability was determined by the number of passes conducted up to a critical flank wear VBc of 0.3 mm. The machined surfaces and the built-up edges (BUEs) formed by quick-stop experiments were analyzed by optical (OM) and scanning electron microscopy (SEM). The cutting process was characterized by the presence of large deep turning feed marks, and by the formation of a built-up edge in the sintered materials at the expense of pores. These are the mechanisms by which porosity impairs machinability. It was determined that the machinability of porous materials is affected by their resistance to plastic deformation. High porosity in sintered steels results in reduced machinability due to lower resistance to plastic deformation, and vice versa. The same holds true for carbon content: increasing the carbon level from 0.3 w/o to 0.5 w/o increases the resistance to plastic deformation of sintered steels and enhances machinability.
*Chief Scientific Worker, **Senior Scientific Worker, Institute of Materials Research of SAS, Watsonova 47, 043 53 Kos˘ice, Slovak Republic; E-mail:
[email protected], ***Professor, Technical University Kos˘ice, Faculty of Manufacturing Technologies, Bayerova 1, 080 01 Pres˘ov, Slovak Republic, ***Professor, Vienna University of Technology, Inst. for Chemical Technologies and Analytics, Getreidemarkt 9/164, A-1060 Vienna, Austria
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FACE TURNING OF PM STEELS: EFFECT OF POROSITY AND CARBON LEVEL
“deformation cutting theory,” 3 but it did not explain the mechanism of the detrimental effect of porosity on the machinability of PM steels. In general, it was found that the machinability and surface finish of sintered steels improved with increasing density.11,12 The present study focused on experimental and theoretical explanations of the mechanisms relevant to the detrimental effect of porosity on machinability of various sintered alloys. The steels were evaluated by face turning under identical cutting conditions. EXPERIMENTAL PROCEDURE For the preparation of the test pieces the powder mixes/steels cited in Table I were used.13 Graphite in amounts of 0.3 w/o and 0.5 w/o, respectively, was added to the powders as natural graphite CR12 (GRAFIT, a.s., Netolice) as was 0.8 w/o HWC (microwax C) as a lubricant. The ringshaped specimens (10 mm ID / 30 mm OD × 15 mm) prepared from the powder mixes with 0.3 w/o graphite addition were compacted to 20 v/o and 15 v/o porosity (green density 6.5 g/cm3 and 6.7 g/cm3) and rings with 0.5 w/o graphite were compacted to 20 v/o, 15 v/o, and 11 v/o porosity (green density 6.5 g/cm 3, 6.7 g/cm 3, and 6.9 g/cm3). The density of the samples was determined by the water displacement method according to ISO 3369. The sintering of the specimens was carried out in a laboratory tube furnace at 1,121°C (2,050°F) for 30 min in dissociated ammonia (dew point -30°C) under a getter (alumina + 5 w/o graphite), to avoid decarburization. The cooling rate over the temperature interval 900°C to 300°C (1,652°F to 572°F) was 0.06°C/s (0.11°F/s). The face-turning test method for assessing the machinability of PM steels shown in Figure 1 has been described as a possible standard test.14 The dry face-turning test was carried out under the following constant conditions: turning speed of
the lathe spindle (i.e., of the workpiece) 1,400 rpm; feed rate f 0.13 mm/rev. (0.005 in./rev.); depth of cut h 0.1 mm (0.004 in.), cutting from the center. The initial cutting speed at the ID of the specimen (10 mm) was 44 m/min (144 ft./min) (vc (min)), and the highest speed at the OD (30 mm) was 132 m/min (433 ft./min) (vc (max)). Triangular hardmetal (HM) indexable inserts (ISO code TPMN110304 TN P20; nose radius rε = 0.4 mm) with a top rake angle γ = +5° were selected with the objective of shortening the test time. The criterion (machining index) for machinability was the number of passes conducted up to a critical flank wear VBc of 0.3 mm which was monitored periodically by optical microscopy. It must be recognized that the parameters for optimum machinability vary with the materials, but, for practical reasons, standard test conditions have to be selected. It should also be stressed that the cutting conditions selected here were not those employed in industrial practice but were more severe in order to arrive at a clearer and faster discrimination between the responses of the different materials. This is the reason for the low cutting speed and for the employment of uncoated hardmetal inserts. The results may be used to generate a machinability ranking between the materials tested under the cutting conditions used, but not for quantitative tool-life assessment in industrial production practice. In this work the following parameters were determined and studied: number of passes in relation to cutting-tool wear; as-sintered hardness; machined-surface characteristics; built-upedge (BUE) formation, all in relation to the level of porosity and carbon level. RESULTS The total porosity, combined carbon content, and as-sintered apparent hardness, as base char-
TABLE I. STEEL COMPOSITIONS (w/o) AND CODES Composition
Base Powder Grade
Fe-C Fe-2 w/o Cu-C Fe-1 w/o Cr-0.3 w/o Mo-0.3 w/o V-C Fe-1.75 w/o Ni-1.5 w/o Cu-0.5 w/o Mo-C Fe-4 w/o Ni-1.5 w/o Cu-0.5 w/o Mo-C Fe-0.85 w/o Mo-C
ASC 100.29, Höganäs AB (Fe) A ASC 100.29, Höganäs AB (Fe-2Cu) B Prealloyed Powder 103 V, Kawasaki (103V) C Distaloy SA, Höganäs AB (Dist.SA) D Distaloy SE, Höganäs AB (Dist.SE) E Astaloy 85Mo, Höganäs (Ast.85Mo) F
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Code
Figure 1. Cutting process in face turning—schematic. Dimensions in mm
Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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FACE TURNING OF PM STEELS: EFFECT OF POROSITY AND CARBON LEVEL
TABLE II. POROSITY (P), COMBINED CARBON CONTENT (CC) AND AS-SINTERED HARDNESS (HV 10) 0.3 w/o C Alloy P (v/o) *Cc (w/o) HV 10
A (Fe)
B (Fe-2Cu)
C (103V)
D (Dist.SA)
E (Dist.SE)
F (Ast.85Mo)
20 15 0.26 51 55
20 15 0.28 54 67
20 15 0.32 64 76
20 15 0.30 99 111
20 15 0.31 121 137
20 15 0.32 107 122
A (Fe)
B (Fe-2Cu)
C (103V)
D (Dist.SA)
E (Dist.SE)
F (Ast.85Mo)
20 15 11 0.51 61 65 81
20 15 11 0.53 68 78 102
20 15 11 0.51 97 109 114
20 15 11 0.50 119 135 153
20 15 11 0.51 140 148 189
20 15 11 0.50 105 127 149
0.5 w/o C Alloy P (v/o) *Cc (w/o) HV10
*The differences in combined carbon content in relation to porosity were in the range ±0.02 w/o acteristics of the steel with 0.3 w/o and 0.5 w/o C admixed graphite, are listed in Table II.
0.3 w/o C and porosity 20 v/o and 15 v/o. Significantly, Figure 4 shows the corresponding data for grades with 0.5 w/o C, and porosity 20
Machinability Figure 2 shows the dependence of cutting tool–flank wear on the number of passes for face turning tests up to VBc = 0.3 mm for steels with 20 v/o and 15 v/o porosity. The results can be compared with the initial wear stage in the toolwear-time diagram for wrought steels,3 and confirm the sensitivity of the face-turning test method with a P20 tool to the steel characteristics, notwithstanding the fact that the tool grade and cutting conditions cannot be optimum for all the steel grades tested. Effect of Porosity and Carbon Content on Machinability Figure 3 shows the machinability of steels with
Figure 3. Machinability (number of passes) of steel grades A–F with 0.3 w/o C and porosity levels of 20 v/o and 15 v/o15
Figure 2. Relation between flank wear (up to VBc = 0.3 mm) and number of passes in face turning with HM P20 inserts for steel grades A–F, porosity 20 v/o and 15 v/o, and carbon contents of 0.3 w/o and 0.5 w/o. A–F grade codes, Table I. First digit: 3 = 0.3 w/o C, 5 = 0.5 w/o C. Second digit: 5 = porosity 20 v/o, 7 = porosity 15 v/o. 62.8 mm3 material removed per pass15
Figure 4. Machinability (number of passes) of steel grades A–F with 0.5 w/o C and porosity levels of 20 v/o, 15 v/o and 11 v/o15
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FACE TURNING OF PM STEELS: EFFECT OF POROSITY AND CARBON LEVEL
v/o, 15 v/o, and 11 v/o. Based on Figure 3, all the steels with a porosity of 15 v/o exhibited superior machinability, though to varying degrees, compared with steels with 20 v/o porosity. The relatively poor machinability of grade A (Fe-0.3 w/o C) is in agreement with the fact that plain Fe-C steels with carbon contents <0.3-0.2 w/o, are known to exhibit poor machinability, due primarily to adhesive effects such as smearing.3,12 The steels with 0.5 w/o C, except grade E, exhibited superior machinability compared with those containing 0.3 w/o C at all three levels of porosity, Figure 4. The unexpected good machinability of grade A (Fe-0.5 w/o C, porosity 15 v/o), and among the alloyed steels, the best of steel D (Distaloy SA) should be analyzed in more detail. The alloy steels D (Distaloy SA) and F (Astaloy 85Mo) with 0.5 w/o C exhibited minor or no differences in machinability in relation to the level of porosity compared with the other grades. The inferior machinability of alloy steel E (Distaloy SE0.5 w/o C) is in agreement with the experience that sintered steel Distaloy SE-0.5 w/o C is extremely difficult to machine.2 Effect of Hardness on Machinability Figure 5 shows the dependence of the number of passes on the as-sintered hardness of the steels. Machinability improved with increasing assintered hardness. This phenomenon is linked to reduced porosity, to the increased carbon content, and to the alloying resulting in increased apparent hardness, Table II and Figures 3 and 4. In Figure 6 essentially a linear relationship is
Figure 5. Machinability (number of passes) of steel grades A–F (0.3 w/o and 0.5 w/o C; 20 v/o, 15 v/o, and 11 v/o porosity) as a function of as-sintered hardness15
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exhibited between the hardness of the machined surfaces and the as-sintered hardness of the steels without marked differences attributable to porosity, carbon content, and alloying, except for steel F (Astaloy 85Mo-0.5 w/o C) with a higher hardness. It may be deduced that this grade of steel exhibited a higher level of plastic deformation and work hardening in the cutting process, compared with the other steels. Characteristics of Machined Surfaces Figure 7 shows the appearance of machined surfaces in the form of turning-feed marks (the width of each turning mark corresponds to the feed) for representative PM steels with 20 v/o porosity after the last pass, and for wrought steel. The surface of grade A steel (Fe-0.3 w/o C) exhibited a markedly different finish, Figure 7(a). This morphology of the machined surface is explained by adhesive effects (smearing) and possibly the adherence of BUE fragments of the steel.3,12,16 The PM alloy steel machined surfaces exhibit the same appearance as that of the wrought steel (pore-free) due to cold deformation and work hardening as a result of the cutting process. Metallographic sections of the machined surfaces of some steels, oriented perpendicularly to the direction of the tool movement (section A–A in Figure 7(e)), are shown in Figure 8. The deep turning-feed marks (14.3 µm to 18.5 µm) measured in Figures 8(b)–8(e), were characteristic of all the sections in the PM steels, except grade A. The machined subsur face of all the steels was deformed and work hardened, as confirmed by the microindentation hardness values, Table III.
Figure 6. Hardness of machined surfaces as a function of as-sintered hardness of steels grades A–F (0.3 w/o and 0.5 w/o C; 20 v/o, 15 v/o, and 11 v/o porosity)
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Figure 7. Machined surfaces of PM steels with 20 v/o porosity and wrought steel: (a) A (Fe-0.3 w/o C), (b) C 103V-0.3 w/o C, (c) D (Distaloy SA0.3 w/o C), (d) E (Distaloy SE-0.3 w/o C), (e) F (Astaloy 85 Mo-0.3 w/o C), (f) C45 E4 ISO 683 wrought steel (hardness 175 HV 10). In (e) vertical arrow points away from center of specimen (decreasing carbon content). SEM/SEI
Typical indentations are shown in Figures 8(c) and 8(d). Figures 8(g) and 8(h) show the uniform turning-feed marks on the metallographic surfaces of two steels at a lower magnification. The section through the machined surface of the wrought steel, Figure 8(f), shows a significant difference in characteristics compared with the PM steels. Only a thin, highly deformed and workhardened layer, the shape of which is similar to the deep feed marks observed in PM steel grades, was formed in these feed marks without deepening. The noses shown on the end of singular feed marks in Figure 8(c) and in Figure 8(e) are the highly deformed material peaks formed at the location of chip detachment from the work material. This enhances the surface roughness with attendant deformation by the tool clearance face and smearing onto the machined surface and the effect appears in the form of veils as shown in Figures 7(b)–7(f). In all the steel grades, the microindentation hardness of the microstructural constituents of the subsurface layers in the turning-feed marks Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
was higher compared with the core values. This is a consequence of surface deformation and work hardening caused by the cutting process. Quick-Stop Experiments Quick-stop experiments freeze the steady-state deformation field in front of the cutting edge, which is the most important region controlling the mechanics of cutting. 17 The frozen quick-stop state also makes it possible to study the formation of the BUE as a part of the chip formation mechanism whose effect lowers the machinability of the material. The BUE is formed by the tool rake face near the cutting edge under extreme strain conditions.3,17–19 In PM steel machining it is necessary to consider in particular the effect of porosity and of heterogeneous microstructure on BUE formation. The quick-stop sections in the two-phase steel A (Fe-0.5 w/o C) and in the multi-phase steel E (Distaloy SE-0.5 w/o C), which has the poorest machinability, at all three levels of porosity for both, are shown in Figure 9, along with the
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Figure 8. Sections of machined surfaces of steels. (a) A (Fe-0.5 w/o C/15 v/o porosity), (b) B (Fe-2 w/o Cu-0.5 w/o C/15 v/o porosity), (c) C (103V-0.3 w/o C/20 v/o porosity), (d) D (Distaloy SA-0.5 w/o C/15 v/o porosity), (e) E (Distaloy SE-0.5 w/o C/15 v/o porosity), (f) C45 wrought steel, (g) A (Fe-0.5 w/o C/15 v/o porosity), (h) F (Astaloy 85 Mo-0.5 w/o C/15 v/o porosity). Optical micrographs, nital etch (except (g), unetched)
TABLE III. MICROINDENTATION HARDNESS IN CORE AND IN TURNING-FEED MARK SUBSURFACE LAYERS OF THE SPECIMENS. A–F GRADE CODES 0.3 w/o C, 20 w/o P Steel
A35
B35
C35
D35
0.5 w/o C, 20 w/o P E35
F35
A55
B55
C55
D55
E55
F55
168–393 230–482
156–399 235–431
162–518 175–553
159–388 406–563
Microindentation Hardness (HV 0.01) Core Mark
82–152 157–254
125–298 181–376
163–365 292–477
145–339 364–536
153–423 225–543
149–340 326–714
0.5 w/o C, 15 w/o P Steel
A57
B57
C57
D57
119–212 168–393
142–310 263–454
0.5 w/o C, 11w/o P E57
F57
A50
E50
Microindentation Hardness (HV 0.01) Core Mark
116–205 164–311
152–264 234–371
163–401 218–452
154–408 211–512
144–449 156–548
145–364 165–507
136–282 157–308
183–521 258–553
P = Porosity wrought steel C45 for comparison. This figure clearly demonstrates intense plastic deformation of the entire region of the material in front of the cutting edge, which shows (Figure 9(c)) that the cut always occurred in pore-free material. The for-
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mation of BUEs, except Figure 9(c) (in a subsequent experiment the BUE was formed), was the result of the intense deformation during chip formation near the cutting-edge tip. The BUE was also formed when cutting PM steel with a sharp Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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Figure 9. Sections through quick-stop experiment samples showing fragments of cutting tool adhereing to BUEs: (a) A (Fe-0.5 w/o C/20 v/o porosity), (b) A (Fe-0.5 w/o C/15 v/o porosity), (c) A (Fe-0.5 w/o C/11 v/o porosity), (d) E (Distaloy SE-0.5 w/o C/20 v/o porosity), (e) E (Distaloy SE-0.5 w/o C/15 v/o porosity), (f) E (Distaloy SE-0.5 w/o C/11 v/o porosity), (g) D (Distaloy SA-0.5 w/o C/11 v/o porosity), (h) D (Distaloy SA-0.5 w/o C/11 v/o porosity), (i) C45 wrought steel. (a) to (e) rounded cutting edge, rε = 0.4 mm. (g) to (i) HM sharp insert. Vc = 80 m/min on 20 mm dia. Optical micrographs, nital etch (except (h), unetched)
HM insert, Figure 9(g). In Figure 9(a) (20 v/o porosity) the arrows indicate a possible effect of porosity on detachment of the chip from the workpiece through the compressed unwelded pores. DISCUSSION The results shown in Figures 2–4 demonstrate the detrimental effect of porosity on the machinability of the materials (ignoring the effect of the microstructure). In the present work, this is explained by two mechanisms operating concurrently under the cutting conditions used. The first one, analyzed on metallographic sections of the machined surfaces (final surface finish), is the formation of the feed marks in wave form, Figure 8. The second one is the formation of the BUEs in quick-stop—in situ, Figure 9. For the explanation of these mechanisms the common orthogonal model of chip formation in Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
cutting wrought-metal materials is applied which, from the physical–chemical aspect is dynamic and complex, Figure 10. This process is linked to the high plastic deformation of the workpiece material ahead of the cutting edge with extremely high deformation rates, under the effect of high stresses. The primary defor mation (flow) zone is induced by the tool motion which compresses the workpiece material immediately ahead of the tool. In this zone the newly formed metal chip is deflected and sheared from the workpiece material. The secondary deformation zone is formed when, during moving along the tool, the chip seizes onto it along the tool rake face.19 The deformation field in the shear plane is uniquely determined by the tool rake angle γn and the shear angle φ. Part of the chip formation process is also the formation of the BUE. These are the zones in which the processes of chip formation and tool
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Figure 10. Orthogonal model of chip, workpiece and tool in chip-formation process in metal cutting.19-21 F = cutting force, Fn(1) = active component, Ft(2) = passive component, h = uncut chip thickness (depth of cut), h1 = chip thickness, shear angle φ = angle between direction of cutting-tool movement and shear plane—schematic
wear in metal cutting occur.20 Plastic deformation generated during cutting, with final consequences on the surface quality of the workpiece and tool life, depends on the stress conditions of the workpiece material and cuttingtool material, and on the geometry of cutting. In a PM cutting pair the workpiece material is characterized by special features compared with the wrought workpiece materials, namely, porosity and a heterogeneous microstructure, which must be considered as the reason for the inferior machinability of PM steels. To explain the mechanism of formation of the turning-feed marks in wave form during cutting on the machined surface of a porous material (Figure 8), utilizing the chip formation model, we elaborate on the scheme shown in Figure 11. The machined surface is the final result of the chipremoval process occurring under the tool clearance face. Thus, Figure 11 shows one characteristic wave-shaped feed mark, the formation of which is illustrated by the seven positions (and angles) of the moving tool (changes in γn and in the direction of Ft). It is necessary, therefore, in each position, to consider the chip formation zones in front of the cutting edge and to characterize the primary factors affecting the cutting process. The primary and extended deformation zone, which extends partly to the detached layer and partly below the tool clearance face into the machined surface layer, is that in which chip for-
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Figure 11. Mechanism of effect of porosity on decreasing machinability of sintered steels in dry face turning along one deep feed mark formed on section of machined surface supported by one section—schematic. Micrograph is Figure 8(e)
Figure 12. Deformation zones in turning ahead of cutting edge at two positions of cutting edge during one revolution of porous sintered workpiece. (a) rake angle γn = 0° and shear angle φmin, maximum depth of the feed mark formed, (b) rake angle +γn and shear angle φmax at start of cutting process. Shaded area is superposed on cutting edge position as in (b) and corresponds to completion of formation of feed mark with rake angle +γn and shear angle φmax
mation occurs under the effect of the cutting force, Figure 10. The degree of deformation depends primarily on the tool rake angle γ and the shear angle φ which copies the rake angle. At an increasing rake angle the shear angle increases. The shear angle can be measured statically by interrupting the cut in the quick-stop experiment, as illustrated in Figure 9(c). The shear angle characterizes the shear plane in which the principal cutting action occurs, and the intensity of the deformation. Along the right side of the shear angle the detachment of the chip occurs and its role in the deformation process in cutting is highlighted. Using an insert with a positive rake angle, and a relatively large shear angle, compression of the chip and deformation are reduced, and chip Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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flow occurs readily across the rake surface. With a lower tool rake angle (even negative values), the shear angle is reduced, the degree of deformation and area are increased and the chip becomes thicker.3,19–21 In the turning of wrought steel, typical rake angles are +8° to +20° with HSS tools and small angles on HM inserts and on other hard tool materials attached to massive toolholders, depending on the workpiece material properties, operation, and cutting conditions. The rake angle in the work described here was +5°.3,19,22 A more-detailed view of the cutting process is depicted in Figure 12. The first position in Figure 12(b) is position 1 in Figure 11 (initiation of cutting on the surface plane adjusted by the depth of cut) and is characterized by the positive tool rake angle +γn, by the shear angle φ, here termed φmax, by the depth of cut h (thickness of the undeformed chip), and by the contact-length chip/tool interface l1. According to Figure 11, during cutting of a porous material, which a priori exhibits lower apparent hardness, and thus is softer than pore-free material of the same composition, the workpiece material is compressed to some depth in the bulk material at the expense of pores and densified by the moving cutting edge (position 1→ position 2). Consequently, the positive rake angle is continuously lowered, in some cases even to negative values (apparently adapting to the machined surface that is formed), the actual angle being affected by work hardening of the machined surface at a sufficient reduction of the shear angle. Compression of the workpiece material to some depth by the edge is limited by the amount of plastic deformation of the matrix material, namely, the amount of plastic deformation manifested in work hardening in cutting. A lowering of the rake angle occurs from an initial positive value +γn to γn = 0° at position 4 in Figure 11. This results in a continuous decrease of the moving edge into the bulk material (i.e., deepening of the feed mark) due to increasing plastic deformation of the deformation zone with attendant work hardening. This condition, shown in Figure 12(a) (thick contours), corresponds to the maximum depth of the feed mark formed ∆hmax. It is characterized by a reduction of the shear angle to φmin, by the increased length of the chip/tool interface l 2 , and by the larger chip thickness h 2 at the constant value h. Consequently, as shown, due to the reduced shear angle, the area of heterogeneous plastic Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
deformation increases and the chip becomes thicker with attendant increase in cutting forces. The changes in the thickness of the chip due to the changes in the shear angle are known and analyzed in the cutting theory.18,19,22,23 The workhardened areas form ahead of the cutting edge and in the machined surface layer under the tool clearance face. This deformation state results in maximum work hardening of the workpiece material ahead of the edge tip at the highest chip thickness and the maximum length of the chip/tool interface. Consequently, in this position of the cutting edge, equilibrium between the cutting force and the cutting resistance is attained approximately in the center of the feed mark. It means that further cutting under these geometric conditions and state is not possible. Continuation of cutting occurs in such a way that the moving tool is pushed back from the work-hardened workpiece material under the effect of the radial component of the cutting force from the deepest feed mark to the surface level defined by the adjusted depth of cut, in this case 0.1 mm. This occurs through the reverse change of the rake angle from 0° to the positive value +γn with a corresponding increase of the shear angle from fmin to fmax. The change in the shear angle during one revolution in the range φmax > φmin < φmax characterizes the fluctuations in the dynamic deformation intensity ahead of the edge, and in fluctuations of the cutting force. This finally results in a continuous decrease in work hardening to a level that permits a further cut in the starting direction on the surface plane, with a corresponding decrease in shearing force, i.e., repeating the process during the next revolution. This corresponds to position 7 (= position 1) shown in Figures 11 and 12(b). Figure 12(a) shows, by superposition of the two states, the mutual effect of rake and shear angle changes on the deformation process occurring ahead of the edge, resulting in a corresponding decrease or increase of the deformation zone area, chip thickness, and contact-length chip/rake face. Simultaneously, the changes in cutting geometry also affect the contact length of the chip/tool interface which, in all cases, is larger than that at the starting “optimum” positive rake angle (l2 > l1); this results in increased friction up to seizure. The increased cutting force required during the entire process causes an increase in the temperature of the cutting zone and, in turn, at the
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chip/tool interface, causing softening of the workpiece material ahead of the cutting edge. The consequences of extensive plastic deformation in porous soft sintered steels ahead of the edge are of the same character as those in ductile wrought materials. They permit extensive plastic deformation of the chip during cutting, which increases heat generation and temperature, and also results in longer, continuous chips that remain in contact with the tool face for a longer time period, thus causing more frictional heat.19 From the mechanism presented for the adverse effect of porosity it can be deduced that the cutting of porous material, characterized by wave feed marks on the machined surface, is performed by tool edges with rake angles lower than the starting (+) angle. It follows that the reason for formation of the wave feed marks on the machined surface during each feed is the lower resistance to plastic deformation of the porous soft sintered material. Materials with higher levels of porosity exhibit lower resistance, and vice versa, and consequently either inferior or superior machinability, taking into account the effect of alloying on the apparent hardness, Figures 2 and 3. For this reason, PM steels exhibit enhanced machinability at lower levels of porosity and thus at a higher apparent hardness. This is in marked contrast to the machinability of wrought steels, the performance of which decreases with increasing hardness: the stronger the material, the higher the resistance to shear stress.24 The mechanism presented here explains not only the adverse effect of porosity on the machinability of PM materials, but also their inferior surface finish. Furthermore, the elimination of porosity ahead of the cutting edge was confirmed, which means that the edge neither enters into the porous material nor suffers intermittent cutting. From these results, it can also be deduced that measures for avoiding plastic deformation in cutting a sintered material with a geometrically defined cutting edge improve machinability. Work hardening of the machined surfaces caused by the cutting process was confirmed by the apparent hardness of the as-machined surface (Figure 6), and by the feed-mark subsurface microindentation hardness values (Table III). The formation of a thin white layer on the machined surface is further proof of plastic deformation and increased temperature in the cutting zone fol-
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lowed by dynamic recrystallization24 caused by the cutting force together with the change of cutting geometry. The formation of a thin white layer was also observed on as-machined surfaces of hard PM materials, regardless of the level of porosity.25 Instability in cutting, manifested as a uniform wave movement of the edge during the feed, and the corresponding fluctuations of the cutting force at less than optimum rake angles and shear angles due to porosity, results in higher cutting forces for machining of a sintered steel compared with a wrought steel. The consequence of these factors resulting from the character of the cutting process of porous sintered steels is an increase in tool wear, inferior machinability, and inferior surface finish of the workpiece. The formation of feed marks in the profile is a result of the interacting effect of the variable cutting force and the resistance to plastic deformation of the porous material. The cutting-edge wave movement occurs at the expense of elastic deformation of the toolholding system. It is possible that the cyclic change of the tool rake angle and the resulting wave-shaped feed marks may be alleviated, to some extent, by higher rigidity of the toolholding system. The changes in the tool rake angle and the formation of the feed marks were also observed in green face turning of Distaloy DH1-0.5 w/o C with a density of 7.28 g/cm3.26 In agreement with previous results, an improvement in machinability was attained with 0.5 w/o C compared with 0.3 w/o C, due to higher resistance to defor mation as a result of increased hardness (Figure 4, Table II). With respect to machinability, the carbon content in a steel is limited by the fraction of pearlite in the microstructure, which contains hard cementite. This explains why 0.5–0.6 w/o C is cited as optimal for good machinability of sintered Fe-C steels.3 The recommended addition of 0.5 w/o Mn,3,27 and small amounts of phosphorous or copper 25 to improve the machinability of PM steels, can also be explained by an increase in the resistance to deformation. However, the effect of different alloying elements on the cutting process of sintered steels should be investigated in more detail. The formation of a BUE during cutting of PM steels is regarded by the authors as a further factor adversely affecting machinability. This is part of the chip-formation process in the cutting of Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
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FACE TURNING OF PM STEELS: EFFECT OF POROSITY AND CARBON LEVEL
metallic materials, not a separate phenomenon. In PM machining, BUE formation is, to a large extent, affected a priori by cutting force fluctuations due to cutting-edge oscillations, i.e., it is linked to the mechanism of formation of the wave feed marks on the machined surface. As shown in Figure 9, in the turning of sintered porous steels BUEs are formed in the area of extensive plastic deformation in the primary flow zone ahead of the edge tip. This means that, due to porosity, the BUE-formation process in sintered material extends further into the primary deformation zone (Figures 9(a), (e), (f), and (h). It is possible that this results in the variable size and high work hardening of the BUE, because at lower shear angles the area of intense plastic deformation increases, and vice versa. 17 The presence of a BUE further changes the shear angle, causing increased instability in the chip for ming process. 20 For these reasons formation of the BUE in the cutting of sintered steels is enhanced by the higher temperature (and resulting softening of the workpiece material) in the cutting zone compared with the cutting of wrought steel. In contrast, the zone of intense plastic deformation in wrought steel machining is formed on the back surface of the chip which, under the cutting conditions used, transforms into a BUE, as shown in Figure 9(i). The adverse effect of BUEs on the machinability of PM steels is also a consequence of the modified cutting-edge geometry. A new large, stubby (geometrically not definable) cutting edge of lower hardness than the tool is formed: in consequence, the cut can occur over longer or shorter time periods. This geometry of the cutting edge, which substitutes for the function of the cutting edge at the detachment of the chip, results in extremely high cutting forces, and in increased friction and temperature. These further accelerate tool wear and damage. The BUEs may adhere to the tool edge at high temperatures in the cutting zone (at ~300°C in cutting wrought steel),18,23 or they frequently break away and are taken off by the chips, or are compressed and smeared on the machined surface by the tool clearance face, adversely affecting the surface finish. If the BUEs adhering to the cutting tool are not observed after cutting, this cannot be taken as proof that they do not form intermittently during the operation. Figure 9 also shows that the chips formed in cutting porous materials are discontinuous and short. This Volume 44, Issue 2, 2008 International Journal of Powder Metallurgy
means that they remove a lower fraction of the heat formed in the process, which increases the temperature of the tool. The intense plastic deformation and resulting work hardening of the BUEs formed in cutting was confirmed by the microindentation hardness measurements. The maximum hardness values, which can be regarded as extreme, delineated by the indentations (white or black) in Figure 9 on the BUEs were: (a) 827, (b) 891, (d) 1,006, (e) 1,105, (f) 1,128, and (g) 831 HV 0.025. The values are much higher than those for wrought steel, namely, 627 HV 0.025 in Figure 9(i), and ~500 HVN.18 The formation of BUEs was observed in the turning of sintered Fe-C steels at porosity levels >5 v/o,28 but without a detailed investigation. In general, the formation of large BUEs was observed in machining soft PM materials resulting in a poor surface finish,29 and in the turning of Fe-1.5 w/o Cu-0.7 w/o C steel at a cutting speed of 60 m/min (197 ft./min).30 The formation of BUEs was also observed in the face turning of hardened Fe-0.8 w/o Ni-2 w/o Cu-0.8 w/o C steel at a cutting speed of 305 m/min (1000 ft./min),31 and prealloyed Fe-3 w/o Cr-0.5 w/o Mo-0.4 w/o C steel at a cutting speed of 300 m/min (984 ft./min).32 Thus, BUE formation in PM machining also occurs at cutting speeds >200 m/min (656 ft./min). This speed is considered an approximate upper limit for the formation of BUEs in wrought steel machining.23 Conclusively our results and analysis show that the cost effectiveness of the cutting process in PM machining can be improved by an appropriate increase in resistance to plastic deformation of the sintered matrix. This must be considered when adjusting the cooling conditions after sintering, and by optimal selection of cutting parameters such as cutting speed, feed rate, depth of cut, cutting tool material, and geometry. CONCLUSIONS 1. The adverse effect of porosity on the machining of PM steels is related to two mechanisms: • the cutting process resulting in the formation of deep feed marks on the machined surface which exhibit uniform waves • the formation of BUEs 2. The reason for the formation of uniform wave turning-feed marks on the machined surface is the lower resistance to plastic deformation of
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the porous workpiece in sintered material compared with pore-free material. Higher porosity in sintered steels results in lower machinability due to a lower resistance to plastic deformation, and vice versa. The same holds true for carbon content. A 0.5 w/o C content increases the apparent hardness of the alloys and increases the resistance to plastic deformation compared with a 0.3 w/o C content. In cutting a porous material, it is compressed to some depth into the bulk workpiece, ahead of the moving cutting edge, at the expense of pores, and is continuously plastically deformed. Increasing plastic deformation during the feed results in increased resistance to the cut and in a reduction of the tool rake angle and the shear angle, due to the change in the direction of tool movement. The characteristic wave-shaped machined surface of a porous material is the result of the interaction between the cutting force and the resistance to plastic deformation of the porous workpiece material. The cutting process occurs as a result of cutting force fluctuations caused by the wave motion of the cutting edge. The formation of wave feed marks on the machined surface also explains, in part, the inferior surface finish created in the turning of sintered materials. The BUEs formed in cutting are characterized by extensive plastic deformation. This occurs ahead of the cutting-edge tip at the expense of pores due to lower resistance to plastic deformation of the workpiece material. The results indicate that BUE formation occurs in the turning of all sintered materials despite the fact that they do not always adhere to the tool permanently. The pores also contribute to detachment of the BUEs from the bulk workpiece material and from the chip. Plastically deformed areas formed ahead of the cutting tool in a porous material observed in quick-stop experiments prove that the cutting edge always enters pore-free work-hardened material. The cutting process in sintered steels that forms the wave feed marks on machined surfaces and the BUEs requires higher cutting forces together with increased friction up to seizure, which generates excessive heat in the cutting zone. Furthermore, the BUEs adversely change the cutting-edge geometry and the tribological conditions at the chip/tool interface,
coupled with a significant increase in temperature. The combination of these factors result in increased tool wear when cutting sintered materials and an inferior surface finish. 7. The effect of porous heterogeneous microstructures on the machinability of sintered steels should be studied further with a particular focus on the extensive plastic deformation that occurs ahead of the cutting edge. ACKNOWLEDGEMENTS This work was supported by the Scientific Grant Agency of ME SR and SAS (Grants No. 2/6209/26, 2/5144/25 and 1/3225/06). REFERENCES 1. J. Capus, “Machined PM Parts: Does the ‘X-Factor’ Hold Key to Success?”, Metal Powder Report, 2004, vol. 59, no. 12, pp. 24–28. 2. P.O. Larsson, “Additives for Machining”, Machinability and Wear, Höganäs Chair Seminar 2006, CD IMR SAS, Kos˘ice and Höganäs AB, Höganäs, Sweden, 2006. pp. 16–21. ˘ alak, M. Selecká and H. Danninger, Machinability of 3. A. S Powder Metallurgy Steels, Second Edition, 2006, Cambridge International Science Publishing, Cambridge, UK. ˘ alak, H. Danninger and M. Selecká, “Factors Affecting 4. A. S Machinability of Sintered Steels”, Proc. PM Auto 05. Int. Congress on PM for Automotive Parts, compiled by A. Arvand and H. Danninger, Isfahan, Mashad Powder Metallurgy, Tehran, 2005, pp. 12–24. 5. D. Raiser, “The Ef fect of Microstructure on the Machinability of P/M Steel”, Int. J. Powder Metall., 2005, vol. 41, no. 4, p. 19. 6. E. Chandler, “Machining of Powder Metallurgy Materials”, ASM Handbook Volume 16: Machining, ASM, Metals Park, Ohio, 1990, pp. 799–891. 7. R.J. Causton, “Machinability of P/M Steels”, Advances in Powder Metallurgy & Particulate Materials, compiled by M. Phillips and J. Porter, Metal Powder Industries Federation, Princeton, NJ, 1995, vol. 2, part 8, pp. 149–170. 8. J.S. Agapiou and M.V. DeVries, “Machinability of Powder Metallurgy Materials”, Int. J. Powder Metall., 1988, vol. 24, no. 1, pp. 47–57. 9. G.A. Chadwick and F.P. Kehoe, “Machinability of Sintered Ferrous Powder Metallurgy Compacts”, 1999, Hi-Tech Metals R & D Ltd. Report, Chandlers Ford, Eastleigh, UK. 10. A. S˘alak, K. Vasilko, and M. Selecká, “Machining of PM Materials by Drilling with Constant Thrust Force”, Proc. Int. Conf. on Deformation and Fracture in Structural PM Materials DF PM 2002, compiled by L’. Parilák and H. Danninger, Stará Lesná, Institute of Materials Research of Slovak Academy of Sciences, Kos˘ice, Slovakia, 2002, vol. 1, pp. 143–151. 11. J.S. Agapiou, G.W. Halldin and M.F. DeVries, “Effect of Porosity on the Machinability of P/M 304L Stainless Steel”, Int. J. Powder Metall., 1989, vol. 25, no. 2, pp. 127–137. 12. S.A. Kvist, “Turning and Drilling of some Typical Sintered
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13. 14.
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20. 21. 22. 23.
Steels”, Powder Metallurgy, 1969, vol.12, no. 24, pp. 538–565. A. S˘ alak, Ferrous Powder Metallurgy, 1995, Cambridge International Science Publishing, Cambridge, UK. ˘ alak, K. Vasilko, M. Selecká and H. Danninger, “New A. S Short T ime Face Tur ning Method for Testing the Machinability of PM Steels”, J. Materials Processing Technology, 2006, no. 176, pp. 62–69. A. S˘ alak, K. Vasilko, H. Danninger, M. Selecká and D. Jakubéczyová, “Effect of Porosity on Machinability of PM Steels Determined by Face Turning Test”, Proc. EURO PM2005, European Powder Metallurgy Association, Shrewsbury, UK, 2005, vol. 3, pp. 205–210. A. S˘ alak, K. Vasilko, H. Danninger, M. Selecká and D. Jakubéczyová, “Effect of Cutting Speed and Tool Grade on Machinability of PM Steels Determined by Face Turning Method”, Powder Metallurgy Progress, 2005, vol. 5, no. 2, pp. 104–114. E.M. Trent, “Metal Cutting and the Tribology of Seizure: I, Seizure in Metal Cutting”, Wear, 1988, vol. 128, pp. 29–45. E.M. Trent, Metal Cutting, 1977, Butterworth & Co, London, UK. J.T. Black, “Introduction to the Machining Process”, Metals Handbook, Machining, Ninth Edition, ASM, Metals Park, Ohio, 1989, vol. 16, pp. 121–135. L.A. Kendall, “Tool Wear and Tool Life”, ibid. reference no. 19, pp. 37–48. H. Weber and T.N. Loladze, Fundamentals of Cutting, VEB Verlag Technik Berlin, Berlin, 1986 (in German). G. Schneider, Jr., Cutting Tool Applications, George Schneider, Jr., Farmington Hills, MI, 2002. K. Vasilko, Analytic Theory of Cutting Process, Technical University Kos˘ice, Faculty of Manufacturing Technologies, Pres˘ov, Slovakia, 2007 (in Slovak).
24. L. Seongey, J. Hwang, R. Shankar, S. Chandrasekar and W.D. Compton, “Large Strain Deformation Field in Machining”, Metallurgical and Materials Transactions, 2006, vol. 27A, pp. 1,622–1,641. 25. M. Gagné and J. Poirier, “Key Parameters for Machining P/M Materials”, Proc. Int. Symposium on Recent Processes in Iron Powder Metallurgy, Tohoyoku University, Sendai, Japan, 1992, pp. 1–17. 26. A. Benner and P. Beiss, “Green Tur ning of War m Compacted PM Steels”, Proc. EURO PM2001, European Powder Metallurgy Association, Shrewsbury, UK, 2001, vol. 3, pp. 59–66. 27. J.M. Capus and C. Fournel, “The Influence of Mix Composition on the Machinability of Sintered Iron”, Modern Developments in Powder Metallurgy, Metal Powder Industries Federation, Princeton, NJ, 1981, vol. 13, pp. 137–145. 28. Y.G. Dorofeyev, Dynamic Hot Pressing of Porous Powder Compacts, Metallurgia, Moscow, Russian Fed. Rep., 1977 (in Russian). 29. J. Hazra, K. Taraman, F.M. Kennedy and F. Badia, “Machinability of Powder Metallurgy Components—A Look at Surface Finish”, Modern Developments in Powder Metallurgy, compiled by H.H. Hausner and P.W. Taubenblat, Metal Powder Industries Federation, Princeton, NJ, 1976, vol. 10, pp. 385–402. 30. N. Usuki, “Machinability of Sintered Steels”, Metal Powder Report, 1996, vol. 51, no. 11, November, pp. 839–843. 31. Z. Zurecki, R. Ghosh and J. H. Frey, “Finish-Turning of Hardened Powder Metallurgy Steel Using Cryogenic Cooling”, Int. J. Powder Metallurgy, 2004, vol. 40, no. 1, pp. 19–31. 32. S. Berg, “Machining of Bainitic and Martensitic Chromium Pre-alloyed PM Steel”, ibid. reference no. 15, pp. 211–216. ijpm
MACHINING & MACHINABILITY OF PM COMPONENTS Compiled by: Christopher P. Adams, Carl Blais and Deepak S. Madan
Available only on a searchable CD-ROM Table of Contents MPIF PUBLICATION, 523 pages, 2006, ISBN 0-9762057-5-0 Item # 3011CD Machining & Machinability of PM Components CD-ROM List Price $140 APMI $126 MPIF Co. $114
Machining & Machinability of PM Components is a collection of technical literature on PM machining and machinability. The publication is a review of the current knowledge on the machining of PM components containing 37 technical papers from sources such as the annual MPIF International Conference on Powder Metallurgy & Particulate Materials, the International Journal of Powder Metallurgy, EPMA World Congress & Exhibition, Powder Metallurgy, and SAE International. The selected papers contribute to identifying the particularities of machining PM components in terms of base material, additives for improved machinability, machining process parameters and case studies. To Order: Fax: (609) 987-8523, Phone: (609) 945-0009, E-mail:
[email protected] or vist us online at www.mpif.org
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MEETINGS AND CONFERENCES
2008 HIP ’08—THE 9TH INTERNATIONAL CONFERENCE ON HOT ISOSTATIC PRESSING May 6–9 Huntington Beach, CA www.hip2008.com 2008 WORLD CONGRESS ON POWDER METALLURGY & PARTICULATE MATERIALS June 8–12 Washington, DC MPIF* 2008 INTERNATIONAL CONFERENCE ON TUNGSTEN, REFRACTORY & HARDMATERIALS VII June 8–12 Washington, DC MPIF* 19TH AEROMAT CONFERENCE & EXPOSITION June 23–26 Austin, TX www.asminternational.org/ae romat
INTERNATIONAL CONFERENCE ON ALUMINUM ALLOYS September 22–26 Aachen, Germany www.dgm.de EURO PM2008 September 29–October 1 Mannheim, Germany www.epma.com/pm2008 MATERIALS SCIENCE & TECHNOLOGY 2008 CONFERENCE & EXHIBITION October 5–9 Pittsburgh, PA www.matscitech.org/2008/ho me.html 5TH INTERNATIONAL POWDER METALLURGY CONFERENCE October 8–12 Ankara, Turkey www.turkishpm.org/5pm2008
BASIC PM SHORT COURSE July 21–23 State College, PA MPIF*
SINTERING 2008 November 16–20 La Jolla, CA www.ceramics.org/sintering2008
PM SINTERING SEMINAR September TBA MPIF*
PMP III THIRD INTERNATIONAL CONFERENCE—PROCESSING MATERIALS FOR PROPERTIES December 7–10 Bangkok, Thailand www.tms.org/meetings/ specialty/pmp08
5TH INTERNATIONAL CONFERENCE ON ADVANCED MATERIALS AND PROCESSING September 3–6 Harbin, China icamp.hit.edu.cn/
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SUPERALLOYS 2008 September 14–18 Champion, PA www.tms.org/Meetings/specialty/superalloys2008/ho me.html
2009 POWDERMET2009: MPIF/APMI INTERNATIONAL CONFERENCE ON POWDER METALLURGY & PARTICULATE MATERIALS June 28–July 1 Las Vegas, NV MPIF* THERMEC 2009: SIXTH INTERNATIONAL CONFERENCE ON ADVANCED MATERIALS AND PROCESSES August 25–29 Berlin, Germany SDMA 2009/ICSF VII—4TH INTERNATIONAL CONFERENCE ON SPRAY DEPOSITION AND MELT ATOMIZATION/7TH INTERNATIONAL CONFERENCE ON SPRAY FORMING September 7–9 Bremen, Germany www.sdma-conference.de/
2010 POWDERMET2010: MPIF/APMI INTERNATIONAL CONFERENCE ON POWDER METALLURGY & PARTICULATE MATERIALS June 27–30 Hollywood (Ft. Lauderdale), FL MPIF* PM2010 WORLD CONGRESS October 10–14 Florence, Italy
*Metal Powder Industries Federation 105 College Road East, Princeton, New Jersey 08540-6692 USA (609) 452-7700 Fax (609) 987-8523 Visit www.mpif.org for updates and registration. Dates and locations may change
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ADVERTISERS’ INDEX
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ACE IRON & METAL CO. INC. ________________(269) 342-0185 ______________________________________________6 ACUPOWDER INTERNATIONAL, LLC ___________(908) 851-4597______www.acupowder.com______________________39 ARBURG GmbH + Co KG ____________________(860) 667-6522______www.arburg.com _________________________12 BÖHLER UDDEHOLM _______________________(603) 883-3101______www.bucorp.com __________________________8 ELNIK SYSTEMS ___________________________(973) 239-6066______www.elnik.com___________________________10 HENKEL TECHNOLOGIES ____________________(315) 637-3086______www.henkel.com _________________________32 HOEGANAES CORPORATION _________________(856) 786-2574______www.hoeganaes.com______INSIDE FRONT COVER INCO SPECIAL PRODUCTS ___________________(201) 848-1022______www.incosp.com __________________________4 NORILSK NICKEL __________________________(+ 7 495) 785 58 08 __www.norilsknickel.com ____________________11 NORTH AMERICAN HÖGANÄS INC. ____________(814) 479-2003______www.nah.com_____________INSIDE BACK COVER QMP ____________________________________(734) 953-0082______www.qmp-powders.com ___________BACK COVER SCM METAL PRODUCTS, INC.________________(919) 544-7996______www.scmmetals.com_______________________3
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