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© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
www.asminternational.org
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ASM Handbook Volume 15 Casting Prepared under the direction of the ASM International Handbook Committee
Editorial Committee Srinath Viswanathan, University of Alabama, Chair Diran Apelian, Worcester Polytechnic Institute Raymond J. Donahue, Mercury Marine Babu DasGupta, National Science Foundation Michael Gywn, ATI Inc John L. Jorstad, J.L.J. Technologies, Inc Raymond W. Monroe, Steel Founders’ Society of America Mahi Sahoo, CANMET Materials Technology Laboratory Thomas E. Prucha, American Foundry Society Daniel Twarog, North American Die Casting Association Steve Lampman, Project Editor Charles Moosbrugger, Editor Eileen DeGuire, Editor Madrid Tramble, Senior Production Coordinator Ann Britton, Editorial Assistant Diane Whitelaw, Production Coordinator Kathryn Muldoon, Production Assistant Scott D. Henry, Senior Product Manager Bonnie R. Sanders, Manager of Production
Editorial Assistance Elizabeth Marquard Heather Lampman WPR Indexing Service
Materials Park, Ohio 44073-0002 www.asminternational.org
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
www.asminternational.org
Copyright # 2008 by ASM InternationalW All rights reserved No part of this book may be reproduced, stored in a retrieval system, or transmitted, in any form or by any means, electronic, mechanical, photocopying, recording, or otherwise, without the written permission of the copyright owner. First printing, December 2008
This book is a collective effort involving hundreds of technical specialists. It brings together a wealth of information from worldwide sources to help scientists, engineers, and technicians solve current and long-range problems. Great care is taken in the compilation and production of this Volume, but it should be made clear that NO WARRANTIES, EXPRESS OR IMPLIED, INCLUDING, WITHOUT LIMITATION, WARRANTIES OF MERCHANTABILITY OR FITNESS FOR A PARTICULAR PURPOSE, ARE GIVEN IN CONNECTION WITH THIS PUBLICATION. Although this information is believed to be accurate by ASM, ASM cannot guarantee that favorable results will be obtained from the use of this publication alone. This publication is intended for use by persons having technical skill, at their sole discretion and risk. Since the conditions of product or material use are outside of ASM’s control, ASM assumes no liability or obligation in connection with any use of this information. No claim of any kind, whether as to products or information in this publication, and whether or not based on negligence, shall be greater in amount than the purchase price of this product or publication in respect of which damages are claimed. THE REMEDY HEREBY PROVIDED SHALL BE THE EXCLUSIVE AND SOLE REMEDY OF BUYER, AND IN NO EVENT SHALL EITHER PARTY BE LIABLE FOR SPECIAL, INDIRECT OR CONSEQUENTIAL DAMAGES WHETHER OR NOT CAUSED BY OR RESULTING FROM THE NEGLIGENCE OF SUCH PARTY. As with any material, evaluation of the material under end-use conditions prior to specification is essential. Therefore, specific testing under actual conditions is recommended. Nothing contained in this book shall be construed as a grant of any right of manufacture, sale, use, or reproduction, in connection with any method, process, apparatus, product, composition, or system, whether or not covered by letters patent, copyright, or trademark, and nothing contained in this book shall be construed as a defense against any alleged infringement of letters patent, copyright, or trademark, or as a defense against liability for such infringement. Comments, criticisms, and suggestions are invited, and should be forwarded to ASM International. Library of Congress Cataloging-in-Publication Data ASM International ASM Handbook Includes bibliographical references and indexes Contents: v.1. Properties and selection—irons, steels, and high-performance alloys—v.2. Properties and selection—nonferrous alloys and special-purpose materials—[etc.]—v.21. Composites 1. Metals—Handbooks, manuals, etc. 2. Metal-work—Handbooks, manuals, etc. I. ASM International. Handbook Committee. II. Metals Handbook. TA459.M43 1990 620.1’6 90-115 SAN: 204-7586 ISBN-13: 978-0-87170-711-6 ISBN-10: 0-87170-711-X
ASM InternationalW Materials Park, OH 44073-0002 www.asminternational.org Printed in the United States of America Multiple copy reprints of individual articles are available from Technical Department, ASM International.
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
www.asminternational.org
Dedicated to the memory of JOSEPH R. DAVIS (1954–2007) Editor, Metals Handbook, 1982–1991
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
www.asminternational.org
Foreword
In this revision of ASM HandbookW Volume 15 on casting, ASM International is indebted to the volunteer efforts of the Volume 15 Editorial Committee and over 150 participants who helped as authors or reviewers. Their professional commitment and efforts represent a continuing devotion to the practice of metalcasting and the publication of peer consensus information on it. Special thanks are extended to Srinath Viswanathan, University of Alabama, for recruiting an outstanding Editorial Committee with Diran Apelian, Worcester Polytechnic Institute; Babu DasGupta, National Science Foundation, Raymond J. Donahue, Mercury Marine; Michael Gywn, ATI Inc.; John L. Jorstad, J.L.J. Technologies, Inc.; Raymond W. Monroe, Steel Founders’ Society of America; Thomas E. Prucha, American Foundry Society; Kumar Sadayappan and Mahi Sahoo, CANMET Materials Technology Laboratory; Edward S. Szekeres, Casting Consultants Incorporated; and Daniel Twarog, North American Die Casting Association. We thank them and the other contributors for this publication.
Dianne Chong President ASM International Stanley C. Theobald Managing Director ASM International
iv
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
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Policy on Units of Measure
By a resolution of its Board of Trustees, ASM International has adopted the practice of publishing data in both metric and customary U.S. units of measure. In preparing this Handbook, the editors have attempted to present data in metric units based primarily on Syste`me International d’Unite´s (SI), with secondary mention of the corresponding values in customary U.S. units. The decision to use SI as the primary system of units was based on the aforementioned resolution of the Board of Trustees and the widespread use of metric units throughout the world. For the most part, numerical engineering data in the text and in tables are presented in SI-based units with the customary U.S. equivalents in parentheses (text) or adjoining columns (tables). For example, pressure, stress, and strength are shown both in SI units, which are pascals (Pa) with a suitable prefix, and in customary U.S. units, which are pounds per square inch (psi). To save space, large values of psi have been converted to kips per square inch (ksi), where 1 ksi = 1000 psi. The metric tonne (kg 103) has sometimes been shown in megagrams (Mg). Some strictly scientific data are presented in SI units only. To clarify some illustrations, only one set of units is presented on artwork. References in the accompanying text to data in the illustrations are presented in both SI-based and customary U.S. units. On graphs and charts, grids corresponding to SI-based units usually appear along the left and bottom edges. Where appropriate, corresponding customary U.S. units appear along the top and right edges. Data pertaining to a specification published by a specification-writing group may be given in only the units used in that specification or in dual units, depending on the nature of the data. For example, the typical yield strength of steel sheet made to a specification written in customary U.S.
units would be presented in dual units, but the sheet thickness specified that specification might be presented only in inches. Data obtained according to standardized test methods for which the standard recommends a particular system of units are presented in the units of that system. Wherever feasible, equivalent units are also presented. Some statistical data may also be presented in only the original units used in the analysis. Conversions and rounding have been done in accordance with IEE/ ASTM SI-10, with attention given to the number of significant digits in the original data. For example, an annealing temperature of 1570 F contains three significant digits. In this case, the equivalent temperature would be given as 855 C; the exact conversion to 854.44 C would not be appropriate. For an invariant physical phenomenon that occurs at a precise temperature (such as the melting of pure silver), it would be appropriate to report the temperature as 961.93 C or 1763.5 F. In some instances (especially in tables and data compilations), temperature values in C and F are alternatives rather than conversions. The policy of units of measure in this Handbook contains several exceptions to strict conformance to IEEE/ASTM SI-10; in each instance, the exception has been made in an effort to improve the clarity of the Handbook. The most notable exception is the use of g/cm3 rather than kg/m3 as the unit of measure for density (mass per unit volume). SI practice requires that only one virgule (diagonal) appear in units formed by combination of several basic units. Therefore, all of the units preceding the virgule are in the numerator and all units following the virgule are in the denominator of the expression; no parentheses are required to prevent ambiguity.
v
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
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Preface
From the preceding Volume 5 (1970) of the 8th Edition Metals Handbook and the 9th Edition Metals Handbook Volume 15 (1988) on casting, this Handbook provides an update on the continuing advances in casting technologies and applications. Casting, as both a science and practical tool in art and technology, is enormously varied in scope. It is impossible to capture the full scope of casting technology in one volume.
on sand casting is expanded relative to the previous edition with the intention of providing a reference that may be helpful as a communication tool between product designers and metalcasters in developing successful and economical products. This Volume consists of 18 sections. The first section introduces the historical development of metal casting, as well as to the advantages of castings over parts produced by other manufacturing processes, their applications, and the current market size of the industry. This includes an article on “Metalcasting Technology and the Purchasing Process” written by Al Spada and the technical staff of the American Foundry Society. Then, the principles and practice of melt processing are described in the next three sections followed by a section on principles of solidification including nucleation kinetics, fundamentals of growth, transformation behavior, and microstructure development. Solid-state processing of casting, such as heat treat treatment and hot isostatic pressing, are also introduced. This is followed by a section on the “Modeling and Analysis of Casting Processes.”
The main focus of this Volume is on the products and processes of foundry (shape) casting, although primary (ingot or continuous casting) of steel and aluminum are also covered. In addition, continuous casting of copper is described in an article, as copper continuous casting was a precursor to steel and aluminum continuous casting in some respects. Some of the articles on melt processing, such as the articles on “Electric Arc Furnace Melting” and “Steel Melt Processing,” also briefly describe primary production of cast metal. Shape casting of metal is dominated by cast iron, which constitutes just over 70% of the worldwide production of castings on a tonnage basis (See Table 2 in the first article “History and Trends of Metal Casting”). This is followed by steel, copper-alloy, and aluminum-alloy castings, which make up about 25% on the worldwide tonnage of casting production. Magnesium and zinc are on the order of 1% or less. The dominance of just a few alloys in shape casting is due to the fact that successful and economic shape casting typically involves alloy compositions near a eutectic. The lower melting points and narrower freezing range of near-eutectic compositions promote better castability.
Like the previous edition, traditional subjects such as patterns, molding and casting processes, foundry equipment, and processing considerations are extensively covered in the next sections. As noted, coverage on sand casting has been consolidated and expanded. For example, the major method of shell molding is described in an article—based on an update of a still largely valid 1970 handbook (Volume 5) article. New updates are also provided on processes growing in use, such as squeeze casting, lost-foam casting, semisolid metal forming, and low-pressure casting. The latter is particularly important in producing quality products, as described by John Campbell in the article “Filling and Feeding Concepts.”
Since the 1988 edition of Volume15, several developments have occurred (see Table 5 in the first article “History and Trends of Metal Casting”). Of these developments, computer technology continues to shorten development time and help simulate the casting process. Automation and robotic technology also has improved the productivity and process control of casting. In terms of processes, semisolid processing, squeeze casting, lost-foam, vacuum molding, and various dies casting technologies continue to improve and finds new applications. These important topics are updated in this Volume. For nonferrous alloys, high-pressure die casting of aluminum is a major area of expansion and update in this volume.
Finally, the last five sections describe the major types of cast alloys in term of processing and the properties and characteristics of cast ferrous and nonferrous alloys. Emphasis is placed on cast iron, cast steel, aluminum, copper, and zinc. The last section covers the quality aspects of cast products and the processing of castings. It is hoped that this Handbook is a useful work of peer-consensus reference information for the producers, designers, and buyers of castings. Many thanks are extended to all the contributors and the editors who worked on this Volume. This publication would not have been possible without their commitment and effort.
In addition, coverage on sand casting is expanded and consolidated in this Volume with major articles on “Green Sand Molding,” No-Bake Sand Molding,” and “Shell Molding and Shell Coremaking.” Bonded sand mold casting, although well-established for many years, is the most widely used method of casting on a tonnage basis. Improvement in methods and materials continue to provide better yields, productivity, and product quality. The sand system is also a major factor in the economics of large-volume, production casting. Coverage
Srinath Viswanathan University of Alabama Chair, Volume 15 Editorial Committee
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© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
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Officers and Trustees of ASM International (2007–2008)
Dianne Chong President and Trustee The Boeing Company Roger J. Fabian Vice President and Trustee Bodycote Thermal Processing Lawrence C. Wagner Immediate Past President and Trustee Texas Instruments Inc (retired) Paul L. Huber Treasurer and Trustee Seco/Warwick Corporation Stanley C. Theobald Secretary and Managing Director ASM International
Trustees
Sue S. Baik-Kromalic Honda of America Christopher C. Berndt Swinburne University of Technology Brady G. Butler University of Utah Danniel P. Dennies The Boeing Company Leigh C. Duren Worcester Polytechnic Institute Pradeep Goyal Pradeep Metals Ltd.
Digby D. Macdonald Penn State University Subhash Mahajan Arizona State University Charles A. Parker Honeywell Aerospace Megan M. Reynolds Washington State University Mark F. Smith Sandia National Laboratories Jon D. Tirpak ATI
Members of the ASM Handbook Committee (2007–2008) Larry D. Hanke (Chair 2006–; Member 1994–) Materials Evaluation and Engineering Inc. Kent L. Johnson (Vice Chair 2006–; Member 1999–) Engineering Systems Inc. Viola L. Acoff (2005–) University of Alabama David E. Alman (2002–2008) National Energy Technology Laboratory Tim Cheek (2004–) DELTA (v) Forensic Engineering Lichun Leigh Chen (2002–) Technical Materials Incorporated Sarup K. Chopra (2007–) Nook Industries Craig Clauser (2005–) Craig Clauser Engineering Consulting Incorporated Craig V. Darragh (1989–)
The Timken Company Jon L. Dossett (2006–) Consultant David U. Furrer (2006–) Rolls-Royce Corporation Lee Gearhart (2005–2008) Moog Inc. Ernest W. Klechka (2006–) Lloyd Register Capstone Inc. Alan T. Male (2003–) University of Kentucky William L. Mankins (1989–) Metallurgical Services Inc. Dana J. Medlin (2005–2008) South Dakota School of Mines and Technology Joseph W. Newkirk (2005–) University of Missouri-Rolla Cory J. Padfield (2006–) Hyundai America Technical Center Inc.
Toby V. Padfield (2004–) ZF Sachs Automotive of America Charles A. Parker (2007–) Honeywell Aerospace Elwin L. Rooy (2007–) Elwin Rooy & Associates Karl P. Staudhammer (1997–) Los Alamos National Laboratory Kenneth B. Tator (1991–) KTA-Tator Inc. George F. Vander Voort (1997–) Buehler Ltd.
Chairs of the ASM Handbook Committee J.F. Harper (1923–1926) (Member 1923–1926) W.J. Merten (1927–1930) (Member 1923–1933) L.B. Case (1931–1933) (Member 1927–1933) C.H. Herty, Jr. (1934–1936) (Member 1930–1936) J.P. Gill (1937) (Member 1934–1937) R.L. Dowdell (1938–1939) (Member 1935–1939) G.V. Luerssen (1943–1947) (Member 1942–1947) J.B. Johnson (1948–1951) (Member 1944–1951) E.O. Dixon (1952–1954) (Member 1947–1955) N.E. Promisel (1955–1961) (Member 1954–1963)
R.W.E. Leiter (1962–1963) (Member 1955–1958, 1960–1964) D.J. Wright (1964–1965) (Member 1959–1967) J.D. Graham (1966–1968) (Member 1961–1970) W.A. Stadtler (1969–1972) (Member 1962–1972) G.J. Shubat (1973–1975) (Member 1966–1975) R. Ward (1976–1978) (Member 1972–1978) G.N. Maniar (1979–1980) (Member 1974–1980) M.G.H. Wells (1981) (Member 1976–1981) J.L. McCall (1982) (Member 1977–1982) L.J. Korb (1983) (Member 1978–1983)
vii
T.D. Cooper (1984–1986) (Member D.D. Huffman (1986–1990) (Member D.L. Olson (1990–1992) (Member 1989–1992) R.J. Austin (1992–1994) (Member W.L. Mankins (1994–1997) (Member M.M. Gauthier (1997–1998) (Member C.V. Darragh (1999–2002) (Member Henry E. Fairman (2002–2004) (Member Jeffrey A. Hawk (2004–2006) (Member Larry D. Hanke (2006–2008) (Member
1981–1986) 1982–) 1982–1988, 1984–1985) 1989–) 1990–2000) 1989–) 1993–) 1997–) 1994–)
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
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Authors and Contributors
University of Birmingham, UK Kevin R. Anderson Mercury Marine
Paul G Campbell Metaullics Systems Division Pyrotek, Inc.
Princewill N. Anyalebechi Padnos College of Engineering & Computing, Grand Valley State University
Robert D Carnahan Thixomat Inc.
Diran Apelian Metal Processing Institute, Worcester Polytechnic Institute
Sujoy Chaudhury Honeywell International, Inc. K. K. Chawla University of Alabama At Birmingham
Alan P. Druschitz University of Alabama at Birmingham Cor van Ettinger Gieterij Doesburg Henry “Ed” Fairman Cincinnati Metallurgical Consultants Edward W. Flynn General Motors
Leigh Chen Technical Materials Incorporated
Jim Frost ACIPCO (American Cast Iron Pipe Company)
Henry Bakemeyer Die Casting Design and Consulting
Yeou-Li Chu Ryobi Die Casting
Richard Fruehan Carnegie Mellon University
Bryan Baker Vulcan Engineering Co.
Craig D. Clauser Craig Clauser Engineering Consulting Incorporated
Rafael Gallo Foseco Metallurgical
Lars Arnberg Norwegian University Of Science & Technology
Steve L. Cockroft The University of British Columbia
Charles-Andre´ Gandin Centre de Mise en Forme des Mate´riaux (CEMEF)
Chris Cooper Vulcan Engineering Co.
George Giles Morganite Crucible Inc
James D. Cotton The Boeing Co.
Peter C. Glaws The Timken Company
Randall Counselman Oldenburg Group Incorporated
Martin E. Glicksman University of Florida
Joe Bigelow CONTECH U. S. LLC.
Jonathan A. Dantzig University of Illinois-Urbana
George Goodrich Stork Climax Research Services
Ron Bird Stainless Foundry & Engineering
Atef Daoud Central Metallurgical Research and Development Institute, Cairo, Egypt
Frank E. Goodwin International Lead Zinc
Stewart A Ballantyne Allvac Allegheny Technologies Company Scott A. Balliett Latrobe Specialty Steel Company Christoph Beckermann University Of Iowa, The Michel Bellet Ecole des Mines de Paris, Centre de Mise en Forme des Mate´riaux (CEMEF)
Malcolm Blair Steel Founders Society of America William Boettinger NIST James A. Brock Sr. Alcoa Harold D. Brody University Of Connecticut Craig C. Brown Stork Technimet, Inc. Zach Brown CONTECH U.S. LLC Andreas Buhrig-Polackzek Gieberei-Institut, Germany Willam A. Butler General Motors (Retired) John Campbell
Craig V. Darragh The Timken Company Babu DasGupta National Science Foundation Bob Dawson Honeywell Aerospace MCOE Raymond Decker Thixomat Incorporated Raymond J Donahue Mercury Marine Daniel Dos Santos Gießerei Institut der RWTH Aachen Jon Dossett Metallurgical Consultant Jean Marie Drezet Ecole Polytechnique Fe´de´rale de Lausanne
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Lee Gouwens CMI Novacast Inc. Ronald Graham ATI Wah Chang Patrick Grant Oxford University Daniel E.Groteke Q.C. Designs Inc. Mustafa Guclu Stanley Associates Inc. Richard Gundlach Stork Climax Research Services Nikhil Gupta Polytechnic University Mike Gwyn ATI, Inc. John Hall
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
Hall Permanent Mold Machines Qingyou Han Purdue University Larry D. Hanke Materials Evaluation and Engineering Inc Justin Heimsch Madison Kipp Corporation
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Makhlouf M. Makhlouf Metal Processing Institute, Worcester Polytechnic Institute William L. Mankins Metallurgical Services Incorporated Stephen J. Mashl Bodycote HIP Andover
Intermet-PCPC Robert Pischel FOSECO Metallurgical Inc. David R. Poirier University of Arizona David Poweleit Steel Founders’ Society of America
Juan Carlos Heinrich The University of New Mexico
Ragnvald H. Mathiesen Norwegian University of Science and Technology
Gregg Holman Wah Chang
Robert J. McInerney BuhlerPrince
Thomas E. Prucha American Foundry Society
Robert A. Horton Precision Metalsmiths Incorporated
Scott McIntyre Grede Foundrie
Donald J. Hurtuk H.C. Starck Inc.
Brad McKee Deloro Stellite Inc.
Peter Quested Industry and Innovation Division, National Physical Laboratory, UK
Alain Jacot Laboratoire de Simulation des Mate´riaux
Harold T. Michels Copper Development Association
John L. Jorstad J.L.J. Technologies, Inc.
Tony Midea Foseco Metallurgical Inc.
Ursula R. Kattner N I S T/Metallurgy Division
Paul H Mikkola Metal Casting Technology Incorporated
Seymour Katz S. Katz Associates
Asbj rn Mo SINTEF Materials and Chemistry
Jay S. Keist Worcester Polytechnic Institute
Raymond W. Monroe Steel Founders’ Society of America
Graham Keough Consarc Corporation
Y.V. Murty Cellular Materials International, Inc.
Matthew Krane Purdue University
Laurentiu Nastac Concurrent Technologies Corporation
Mary Beth Krusiak Sand Technology Co. LLC
David V. Neff Metaullics Systems Division Pyrotek, Inc.
T.A. Kuhn Alcoa Primary Metals
Charles Nelson Morris Bean and Company
Wilfried Kurz Ecole Polytechnique Federale De Lausanne (EPFL) Switzerland
Itsuo Ohnaka Osaka Sangyo University, Japan
Selcuk Kuyucak CANMET Materials Technology Laboratory Vic La Fay Hill & Griffith Company Robin A. Lampson Retech Systems LLC
Cory Padfield American Axle & Manufacturing Qingyue Pan SPX Corporation
William L. Powell ThyssenKrupp-Waupaca (Retired)
Wayne Rasmussen Consultant Martin Reeves Striko-Dynarad Allen Richardson SPX Corporation Pradeep Rohatgi University of Wisconsin-Milwaukee Elwin L. Rooy Elwin Rooy and Associates Gary Ruff SMW Automotive Corporation Kumar (Muthukumarasamy) Sadayappan CANMET Materials Technology Laboratory Mahi Sahoo CANMET Materials Technology Laboratory Kenneth Savage Magnesium Elektron Gary Scholl Metal Casting Technology Incorporated Ted Schorn Enkei America, Inc. Benjamin F. Schultz University of Wisconsin-Milwakee David Schwam Case Werstern Reserve University
Charles A. Parker Honeywell Aerospace
William D. Scott AAA Alchemy
David Lee Howmet Corp.
Kent Peaslee Missouri University of Science and Technology (Missouri S&T)
Ben Q. Li University of Michigan Dearborn
Pehlke, Robert D. Univ of Michigan
Sumanth Shankar McMaster University
Alan Luo General Motors R&D Center
John H. Perepezko University of Wisconsin, Madison
Geoffrey K. Sigworth Alcoa Primary Metals
Dan Maas Ex-One
Ralph Perkul Foundry Solutions and Design LLC
Steve Sikkenga Cannon-Muskegon Corporation
James F. Major Alcan International Ltd
Gordon H. Peters
Brian V. Smith
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John Serra Carpenter Brothers Inc.
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
General Motors Corporation, Powertrain Division Jerry Sokolowski University of Windsor Al Spada American Foundry Society Karl Staudhammer Los Alamos National Laboratory Doru M. Stefanescu The Ohio State University Ingo Steinbach RWTH-Aachen, ACCESS e. V. Karthik Subbiah SPX Product Solutions Group Erin H. Sunseri Iowa State University Tony Suschil Foseco Inc Richard Sutherlin ATI Wah Chang Edward S.Szekeres Casting Consultants Incorporated Brian G. Thomas
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University of Illinois Michael L Tims Concurrent Technologies Corporation David W. Tripp TIMET
Srinath Viswanathan University of Alabama Birmingham Robert C. Voigt Penn State University Vaughan R. Voller University of Minnesota
Rohit Trivedi AMES Laboratory Paul Trojan University of Michigan-Dearborn John Troxler Rolls Royce
Tom Waring ME Elecmetal Sufei Wei Centrifugal Casting Machine Company Incorporated David White Schaefer Furnace
Nao Tsumagari American Showa, Inc. Daniel Twarog North American Die Casting Association Derek E. Tyler Olin Metals Thin Strip (retired) Don Tyler General Aluminum Manufacturing Company Stephen Udvardy North American Die Casting Association Juan J. Valencia Concurrent Technologies Corporation
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Lawrence Whiting CANMET Materials Technology Laboratory Richard Williams FOSECO Metallurgical Inc. Greg G.Woycik Woycik Consulting Oscar Yu RTI International Metals, Inc. Keith Zhang Teck Cominco Metals Ltd.
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
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Contents
Applications and Specification Guidelines. . . . . . . . . . . . . . . . . . . 1 Raymond W. Monroe, Steel Founders’ Society of America; Thomas E. Prucha, American Foundry Society; Daniel Twarog, North American Die Casting Association History and Trends of Metal Casting. . . . . . . . . . . . . . History of Metal Casting . . . . . . . . . . . . . . . . . . . . Metal Casting Methods. . . . . . . . . . . . . . . . . . . . . . Production Trends . . . . . . . . . . . . . . . . . . . . . . . . . Market Trends . . . . . . . . . . . . . . . . . . . . . . . . . . . . Metalcasting Technology and the Purchasing Process . . . . The Purchasing Process . . . . . . . . . . . . . . . . . . . . . Define Requirements and Develop a Purchasing Plan. Requesting and Evaluating Quotations/Bids . . . . . . . Select a Supplier and Negotiate Contract Terms . . . . Contract Fulfillment and Continuous Improvement . . Considerations in Purchasing Metal Castings . . . . . .
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Principles of Liquid Metal Processing . . . . . . . . . . . . Chemical Thermodynamics and Kinetics . . . . . . . . . . . . Chemical Thermodynamics . . . . . . . . . . . . . . . . . . Chemical Kinetics . . . . . . . . . . . . . . . . . . . . . . . . Kinetics of Alloys Additions . . . . . . . . . . . . . . . . . Thermodynamic Properties of Iron-Base Alloys . . . . . . . Purification of Ferrous Melts. . . . . . . . . . . . . . . . . Thermodynamics of Ferrous Systems . . . . . . . . . . . Multicomponent Iron-Carbon Systems . . . . . . . . . . Structural Diagrams . . . . . . . . . . . . . . . . . . . . . . . Thermodynamic Properties of Aluminum-Base and Copper-Base Alloys . . . . . . . . . . . . . . . . . . . . Activities and Thermal Properties . . . . . . . . . . . . . Ineraction Coefficients . . . . . . . . . . . . . . . . . . . . . Thermal Properties for Hypothetical Standard State . Phase Diagrams . . . . . . . . . . . . . . . . . . . . . . . . . . Gases in Metals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Cast Iron. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Aluminum and Its Alloys . . . . . . . . . . . . . . . . . . . Copper and Its Alloys . . . . . . . . . . . . . . . . . . . . . Magnesium and Its Alloys . . . . . . . . . . . . . . . . . . Overcoming Gas Porosity . . . . . . . . . . . . . . . . . . . Inclusion-Forming Reactions . . . . . . . . . . . . . . . . . . . . Inclusion Types . . . . . . . . . . . . . . . . . . . . . . . . . . Physical Chemistry . . . . . . . . . . . . . . . . . . . . . . . Control of Inclusions . . . . . . . . . . . . . . . . . . . . . . Inclusions in Ferrous Alloys . . . . . . . . . . . . . . . . . Inclusions in Nonferrous Alloys . . . . . . . . . . . . . .
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56 56 56 58 58 64 64 67 70 71 73 74 74 75 76 77 81
Melting and Remelting . . . . . . . . . . Electric Arc Furnace Melting . . . . . . . Power Supply . . . . . . . . . . . . . . Arc Furnace Components . . . . . . Weighing and Chemical Analysis Scrap and Alloy Storage . . . . . . .
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Acid Melting Practice . . . . . . . . . . . . . . . . . Basic Melting Practice . . . . . . . . . . . . . . . . . Newer Technologies. . . . . . . . . . . . . . . . . . . Cupola Furnaces . . . . . . . . . . . . . . . . . . . . . . . . . Progressive Refinements in Cupola Equipment Cupola Construction and Operation . . . . . . . . Charge Materials . . . . . . . . . . . . . . . . . . . . . Cupola Control Principles . . . . . . . . . . . . . . . Control Tests and Analyses. . . . . . . . . . . . . . Specialized Cupolas . . . . . . . . . . . . . . . . . . . Induction Furnaces . . . . . . . . . . . . . . . . . . . . . . . Types of Furnaces . . . . . . . . . . . . . . . . . . . . Electromagnetic Stirring . . . . . . . . . . . . . . . . Power Supplies . . . . . . . . . . . . . . . . . . . . . . Water Cooling Systems . . . . . . . . . . . . . . . . Lining Material . . . . . . . . . . . . . . . . . . . . . . Melting Operations . . . . . . . . . . . . . . . . . . . Vacuum Induction Melting . . . . . . . . . . . . . . . . . Process Description . . . . . . . . . . . . . . . . . . . VIM Metallurgy . . . . . . . . . . . . . . . . . . . . . Production of Nonferrous Materials . . . . . . . . Electroslag Remelting . . . . . . . . . . . . . . . . . . . . . ESR Furnaces . . . . . . . . . . . . . . . . . . . . . . . Ingot Solidification. . . . . . . . . . . . . . . . . . . . Process Variations . . . . . . . . . . . . . . . . . . . . Steel ESR . . . . . . . . . . . . . . . . . . . . . . . . . . Superalloy ESR . . . . . . . . . . . . . . . . . . . . . . Vacuum Arc Remelting . . . . . . . . . . . . . . . . . . . . Process Description . . . . . . . . . . . . . . . . . . . Solidification . . . . . . . . . . . . . . . . . . . . . . . . Process Variables. . . . . . . . . . . . . . . . . . . . . Superalloy VAR . . . . . . . . . . . . . . . . . . . . . Titanium VAR . . . . . . . . . . . . . . . . . . . . . . Skull Melting . . . . . . . . . . . . . . . . . . . . . . . . . . . Vacuum Arc Skull Melting . . . . . . . . . . . . . . Furnaces . . . . . . . . . . . . . . . . . . . . . . . . . . . Induction Skull Melting . . . . . . . . . . . . . . . . Electron Beam Melting . . . . . . . . . . . . . . . . . . . . Process Description . . . . . . . . . . . . . . . . . . . Drip Melting . . . . . . . . . . . . . . . . . . . . . . . . Cold Hearth Melting . . . . . . . . . . . . . . . . . . Plasma Melting and Heating . . . . . . . . . . . . . . . . Plasma Torches . . . . . . . . . . . . . . . . . . . . . . Furnace Equipment . . . . . . . . . . . . . . . . . . . Atmosphere Control . . . . . . . . . . . . . . . . . . . Melting Processes . . . . . . . . . . . . . . . . . . . . Crucible Furnaces . . . . . . . . . . . . . . . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . . . . . . Furnace Types . . . . . . . . . . . . . . . . . . . . . . . Design and Operation. . . . . . . . . . . . . . . . . . Reverberatory and Stack Furnaces . . . . . . . . . . . . Furnace Types . . . . . . . . . . . . . . . . . . . . . . . Reverberatory Furnace Practice . . . . . . . . . . . Molten Metal Circulation . . . . . . . . . . . . . . .
3 4 8 9 14 16 16 17 21 22 22 22
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91 95 98 99 99 100 103 106 106 107 108 109 109 111 111 113 116 116 119 122 124 125 128 129 129 131 132 132 134 135 137 137 139 139 139 140 142 142 143 145 149 149 149 151 151 151 155 155 156 158 160 161 163 165
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
www.asminternational.org
Molten Metal Processing and Handling . . . . . . . . . . . . . Transfer and Treatment of Molten Metal—An Introduction . Launders. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Tundishes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Molten-Metal Pumps . . . . . . . . . . . . . . . . . . . . . . . . Pouring and Dosing . . . . . . . . . . . . . . . . . . . . . . . . . Degassing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Aluminum Foundry Degassing . . . . . . . . . . . . . . . . . Degassing of Magnesium . . . . . . . . . . . . . . . . . . . . . Degassing of Copper Alloys . . . . . . . . . . . . . . . . . . . Molten-Metal Filtration . . . . . . . . . . . . . . . . . . . . . . . . . . Sources of Inclusion. . . . . . . . . . . . . . . . . . . . . . . . . Removal of Inclusions . . . . . . . . . . . . . . . . . . . . . . . Filter Characteristics . . . . . . . . . . . . . . . . . . . . . . . . Types of Molten-Metal Filters . . . . . . . . . . . . . . . . . . In-Mold Filtration . . . . . . . . . . . . . . . . . . . . . . . . . . In-Furnace Filtration of Molten Aluminum . . . . . . . . . Filtered Metal Quality . . . . . . . . . . . . . . . . . . . . . . . Steel Melt Processing . . . . . . . . . . . . . . . . . . . . . . . . . . . Furnaces and Refractories . . . . . . . . . . . . . . . . . . . . . Converter Metallurgy . . . . . . . . . . . . . . . . . . . . . . . . Ladle Metallurgy . . . . . . . . . . . . . . . . . . . . . . . . . . . Aluminum Fluxes and Fluxing Practice . . . . . . . . . . . . . . . Aluminum Fluxing. . . . . . . . . . . . . . . . . . . . . . . . . . Categories of Fluxes . . . . . . . . . . . . . . . . . . . . . . . . Foundry Practices . . . . . . . . . . . . . . . . . . . . . . . . . . Modification of Aluminum-Silicon Alloys . . . . . . . . . . . . . Heat Treated Castings . . . . . . . . . . . . . . . . . . . . . . . Porosity and Shrinkage Formation . . . . . . . . . . . . . . . Modification Mechanisms . . . . . . . . . . . . . . . . . . . . . Eutectic Nucleation . . . . . . . . . . . . . . . . . . . . . . . . . Eutectic Grain Structure . . . . . . . . . . . . . . . . . . . . . . Recommended Foundry Practices . . . . . . . . . . . . . . . Growth of Silicon Eutectic . . . . . . . . . . . . . . . . . . . . Effect of Other Elements . . . . . . . . . . . . . . . . . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Grain Refinement of Aluminum Casting Alloys . . . . . . . . . Measurement of Grain Size. . . . . . . . . . . . . . . . . . . . Mechanisms of Grain Refinement . . . . . . . . . . . . . . . Boron as a Grain Refiner . . . . . . . . . . . . . . . . . . . . . Best Practices for Grain Refinement. . . . . . . . . . . . . . Benefits of Grain Refinement . . . . . . . . . . . . . . . . . . Concluding Remarks . . . . . . . . . . . . . . . . . . . . . . . . Refinement of the Primary Silicon Phase in Hypereutectic Aluminum-Silicon Alloys . . . . . . . . . . . . . . . . . . . . . Importance of Primary Silicon Refinement . . . . . . . . . Accomplishing Primary Silicon Refinement . . . . . . . . Principles of Solidification . . . . . . . . . . . . . . . . Diran Apelian, Worcester Polytechnic Institute Thermodynamics and Phase Diagrams . . . . . . . . Thermodynamics . . . . . . . . . . . . . . . . . . . . Phase Diagrams . . . . . . . . . . . . . . . . . . . . . Solidification . . . . . . . . . . . . . . . . . . . . . . . Nucleation Kinetics and Grain Refinement . . . . . Thermodynamics of Solidification . . . . . . . . Nucleation Phenomena . . . . . . . . . . . . . . . . Inoculation Practice . . . . . . . . . . . . . . . . . . Grain Refinement Models . . . . . . . . . . . . . . Master Alloy Processing . . . . . . . . . . . . . . . Transport Phenomena During Solidification . . . . . Fundamentals . . . . . . . . . . . . . . . . . . . . . . Transport and Microstructure . . . . . . . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . Plane Front Solidification . . . . . . . . . . . . . . . . . Transients and Steady State . . . . . . . . . . . . Morphological Stability . . . . . . . . . . . . . . .
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171 173 174 174 175 178 185 185 188 189 194 194 196 197 199 199 201 204 206 206 208 212 230 231 233 235 240 241 242 243 244 245 247 249 249 250 255 255 255 257 258 260 260
Rapid Solidification Effects. . . . . . . . . . . . . . . . . . . . . . Solidification Microstructure Selection Maps . . . . . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Non-Plane Front Solidification . . . . . . . . . . . . . . . . . . . . . . . Cellular Microstructures . . . . . . . . . . . . . . . . . . . . . . . . Dendrite Growth . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Nonplanar Eutectic Growth. . . . . . . . . . . . . . . . . . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Eutectic Solidification—An Introduction . . . . . . . . . . . . . . . . Binary Eutectic Phase Diagram . . . . . . . . . . . . . . . . . . . Eutectic Structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . Solidification of Eutectic Alloys—Aluminum-Silicon . . . . . . . The Aluminum-Silicon Equilibrium Phase Diagram . . . . . Classification of the Aluminum-Silicon Eutectic Structure Microstructure of the Aluminum-Silicon Eutectic . . . . . . Theories of Solidification of the Aluminum-Silicon Eutectic . . . . . . . . . . . . . . . . . . Theories of Chemical Modification of the Aluminum-Silicon Eutectic . . . . . . . . . . . . . . . . . . Solidification of Eutectic Alloys—Cast Iron. . . . . . . . . . . . . . Structure of Liquid Iron-Carbon Alloys . . . . . . . . . . . . . Nucleation of Eutectic in Cast Iron . . . . . . . . . . . . . . . . Nucleation of Primary Phases in Cast Iron . . . . . . . . . . . Growth of Eutectic in Cast Iron. . . . . . . . . . . . . . . . . . . Cooling Curve Analysis . . . . . . . . . . . . . . . . . . . . . . . . Peritectic Solidification . . . . . . . . . . . . . . . . . . . . . . . . . . . . Peritectic Solidification. . . . . . . . . . . . . . . . . . . . . . . . . Peritectic Reaction . . . . . . . . . . . . . . . . . . . . . . . . . . . . Peritectic Transformation and Direct Precipitation . . . . . . Primary Metastable Precipitation of Beta . . . . . . . . . . . . Peritectic Cascades . . . . . . . . . . . . . . . . . . . . . . . . . . . Peritectic Transformations in Multicomponent Systems . . Microsegregation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Solidification Models . . . . . . . . . . . . . . . . . . . . . . . . . . Binary Isomorphous Systems, k > 1: Titanium-Molybdenum . . . . . . . . . . . . . . . . . . . . . Binary Eutectic Systems, k > 1: Hypoeutectic Aluminum-Copper Alloys . . . . . . . . . . . . . . . . . . . Binary Eutectic Systems: Hypoeutectic Aluminum-Silicon Alloys . . . . . . . . . . . . . . . . . . . Binary Peritectic Systems: Copper-Zinc a-Brass . . . . . . . Multicomponent, Multiphase Eutectic Systems: Hypoeutectic Al-Si-Cu-Mg Alloys . . . . . . . . . . . . . . . Multicomponent Multiphase Steel: Fe-C-Cr . . . . . . . . . . Multicomponent, Multiphase Nickel-Base Superalloy. . . . Macrosegregation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Macrosegregation Induced by Flow of the Liquid . . . . . . Macrosegregation Induced by Movement of the Solid . . . Interpretation and use of Coolng Curves—Thermal Analysis . . Quantitative Thermal Analysis . . . . . . . . . . . . . . . . . . . Deviation from Equilibrium . . . . . . . . . . . . . . . . . . . . . Differential Thermal Analysis . . . . . . . . . . . . . . . . . . . . Determining Liquidus and Solidus Temperature . . . . . . . X-Ray Imaging of Solidification Processes and Microstructure Evolution . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . X-Ray Imaging . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . In-House Laboratory X-Ray Imaging of Solidification Processes . . . . . . . . . . . . . . . . . . . . . X-Ray Imaging Studies of Solidification Processes with Synchrotron Radiation . . . . . . . . . . . . . . . . . . . . Further Reading. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . The Oxide Film Defect—The Bifilm. . . . . . . . . . . . . . . . . . . Bifilms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Entrainment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Furling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Unfurling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Reliability and Reproducibility of Mechanical Properties .
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© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
www.asminternational.org
Central Role of Bifilm Defects in Fracture . . . . . . . . . . . Shrinkage Porosity and Gas Porosity. . . . . . . . . . . . . . . . . . . Shrinkage Porosity . . . . . . . . . . . . . . . . . . . . . . . . . . . . Gas Porosity . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Factors Affecting Porosity Formation . . . . . . . . . . . . . . . Castability—Fluidity and Hot Tearing. . . . . . . . . . . . . . . . . . Fluidity. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Hot Tearing. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Semisolid Metal Processing . . . . . . . . . . . . . . . . . . . . . . . . . Flow Behavior of SSM. . . . . . . . . . . . . . . . . . . . . . . . . Semisolid Microstructural Formation . . . . . . . . . . . . . . . Semisolid Processes: Thixocasting versus Rheocasting . . . Spray Casting. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Principles . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Atomization and the Droplet Spray . . . . . . . . . . . . . . . . Deposition and Grain Mulitplication . . . . . . . . . . . . . . . Microsegregation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Macrosegregation and Coarsening . . . . . . . . . . . . . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Rapid Solidification . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Microstructural Modification . . . . . . . . . . . . . . . . . . . . . Solidification During Casting of Metal-Matrix Composites . . . Incorporation of Reinforcements . . . . . . . . . . . . . . . . . . Reinforcement-Metal Wettability . . . . . . . . . . . . . . . . . . Solidification Processing of MMC . . . . . . . . . . . . . . . . . Solidification Fundamentals . . . . . . . . . . . . . . . . . . . . . Modeling of Particle-Pushing Phenomenon . . . . . . . . . . . Nanocomposites. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Conclusions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Solidification Research in Microgravity. . . . . . . . . . . . . . . . . Availability of Reduced Gavity: Past, Present, and Future Solidification Research in Microgravity . . . . . . . . . . . . . Key Solidification Experiments . . . . . . . . . . . . . . . . . . . The Future of Microgravity Solidification . . . . . . . . . . . . Homogenization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Sources of Inhomogeneity. . . . . . . . . . . . . . . . . . . . . . . Benefits of Homogenization Treatment. . . . . . . . . . . . . . Characterization Methods . . . . . . . . . . . . . . . . . . . . . . . Computational Modeling . . . . . . . . . . . . . . . . . . . . . . . Heat Treatment. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Designations of Heat Treatment . . . . . . . . . . . . . . . . . . Solution Heat Treatment. . . . . . . . . . . . . . . . . . . . . . . . Quenching . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Aging . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Emerginhg Heat Treating Technologies . . . . . . . . . . . . . Microstructural Changes due to Heat Treatment of Cast Alloys . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Hot Isostatic Pressing of Castings. . . . . . . . . . . . . . . . . . . . . History of HIP. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Reasons for Using HIP . . . . . . . . . . . . . . . . . . . . . . . . . Overview of a HIP System . . . . . . . . . . . . . . . . . . . . . . Effect of HIP on Mechanical Properties . . . . . . . . . . . . . Effect of Inclusions on As-HIPed Properties . . . . . . . . . . Effect of HIP on the Shape and Structure of Castings . . . Problems Encountered in HIP . . . . . . . . . . . . . . . . . . . . Economics of HIP . . . . . . . . . . . . . . . . . . . . . . . . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Modeling and Analysis of Casting Processes. . . . . . . . Srinath Viswanathan, University of Alabama Numerical Methods for Casting Applications . . . . . . . . . Fundamentals of CFD . . . . . . . . . . . . . . . . . . . . . Numerical Solution of the Fluid-Flow Equations . . . Grid Generation for Complex Geometries. . . . . . . . Casting Applications . . . . . . . . . . . . . . . . . . . . . . Modeling of Transport Phenomena and Electromagnetics Conservation of Thermal Energy . . . . . . . . . . . . . .
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406 408 408 409 409 411 412 413 414 415 415
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419 419 420 420 422 425 425
Conservation of Solute . . . . . . . . . . . . . . . . . . . . . . . . . Equations Cast in Advection-Diffusion Form . . . . . . . . . Basic Equations for Mass and Momentum Conservation . Modeling the Momentum in the Solid + Liquid Mushy Region . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Modeling Turbulence . . . . . . . . . . . . . . . . . . . . . . . . . . Phenomena Restraining or Driving Flow . . . . . . . . . . . . Dealing with Fronts and Free Surfaces . . . . . . . . . . . . . . Simulation of Electromagnetic Effects in Castings. . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Direct Modeling of Structure Formation . . . . . . . . . . . . . . . . Introduction. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Direct Microstructure Simulation Using the Phase Field Method. . . . . . . . . . . . . . . . . . . . . . . . Direct Grain Structure Simulation Using the Cellular Automation Method . . . . . . . . . . . . . . . . . Coupling of Direct Structure Simulation at Macroscopic Scale . . . . . . . . . . . . . . . . . . . . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Modeling of Microsegregation and Macrosegregation . . . . . . . Microsegregation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Microsegregation with Diffusion in the Solid . . . . . . . . . Continuum Model of Macrosegregation . . . . . . . . . . . . . Modeling of Stress, Distortion, and Hot Tearing . . . . . . . . . . Governing Equations . . . . . . . . . . . . . . . . . . . . . . . . . . Thermomechanical Coupling . . . . . . . . . . . . . . . . . . . . . Numerical Solution . . . . . . . . . . . . . . . . . . . . . . . . . . . Model Validation. . . . . . . . . . . . . . . . . . . . . . . . . . . . . Example Applications . . . . . . . . . . . . . . . . . . . . . . . . . Hot-Tearing Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . Conclusions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Practical Issues In Computer Simulation of Casting Processes . Setting Simulation Objectives, Modeling, and Selection of Simulation Code . . . . . . . . . . . . . . . . . . . . . . . . Modeling of Shape and Phenomena . . . . . . . . . . . . . . . . Initial and Boundary Conditions and Physical Properties . Enmeshing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Evaluation of Simulation Results . . . . . . . . . . . . . . . . . . Estimation of Structure and Properties . . . . . . . . . . . . . . Optimization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Thermophysical Properties . . . . . . . . . . . . . . . . . . . . . . . . . . Sources and Availability of Reliable Data . . . . . . . . . . . Limitations and Warning on the Use of Data . . . . . . . . . Methods to Determine Thermophysical Properties . . . . . . Specific Heat Capacity and Enthalpy of Transformation . . Enthalpy of Melting, Solidus, and Liquidus Temperatures Coefficient of Thermal Expansion . . . . . . . . . . . . . . . . . Density . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Surface Tension. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Viscosity. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Electrical and Thermal Conductivity . . . . . . . . . . . . . . . Emissivity. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Typical Themophysical Properties Ranges of Some Cast Alloys . . . . . . . . . . . . . . . . . . . . . . . . . Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Principles and Practices of Shape Casting . . . . . . . Shape Casting Processes—An Introduction . . . . . . . . Patterns and Patternmaking . . . . . . . . . . . . . . . . . . . Types of Patterns . . . . . . . . . . . . . . . . . . . . . . . Additional Features of Patterns . . . . . . . . . . . . . Pattern Allowances . . . . . . . . . . . . . . . . . . . . . Pattern Materials . . . . . . . . . . . . . . . . . . . . . . . Selection of Pattern Type . . . . . . . . . . . . . . . . . Trends in Pattern Design and Manufacture . . . . . Casting Practice—Guidelines for Effective Production of Reliable Castings . . . . . . . . . . . . . . . . . . . . .
xiii
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428 428 429 429 431 432 435 435
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441 443 445 445 445 446 449 449 451 451 453 454 456 458 462
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462 462 463 464 465 466 466 468 468 468 468 470 470 471 471 471 474 474 474
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483 485 488 488 490 491 492 494 495
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© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
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Rule 1: Good-Quality Melt . . . . . . . . . . . . . . . . Rule 2: Liquid Front Damage . . . . . . . . . . . . . . Rule 3: Liquid Front Stop. . . . . . . . . . . . . . . . . Rule 4: Bubble Damage . . . . . . . . . . . . . . . . . . Rule 5: Core Blows . . . . . . . . . . . . . . . . . . . . . Rule 6: Shrinkage Damage . . . . . . . . . . . . . . . . Rule 7: Convection Damage . . . . . . . . . . . . . . . Rule 8: Segregation . . . . . . . . . . . . . . . . . . . . . Rule 9: Residual Stress. . . . . . . . . . . . . . . . . . . Rule 10: Location Points . . . . . . . . . . . . . . . . . Filling and Feeding System Concepts . . . . . . . . . . . . Entrainment Defects in Gravity-Poured Castings . Traditional Filling Designs . . . . . . . . . . . . . . . . Prepriming Techniques . . . . . . . . . . . . . . . . . . . Naturally Pressurized Filling System . . . . . . . . . Filters. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Performance of Alternative Filling Systems . . . . Recommended Foundry Practice . . . . . . . . . . . . Processing and Finishing of Castings . . . . . . . . . . . . Shakeout and Core Knockout . . . . . . . . . . . . . . Cleaning Operations. . . . . . . . . . . . . . . . . . . . . Automating Gate Removal and Grinding . . . . . . Flame Cutting . . . . . . . . . . . . . . . . . . . . . . . . . Blast Cleaning of Castings . . . . . . . . . . . . . . . . Inspection . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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Expandable-Mold Casting Processes with Permanent Patterns. . . . . . . . . . . . . . . . . . . John L. Jorstad, J.L.J. Technologies, Inc Introduction—Expendable Mold Processes with Permanent Patterns . . . . . . . . . . . . . Aggregates and Binders for Expendable Molds Mold Media . . . . . . . . . . . . . . . . . . . . . Foundry Sands. . . . . . . . . . . . . . . . . . . . Clays for Green Sand Molding . . . . . . . . Inorganic Binders . . . . . . . . . . . . . . . . . Organic Binders. . . . . . . . . . . . . . . . . . . Other Expendable Mold Media . . . . . . . . Green Sand Molding . . . . . . . . . . . . . . . . . . . Sand Molding . . . . . . . . . . . . . . . . . . . . System Control and Formulation . . . . . . . Influence of Molding Equipment . . . . . . . Maintaining Sand System Quality . . . . . . Molding Methods. . . . . . . . . . . . . . . . . . Green Sand Media Preparation . . . . . . . . Molding Problems . . . . . . . . . . . . . . . . . Mold Finishing . . . . . . . . . . . . . . . . . . . Shakeout. . . . . . . . . . . . . . . . . . . . . . . . Sand/Casting Recovery. . . . . . . . . . . . . . Sand Reclamation . . . . . . . . . . . . . . . . . No-Bake Sand Molding . . . . . . . . . . . . . . . . . No-Bake Sand Processing . . . . . . . . . . . . CO2 -Cured Sodium Silicate . . . . . . . . . . Air-Setting Sodium Silicates . . . . . . . . . . Air-Setting Organic Binders . . . . . . . . . . Gas-Cured Organic Binders. . . . . . . . . . . Sand Reclamation and Reuse . . . . . . . . . Coremaking . . . . . . . . . . . . . . . . . . . . . . . . . Binders for Sand Coremaking . . . . . . . . . Basic Principles of Coremaking. . . . . . . . Core Sands and Additives. . . . . . . . . . . . Compaction. . . . . . . . . . . . . . . . . . . . . . Curing of Compacted Cores . . . . . . . . . . Core Coatings (Core Washes) . . . . . . . . . Assembly and Core Setting . . . . . . . . . . . Design Considerations . . . . . . . . . . . . . . Coring of Tortuous Passages . . . . . . . . . .
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497 497 499 500 501 501 502 503 504 505 506 506 506 507 508 511 511 512 513 513 516 517 519 519 520
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525 528 528 529 535 538 539 546 549 549 551 552 553 554 558 560 560 561 562 564 567 567 569 572 573 576 577 581 581 582 583 584 585 587 588 589 592
Green Sand versus Dry Sand Cores . . . . . Cores in Permanent Mold Castings . . . . . Cores in Investment Castings . . . . . . . . . Designing to Eliminate Cores . . . . . . . . . Coring versus Drilling . . . . . . . . . . . . . . Shell Molding and Shell Coremaking . . . . . . . Applicability . . . . . . . . . . . . . . . . . . . . . Properties and Selection of Shell Sands . . Resins . . . . . . . . . . . . . . . . . . . . . . . . . Preparation of Resin-Sand Mixtures . . . . . Sand Reclamation . . . . . . . . . . . . . . . . . Control Testing of Resin-Sand Properties . Patterns and Core Boxes. . . . . . . . . . . . . Production of Molds and Cores . . . . . . . . Mold Assembly . . . . . . . . . . . . . . . . . . . Casting Process . . . . . . . . . . . . . . . . . . . Dimensional Accuracy . . . . . . . . . . . . . . Causes and Prevention of Casting Defects Comparisons with Green Sand Molding . . Slurry Molding . . . . . . . . . . . . . . . . . . . . . . . Plaster Molding . . . . . . . . . . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . Plaster Mold Compositions . . . . . . . . . Patterns and Core Boxes . . . . . . . . . . . Mold-Drying Equipment . . . . . . . . . . . Metals Cast in Plaster Molds . . . . . . . . Conventional Plaster Mold Casting. . . . Match Plate Pattern Molding . . . . . . . . The Antioch Process . . . . . . . . . . . . . . Foamed Plaster Molding Process . . . . . Ceramic Molding . . . . . . . . . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . Shaw Process . . . . . . . . . . . . . . . . . . . Unicast Process . . . . . . . . . . . . . . . . . No-Bond Sand Molding. . . . . . . . . . . . . . . . . Magnetic Molding . . . . . . . . . . . . . . . . . Vacuum (V-Process) Molding . . . . . . . . .
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Expandable-Mold Casting Processes with Expendable Patterns . . . . . . . . . . . . . . . . . . . John L. Jorstad, J.L.J. Technologies, Inc Introduction—Expendable Mold Processes with Expendable Patterns . . . . . . . . . . . . . . . . . Lost Foam Casting . . . . . . . . . . . . . . . . . . . . . Process Technique . . . . . . . . . . . . . . . . . . Processing Parameters . . . . . . . . . . . . . . . Advantages of Lost Foam Casting . . . . . . . Casting Quality . . . . . . . . . . . . . . . . . . . . Investment Casting . . . . . . . . . . . . . . . . . . . . . Pattern Materials . . . . . . . . . . . . . . . . . . . Patternmaking . . . . . . . . . . . . . . . . . . . . . Pattern and Cluster Assembly . . . . . . . . . . Ceramic Shell Mold Manufacture . . . . . . . Manufacture of Ceramic Cores . . . . . . . . . Pattern Removal . . . . . . . . . . . . . . . . . . . Mold Firing and Burnout . . . . . . . . . . . . . Melting and Casting. . . . . . . . . . . . . . . . . Postcasting Operations . . . . . . . . . . . . . . . Inspection and Testing . . . . . . . . . . . . . . . Design Advantages of Investment Castings . Design Recommendations . . . . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . . . . Special Investment Casting Processes. . . . . Replicast Molding. . . . . . . . . . . . . . . . . . . . . . Process Details . . . . . . . . . . . . . . . . . . . . Process Capabilities . . . . . . . . . . . . . . . . .
xiv
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593 593 594 595 596 598 599 599 600 601 603 603 604 604 607 610 611 614 615 617 617 617 617 618 618 618 619 620 620 621 622 622 622 624 628 628 628
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637 640 640 642 645 645 646 646 648 650 651 654 654 654 655 656 656 657 657 657 657 662 662 663
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
Centrifugal Casting . . . . . . . . . . . . . . . . . John L. Jorstad, J.L.J. Technologies, Inc Centrifugal Casting . . . . . . . . . . . . . . . . . . Centrifugal Casting Methods . . . . . . . . Equipment . . . . . . . . . . . . . . . . . . . . Molds. . . . . . . . . . . . . . . . . . . . . . . . Defects in Centrifugal Castings . . . . . . Applications . . . . . . . . . . . . . . . . . . . Horizontal Centrifugal Casting . . . . . . . . . . Equipment . . . . . . . . . . . . . . . . . . . . Casting Process . . . . . . . . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . Vertical Centrifugal Casting. . . . . . . . . . . . Mold Design . . . . . . . . . . . . . . . . . . . Process Details . . . . . . . . . . . . . . . . .
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Permanent Mold and Semipermanent Mold Processes . . . . . . . . . . . . . . . . . . . . . . . . John L. Jorstad, J.L.J. Technologies, Inc Permanent Mold Casting . . . . . . . . . . . . . . . . . . Gravity Casting Methods . . . . . . . . . . . . . . Advantages and Disadvantages . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . . . . . Casting Design . . . . . . . . . . . . . . . . . . . . . Mold Design . . . . . . . . . . . . . . . . . . . . . . . Mold Materials . . . . . . . . . . . . . . . . . . . . . Mold Coatings. . . . . . . . . . . . . . . . . . . . . . Mold Life . . . . . . . . . . . . . . . . . . . . . . . . . Mold Temperature . . . . . . . . . . . . . . . . . . . Control of Mold Temperature . . . . . . . . . . . Dimensional Accuracy . . . . . . . . . . . . . . . . Surface Finish . . . . . . . . . . . . . . . . . . . . . . Low-Pressure Die Casting . . . . . . . . . . . . . . . . . Conventional Low-Pressure Casting . . . . . . . Counterpressure Casting . . . . . . . . . . . . . . . Vacuum Riserless/Pressure Riserless Casting Low Pressure Counter Gravity Casting . . . . . . . . Mold Filling . . . . . . . . . . . . . . . . . . . . . . .
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Die Casting Tooling . . . . . . . . . . . . . . . . . . . . . . . Die Casting Machines and Dies. . . . . . . . . . . . Product Design and the Process. . . . . . . . . . . . Tooling and Metal Flow . . . . . . . . . . . . . . . . . Heat Removal . . . . . . . . . . . . . . . . . . . . . . . . Materials Selection for Tooling . . . . . . . . . . . . Automation in High-Pressure Die Casting . . . . . . . . Die Casting Cycle . . . . . . . . . . . . . . . . . . . . . Automatic Pouring . . . . . . . . . . . . . . . . . . . . . Injection Process . . . . . . . . . . . . . . . . . . . . . . Solidification and Heat Removal . . . . . . . . . . . Casting Extraction and Insert Loading . . . . . . . Automatic Spray . . . . . . . . . . . . . . . . . . . . . . Finishing Automation . . . . . . . . . . . . . . . . . . . Die Casting Internal Quality Check Automation
667 667 669 670 670 671 674 674 675 678 680 680 684
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High-Pressure Die Casting. . . . . . . . . . . . . . . . . . . . . . . . . John L. Jorstad, J.L.J. Technologies, Inc High-Pressure Die Casting . . . . . . . . . . . . . . . . . . . . . . . . . . Die Casting Alloys and Processes . . . . . . . . . . . . . . . . . Advantages of High-Pressure Die Casting. . . . . . . . . . . . Disadvantages of High-Pressure Die Casting. . . . . . . . . . Product Design for the Process . . . . . . . . . . . . . . . . . . . Metal Injection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Hot Chamber Die Casting . . . . . . . . . . . . . . . . . . . . . . . . . . Melting Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Injection Components. . . . . . . . . . . . . . . . . . . . . . . . . . Distinctions Between Hot and Cold Chamber Processes . . Gate and Runner Design. . . . . . . . . . . . . . . . . . . . . . . . Temperature Control . . . . . . . . . . . . . . . . . . . . . . . . . . Ejection and Postprocessing . . . . . . . . . . . . . . . . . . . . . Cold Chamber Die Casting . . . . . . . . . . . . . . . . . . . . . . . . . Machine Components . . . . . . . . . . . . . . . . . . . . . . . . . . Process Parameters. . . . . . . . . . . . . . . . . . . . . . . . . . . . Squeeze Casting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Squeeze Casting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Casting and Tooling Design for Squeeze Casting . . . . . . Factors Affecting Squeeze-Cast Products . . . . . . . . . . . . Squeeze-Cast Applications . . . . . . . . . . . . . . . . . . . . . . Commonly Used Alloys, Properties, and Microstructures . Vacuum High-Pressure Die Casting . . . . . . . . . . . . . . . . . . . Conventional Die Casting . . . . . . . . . . . . . . . . . . . . . . . Vacuum Die Casting . . . . . . . . . . . . . . . . . . . . . . . . . . High-Vacuum Die Casting . . . . . . . . . . . . . . . . . . . . . .
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689 690 690 691 691 692 693 695 697 697 698 698 699 700 700 704 706 709 709
. . . 713 . . . . . . . . . . . . . . . . . . . . . . . . . .
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715 715 716 717 717 717 719 719 719 720 721 721 722 724 724 725 727 727 727 728 729 730 732 732 732 732
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Semisolid Casting . . . . . . . . . . . . . . . . . . . . . . . . . John L. Jorstad, J.L.J. Technologies, Inc Semisolid Casting — Introduction and Fundamentals . Advantages of SSM Processing . . . . . . . . . . . . . SSM Processing Routes . . . . . . . . . . . . . . . . . . Thixocasting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Bar-Making Processes . . . . . . . . . . . . . . . . . . . Billet Sawing and Volume Control . . . . . . . . . . Reheating and Billet Handling. . . . . . . . . . . . . . Rheological Tests . . . . . . . . . . . . . . . . . . . . . . Basics of Semisolid Tooling Design. . . . . . . . . . Casting Process and Parameters. . . . . . . . . . . . . Rheocasting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Thixomolding. . . . . . . . . . . . . . . . . . . . . . . . . . . . . Process Description . . . . . . . . . . . . . . . . . . . . . Emerging Developments . . . . . . . . . . . . . . . . . . Mechanical Properties and Microstructure . . . . .
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Cast Irons . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Thomas E. Prucha, American Foundry Society Introduction to Cast Irons . . . . . . . . . . . . . . . . . . . Classification of Cast Irons . . . . . . . . . . . . . . . Composition of Cast Irons . . . . . . . . . . . . . . . Solidification of Cast Irons . . . . . . . . . . . . . . . Graphite Morphologies . . . . . . . . . . . . . . . . . . Matrix Structure Development. . . . . . . . . . . . . Physical Properties of Cast Irons . . . . . . . . . . . Thermal Properties. . . . . . . . . . . . . . . . . . . . . Conductive Properties. . . . . . . . . . . . . . . . . . . Magnetic Properties . . . . . . . . . . . . . . . . . . . . Acoustic Properties . . . . . . . . . . . . . . . . . . . . Heat Treatment . . . . . . . . . . . . . . . . . . . . . . . Welding of Cast Irons . . . . . . . . . . . . . . . . . . Machining and Grinding . . . . . . . . . . . . . . . . . Coatings . . . . . . . . . . . . . . . . . . . . . . . . . . . . Materials Selection . . . . . . . . . . . . . . . . . . . . Cast Iron Foundry Practices . . . . . . . . . . . . . . . . . . Cast Iron Melting Practice . . . . . . . . . . . . . . . Gray Iron Foundry Practice. . . . . . . . . . . . . . . Ductile Iron Foundry Practice . . . . . . . . . . . . . Compacted Graphite Iron Foundry Practice. . . . Malleable Iron Foundry Practice . . . . . . . . . . . Foundry Practice for High-Nickel Ductile Irons Foundry Practice High-Silicon Ductile Irons . . . Foundry Practice for High-Silicon Gray Irons . . Foundry Practice for High-Alloy White Irons . . Gray Iron Castings . . . . . . . . . . . . . . . . . . . . . . . . Classes of Gray Iron . . . . . . . . . . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . . . . . . . Castability. . . . . . . . . . . . . . . . . . . . . . . . . . .
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734 734 735 737 740 742 747 747 748 749 751 753 754 755 757
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761 761 762 764 764 766 767 768 769 770 773 773 777 777 778 780
. . . . . . . . . . 783 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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785 785 788 791 792 794 794 795 797 800 801 802 805 806 809 809 812 812 815 820 826 829 830 830 830 831 835 835 835 835
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
Microstructure . . . . . . . . . . . . . . . . . . . . . . . . . . Section Sensitivity . . . . . . . . . . . . . . . . . . . . . . . Prevailing Sections. . . . . . . . . . . . . . . . . . . . . . . Test Bar Properties . . . . . . . . . . . . . . . . . . . . . . Specifications . . . . . . . . . . . . . . . . . . . . . . . . . . Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Fatigue Limit in Reversed Bending . . . . . . . . . . . Pressure Tightness . . . . . . . . . . . . . . . . . . . . . . . Impact Resistance . . . . . . . . . . . . . . . . . . . . . . . Machinability . . . . . . . . . . . . . . . . . . . . . . . . . . Wear . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Resistance to Wear . . . . . . . . . . . . . . . . . . . . . . Elevated-Temperature Properties . . . . . . . . . . . . . Dimensional Stability . . . . . . . . . . . . . . . . . . . . . Effect of Shakeout Practice. . . . . . . . . . . . . . . . . Alloying to Modify As-Cast Properties. . . . . . . . . Heat Treatment . . . . . . . . . . . . . . . . . . . . . . . . . Physical Properties. . . . . . . . . . . . . . . . . . . . . . . Ductile Iron Castings . . . . . . . . . . . . . . . . . . . . . . . . Classes and Grades of Ductile Iron . . . . . . . . . . . Factors Affecting Mechanical Properties. . . . . . . . Hardness Properties . . . . . . . . . . . . . . . . . . . . . . Tensile Properties . . . . . . . . . . . . . . . . . . . . . . . Shear and Torsional Properties . . . . . . . . . . . . . . Compressive Properties . . . . . . . . . . . . . . . . . . . Fatigue Properties . . . . . . . . . . . . . . . . . . . . . . . Fracture Toughness . . . . . . . . . . . . . . . . . . . . . . Physical Properties. . . . . . . . . . . . . . . . . . . . . . . Austempered Ducfile Iron (ADI) . . . . . . . . . . . . . Compacted Graphite Iron Castings . . . . . . . . . . . . . . . Graphite Morphology . . . . . . . . . . . . . . . . . . . . . Chemical Composition . . . . . . . . . . . . . . . . . . . . Castability. . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mechanical Properties and Corrosion Resistance . . Applications . . . . . . . . . . . . . . . . . . . . . . . . . . . Malleable Iron Castings. . . . . . . . . . . . . . . . . . . . . . . Melting Practices . . . . . . . . . . . . . . . . . . . . . . . . Microstructure . . . . . . . . . . . . . . . . . . . . . . . . . . Current Production Technologies . . . . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . . . . . . . . . Ferritic Malleable Iron . . . . . . . . . . . . . . . . . . . . Pearlitic and Martensitic Malleable Iron . . . . . . . . White Iron and High-Alloyed Iron Castings. . . . . . . . . Nickel-Chromium White Irons . . . . . . . . . . . . . . High-Chromium White Irons. . . . . . . . . . . . . . . . High-Alloy Graphitic Irons . . . . . . . . . . . . . . . . . . . . High-Silicon Irons for High-Temperature Service . Austenitic Nickel-Alloyed Gray and Ductile Irons . High-Silicon Irons for Corrosion Resistance . . . . .
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Steel Castings . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Raymond W. Monroe, Steel Founders’ Society of America; Thomas E. Prucha, American Foundry Society Steel Ingot Casting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Pouring Method. . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ingot Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Hot Tops . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ingot Molds and Stools . . . . . . . . . . . . . . . . . . . . . . . Ingot Solidification. . . . . . . . . . . . . . . . . . . . . . . . . . . Ingot Stripping . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Solidification Simulation. . . . . . . . . . . . . . . . . . . . . . . Steel Continuous Casting. . . . . . . . . . . . . . . . . . . . . . . . . . Historical Aspects of Continuous Casting of Steel . . . . . General Description of the Process . . . . . . . . . . . . . . . Plant Layout . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Near-Net Shape Casting . . . . . . . . . . . . . . . . . . . . . . . Tundish Metallurgy . . . . . . . . . . . . . . . . . . . . . . . . . .
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836 837 838 839 840 841 843 844 846 846 848 848 849 849 850 851 851 853 ... 856 857 859 859 861 861 862 865 867 869 872 872 872 873 874 879 884 884 885 887 888 889 890 896 896 899 904 904 906 907
. . . . 909 . . . . . . . . . . . . . .
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911 911 912 913 913 914 916 916 918 918 918 919 919 920
Pouring Stream Protection . . . . . . . . . . . . . . . . . . Molds for Continuous Casting of Steel. . . . . . . . . . Productivity Improvements . . . . . . . . . . . . . . . . . . Quality Improvements . . . . . . . . . . . . . . . . . . . . . Horizontal Continuous Casting . . . . . . . . . . . . . . . Future Developments . . . . . . . . . . . . . . . . . . . . . . Shape Casting of Steel . . . . . . . . . . . . . . . . . . . . . . . . Castability of Steel. . . . . . . . . . . . . . . . . . . . . . . . Melting Furnaces . . . . . . . . . . . . . . . . . . . . . . . . . Metlting Practice . . . . . . . . . . . . . . . . . . . . . . . . . Pouring Practice . . . . . . . . . . . . . . . . . . . . . . . . . Foundry Factors in Design . . . . . . . . . . . . . . . . . . Design Factors. . . . . . . . . . . . . . . . . . . . . . . . . . . Avoidance of Hot Spots . . . . . . . . . . . . . . . . . . . . Thin-Wall Steel Castings . . . . . . . . . . . . . . . . . . . Tolerances . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Size of Cores . . . . . . . . . . . . . . . . . . . . . . . . . . . Casting in Graphite Molds . . . . . . . . . . . . . . . . . . Steel Castings Properties . . . . . . . . . . . . . . . . . . . . . . . Carbon Steels . . . . . . . . . . . . . . . . . . . . . . . . . . . Low-Alloy Steels . . . . . . . . . . . . . . . . . . . . . . . . . Wear-Resistant Steels . . . . . . . . . . . . . . . . . . . . . . Corrosion-Resistant Steels. . . . . . . . . . . . . . . . . . . Heat-Resistant Steels . . . . . . . . . . . . . . . . . . . . . . Common Alloys . . . . . . . . . . . . . . . . . . . . . . . . . Carbon and Low-Alloy Mechanical Properties . . . . Mechanical Properties of Wear-Resisting Steels . . . Mechanical Properties of Corrosion-Resisting Steels Mechanical Properties of Heat-Resisting Steels . . . . Ferrite in Cast Stainless Steels . . . . . . . . . . . . . . . Carbon and Low-Alloy Heat Treatments . . . . . . . . High-Alloy Heat Treatment. . . . . . . . . . . . . . . . . . Corrosion-Resistant Applications . . . . . . . . . . . . . . Heat-Resistant Applications . . . . . . . . . . . . . . . . . Selection and Evaluation of Steel Castings . . . . . . . . . . Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Manufacturing . . . . . . . . . . . . . . . . . . . . . . . . . . . Nondestructive Examination . . . . . . . . . . . . . . . . . Purchasing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Research and Development . . . . . . . . . . . . . . . . . .
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921 922 923 924 924 924 926 926 928 930 933 934 936 938 943 946 947 948 949 949 950 950 951 952 953 953 960 961 966 968 970 971 972 973 975 975 980 982 986 986
Casting of Nonferrous Alloys . . . . . . . . . . . . . . . . . . . . . Nonferrous Casting — An Introduction . . . . . . . . . . . . . . . Nickel Alloys . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Titanium Alloys . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Lead. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Dross, Melt Loss, and Fluxing of Light Alloy Melts. . . . . . Dross Formation . . . . . . . . . . . . . . . . . . . . . . . . . . . Influence of Melter Type on Dross Generation . . . . . . Influence of Charge Materials on Melt Loss . . . . . . . . Influence of Operating Practices on Melt Loss . . . . . . Economic Implications of Dross . . . . . . . . . . . . . . . . In-Plant Enhancement or Recovery of Dross Metallics . Ways to Reduce Dross Formation . . . . . . . . . . . . . . . Aluminum Alloy Ingot Casting and Continuous Processes. . Ingot Forms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Molten-Metal Processing . . . . . . . . . . . . . . . . . . . . . Ingot Casting Processes . . . . . . . . . . . . . . . . . . . . . . Continuous Processes . . . . . . . . . . . . . . . . . . . . . . . . Solidification in the DC Process . . . . . . . . . . . . . . . . Postsolidification Processes . . . . . . . . . . . . . . . . . . . . Safety. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Aluminum Shape Casting . . . . . . . . . . . . . . . . . . . . . . . . Melting and Melt Handling . . . . . . . . . . . . . . . . . . . . Casting Process Selection . . . . . . . . . . . . . . . . . . . . . Pressure Die Casting of Aluminum . . . . . . . . . . . . . . Premium Engineered Castings . . . . . . . . . . . . . . . . . .
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987 989 989 989 990 992 992 993 994 994 998 998 999 1001 1001 1001 1002 1004 1006 1007 1007 1009 1009 1010 1013 1016
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© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
Copper Continuous Casting . . . . . . . . . . . . . . History . . . . . . . . . . . . . . . . . . . . . . . . . Vertical Continuous Casting . . . . . . . . . . Horizontal Continuous Casting . . . . . . . . Strip Casting . . . . . . . . . . . . . . . . . . . . . Wheel Casting . . . . . . . . . . . . . . . . . . . . Ohno Continuous Casting Process . . . . . . Casting of Copper and Copper Alloys . . . . . . . Castability. . . . . . . . . . . . . . . . . . . . . . . Melting and Melt Control . . . . . . . . . . . . Fluxing of Copper Alloys . . . . . . . . . . . . Degassing of Copper Alloys . . . . . . . . . . Deoxidation of Copper Alloys . . . . . . . . . Grain Refining of Copper Alloys . . . . . . . Filtration of Copper Alloys . . . . . . . . . . . Melt Treatments for Group I to III Alloys Production of Copper Alloy Castings . . . . Casting Process Selection . . . . . . . . . . . . Gating . . . . . . . . . . . . . . . . . . . . . . . . . Feeding . . . . . . . . . . . . . . . . . . . . . . . . Casting of Zinc Alloys . . . . . . . . . . . . . . . . . Control of Alloy Composition . . . . . . . . . Melt Processing . . . . . . . . . . . . . . . . . . . Die Casting of Zinc Alloys . . . . . . . . . . . Other Casting Processes for Zinc Alloys. .
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Nonferrous Alloy Castings . . . . . . . . . . . . . . . . . . . . . . . Mahi Sahoo, CANMET Materials Technology Laboratory Aluminum and Aluminum Alloy Castings . . . . . . . . . . . . . Chemical Compositions . . . . . . . . . . . . . . . . . . . . . . Effects of Alloying and Impurity Elements . . . . . . . . . Structure Control . . . . . . . . . . . . . . . . . . . . . . . . . . . Modification and Refinement of Aluminum-Silicon Alloys . . . . . . . . . . . . . . . . . Foundry Alloys for Specific Casting Applications . . . . Heat Treatment . . . . . . . . . . . . . . . . . . . . . . . . . . . . Welding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Properties of Aluminum Casting Alloys . . . . . . . . . . . Copper and Copper Alloy Castings . . . . . . . . . . . . . . . . . . Alloying Additions. . . . . . . . . . . . . . . . . . . . . . . . . . Alloying Systems and Specifications . . . . . . . . . . . . . Physical Metallurgy . . . . . . . . . . . . . . . . . . . . . . . . . Types of Copper Castings. . . . . . . . . . . . . . . . . . . . . Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Heat Treatment . . . . . . . . . . . . . . . . . . . . . . . . . . . . Welding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Cupronickels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Zinc and Zinc Alloy Castings . . . . . . . . . . . . . . . . . . . . . Alloying Additions. . . . . . . . . . . . . . . . . . . . . . . . . . Physical Metallurgy . . . . . . . . . . . . . . . . . . . . . . . . . Die Casting and Aging . . . . . . . . . . . . . . . . . . . . . . . Specialty Alloys . . . . . . . . . . . . . . . . . . . . . . . . . . . Finishing of Die Castings . . . . . . . . . . . . . . . . . . . . . Metalworking and Machining . . . . . . . . . . . . . . . . . . Joining . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Applications for Zinc Die Castings . . . . . . . . . . . . . . Magnesium and Magnesium Alloys . . . . . . . . . . . . . . . . . Magnesium Alloys . . . . . . . . . . . . . . . . . . . . . . . . . . General Applications . . . . . . . . . . . . . . . . . . . . . . . . Melting Furnaces and Auxiliary Pouring Equipment . . Melting Procedures and Process Parameters . . . . . . . . Sand Casting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Investment Casting. . . . . . . . . . . . . . . . . . . . . . . . . . Permanent Mold Casting . . . . . . . . . . . . . . . . . . . . . Die Casting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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1019 1019 1020 1022 1023 1024 1024 1026 1026 1027 1029 1032 1035 1037 1037 1038 1039 1040 1041 1044 1049 1049 1049 1051 1054
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1059 1059 1059 1063
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1066 1069 1071 1077 1077 1085 1085 1087 1088 1091 1091 1091 1091 1091 1095 1095 1096 1097 1097 1098 1098 1099 1097 1099 1100 1102 1102 1103 1106 1109 1109 1109
Direct Chill Casting . . . . . . . . . . . . . . . . . . . . . High-Temperature Alloys . . . . . . . . . . . . . . . . . Welding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Cobalt and Cobalt Alloy Castings. . . . . . . . . . . . . . . Physical Metallurgy . . . . . . . . . . . . . . . . . . . . . Foundry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Postcasting Treatment . . . . . . . . . . . . . . . . . . . Applications . . . . . . . . . . . . . . . . . . . . . . . . . . Nickel and Nickel Alloy Castings. . . . . . . . . . . . . . . Compositions . . . . . . . . . . . . . . . . . . . . . . . . . Structure and Property Correlations . . . . . . . . . . Melting Practice and Metal Treatment . . . . . . . . Foundry Practice . . . . . . . . . . . . . . . . . . . . . . . Pouring Practice . . . . . . . . . . . . . . . . . . . . . . . Gating Systems . . . . . . . . . . . . . . . . . . . . . . . . Risers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Welding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Heat Treatment . . . . . . . . . . . . . . . . . . . . . . . . Specific Applications . . . . . . . . . . . . . . . . . . . . Titanium and Titanium Alloy Castings . . . . . . . . . . . Historical Perspective of Casting Technology . . . Physical Metallurgy . . . . . . . . . . . . . . . . . . . . . Alloy Ti-6Al-4V . . . . . . . . . . . . . . . . . . . . . . . Casting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Postcasting Practice . . . . . . . . . . . . . . . . . . . . . Product Applications . . . . . . . . . . . . . . . . . . . . Zirconium and Zirconium Alloy Castings . . . . . . . . . Zirconium Reactivity Considerations . . . . . . . . . Melting Processes . . . . . . . . . . . . . . . . . . . . . . Casting Processes. . . . . . . . . . . . . . . . . . . . . . . Processing. . . . . . . . . . . . . . . . . . . . . . . . . . . . Physical Properties of Zirconium Castings . . . . . Impact Properties of Zirconium Castings . . . . . . Synthesis and Processing of Cast Metal-Matrix Composites and Their Applications . . . . . . . . . . Classification of MMCs . . . . . . . . . . . . . . . . . . Solidification Processing of MMCs and Possible Effects of Reinforcement on Solidification. . Property Motivation for Using MMCs . . . . . . . . Components of MMCs Currently in Use. . . . . . . Development of Select Cast MMCs . . . . . . . . . . Research Imperatives . . . . . . . . . . . . . . . . . . . .
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1112 1112 1112 1114 1114 1116 1117 1117 1119 1119 1121 1123 1124 1124 1124 1125 1125 1125 1126 1128 1129 1129 1132 1134 1137 1138 1143 1143 1144 1144 1145 1147 1147
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Quality Assurance, Nondestructive Evaluation, and Failure Analysis. . . . . . . . . . . . . . . . . . . . . . . Babu DasGupta, National Science Foundation Approaches to Measurement of Metal Quality. . . . . . Metal Cleanliness Assessment Techniques . . . . . . . . Nondestructive Testing of Components . . . . . . . . . . . . . . Liquid Penetrant Inspection. . . . . . . . . . . . . . . . . . . Radiographic Inspection . . . . . . . . . . . . . . . . . . . . . Fluoroscopic Inspection and Automated Defect Recognition . . . . . . . . . . . . . . . . . . . . . Ultrasonic Inspection . . . . . . . . . . . . . . . . . . . . . . . Eddy Current Inspection . . . . . . . . . . . . . . . . . . . . . Process-Controlled Resonant Testing (Contributed by Quasar International, Inc.) . . . . . . . . . . . . . . . . Leak Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Electrical Conductivity Measurement . . . . . . . . . . . . Casting Fracture Characteristics . . . . . . . . . . . . . . . . . . . General Fracture Examination . . . . . . . . . . . . . . . . . Fracture Features . . . . . . . . . . . . . . . . . . . . . . . . . . Casting Failure Analysis Techniques and Case Studies . . . Introduction. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Background Information . . . . . . . . . . . . . . . . . . . . .
xvii
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1150 1155 1156 1160 1162
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1167 1167 1174 1174 1174
. . . . . 1174 . . . . . 1174 . . . . . 1175 . . . . . . . . .
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1175 1177 1177 1178 1178 1179 1183 1183 1183
© 2008 ASM International. All Rights Reserved. ASM Handbook Volume 15: Casting (#05115G)
Planning and Sample Selection . . . . . . . . . . . . . . . . Preliminary Examination . . . . . . . . . . . . . . . . . . . . Material Evaluation . . . . . . . . . . . . . . . . . . . . . . . . Data Analysis and Report Preparation . . . . . . . . . . . Case Studies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Common Defects in Various Casting Processes . . . . . . . . Case Studies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Repair Welding of Castings . . . . . . . . . . . . . . . . . . . . . . Repair Welding of Ferrous Castings. . . . . . . . . . . . . Repair Welding of Nonferrous Castings . . . . . . . . . . Quality Planning Tools and Procedures . . . . . . . . . . . . . . Advanced Product Quality Planning (APQP). . . . . . . Maintenance, Repair, Alterations, and Storage of Patterns and Tooling. . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
www.asminternational.org
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1184 1184 1184 1187 1187 1192 1199 1203 1203 1205 1207 1207
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Ownership . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Maintenance Responsibility. . . . . . . . . . . . . . . . . . . . . Common Failure Mechanisms . . . . . . . . . . . . . . . . . . . Maintenance Interval . . . . . . . . . . . . . . . . . . . . . . . . . Maintenance Records . . . . . . . . . . . . . . . . . . . . . . . . . Preventive Maintenance Program. . . . . . . . . . . . . . . . . Repair . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Alterations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Storage—Physical Protection, Cataloging, and Tracking. Reference Information. . . Metric Conversion Guide. . Abbrieviations & Symbols . Index. . . . . . . . . . . . . . . .
xviii
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1213 1213 1213 1213 1214 1214 1214 1214 1215
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1217 1219 1222 1225
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 3-15 DOI: 10.1361/asmhba0005186
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
History and Trends of Metal Casting Thomas E. Prucha, American Foundry Society Daniel Twarog, North American Die Casting Association Raymond W. Monroe, Steel Founders’ Society of America
CASTING is one of the oldest manufacturing methods known to humankind and a very direct method of producing metal parts. It involves pouring molten metal into a cavity that is close to the final dimensions of the finished form, and it is used to manufacture many types of complex components of any metal, ranging in weight from less than an ounce to single parts weighing several hundred tons. Typical uses of cast components include infrastructure and structural components, water distribution systems (pipes, pumps, and valves), automotive Table 1
components (engine blocks, brakes, steering and suspension components, etc.), prosthetics, jewelry, and gas turbine engine hardware. In one study (Ref 1), castings were estimated to be used in 90% or more of all manufactured goods and in all capital goods machinery used in manufacturing. Some of the major market areas of castings are listed in Table 1 (Ref 2). Economically, casting processes are capable of producing highly reliable, cost-effective components ranging from low-volume, singlepart prototype production runs to economies
of scale for millions of parts. In terms of component design, casting also permits the formation of streamlined, intricate, integral parts with varied shape and size capabilities for designers. Metal casting is very flexible in terms of configuration design, and if a pattern can be made for a part, it can be cast. For example, the flexibility of metal casting, particularly sand molding, may permit the use of difficult design techniques, such as undercuts and curved or reflex contours that are not possible with other high-production processes. Tapered
Other major markets include machine tools, construction
Major markets for metal castings
Ferrous castings Gray iron Ingot molds Construction castings Motor vehicles Farm equipment Engines Refrigeration and heating Construction machinery Valves Soil pipe Pumps and compressors Pressure pipe Other major markets include machine tools, mechanical power transmission equipment, hardware, home appliances, mining machinery, and oil and natural gas pumping and processing equipment(a) Malleable iron Motor vehicles Valves and fittings Construction machinery Railroad equipment Engines Mining equipment Hardware Other major markets include heating and refrigeration, motors and generators, fasteners, ordnance, chains, machine tools, and general industrial machinery(a) Ductile iron Pressure pipe Motor vehicles
Farm machinery Engines Pumps and compressors Valves and fittings Metalworking machinery Construction machinery Other major markets include textile machinery, woodworking and paper machinery, mechanical power and transmission equipment, motors and generators, refrigeration and heating equipment, and air conditioning(a)
Steel Railroad equipment Construction equipment Mining machinery Valves and fittings General and special industrial machinery Motor vehicles Metalworking machinery Other major markets include steel manufacturing, spring goods, heating and air conditioning, recreation equipment, industrial materials handling equipment, ships and boats, and aircraft and aerospace(a) Nonferrous castings Aluminum Auto and light truck Aircraft and aerospace Other transportation Engines Household appliances Office machinery Power tools Refrigeration, heating, and air conditioning
equipment, mining equipment, farm machinery, electronic and communication equipment, power systems, and motors and generators(a) Magnesium Power tools Sporting goods Anodes Automotive Other major markets include office machinery, health care, and aircraft and aerospace(a)
Copper base Valves and fittings Plumbing brass goods Electrical equipment Pumps and compressors Power transmission equipment General machinery Transportation equipment Other major markets include chemical processing, utilities, desalination, and petroleum refining(a)
Zinc Automotive Building hardware Electrical components Machinery Household appliances Other major markets include scientific instruments, medical equipment, radio and television equipment, audio components, toys, and sporting goods(a)
(a)Ranked in approximate order of tonnage shipped, but in some cases, the total of other major markets is larger as a whole than the individual markets listed. Source: Ref 2
4 / Applications and Specification Guidelines sections with thickened areas for bosses and generous fillets are routine. The inherent design freedom of metal casting also allows the designer to combine what would otherwise be several parts of a fabrication into a single piece. This is significant when exact alignment must be held, as in high-speed machinery, machine tool parts, or engine end plates and housings that carry shafts. Various surface finishes on a part are also possible, ranging from a rougher as-cast surface of sand-molded casting to the smoother surfaces obtained through shell molding, investment casting, or other casting methods. In general, casting can provide various functional advantages for component design that include: Design both internal and external contours
independently to almost any requirement
Place metal in exact locations where it is
needed for rigidity, wear, corrosion, or maximum endurance under dynamic stress Produce a complex part as a single, dependable unit Readily achieve an attractive appearance This article briefly introduces the history of metal casting, general capabilities and methods of casting, and industry trends. In terms of value and volume, metal casting ranks second
only to steel rolling in the metal-producing industry. Metal casting is one of the ten largest industries when rated on a value-added basis, and worldwide casting shipments have continued to increase in recent years. Production statistics for 2005 are summarized in Table 2 (Ref 3), with comparison to 2004 production statistics from Modern Casting’s “Annual Census of World Casting Production.”
History of Metal Casting The casting of metal is a prehistoric technology. Exactly when the casting of metals began is not known, but it appears relatively late in the archaeological record. Archaeologists give the name Chalcolithic to the period in which metals were first being mastered, and they date this period, which immediately preceded the Bronze Age, very approximately to between 5000 and 3000 B.C. (Table 3, Ref 4). There were many earlier fire-using technologies, collectively called pyrotechnology, that provided a basis for the development of metal casting. Among these were the heat treatment of stone to make it more workable, the burning of lime to make plaster, and the firing of clay to produce ceramics. At first, it did not include
smelting, because the metal of the earliest castings appears to have been native. Cast iron appeared in China in approximately 600 B.C. Its use was not limited to strictly practical applications, and there are many examples of Chinese cast iron statuary. Most Chinese cast irons were unusually high in phosphorus and, because coal was often used in smelting, high in sulfur as well. These irons therefore have melting points that are similar to those of bronze, and when molten are unusually fluid. The iron castings, like the Chinese cast bronzes, are often remarked upon for the thinness of their wall sections. From these early beginnings, development and application of metal casting continued (Table 4, Ref 5). During the wars of medieval times, metalcasters produced cannons. In peace, they recast the metal into bells. In the Middle Ages, church bells were cast by priests, abbots, or bishops who also were trained metal founders. As the metal was melted, the brethren stood around the furnace intoning psalms and prayers. The molten metal was then blessed, and divine protection was asked for the bell, which usually bore the name of a saint (Ref 5). During the 1800s, there were significant advances in technology. Metallurgy also began
Table 2 World casting production for 2004–2005 40th census of world casting production for 2005, metric tons Country
Austria Belgium Brazil Canada China Croatia Czech Republic Denmark Finland France Germany Great Britain Hungary India Italy Japan Korea Lithuania Mexico Netherlands(a) Norway Poland Portugal Romania Russia Slovenia(a) Slovakia South Africa Spain Sweden Switzerland Taiwan Thailand Turkey Ukraine (i) United States Totals
Gray iron
47,500 70,700(a) 2,460,000(b)(d) 483,000(a)(b)(d) 12,303,963 25,306 269,493 46,932 47,817 896,500 2,469,058 531,000 35,307 4,116,000 923,700 2,782,509 960,100 14,300 600,000 63,100 16,348 420,700 27,285 81,578 3,480,000 75,900 41,520 160,000 491,600 171,700(b) 26,400 827,932 170,000 567,000 626,610 4,457,905 40,788,763
Ductile iron
131,000 15,600(a)(b) ... ... 5,838,753 16,252 49,524 40,333 60,123 957,700 1,487,234 362,000 32,200 618,000 512,200 1,919,435 565,200 200 270,000 78,241 49,207 112,500 69,792 10,500 720,000 27,800(b) 8,760 65,000 551,600 83,300 40,400 220,413 70,000 327,000 40,000 4,241,088 19,591,355
Malleable iron
17,500 ... ... ... 514,129 50 5,435 0 0 1,400 52,713 12,000 10 56,000 8,400 57,851 46,500 ... 2,000 6,209 0 21,740 0 926 280,000 ... ... 10,000 17,400 ... ... ... 30,000 8,000 10,000 75,296 1,233,559
Steel
18,000 42,100 293,891 117,600(a) 3,224,374 2,219 109,680 0 18,532 113,700 200,409 112,000 5,385 805,000 73,324 276,589 149,600 30 75,000 438 3,723 59,560 12,096 37,900 1,200,000 22,700 4,300 136,000 83,200 23,400 ... 73,619 30,000 125,000 266,060 1,287,295 9,002,724
Copper base
5,459 ... ... 18,585(a) 416,097 1,120 1,724 1,126 3,999 26,264 84,386 15,000 2,110 0 16,800 97,794 23,200 5 180,000 ... 4,667 6,300 8,700 2,700 160,000 6,727(f) 2,160 14,900 7,886 11,200 2,689 41,959 28,600 16,000 11,000 292,113 1,511,270
Aluminum
103,347 19,428(c) 240,457(e) 324,973 1,886,000(c) 15,038 80,252(c) 4,000 6,565 318,445(c) 727,139 206,000(c) 75,720 516,000 857,500 1,477,349(c) 144,300 56 660,000 ... 22,926 174,300 20,250 20,000 920,000 30,183 26,260 56,000 139,312(c) 44,400 17,785 250,366 100,000 66,500 20,500 2,080,174 11,718,025
Magnesium
6,580 ... ... ... ... 0 ... 0 0 ... 27,283 ... 3,050 0 12,000 ... 10,300(f) ... 500 ... 0 30 150 10 70,000 ... ... 0 ... 1,700 ... 6,427 ... 500 ... 100,697 239,227
Zinc
12,997 ... ... ... 237,889 502 2,656 1 434 25,066 56,439 21,000 3,050 0 70,000 36,216 ... ... 350 ... 0 8,600 900 700 20,000 ... 1,800 2,600 16,051 5,900 1,616 69,316 16,900 12,700 ... 312,978 936,661
Other nonferrous
... ... ... ... ... 1,300 279 5.022 0 3,461 3,346 2,000 0 0 67,600 8,102 0 ... ... ... ... 770 120 5 50,000 ... 5 0 433 ... ... 4,418 ... ... ... 48,987 195,848
Total
342,383 124,507(a) 2,968,600 863,373(a) 24,421,205 61,787 519,043 97,414 137,470 2,342,536 5,108,007 1,261,000 156,832 6,111,000 2,541,524 6,655,845 1,899,200 14,591 1,787,850 147,988(g) 96,871 804,500 139,293 154,319 7,620,000(h) 163,310 84,805 444,500 1,307,482 341,600 88,890 1,494,450 445,500 1,122,700 974,170 12,896,533 85,741,078
Totals for 2004
325,205 124,507 2,829,916 863,373 22,420,452 56,449 522,400 95,949 133,325 2,465,617 4,984,473 1,273,000 164,892 4,623,000 2,441,282 6,386,449 1,857,300 18,715 2,185,200 147,988 86,757 804,500 122,750 207,023 6,300,000 163,310 4,800 444,500 1,309,249 321,100 104,077 1,451,768 235,850 982,000 974,170 12,314,121 79,745,467
Note: “. . .” indicates unreported data. (a) 2004 tonnage. (b) Includes malleable iron. (c) Includes magnesium. (d) Includes ductile iron. (e) Includes all nonferrous. (f) Includes zinc. (g) Includes only ferrous. (h) Includes 720,000 tons of special iron. (i) 2002 tonnage. Source: Modern Casting
History and Trends of Metal Casting / 5 Table 3
Chronological list of developments in the use of materials
Date
Development
9000 B.C. 6500 B.C. 5000–3000 B.C.
3000–1500 B.C.
3000–2500 B.C. 2500 B.C. 2400–2200 B.C. 2000 B.C. 1500 B.C.
Earliest metal objects of wrought native copper Earliest life-size statues, of plaster Chalcolithic period: melting of copper; experimentation with smelting Bronze Age: arsenical copper and tin bronze alloys Lost-wax casting of small objects Granulation of gold and silver and their alloys Copper statue of Pharoah Pepi I Bronze Age Iron Age (wrought iron)
Location
Near East Jordan Near East
Near East
Near East Near East Egypt Far East Near East
Date
700–600 B.C. 600 B.C. 224 B.C.
Development
Etruscan dust granulation Cast iron Colossus of Rhodes destroyed 200–300 A.D. Use of mercury in gilding (amalgam gilding) 1200–1450 A.D. Introduction of cast iron (exact date and place unknown) Circa 1122 A.D. Theophilus’s On Divers Arts, the first monograph on metalworking written by a craftsman 1252 A.D. Diabutsu (Great Buddha) cast at Kamakura Circa 1400 A.D. Great Bell of Beijing cast 16th century Sand introduced as mold material
Location
Italy China Greece
Date
Development
1709
Cast iron produced with coke as fuel, Coalbrookdale Boring mill or cannon developed Great Bell of the Kremlin cast Cast steel developed by Benjamin Huntsman Cast iron used as architectural material, Ironbridge Gorge Zinc statuary Electrodeposition of copper
1715 Roman world Europe
1735 1740
Germany
1779
Japan
1826 1838
China France
1884
Electrolytic refining of aluminum
Location
England
Switzerland Russia England England
France Russia, England United States, France
Source: Ref 4
Table 4
Timeline of casting developments to 1899
B.C. 9000 B.C.—Earliest metal objects of wrought native copper are produced in the Near East. 3200 B.C.—The oldest casting in existence, a copper frog, is cast in Mesopotamia. 3000 B.C.—Early metalcasters cast bronze tools and weapons in permanent stone molds. 3000–2500 B.C.—Small objects are cast via lost-wax (investment casting) process in the Near East. 1500 B.C.—Wrought iron is discovered in the Near East. 600 B.C.—The first cast iron object, a 270 kg (600 lb) tripod, is cast by the Chinese. 233 B.C—Iron plowshares are cast. 200 B.C.—Oldest iron castings still in existence are produced during the Han Dynasty. A.D. 500—Cast crucible steel is produced in India. 1200s—Loam or sweep molding is used by European metalcasters to cast bells for cathedrals. 1252—The colossal statue, the Great Buddha at Kamakura, Japan, is cast in high-lead tin bronze. The project began in the 700s, and its head alone weighed 140 tons. 1313—The first cannon is cast in bronze by a monk in the city of Ghent. 1400s—During the siege of Constantinople, heavy guns are cast from bronze “on the spot,” virtually under the walls of the besieged city. Movable, cast lead type for printing presses revolutionizes the world’s methods of communication. 1480—Vannoccio Biringuccio (1480–1539), the first true metalcaster and the “father of the metal casting industry,” is born. The founder of the Vatican, his De La Pirotechnia is the first written account of proper metal casting practice. 1500s—Sand is introduced as a molding material in France. 1600s 1612—Mined from under the sea, seacoal is mentioned for the first time by German metalcaster and inventor Simon Sturtevant. 1619—North America’s first iron furnace is built at Falling Creek, Va., on a branch of the James River, 100 km (60 miles) from the Jamestown colony. Three years later, Native Americans destroy it during a raid. 1645—Earliest recorded use of the term foundry appears in the Oxford English Dictionary in its variant “founderie.” 1646—America’s first iron metal casting facility (and second industrial plant), Saugus Iron Works, near Boston, pours the first American metal casting, the Saugus pot. The Saugus River site was selected by Richard Leader and was built to produce iron products for Massachusetts and England. 1661—First U.S. copper deposits are discovered by Gov. Winthrop in Middletown, Conn. Source: Ref 5; see also Selected References
1700s 1709—Two important developments by Abraham Darby from Coalbrookdale, England, improve casting methods. He developed the first true metal casting facility flask to modernize molding practices (which had been carried out in pits on the floor by use of pattern boards tied together or in crude box frames). He would later initiate the use of coke as a furnace fuel for iron production. 1722—A.F. de Reamur, recognized as the world’s first metallurgical chemist, develops malleable iron, known today as European whiteheart malleable. 1750—Benjamin Huntsman reinvents the cast crucible steel process in England, a process that disappeared after first being developed in India. The English parliament prohibits the refining of pig iron or the casting of iron in the American colonies, contributing to the start of the American Revolution. 1756—Coalbrookdale’s Richard Reynolds oversees the invention of cast iron tram-road rails, replacing wooden rails. 1775—Revolutionary patriot Paul Revere, who operated a bell-and-fittings foundry in Boston, rides from Boston to Lexington warning colonists of the British invasion. 1776—Metalcasters Charles Carroll, James Smith, George Taylor, James Wilson, George Ross, Philip Livingston, and Stephen Hopkins sign the American Declaration of Independence. 1779—First iron bridge ever erected (above England’s Severn River) was cast and constructed at Coalbrookdale Works. 1794—John Wilkinson of England invents the first metalclad cupola furnace, using a steam engine to provide the air blast. 1797—First cast plow in the United States is invented by Charles Newbold of Sauk, N.J. 1800s 1809—Centrifugal casting is developed by A.G. Eckhardt of Soho, England. 1815—The cupola furnace is introduced to the United States (Baltimore). 1817—The first iron water line in the United States, 120 m (400 ft) long, is laid in Philadelphia. 1818—The first U.S. cast steel is produced by the crucible process at historic Valley Forge Foundry. 1825—Aluminum, the most abundant metal in the Earth’s crust, is isolated from aluminum chloride by Denmark’s Hans Oerstad. 1830s—Seth Boyden of Newark, N.J., produces the first blackheart malleable iron in the United States. 1832—Nickel-bronze is produced commercially in England. 1837—The first reliable molding machine on the market is made and used by the S. Jarvis Adams Co., Pittsburgh. 1847—Cast steel guns are made by Krupp Works in Germany. Asa Whitney, Philadelphia, obtains a patent on
a process for annealing chilled-iron car wheels cast with chilled tread and flange. 1847—John Deere commissions Jones and Quiggs Steel Works, Pittsburgh, to cast and roll a steel plow, which it accomplishes at one-half the previous cost of the product. 1849—A manually operated die casting machine is patented to supply rapidly cast lead type for newspapers. 1850—Drop-bottom cupola is developed. 1863—Metallography is developed by Henry C. Sorby of Sheffield, England, enabling metalcasters to polish, etch, and microscopically examine metal surfaces for physical analysis. 1867—James Nasmythe, inventor of the steam hammer, develops a gear-tilted safety ladle to prevent pouring accidents. 1870—Sandblasting is developed for large castings by R.E. Tilghman of Philadelphia. 1874—The Colliau cupola, the first commercially made cupola in America, is introduced. 1876—The first authenticated aluminum castings were produced by Col. William Frishmuth at his Philadelphia foundry. Assembled to produce an engineer’s transit, these castings were made from $1/oz chemically produced aluminum. Manganese-bronze patent is granted to Parsons in England. Tobin Bronze begins developing manganesebronze in the United States. 1880–1887—W.W. Sly, Cleveland, develops the first casting cleaning mill, greatly reducing hand-chipping and grinding to allow a custom-finished product. 1884—The first architectural application of aluminum, a cast aluminum pyramid that was produced by Frishmuth, is mounted on the tip of the Washington Monument. 1886—Charles M. Hall, a 22 year old student at Oberlin College, discovers a process of aluminum reduction through electrolysis. The invention replaced chemical reduction and lowered the metal cost (from $15/lb in 1884 to $0.50/lb in 1890), spurring a new industry of aluminum applications. 1887—Eli Millett invents a core oven for drying small cores in individual drawers. 1890—The first motor-driven mold conveyor is installed, integrating molding, pouring, and cooling operations. 1897—Iowa dentist B.F. Philbrook adapts the lost-wax investment casting process for producing dental inlays, the first nonart application of the process in the modern metal casting age. 1898—Poulson and Hargraves (U.K.) produce the first sand molds bonded with sodium silicate. Germany’s Imperial Navy recommends copper-nickel alloys containing 4–45% Ni for saltwater piping system. 1899—Electric arc furnace, developed by France’s Paul Heroult, begins commercial production.
6 / Applications and Specification Guidelines to achieve prominence in 1889 when nickel was alloyed to make steel stronger. In terms of casting, centrifugal casting was developed by A.G. Eckhardt of Soho, England, in 1809. This method involves the pouring of molten metal into a rapidly rotating metallic mold. The method was soon adopted by the pipe foundries and was first used in Baltimore, Maryland, in 1848. Sir Henry Bessemer, famed for his converter, used centrifugal casting to remove gases and was the first to pour two or more metals into a single rotating mold. The centrifugal casting of steel was first attempted in 1898 at the plant of the American Steel Foundries in St. Louis, Missouri. Railroad car wheels were spun cast in 1901 at a rotation speed of 620 rpm. Following the early development of the centrifugal method, a permanent mold method known as slush casting was introduced. Slush casting is a process in which molten metal is poured into a split metal mold (generally made of bronze) until the mold is filled; then, the mold is immediately inverted, and the metal that is still liquid is allowed to run out. The time required for this casting operation is sufficient to freeze a metal shell in the mold, corresponding to the shape of the cavity wall. The thickness of the wall of the casting depends on the time interval between the filling and the inverting of the mold, as well as on the chemical and physical properties of the alloy and the temperature and composition of the mold. Usually, lead and zinc alloy castings are
Table 5
produced by slush casting. The process is limited to the production of hollow castings and was used to produce lamp bases (Ref 6). Manually operated die casting machines were patented as early as 1849 (Sturgiss) and 1852 (Barr) in an effort to satisfy the insatiable demands of a growing reading public by way of rapidly cast lead type. These early inventions led to Ottmar Mergenthaler’s Linotype, an automatic casting machine in which molten lead is forced by piston stroke into a metal mold. The first die casting machine bearing the Linotype name was patented in 1905 by H.H. Doehler. Two years later came E.B. Wagner’s casting machine, a prototype of the now familiar hot chamber die casting machine. It was first used on a large scale during World War I for binocular and gas mask parts. Zinc alloys were used for die casting as early as 1907 but were not competitive until Price & Anderson developed the Zamak die casting alloy in 1929. In 1851, James Bogardus’s factory in Chicago, which was constructed with cast iron supports, opened the way for what many art historians considered to be America’s only original contribution to the arts of the world— the skyscraper. One of the oldest casting techniques, investment (lost-wax) casting, was also rediscovered in 1897 by B.F. Philbrook of Iowa, who used it to cast dental inlays. Industry paid little attention to this sophisticated process until the urgent military demands of World War I overtaxed the machine tool industry. Shortcuts were then needed to provide finished tools and
precision parts, avoiding time-consuming machining, welding, and assembly. Other efforts at this time included the first fully automated plant in the United States (one of the first in the world), which was a Rockford, Illinois, foundry that cast hand grenades for the U.S. Army in 1918. The first separation of aluminum occurred in 1825, but important engineering applications did not occur until the cost of aluminum steadily declined as a result of smelting process improvements and scale of operations through the end of the 19th century. Aside from additions for aluminum-bronze, zinc-aluminum alloys, and the deoxidation of steel, castings were the first important commercial market for aluminum. The process was simple, low cost, and addressed the limited range of product interests. At first, applications were limited to curiosities such as house numbers, hand mirrors, combs, brushes, tie clasps, cuff links, hat pins, and decorative lamp housings that emphasized the low density, silvery finish, and novelty of the new metal. Lightweight, corrosion-resistant and heat-conductive aluminum cookware could be produced using the same patterns used to cast iron and brass pots, pans, and kettles (Ref 7).
The 20th Century Casting technology during the 20th century advanced in many ways (Table 5), with “more casting progress since World War II than in the previous 3000 years” (Ref 8). For more than
Timeline of casting developments in the 20th century
1900s 1900—Brinnell hardness test machines introduced. Aluminum-bronze in regular production in the United States. Early 1900s—First patent for low-pressure permanent mold casting process issued to England’s E.H. Lake. 1901—American Steel Foundries (St. Louis) produces the first centrifugal cast rail wheels. 1903—The Wright Brothers’ first successful machinepowered aircraft contains a cast aluminum block and crankcase (together weighing 69 kg, or 152 lb), produced either at Miami Brass Foundry or the Buckeye Iron and Brass Works. 1905—H.H. Doehler patents the die casting machine. 1906—The first electric arc furnace is installed in the United States at Halcomb Steel Co., Syracuse, N.Y. First lowfrequency induction furnace is installed at Henry Diston & Sons, Tacony, Pa. 1907—Alfred Wilm discovers that the properties of cast aluminum alloys can be enhanced through heat treating and artificial aging. 1908—Stockham Homogenous Sand Mixer Co., Piqua and Newark, Ohio, releases the sand cutter. 1910s 1910—Matchplates are developed, fostering the viability of jolt-squeeze machines. 1911—Metallurgical microscope is obtainable. First electric arc furnace for metal casting is installed at Treadwell Engineering Co., Easton, Pa. 1912—The first muller with individually mounted revolving mullers of varying weights is marketed by Peter L. Simpson. Sand slinger is invented by E.O. Beardsley & W.F. Piper, Oregon Works.
1915—Experimentation begins with bentonite, a colloidal clay of unusually high green and dry strength. Ajax Metal Co., Philadelphia, installs first low-frequency induction furnace for nonferrous melting. 1916—Dr. Edwin Northrup, Princeton University, invents the coreless induction furnace. 1917—Alcoa completes a great deal of early development work in aluminum as World War I generates a great demand for high-integrity castings for aircraft engines. 1918—The first fully automated metal casting facility in Rockford, Ill., casts hand grenade hulks for the U.S. Army. 1920s 1921—Modification of the silicon structure in aluminum begins as Pacz discovers that adding metallic sodium to molten aluminum just prior to pouring greatly improves ductility. Copper-silicon alloys are prepared in Germany as a substitute for tin-bronzes. 1924—Henry Ford sets a production record of 1 million autos in 132 working days. Automotive manufacturing will grow to consume one-third of casting demand in the United States. 1925—X-ray radiography is established as a tool for checking casting quality. By 1940, all military aircraft castings require x-ray inspection prior to acceptance. American Brass, Waterbury, Conn., installs first mediumfrequency induction furnace in the United States. 1928—Alcoa develops the first aluminum vehicle wheel, a sand-cast 355 alloy designed for truck trailers. 1930s 1930—First high-frequency coreless electric induction furnace is installed at Lebanon Steel Foundry, Lebanon, Pa. Spectrography is pioneered by University of Michigan (continued)
Source: Ref 5; see also Selected References
professors for metal analysis. Davenport and Bain develop the austempering process for iron castings. 1937—Applied Research Laboratories’ founder Maurice Hasler produces the first grating spectrograph for the Geological Survey of California. Spectrometers begin finding their way into foundries by the late 1940s, replacing the previous practice of metallurgists estimating chemical compositions with a spectroscope and welder’s arc. The austempered microstructure in cast iron is recognized. 1940s 1940—Chvorinov develops the relationship between solidification time and casting geometry. Early 1940s—Statistical process control is first employed as a quality control tool in U.S. machine shops, principally to control dimensional tolerances. Inoculation of gray iron becomes common as high-quality cast irons replace scarce steel. 1941—U.S. Lt. Col. W.C. Bliss tells the American Foundrymen’s Association St. Louis Chapter that “the side which maintains the larger production of war goods is going to win the war.” The War Production Board reports later that each U.S. soldier requires 2200 kg (4900 lb) of steel compared to 40 kg (90 lb) in World War I. 1942—The use of synthetic sands increases as a replacement for many war materials. 1943—Keith Millis, a 28 year old metallurgist working at the International Nickel Company searching for a replacement to chrome due to interrupted supply, discovers that magnesium alloy in molten iron produces a spheroidal graphite structure. In 1949, he, Albert Gagnebin, and Norman Pilling would receive a U.S. patent on ductile iron production via magnesium treatment.
History and Trends of Metal Casting / 7 Table 5 (continued) 1940s 1944—The first heat-reactive, chemically cured binder is developed by Germany’s Johannes Croning to rapidly produce mortar and artillery shells for Axis troops during World War II. Two years after the war, his shell process is discovered among other inventions in the German patent office and made public. Croning is recognized with an American Foundrymen’s Society (AFS) Gold Medal in 1957 for his invention. 1946—Allied investigators uncover German foundry research on high-temperature alloys. Having “heard” of the Croning discovery (prior to the release of the report), Ford’s Ed Ensign and E.I. Valyi, Navy Bureau of Ships, attempted to replicate the process and produced shellmolded castings at the Midwest Foundry, Coldwater, Mich. Late 1940s—Thermal sand reclamation is applied to core sands and, to a limited degree, clay-bonded sands. 1948—The first nonlaboratory ductile iron casting is produced at Jamestown Malleable Iron Co., Jamestown, N.Y., as a 168 cm (66 in.) test bar is poured. Industry’s first ductile iron pipe is cast at Lynchburg Foundry, Lynchburg, Va. 1949—Keel blocks, diesel engine parts, a pressure cylinder, an 20 cm (8 in.) cube, and two cylinder liners become the first commercial castings of ductile iron at CooperBessemer, Grove City, Pa. Development of magnesiumferrosilicon makes ductile iron treatment far easier. Buffalo Pipe & Foundry Co., Tonawanda, N.Y., is the first U.S. operation to pour castings using Croning’s shell process. 1950s Early 1950s—Experimentation in high-pressure molding begins as metalcasters begin to increase the air pressure in air squeeze-molding machines to increase mold hardness (density). Fast-drying core oils are introduced. The pneumatic scrubber is developed to reclaim clay-bonded sands. Several wet reclamation systems also are in operation. 1951—Ford Motor Co., Dearborn, Mich., converts 100% of its crankshaft production to ductile iron. 1952—D-process is developed for making shell molds with fine sand and fast-dry oil by Harry Dietert. Sodiumsilicate/CO2 system is introduced. 1953—Hotbox system of making and curing cores in one operation is developed, eliminating the need for dielectric drying ovens. 1954—The CO2 process, a novel mold and coremaking process, is introduced from Germany. Working closely with General Motors, B&P develops a method for coating individual sand particles with resin binder. It also introduces a coreblower capable of producing cores with resin-coated shell sand—a modification to the Croning process. 1955—Ductile iron pipe is introduced to the marketplace by People’s Gas of Chicago, the first to install ductile iron gas mains. Mid-1950s—The squeeze-casting process originates in Russia. 1956—The first Betatron is installed in a U.S. foundry at ESCO Corp., Portland, Ore., for radiography of heavy steel castings. 1957—The vertically parted flaskless green sand molding production machine is invented by Vagn Aage Jeppesen, a 40 year old professor at the Technical University of Denmark. He was granted a patent in 1959, which was purchased by Dansk Industri Syndikat in 1961. 1958—Harold F. Shroyer obtains a patent for the full mold process, a process developed by artists in which simple patterns and gating systems are carved from expanded polystyrene and placed into a green sand mold. The process, known today as lost-foam casting (using loose, unbonded sand), is patented a short time later. Phenolic and furan acid-catalyzed no-bake binder systems are introduced. Ductile iron desulfurization via shaking ladles is developed in Sweden. 1959—General Electric uses the transient heat-transfer digital computer program and successfully applies the finite difference method to heavy steel casting production.
1960s 1960—Furan hotbox binders are developed for core production. Deep bed filters are used commercially for aluminum casting at Alcoa and British Aluminum in the United Kingdom. Compactibility and methylene blue clay tests are developed for green sand control. 1961—Alcohol-borne shell coating process is introduced (warm-coated). 1962—New CO2 sand testing method is introduced for sands bonded with sodium silicate and cured with CO2. Beardsley & Piper’s Al Hunter, Bob Lund, and Angello Bisinello develop the first automated green sand molding machine. In their design, the cope and drag are side-blown simultaneously and then hydraulically squeezed. The birth of automated matchplate molding reportedly improved metal casting productivity by levels as much as 60% in a short amount of time. Phenolic hotbox binders are introduced. 1963—Shell flake resin is introduced, eliminating the need for solvents. 1964—Dell & Christ’s paper on mold inoculation spurs the development of many of today’s forms of mold and late stream inoculation. The first vertically parted green sand machine (max. 240 molds/h) is delivered to United Danish Iron Foundries in Frederiksvaerk, Denmark. Early adopters report man-hour per ton improvements on the order of 50%. 1965—Oil urethane no-bake binder systems are used for cores and molds. General Electric’s Jim Henzel and Jack Keverian predict freezing patterns in large steel castings via computer. Cast metal-matrix composites are first poured at International Nickel Co., Sterling Forest, N.Y. 1968—The coldbox process is introduced by Larry Toriello and Janis Robins and introduced to the metal casting industry by Ashland Chemical Co. for high-production coremaking. Germany’s Daimler-Benz foundry in Mannheim is the first to run the process for automotive parts. John Deere Silvas Foundry, Moline, Ill., is the first to use the process for mass production in North America. 1969—The Chevrolet Vega is introduced by General Motors, featuring the first all-aluminum block with no cast iron cylinder liners. A total of 2.5 million blocks were produced during the vehicle life cycle. Late-1960s—Scanning electron microscope is invented in England. Thermal analysis begins to be used in iron foundries for the rapid determination of carbon equivalent and phosphorus contents, making it possible to study the transformation of an alloy during cooling. Manganese Bronze & Brass Co. and J. Stone & Co. join to promote nickel-aluminum bronze propellers. 1970s 1970—The sodium-silicate/ester catalyzed no-bake binder system is introduced for cores and molds. Safety-critical ductile iron steering knuckles are introduced on Chevrolet’s Cadillac. A new phenolic urethane no-bake process is introduced by Ashland Chemical, replacing oil sand dump box operations and significantly reducing energy requirements for core/mold production throughout the 1970s. Diran Apelian’s doctoral work at the Massachusetts Institute of Technology (MIT) leads to the development of foam filters for metal casting by Olin Metals. Commercial ceramic foam filters will be in use in metal casting facilities by 1974. 1971—The vacuum-forming or V-process molding method of using unbonded sand with a vacuum is developed in Japan. MIT doctoral candidate David Spencer performs experiments leading to semisolid molding, a process in which a partially frozen metal (fluidity similar to machine oil) can be poured into a die cavity. Following advancement in metal slurry consistencies with Professor Mert Flemings, Newton Diecasting Co., New Haven, Conn., produces the first semisolid castings. 1972—A 0.5 kg (1 lb) crankshaft for a refrigerator compressor produced at Wagner Castings (designed and engineered by Tecumseh) becomes the first productionvolume austempered ductile iron (ADI) component. Wagner also produces the first as-cast ductile iron ductile rods for passenger cars. The Canada Center for Mineral and Energy Technology (CANMET) uses real-time (continued)
Source: Ref 5; see also Selected References
1970s radiography to study the flow of steel in molds. Hitchiner Manufacturing, Milford, N.H., patents countergravity (vacuum) casting process. 1973—The first U.S. foundry argon oxygen decarburization unit is installed at ESCO Corp. 1974—Fiat introduces the in-mold process for ductile iron treatment. The phenolic urethane no-bake binder system is introduced for mold production. Mid-1970s—Alcoa and Union Carbide commence rotary degassing for wrought aluminum. Reading Foundry Products would apply this technology to aluminum foundries in the mid-1980s. Digital codes are developed to simulate solidification and fluid flow analysis. Ultrasonic verification of ductile iron nodularity is developed. Metal casting facilities examine new beneficial reuse routes for spent foundry sand, leading to applications such as cement and paving products, bricks, and flowable fill. 1976—Foote Mineral Co. and the British Cast Iron Research Association (U.K.) develop compacted graphite iron. Acidslag cupola practices plus external desulfurization with CaC2 begin to replace basic slag cupolas. 1977—General Motors installs ADI rear differential sets in passenger cars. The alumina phosphate no-bake binder system, an inorganic nonsilicate binder, is introduced for mold production. 1978—The furan/SO2 binder system is developed for core and mold making. Polyurethane no-bake binder system is introduced for aluminum applications. 1980s Early 1980s—Tundish ladle is embraced by industry as favored practice of nodularizing ductile iron. 3-D relational parameters are developed for Computer-aided design solid models. 1981—High-production lost-foam casting begins at General Motors’ Massena, N.Y., plant for aluminum cylinder heads. 1982—Warmbox binder system is introduced. 1983—Air impulse molding process is developed. Free radical cure/SO2 binder system is introduced. 1984—Charles Hull applies for a patent on stereolithography process. Other rapid prototyping techniques emerge shortly after. Phenolic ester no-bake binder is introduced. Thermal analysis makes breakthrough in molten aluminum processing for determination of grain refinement and silicon modification. 1985—Phenolic ester coldbox binder is developed. New automaker Saturn makes a strategic decision to select lostfoam process for its aluminum cylinder blocks and heads and ductile iron crankshafts and differential cases. Mid-1980s—Computer solidification software is commercialized. Amine recycling is introduced to enhance the environmental benefits of the amine-cured coldbox process. Ube Machinery introduces first squeeze-casting equipment. Aikoh’s (Japan) flux injection technology is initiated into U.S. aluminum foundry market. Late 1980s—3-D visualization techniques are developed. CaO/CaF2 desulfurization of cupola-melted ductile base iron begins to replace CaC2 method. Lanxide, Newark, Del., develops pressureless metal infiltration process for particulate-reinforced metal bodies. Magnesium wire injection method for ductile iron treatment is first tested. Casting solidification modeling software gains acceptance, allowing metal casting facilities to optimize quality, production, and cost prior to actually pouring a casting. 1988—Rapid prototyping and Computer-aided design/ computer-aided manufacturing technologies combine in a breakthrough to shorten tooling development time. Ford adapts Cosworth process precision sand casting process for high production. Metaullics Systems combines flux injection/rotary degassing technologies for aluminum processing. 1989—IMI (Yorkshire, U.K.) begins experimenting with bismuth as a lead substitute in copper alloys. 1990s 1990—Equipment for semisolid casting is introduced by Alumax Engineered Materials and Buhler, Inc. Foseco patents a direct-pour system that permits casting production without conventional gating/risering. Major automotive application comes in 1995 with CMI International’s upper intake
8 / Applications and Specification Guidelines Table 5
(continued)
manifold. Precision sand casting and casting quality for engine blocks are improved in mass production by major automotive companies with the Cosworth and Zeus processes for aluminum and the Loramendi Key Core process for precision sand-cast iron applications. 1991—“Dry ice” CO2 process is developed for cleaning coreboxes and foundry tooling. A noncontact gage for accurate dimensional analysis of lost-foam patterns and sand cores is developed through the AFS Lost-Foam Consortium. Eight years later, the consortium develops an instrument to measure the gas permeability of lost-foam pattern coatings (which controls flow of metal and has a dominant effect on casting quality). 1993—First foundry application of a plasma ladle refiner (melting and refining in one vessel) occurs at Maynard Steel Casting Co., Milwaukee. 1994—Use of low-expansion synthetic mullite sand for lost foam is patented by Brunswick Corp., Lake Forest, Ill., to enable precision casting of large components. Delphi Chassis Systems and Casting Technology Co. begin a program on squeeze-cast aluminum front knuckle, the
first high-volume production (1.5 million cars) two-cavity squeeze-cast aluminum conversion of its kind. Mid-1990s—Microstructure simulation is developed, contributing to better understanding of metallurgy effects and the prediction and control of mechanical properties in castings. Semisolid casting makes commercial inroads and market penetration. 1996—Cast metal-matrix composites (brake rotors) are used for the first time in production model automobiles, the Lotus Elise. Environmentally friendly (fluorine-free) fluxes are developed at the Worcester Polytechnic Institute and are commercialized. General Motors Corp. introduces GMBond, a water-soluble biopolymer-based core sand binder that is nontoxic and recyclable. 1997—AFS Consortium research at CANMET, Ottawa, Canada, results in the development and commercialization of lead-free copper alloys using bismuth and selenium. Following tests at Germany’s Aachen University, American Cast Iron Pipe Co., Birmingham, Ala., builds a continuously operated electric arc furnace for cast iron production.
Late 1990s—Stress and distortion simulation introduces the benefits of controlling casting distortion, reducing residual stresses, eliminating hot tears and cracking, minimizing mold distortion, and increasing mold life. 2000s 2001—The National Aeronautics and Space Administration (NASA) and the Dept. of Energy/Office of Industrial Technologies released a physics-based software tool to accurately predict the filling of expanded polystyrene patterns and sand cores as numerous variables are changed. Mercury Marine installs North America’s first pressurized lost-foam casting line at its new facility in Fond du Lac, Wis. 2002—Math-based modeling is used to develop a new sand core blowing simulation capability for process modeling software. 2003—The AFS Magnesium Division completes a research project proving that magnesium could be cast via the lost-foam casting process. Similar trials confirmed the castability of magnesium via the V-process.
phenolics led the way to urea and the dielectric process and then to furans and urea-free resins. The continued development of binders for the production of chemically bonded cores and molds is being directed toward increasing productivity as well as achieving the dimensional repeatability necessary to meet the new challenges of net shape and near-net shape casting requirements (Table 6). Many patterns were made of epoxy resins and polyurethane and other expendables such as polystyrene. Current understanding of sand casting now addresses the fundamentals of clay mineralogy, sand preparation, sand compaction, and the physical properties of molding and core sands. A better understanding of molding sand technology has resulted in sands of a higher degree of uniformity being prepared for the repetitive green sand (clay-bonded) molding sand. This high degree of achievement could only be possible with the great advances in sand testing produced by the foremost researchers and developers of sand testing instrumentation. Permanent-mold methods, which preceded the loam mold and the sand mold by centuries, also progressed for die casting (high pressure, low pressure, and gravity), centrifugal casting (vertical and horizontal), and hybrid processes such as squeeze casting and semisolid metal casting. Other contributions to the continued advancement of metal casting include:
Cast metal-matrix composites Argon oxygen decarburization for steel refining
Source: Ref 5; see also Selected References
Table 6 Development of core and mold processes Process
Core oil Shell: liquid and flake Silicate/carbon dioxide Airset oils Phenolic acid-catalyzed no-bake Furan acid-catalyzed no-bake Furan hot box Phenolic hot box Oil urethane no-bake Phenolic/urethane/amine cold box Silicate ester-catalyzed no-bake Phenolic urethane no-bake Alumina phosphate no-bake Furan sulfur dioxide Polyol urethane no-bake Warm box Free radical cure sulfur dioxide Phenolic ester no-bake Phenolic ester cold box
Approximate time of introduction
1950 1950 1952 1953 1958 1958 1960 1962 1965 1968 1970 1974 1977 1978 1978 1982 1983 1984 1985
Source: Ref 9
400 years, foundry processes and materials often relied on the methods developed by Vannoccio Biringuccio, the 16th century “father of the foundry industry,” who recommended using the dregs of beer vats and human urine as binders for molding sand, both of which were in use well into the 20th century. Until the 1920s, sand testing consisted of squeezing a handful of sand to judge its ability to compact and stick together. Early in that period, a sand research committee of the American Foundrymen’s Society began to develop sand test methods. By 1924, standards were established that covered the various properties of molding sands. Since World War II, experimentation also accelerated in organic and chemical sand binders for the thermosetting of molds and cores. Beginning with the Croning process (shell process),
Foundry automation A scientific approach to the gating and riser
ing of castings using computer simulation of solidification Evaporative (lost)-foam casting and semisolid casting Computer-aided design and manufacture of castings Integrated foundry systems The melting of metals using the plasma arc cupola
Metal Casting Methods Metal casting is unique among metal forming processes, with a variety of casting processes (Table 7) (with Fig. 1, Ref 10). Very broadly, casting processes fall into two groups (Fig. 2): expendable-mold and permanent-mold processes. Each manufacturing process has the capability of producing a part with a certain range of tolerance, surface finish, and complexity. Manufacturing costs depend on the required dimensional tolerance, surface detail, and the quantity of parts required. For each process, there is a minimum batch size below which it is not economical to go because of costs of tooling, fixtures, and equipment. Also related to part cost is the production rate or the cycle time, the time required to produce one part. The most commonly used casting processes are evaluated with respect to these characteristics in Table 8 (Ref 11). Virtually any metal that can be melted can and is being cast. Choice of process and mold material (Table 9) is greatly influenced by the melting point of the alloy. A vital issue is avoidance of voids (due to insufficient feeding of melt) and porosity (due to dendritic solidification and the presence of gases). Success depends, to a large extent, on the skill of the mold designer, but there are very strong influences of the metal (alloy) and mold material. Even with the best design, solidification on the mold walls closes passage of melt at a certain distance, and this leads to a minimum allowable section thickness. There are large differences in the fluidity of alloys, and therefore, minimum allowable section thickness depends on the alloy being cast. Lower minimum thickness can be allowed with zinc, aluminum, and cast
History and Trends of Metal Casting / 9 Table 7
General characteristics of casting processes
See Fig. 1 for shape descriptions Casting process Characteristic
Part Material (casting) Porosity and voids(a) Shape(b) See Fig. 1 for shapes
Green sand
Resin-bonded sand
Plaster
Lost foam
All C–E All
All D–E All
Zn to Cu D–E All
Al to cast iron C–E All
All E All
Minimum section, mm (in.) Minimum core diameter, mm (in.) Surface detail(a)
0.01–300,000 (0.02–660,000) 3–6 (0.12–2.4) 4–6 (0.16–0.24) C
0.01–100 (0.02–220) 2–4 (0.08–0.16) 3–6 (0.12–0.24) B
0.01–1000 (0.02–2200) 1 (0.04) 10 (0.4) A
0.01–100 (0.02–220) 2–4 (0.08–0.16) 4–6 (0.16–0.24) C
Cost Equipment(a) Die (or pattern)(a) Labor(a) Finishing(a)
C–E C–E A–C A–C
C B–C C B–D
C–E C–E A–B C–D
Production Operator skill(a) Lead time Rates (piece/h mold) Minimum quantity
A–C Days 1–20 1–100
C Weeks 5–50 100
A–B Days 1–10 10
Size, kg (lb)
Permanent mold
Investment
Die
Zn to Cu A–C Not T3, 5, F5
0.01–100 (0.02–220) 1 (0.04) 0.5–1 (0.02–0.04) A
Zn to cast iron B–C Not T3, 5, F5 with solid core 0.1–100 (0.2–220) 2–4 (0.08–0.16) 4–6 (0.16–0.24) B–C
B–C B–C C C–D
C–E B–C A–B C–D
B B C B–D
A A E C–E
C Weeks–months 1–20 500
A–B Hours–weeks 1–1000 10–1000
C Weeks 5–50 1000
C–D Weeks–months 20–200 100,000
<0.01–50 (0.02–110) 0.5–1 (0.02–0.04) 3 (Zn: 0.8) (0.12; Zn: 0.03) A–B
(a) Comparative ratings, with A indicating the highest value of the variable, E the lowest (e.g., investment casting gives very low porosity, produces excellent surface detail, involves moderate to low equipment cost, medium to high pattern cost, high labor cost, medium to low finishing cost, and high operator skill. It can be used for low or high production rates and requires a minimum quantity of 10 to 1000 to justify the cost of the pattern mold). (b) From Fig. 1. Source: Adapted from Ref 10
iron than with steel. Streamlined flow also is helpful; thus, sharp corners are avoided. Thicker sections solidify last and must be fed adequately. Metals and alloys have different castability characteristics based on their combination of liquid-metal properties and solidification characteristics that promotes accurate and sound castings. Castability factors include fluidity, shrinkage, and resistance to hot cracking: Shrinkage is whether the alloy has a high
tendency to form shrinkage porosity.
Resistance to hot cracking is the ability of
the alloy to withstand stresses developed by the contraction while cooling from the hotshort temperature range. Fluidity is the ability of the liquid alloy to flow readily in the mold and to fill thin sections. Fluidity is usually evaluated by pouring the metal into a narrow spiral mold cavity and measuring the length of penetration of the metal before it solidifies. Fluidity is inversely related to the temperature range over which solidification occurs. Thus, pure metals and eutectic alloys, with short freezing range, have greater fluidity Because the viscosity of molten metals is very low, fluidity is determined more by the solidification dynamics of the metal and the mold. A high surface tension of the liquid metal reduces fluidity, and oxide films forming on the surface of the molten metal thus have a significantly adverse effect on fluidity. Insoluble inclusions also reduce fluidity. Castability ratings from the three factors are often described on a scale of 1 to 5 for each of the important factors. Table 10 shows typical
applications for metal alloys and relative ratings for castability, weldability, and machinability. As an example of a castability comparison in terms of fluidity, Fig. 3 shows three fluidity spiral samples of gray, compacted graphite, and ductile (nodular) iron. They were poured from the same base metal at the same temperature and have comparable chemical compositions. They differ only in their melt treatment. Gray (flake graphite) iron shows the best fluidity, and ductile iron shows the worst. Compacted graphite iron occupies an intermediate position. However, because compacted irons are stronger than gray iron with the same carbon equivalent (CE), a higher CE can be used to obtain the same strength, allowing greater fluidity and easier pouring of thin sections.
Production Trends The use of metal castings is pervasive throughout the economies of all developed countries, both as components in finished manufactured goods and as finished durable goods (Ref 1). As indicated earlier, castings are used in 90% of all manufactured goods and in all capital goods machinery used in manufacturing. They are also extensively used in transportation, building construction, municipal water and sewer systems, oil and gas pipelines, and a wide variety of other applications. In the broadest sense, foundries are categorized into two general groups: ferrous foundries and nonferrous foundries. For marketing purposes, foundries are also categorized according to the nature of their operations as being either captive (producing castings for their own use)
or jobbing (producing castings for sale). The market is sometimes further broken down by major casting processes when they can be readily identified or are particularly significant. This is done more often in the case of nonferrous castings; in this case, the product is usually categorized as being produced in sand, permanent mold, or die cast. Because 90% or more of all iron and steel castings are produced in some form of sand medium, distinguishing the process used is not quite as significant. Ferrous castings shipments are usually classified by market category. For example, iron castings are generally categorized as engineered (designed for specific, differentiated customers) and nonengineered (produced in large volumes of interchangeable units, usually consisting of ingot molds and pressure and soil pipe). The diversity among the various foundries makes it difficult to determine the exact structure of the industry. For example, it is not unusual for a single operating foundry to produce a variety of metals and alloys, both ferrous and nonferrous, in the same plant. Some also use a variety of processes in their operations. Many aluminum foundries, for instance, use both sand and permanent-mold processes, and some even produce die castings in the same facility. In addition, the foundry industry consists of a variety of large and small facilities. Table 11 gives recent census results on the number of casting facilities in major industrialized countries. Ferrous castings easily make up the largest production sector of the foundry industry in terms of volume, with gray iron outdistancing the other major metals constituting this group (Table 2). In comparison, iron and steel
10 / Applications and Specification Guidelines Increasing spatial complexity
Type (abbreviation)
0 Uniform cross section
1 Change at end
2 Change at center
3 Spatial curvature
4 Closed one end
5 Closed both ends
6 Transverse element
7 Irregular (complex)
Round (R)
Bar (B)
Section, open (S) Semiclosed (SS)
Tube (T)
Flat (F)
Spherical (Sp)
Fig. 1
The choice of possible manufacturing methods is aided by classifying shapes according to their geometric features. Source: Ref 10
Shape-casting processes
Expendable mold
Permanent pattern
Sand molding Clay Green Skin-dry Dry Carbon dioxide Cement Oil Resin Cold box Hot box Shell Flaskless
Fig. 2
Permanent mold
Expendable pattern
Slurry molding Plaster Ceramic
Semipermanent (sand cores)
Gravity Slush
Permanent (metal cores)
Low pressure
Hot chamber Sand (full-mold)
Slurry Lost wax Centrifuged
True centrifugal (outer mold)
Semicentrifugal (complete mold)
General classification of shape-casting processes. Source Ref 10
Centrifuged (complete mold)
Squeeze
Die
Cold chamber
foundries in North America shipped an average of 15.5 million tons of castings annually in the years from 1947 to 1986. The following decade (1967 to 1976) saw a significant upturn as the average annual shipments of iron and steel castings rose to slightly more than 17.5 million tons. Iron casting shipments have averaged approximately 14 million tons per year during the last four decades. From 1977 to 1986, ferrous casting shipments from American foundries fell to approximately 9.5 million tons by 1982 (Fig. 4). In this same time period, casting production around the world has shown a significant decline, generally reflecting the state of manufacturing. World production of ferrous castings fell some 27%, from 64.6 million tons in 1979 to 47.3 million tons in 1982 (Table 12). Nonferrous shipments followed a similar path in this time frame (Fig. 5). Recent Production Trends (1995 to 2005). Currently, metal casting production is dominated by China, along with large growth rates
History and Trends of Metal Casting / 11 Table 8
Casting processes and their attributes Surface roughness
Process
Dimensional accuracy
Complexity
Production rate
Production run
Relative cost
Size (projected area)
Pressure die casting Centrifugal casting Sand casting Shell mold casting Investment casting Powder metallurgy
L M H L L L
H M M H H H
H M M H H H
H/M L L H/M L H/M
H M/L H/M/L H/M H/M/L H
H H/M H/M/L H/M H/M H/M
M/L H/M/L H/M/L M/L M/L L
Key: H M
>250 >63 and <250
High Medium
>100 >10 and <100
>5000 >100 and <5000
High Medium
>0.5 >0.02 and <0.5
L Units
<63 min.
<0.005 >0.005 and <0.05 >0.05 in.
Low ...
<10 Parts/h
<100 Parts
Low ...
<0.02 m2
Source: Adapted from Ref 11
Table 9
Process and mold material options for casting
Process
Sand casting Investment casting Die casting
Cast iron
Carbon steel
Alloy steel
Stainless steel
Aluminum and aluminum alloys
Copper and copper alloys
Zinc and zinc alloys
Magnesium and magnesium alloys
Titanium and titanium alloys
Nickel and nickel alloys
Refractory metals
A B
A A
A A
A A
A A
A A
B B
A B
B B
A A
B B
C
C
C
C
A
B
A
A
C
C
C
A, normal practice; B, less-common practice; C, not applicable. Source: Adapted from Ref 12
Table 10 Typical applications for castings and general rating of castability, machinability, and weldability Note that these are general overall ratings for a class of alloys. Within each class, different alloys may vary considerably with respect to specific castability characteristics of resistance to hot cracking, fluidity, and shrinkage tendency.
Table 11 Number of operating metal casting facilities by nation in 2005
Casting alloy type
Country
Applications
Aluminum Copper Gray iron Magnesium Malleable iron Nickel Nodular iron Steel (carbon and low alloy) Steel (high alloy) White cast iron Zinc
Castability Machinability Weldability
Pistons, clutch housings, exhaust manifolds Pumps, valves, marine propellors Engine blocks, gears, brake disks and drums, machine bases Crankcases, transmission housings Farm and construction machinery, heavy-duty bearings, railroad rolling stock Gas-turbine blades, pump and valve components for chemical plants Crankshafts, heavy-duty gears Die blocks, heavy-duty gear blanks, railroad wheels
1–2 1–3 1 1–2 2
1–2 1–2 2 1 2
3 3 5 2 5
3
3
3
2 3
2 2–3
5 1
Gas-turbine housings, pump and valve components, rockcrusher jaws Ball mill liners, shotblasting nozzles, railroad brake shoes, crushers Door handles, radiator grilles, carburetor bodies
3
3
1
2
4
4
1
1
5
1, excellent; 2, good; 3, fair; 4, very poor; 5, difficult. Source: Ref 13
FG iron
CG iron
SG iron
10 cm
Fig. 3
Spiral fluidity samples poured from gray (flake graphite), compacted graphite, and ductile (spheroidal graphite) iron. Source: Ref 14
Austria Belgium Brazil Canada(a) China Croatia Czech Republic Denmark Finland France Germany Great Britain Hungary India Italy Japan Korea Lithuania Mexico Netherlands(a)(b) Norway Poland Portugal Romania Russia(e) Slovenia(e) Slovakia South Africa Spain Sweden Switzerland Taiwan(a) Thailand Turkey Ukraine(e) United States
Iron
Steel
Nonferrous
Total
12 18 467 60 17,000 16 106 10 16 115 212 218 41 ... 180 457 499 8 380 16 7 173(c) 50 58 ... 16 12 95 67 36 19 520 230 857 400 619
4 9 139 24 4,700 5 32 0 6 38 53 46 27 ... 27 78 140 3 10 0 3 36(c) 10 41 ... 4 7(f) 42 35 15 0 45 26 73 233 262
29 9 658 66 4,300 23 56 11 15 352 365 216 143 ... 870 1,173 208 5 250 5 11 245(d) 38 65 ... 33 32 119 57 84 32 350 220 361 437 1,499
45 36 1,264 150 26,000 44 194 21 37 505 630 480 211 4,500 1,077 1,708 847 16 640 21 21 454 98 164 1,900 53 51 256 159 135 51 915 476 1,291 960 2,380
Note: “. . .”indicates unreported data.(a) 2004 data. (b) Includes only members of AVNeG (General Assoc. of Dutea Foundries). (c) Includes approximately 24 metal casting facilities making both iron and steel. (d) Includes approximately 40 metal casting facilities working with both ferrous and nonferrous metals. (e) 2002 data. (f) Includes investment casting facilities. Source: Modern Casting
12 / Applications and Specification Guidelines 25 000
20 000 Thousands of tons
Table 12 Ferrous casting production from 1982 to 1986
Gray iron Ductile iron Malleable iron Steel Total
15 000
Figures given in thousands of metric tons Year Country
United States Japan West Germany France United Kingdom Brazil Canada(a) Korea Taiwan
10 000
5 000
0 78
Fig. 4
79
80
81
82 83 Year
84
85
86
87
1982
1983
1984
1985
1986
8,651,455 5,735,001 3,501,299 2,230,264 1,621,700 1,207,700 671,913 639,000 401,990
9,313,866 5,515,999 3,311,600 2,014,387 1,549,100 986,142 781,722 762,000 492,150
10,856,200 5,128,896 3,387,300 1,855,634 1,490,500 ... 981,730 805,000 599,122
9,991,393 5,240,261 3,499,900 1,902,608 ... 1,407,324 856,077 848,000 637,660
8,564,725 4,860,591 3,452,000 1,797,960 1,222,000 1,543,147 865,777 934,000 778,930
(a) Canada’s tonnage is estimated to be 75% of actual. Source: Ref 15
U.S. ferrous casting shipments from 1978 to 1987. Source: Ref 2
24 2500
21 Aluminum base Copper base
18
Zinc base Total
Castings, millions of tons
Thousands of tons
2000
China
1500
1000
500
U.S.
15 12
Russia
9
Japan
6 0 77
78
79
80
81 82 Year
83
84
85
3
86
Germany
India
0 1998
U.S. nonferrous casting shipments from 1977 to 1986. Source: Ref 2
in India (Fig. 6). The market for castings is likely to shift more to Asia and, to a lesser extent, to Central and Eastern Europe. However, despite the contraction of industry production bases in the West, high-quality manufacturing should continue to meet the needs of home markets and the demands of more productivity, energy efficiency, and environmental sustainability. German metal casting sites are reported to have the highest productivity in 2005, followed by the United States and France (Fig. 7). From 1995 to 2005, global output in the foundry industry grew by some 25%, topping 85 million tons. This is due in part to strong growth in its traditional markets but also to increasing use of these materials in a wide variety of other applications. These include the use of cast iron products in wind farms and large motors for power generation and ships, as well as the increasing popularity of light materials in automotive engineering, especially aluminum. During this period, output of cast steel continued to grow at an average pace, although there was something of a downward trend until 2004. Cast steel tends to be used less often in automotive and mechanical engineering than in
1999
2000
2001
2002
2003
2004
2005
6.0 5.5 5.0
Castings, millions of tons
Fig. 5
4.6 4.0 India
3.5
Brazil
3.0
France
2.5 Italy
2.0 1.5
Korea
1.0 1998
Fig. 6
1999
2000
2001
2002
2003
2004
Tonnage of major casting-producing nations from 1998 to 2005. Source: Ref 3
2005
History and Trends of Metal Casting / 13
8,108
Productivity per metal casting site
5,419
Germany (630 sites)
4,639 France (505 sites) Russia (1900 sites)
4,010 3,097 2,360 2,348 2,242 1,358 939 0
Fig. 7
U.S. (2380 sites)
Japan (1708 sites)
Italy (1077 sites) Brazil (1264 sites) Korea (847 sites)
India (4500 sites)
China (24,000 sites) 1
2
3
4
5
6
7
8
9
Productivity per metal casting site in 2005 for major industrial nations. See Table 11 for breakdown of site count. Source: Ref 3
1995 20%
United States
17%
China
16%
Commonwealth of Independent State
10%
Japan
5%
India
6%
Germany
26%
Other countries
30%
China
14%
United States
8%
India
8%
Japan
7%
Russia
7%
Germany
26%
Other countries
2005
Fig. 8
National shares in cast iron production. Source: Ref 3
power plant and factory construction. It benefited from robust growth in its main markets in 2005.
Cast Iron. China dominates cast iron production (Fig. 8). In terms of volume, cast iron and cast steel account for the bulk of foundry
industry production, comprising approximately 80% of total output. Global production has increased since 1995 by some 15%, with Asia recording the fastest rate of growth: by 2005 the region was already responsible for over half of world output (Fig. 8). Conversely, production in North and South America and Europe declined slightly. However, the situation varies considerably from country to country. In North America, production of cast iron in the United States and Canada fell by approximately 15%, likewise reducing their share of the world market (Fig. 8). Production volume has increased again since 2002, although it still remains below the 1995 figure. Even so, the United States continues to be the world’s second-largest producer. By contrast, Latin American countries such as Brazil and Mexico boosted their output by over 80%. From 1995 to 2005, China doubled its share in the global market. In Japan, on the other hand, production volume dropped by nearly 10%. Growing at a fast pace, output in India rose by approximately 80%, surpassing Japanese production for the first time in 2005. Together, these three countries account for approximately 90% of Asian production of cast iron. Nonferrous Castings. Nonferrous metal casting have sustained an upward trend. Global output of nonferrous metal casting products grew by some 75% from 1995 to 2005 (Fig. 9). Although production in regions around the globe gained ground, growth in North and South America (50%) lagged behind the rest of the world, where the volume of production more than doubled. In the Western hemisphere, the United States accounted for a good two-thirds of production, its 17% rate of growth pulling down the regional average. Copper and copper alloys as well as zinc products contracted by 6 and 20%, respectively, whereas output of all other materials expanded. Particularly notable is the robust increase in aluminum production in Canada, where Alcan Inc., one of the world’s top aluminum producers, boosted capacity. In Mexico, copper output increased tenfold, evidently due to the transfer of production from its northern neighbor. The volume of production in Asia increased by over 130%, driven first and foremost by rapid expansion in China, where output nearly tripled, and India, which likewise succeeded in significantly widening its share of the world market. Although production in Japan also rose, the country’s relative importance in the world market diminished. Aluminum continues to dominate nonferrous metal casting. Worldwide, aluminum accounts for approximately 80% of nonferrous metal casting, trailed by copper and copper alloys with 10%; the share of zinc is slightly over 6%. Production of aluminum cast parts doubled during the period under review, while other metals posted total gains of “only” approximately 10%. This is due primarily to greater use of aluminum in the automotive sector. Owing to its low specific weight, aluminum is an increasingly common
14 / Applications and Specification Guidelines In million tons 14 12 10 8 6 4 2 0 1995
2000
2001
2002
2003
2004
2005
Asia Europe North and South America
Fig. 9
Trends in production of nonferrous metal castings. Source: Ref 3
1995 30%
United States
19%
Japan
9%
China
8%
Germany
8%
Italy Commonwealth of Independent State
3% 1%
India
22%
Other countries
18%
United States
16%
China
13%
Japan
8%
Russia
7%
Italy
6%
Germany
4%
India
28%
Other countries
2005
Fig. 10
National shares in production of aluminum castings. Source: Ref 3
material in lightweight engineering. In the aviation industry, too, it continues its rapid ascent. The relative share of the traditional industrial nations in global aluminum production declined markedly from 1995 to 2005 (Fig. 10). While the United States and Japan increased their production volumes by approximately 30%, China’s output practically tripled. India, starting from a very slender base, also expanded its share massively. By 2005, China’s
production volume had nearly reached that of the United States, which it is likely to over-take during the next two years.
Market Trends The two largest sectors of metal castings are the automotive industry and machinery manufacturers. In the automotive industry, the
world’s major industrial economies account for approximately 85% of global car production. However, vehicle output in China and the nations of Central and Eastern Europe has been growing at a much faster pace. The next ten years are likely to see a continuation of this trend. In 2015, the world’s traditional car-making nations will probably be responsible for just 70% of global production; China’s share is likely to increase from approximately 10% to a more than 15% (Ref 16). Another important market for castings is the machinery/instrumentation sector, with products ranging from massive hubs, housings, and shafts for power generators to miniscule components for semiconductors. The world’s two leading makers of machines are the United States and Japan, which both lost ground from 2000 to 2005. In context, German manufacturers widened their share of the world market (Fig. 11). The Chinese sector is also expected to grow from its fourth place ranking. Although China remains one of the world’s biggest importers of plant and machinery, it is increasing its presence as an exporter as well. In 2005, China’s machine-tools sector (with annual sales of approximately £5 billion) exceeded that of Italy. Even with a distinct slowdown in the growth of imports, foreign suppliers meet 60% of China’s total demand for machine tools.
REFERENCES 1. “Competitive Assessment of the U.S. Foundry Industry,” USITC Publication 1582, U.S. Department of Commerce, Sept 1984, p xiii 2. D.P. Kanicki, Casting Advantages, Applications, and Market Size, Casting, Vol 15, Metals Handbook, 9th ed., ASM International, 1988, p 37–46 3. Annual Census of World Casting Production, Mod. Cast., Dec 2005 and 2006 4. M. Goodway, History of Casting, Casting, Vol 15, Metals Handbook, 9th ed., ASM International, 1988, p 15–23 5. Timeline of Casting Technology, Mod. Cast. Cast Expo Issue, May 2005 6. G.L. Werley, Slush Casting, Forging and Casting, Vol 5, Metals Handbook, 8th ed., American Society for Metals, 1970, p 334 7. E. Rooy, A History of Aluminum Foundry Alloy Development, AFS International Conference on Structural Aluminum Castings: Recent Advancements for the Design and Manufacturing of Reliable Structural Aluminum Castings, Nov 2–4, 2003 (Orlando, FI), p 87–98 8. B.L. Simpson, in History of the Metalcasting Industry, American Foundrymen’s Society, 1948 9. E.L. Kotzin, Development of Foundry Technology in the United States, Casting, Vol 15, Metals Handbook, 9th ed., ASM International, 1988, p 15–23
History and Trends of Metal Casting / 15 SELECTED REFERENCES (1980–1990) AFS/CMI Conference on Green Sand TechUnited States
Japan
Germany
China Italy
France Great Britain
South Korea billion
0
50
100
150
200
250
300
350
2000 2005
Fig. 11
Global engineering sales of machinery/machine-tool sector. Source. Ref 17
10. J.A. Schey, Introduction to Manufacturing Processes, 2nd ed., McGraw-Hill, 1987 11. E.B. Magrab, Integrated Product and Process Design and Development, CRC Press, Inc., 1997 12. G. Boothroyd, P. Dewhurst, and W. Knight, Product Design for Manufacture and Assembly, Marcel Dekker, Inc., 1994 13. S. Kalpakjian, Manufacturing Processes for Engineering Materials, 2nd ed., Addison-Wesley, 1991, p 262 14. D.M. Stefanescu, R. Hummer, and E. Nechtelberger, Compacted Graphite Irons, Casting, Vol 15, ASM Handbook, ASM International, 1988 15. Census of World Casting Production, Mod. Cast., American Foundrymen’s Society, 1987 16. Foundry Industry, IKB Deutsche Industriebank, June 2007 17. Verband Deutscher Maschinen und Anlagenbau, German Engineering Federation SELECTED REFERENCES (published after 1990) A Tribute to Keith Millis and the Unveiling of
Ductile Iron 50 Years Ago, Mod. Cast., Oct 1998
J.R. Bodine, 1999 Charles Edgar Hoyt Memo-
rial Lecture: From a Monument to the Vega: The Journey of the Aluminum Casting Industry, Mod. Cast., May 1999, p 60–63 R.M. Cooper, The History of the Soho Foundry, GEC Rev., Vol 11 (No. 2), 1996, p 114–119 J. Gerin Sylvia, Cast Metals Technology, AFS, 1990 F. Habashi, Milestones in the History of Metallurgy, BUMA-V: Fifth International Conference on the Beginnings of the Use of Metals and Alloys, April 21–24, 2002 (Gyeongju, Korea), p 53–60 E.L. Kotzin, 1997 Hoyt Memorial Lecture: Lost Wax to Lost Foam: Reflections on the Past, Present and Future, Mod. Cast., July 1997, p. 42–45 Modern Casting’s A Look Back at the 20th Century Series, 2000–2001 Proceedings of 1992 Coreless Induction Melting Conference, AFS, 1992 N. Wukovich, A Short History of the Steel Foundry, Trans. Am. Foundrymen’s Soc., Vol 104, April 20–23, 1996 (Philadelphia, PA), p 643–649 1991 AFS International Sand Reclamation Conference, AFS, 1991 1995 Lost Foam Technology and Applications Conference Proceedings, AFS, 1995
nology—Productivity for the ’80s, AFS, 1983 N.N. Bubtsov, History of Foundry Practice in USSR, Moscow, 1962; trans., Washington, D.C., 1975 Casting, Vol 15, Metals Handbook, 9th ed., ASM International, 1988 J. Foster, “The Iron Age Moulds from Gussage All Saints,” Occasional Paper 12, British Museum, London, 1980 Foundry Trade J., Sept 11, 1986 High-Pressure Molding, 2nd ed., AFS, 1987 E.L. Kotzin, Metalcasting & Molding Processes, AFS, 1981 J.P. LaRue, Basic Metalcasting, AFS, 1989 C.S. Smith, The Early History of Casting, Molds, and the Science of Solidification, A Search for Structure: Selected Essays on Science, Art, and History, MIT Press, Cambridge, MA, and London, 1981, p 127–173 R.F. Tylecote, History of Metallurgy, The Metals Society, London, 1976 R.F. Tylecote, The Early History of Metallurgy in Europe, Longman, London, 1987 C.F. Walton and T.J. Opar, Iron Castings Handbook, Iron Castings Society, Inc., 1981 T.A. Wertime and J.D. Muhly, The Coming of the Age of Iron, Yale University Press, New Haven, 1980 T.A. Wertime and S.F. Wertime, Early Pyrotechnology: The Evolution of the First FireUsing Industries, Washington, D.C., 1982
SELECTED REFERENCES (published prior to 1980) L. Aitchison, A History of Metals, New York,
1960
Aluminum: Design and Application, Vol II,
American Society for Metals, 1967
J.A. Gitzen, Progress in the Development
and Utilization of Additives in the Foundry Industry, AFS, 1957 P. Knauth, The Metalsmiths, New York, 1974 C.A. Sanders and D.C. Gould, History Cast in Metal, AFS, 1976 B.L. Simpson, Development of the Metal Castings Industry, American Foundrymen’s Association, Chicago, 1948 B.L. Simpson, History of the Metalcasting Industry, AFS, 1948
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 16-28 DOI: 10.1361/asmhba0005188
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Metalcasting Technology and the Purchasing Process American Foundry Society Technical Department
CAST METAL COMPONENTS are not items that are commonly bought out of catalogs; most often they are tailor-made objects, designed and produced to the specific needs of a particular customer. As such, cast components cannot and should not be treated as low-value, high-volume secondary commodity items, where price is the sole purchase criterion. Cast components should be considered highvalue manufacturing products, where the metalcasting facility is an extension of the buyer’s manufacturing capability, involving long-term relationships and open, on-going communication at all levels. As high-value items, the purchasing criteria for cast components include a balance of quality, cost, service, and delivery. Many original equipment manufacturing (OEM) component buyers are recognizing that a carefully considered purchasing policy is the key to bestvalue procurement of metal castings. The policy should define the buyer-seller relationship, based on the following principles:
Long-term relationships based on shared
benefits and best value for all parties
Qualified cast component suppliers who
meet production quality standards, deliver on schedule, provide expertise and knowledge to support component design and improvement, and provide best-value pricing Open and continuous communication at executive, purchasing, and technical levels to support concurrent engineering/design, design for manufacturability, and continuous improvement Over the last 20 years, significant change has occurred in the industrial purchasing environment—the old purchasing rules and situations no longer apply. Today’s (2008) manufactured product market is marked by six major changes that impact both the buyers and the sellers: Globalization with new competition and new
markets: Today (2008), information, goods, and services move around the globe at an unprecedented speed. The OEM purchasing
managers can find sources for manufactured products in Malaysia, Mexico, or Poland as easily as they find a source in Ohio, Georgia, or California. The flip side is that a producer in Wisconsin can market components in Europe, Japan, and Brazil. Products can be delivered the next day from around the world. Best-value purchasing: The competitive market is driving everyone to focus on providing the best value to the final customer. Best value means that buying decisions are made not just on cost but also on a balance of quality, service, price, and availability/delivery. Recognizing how subcontractors impact final cost and quality, smart OEM manufacturers purchase their components and supplies from vendors who consistently can meet the best-value requirements. Value-added services or “one-stop shopping” for castings is common. Outsourcing: In the old days, many OEM manufacturers used internal production divisions to supply components. Today (2008), more OEM companies are recognizing that it is smarter to outsource their components through strategic partnerships with sources that have the cutting-edge expertise and advanced technology to provide a best-value product. Supplier management and quality assurance: When the quality and cost of a manufacturer’s end product depend critically on components from an outside supplier, the manufacturer will take a supply-management approach to working with the source. In supply management, buyers work with all their purchasing sources, developing long-term relationships, providing oversight on quality assurance, and working cooperatively with the sources in product-improvement teams. Continuous improvement: Nothing stands still for long anymore. In a competitive market, if a company does not improve its product and services, its competitors will improve theirs and win the contract. The supplier must work with the customer to find
and exploit opportunities to improve quality, reduce cost, and cut cycle/production time. That kind of innovative effort applies to working on current production components as well as on new product designs. Information availability and velocity: The internet and electronic commerce have changed the way that information moves and circulates. Both buyers and sellers are expanding their electronic commerce reach in locating and assessing sources and opportunities for purchases/sales. Electronic marketplaces serve as locations for posting requests for quotes, advertising production capabilities, and finding matches between requirements and services. In a similar manner, suppliers are posting information about their capabilities and expertise on the internet, providing instant information to potential customers. Electronic commerce extends beyond the marketing function. All types of information (orders, contracts, technical, and financial) move electronically rather than by paper and mail. The speed of information movement is measured in minutes rather than days, cutting cycle time, reducing data errors, and overcoming geographic separation between buyers and suppliers. The metalcasting market is not isolated from these changes. Cast component buyers and sellers recognize that they must respond to the new market imperatives if they are to stay in business and remain competitive.
The Purchasing Process Purchasing is a critical process in any business. A company purchases raw materials, subcomponents, and support services that are essential to the production and delivery of the finished product to their customer. There is tremendous value in developing a purchasing process that is tailored to a the, specific needs of a business and that obtains components and services from outside sources in an effective, efficient, and timely manner.
Metalcasting Technology and the Purchasing Process / 17
Fig. 1
Four basic steps of purchasing process
Whether it is a product or service, the purchasing process consists of four basic steps (Fig. 1): 1. Define requirements and develop a purchasing plan. 2. Request and evaluate bids from potential sources. 3. Select a source and negotiate contract terms. 4. Carry out the contract and pursue continuous improvement. These basic steps of the purchasing process are described in the next four sections of this article. It is also important to remember that purchasing does not end with the selection of a source and contract agreement. The process is not complete until the required product/service is received and meets the requirements for quality, cost delivery, and service. Purchasing continues through product delivery (Fig. 1). The following sections provide detailed guidance on purchasing cast components, going through each step and describing specific issues and approaches that have proved useful in purchasing castings.
Define Requirements and Develop a Purchasing Plan The first step in any purchasing process is to define the product requirements and develop a purchasing plan. For cast components, product requirements are based on technical performance and targets for quality, quantity, schedule and cost. With this information, a complete and accurate request for quotation can be drafted and distributed to metalcasting facilities for consideration. To be a smart buyer, one needs more than just a requirements list. A purchasing plan that outlines and defines the following is also required: Objectives and issues in the purchase Purchase requirements How sources will be identified, screened,
and selected How the purchase will be contracted and managed for quality, cost, and schedule This plan may be simple and straightforward for small, low-value purchases; it may be very detailed and extensive for a high-value, critical purchase.
Request a Quotation Purchasing cast components is generally not a simple matter of transferring pattern numbers and quantity desired from the engineering bill of materials to a purchase order. Careful consideration of technical quality, quantity, and schedule factors is necessary to ensure that all pertinent and important information is supplied to the metalcasting facility. When necessary information is not provided, this slows the procurement and results in an inaccurate cost estimate or in a cast component that does not fully meet requirements. When possible, the metalcasting facility should be supplied with Computer-aided design files or blueprints for both the as-cast and the machined part. If the metalcasting facility is to provide the most suitable and useful cast components, it is important that the buyer indicate the areas to be machined and the locating points for machining. In addition, the buyer must know and understand the critical factors in specifying cast components and how those factors impact production and costs in the metalcasting facility. A request-for-quotation (RFQ) form is used to good advantage by many metalcasting facilities and cast component buyers. It provides a convenient way to develop the component requirements in a comprehensive and detailed approach. A model RFQ form is provided in Fig. 2. The RFQ form supplies the metalcasting facility with complete information on component requirements and enables the metalcasting facility to make an accurate casting price quotation for the most suitable casting for a particular application.
The Purchasing Plan With the detailed requirements for the casting defined in a draft RFQ, the buyer can develop a purchasing plan (Fig. 3.) that specifies: Major purchasing objectives Purchasing strategy Deliverables expected from the metalcast-
ing facility at each stage of the purchasing process Desired purchaser-metalcasting facility working relationship to ensure that quality, schedule, and cost objectives are met Specific selection criteria Purchasing tasks, responsibilities, and schedule
Purchasing Objectives. The buyer must clearly define the overall requirements and priorities in purchasing the cast component. Those requirements must be described in hard numbers that are clearly understandable and easy to communicate. The overall requirements package will include the engineering requirements as well as cost targets, trade-off considerations, and the company supply-chain management strategy. Each metalcasting purchase will have different answers to the following questions: What are the quality and performance require-
ments for the cast component?
How many cast components will be purchas-
ed over a certain period?
What is the cost target for the cast compo-
nents?
What is the delivery schedule for the cast com-
ponents?
How critical is the casting to the project/busi-
ness in terms of performance, risk, schedule, and cost? How will the different objectives (quality, performance, cost, and schedule) be balanced and prioritized? This requires a trade-off analysis (either informal or formal). How does the cast metal component purchase fit into the overall supply-chain strategy for the company? The first priority is to meet the quality, quantity, service, schedule, and cost requirements for a given cast part. A cast component order is seldom a one-time purchase for a company. More often, cast components are integral to a company product line, and the purchase must be considered in the context of a company overall supply-chain strategy. A corporate supplychain strategy considers how to maximize the value and efficiency of all purchases through coordination and integration, rather than simply maximizing the value of individual purchases. Purchasing Strategy. Purchasing is nothing less than asset management, where the assets are suppliers and the goal is to maximize performance and value while minimizing risk across the casting purchase process. As with any investment, the buyer should develop a strategy that will meet objectives with a balance between risk and return. For a purchasing strategy, the buyer must ask the following questions in light of the particular requirements for a cast component: How much benefit is there in finding the best
possible source?
What is the level of risk in purchasing from
a nonperforming source?
What are the cost drivers for the component,
and where can the metalcasting facility have a positive or negative impact on those drivers? What investment of time and resources is necessary to identify a suitable source versus a best-possible source? How does the purchase of a given cast metal component fit into the overall purchasing
18 / Applications and Specification Guidelines
Fig. 2
Casting request-for-quotation form
Metalcasting Technology and the Purchasing Process / 19
Fig. 3
Basic elements of a purchasing plan and strategy
strategy for castings within a division or company? How will the company work with the supplier to maximize value during all stages of the purchase cycle?
cast components will include more than just the receipt of qualified castings at the shipping dock. It also should include details for technical partnership services during the entire purchasing process that include:
Answers to these questions will provide a framework for determining the value of a well-qualified metalcasting facility source to the success of a given project/program as well as the overall supply-chain management effort. In that framework, the purchasing agent then must consider the strategic questions for a particular purchase:
Initial design services by the metalcasting
Will the cast part be purchased from a single
Working Relationship Requirements. The working relationship between the customer and the metalcasting facility is a critical factor in defining a purchasing strategy that will provide best value for the casting purchase. The buyer must recognize from the first step that a cooperative team effort between the production design engineer and the metalcasting facility engineer is the key to a cast component design and production effort that meets the desired quality, cost, and schedule requiements. The buyer should work with the production design engineer during the initial purchase plan to define the goals, tasks, level of effort, and schedule required for a successful buyer/metalcasting facility working plan. The plan must consider consultation, technical services, and information exchange at each stage of the purchasing process. This plan should include:
source or multiple sources? What levels of metalcasting facility capability and production experience are necessary to successfully produce and deliver the casting on time and at cost? Do current casting suppliers have the capability and capacity to meet the new purchasing requirement? How will the performance of the supplier be evaluated in terms of quality, schedule, and cost? How will the supplier be integrated into the initial design process and future productimprovement efforts? What is the plan for developing and maintaining an open exchange of information between the purchaser and the supplier? How will the benefits, risks, and costs of the purchase be shared between the purchaser and the supplier? What kind of contract (firm-fixed price, cost plus fee, contract with performance/penalty fees) will best meet the purchase objectives and the overall purchasing strategy?
Deliverables Expected. If cast components were simple commodities that were bought off the shelf, then a contract would require only the delivery of a finished qualified component at a certain location, on a specific schedule, in a known quantity, and for a given price. However, each casting is unique, and each metalcasting facility has different capabilities, skills, and experience to bring to the table. Engineering service and supplier partnerships are critical factors in the design, production, and delivery of reliable, quality cast parts. For that reason, a deliverables list in a contract for
Fig. 4
Selection criteria of a purchasing plan
facility
Scheduled information exchange on quality Schedule and cost achievements and issues Quality planning and assurance during
production
Quality and cost-improvement initiatives
with shared benefits during production
Initial casting design (for performance and
manufacturability)
Tooling design and construction Casting development and qualification Component production with value enhan-
cement Selection Criteria. A purchasing plan always will include the criteria for evaluating, prioritizing, and selecting the bids from the responding metalcasting facilities. The criteria for qualifying and/or selecting a metalcasting facility and its bid for a specific casting fall into five categories (Fig. 4): Production facilities (capabilities, capacities,
and expertise)
Direct experience and past performance with
the required alloy, similarly sized and
complex parts, and the desired/ recommended casting method Production organization and capabilities (design and engineering support, testing and evaluation capabilities) Quality-assurance procedures and capabilities Quality, delivery, and cost (goals and reliability) Based on the purchasing objectives, the defined strategy, and the detailed deliverables and requirements, the purchasing plan will specify the criteria in detail along with weighting factors for evaluating the bids. Depending on the complexity, critical value, and risk for a given cast part, the evaluation approach for bids may take two general forms: For low-risk castings, qualify bidders as go/no
go against the different criteria, eliminate unqualified bidders, and then rank qualified bidders by quoted cost. For complex, critical components, grade bidders against the designated criteria (not qualified, qualified, or exceptionally qualified), use weighting factors for the different criteria, and then rank bidders high/low by summed and weighted criteria. The OEM buyer has a responsibility to assess the capability of a metalcasting facility to meet the purchase requirement. Assessing the capability of a metalcasting facility to produce high-strength structural cast components is difficult when the metalcasting facility is new to the buyer and does not have a track record based on previous purchases. A critical review of a metalcasting facility facilities, organizational structure, and process documentation is necessary to judge its capabilities. The capability survey is usually combined with qualitycontrol and business assessments before any effort to use the metalcasting facility as a supplier occurs. A questionnaire to address these issues is given in Fig. 5.
20 / Applications and Specification Guidelines The reviewer/assessor should write a capability report on the facility and the type of castings it is capable of providing. In considering and evaluating the five selection factors, the assessor should use the following guidelines to focus inquiries and determine the capabilities of the metalcasting facility. Metalcasting facilities making structural castings must have the proper equipment to attain the necessary quality. The production capability (and available capacity) of the facility should be consistent with the production quantity and rate required by the contract. A capability assessment should be made before giving any metalcasting facility a purchase order. Costs and schedules may be adversely affected if the capability assessment is not performed conscientiously. A user survey team approach often is necessary; the team usually is composed of engineering, quality-control, and purchasing personnel familiar with metalcasting facility operations. The survey usually is not repeated unless there are changes in the metalcasting facility, personnel, or organization or if a significant time has elapsed since the metalcasting facility last supplied components to the user. Metalcasting facility records should be available for similar castings that were successfully produced to comparable requirements. This demonstrated experience is an important factor in qualifying and selecting a metalcasting facility that can meet the quality and production requirements for cast metal parts. Metalcasting facility records should include mechanical test results and photographs or drawings of cast components. These results should be verified by periodic destructive tests. Separate test bar and appendage data are not adequate proof of capability. Metalcasting facilities that have not had experience producing the size, complexity, or quality of the required castings will need time to develop this expertise. Often, metalcasting facilities without experience are not aware of the requirements to cast a particular type of component. To gain the experience needed, they should start with small, less complex parts. Complex, high-performance metal castings must have high radiographic quality to be capable of exhibiting the necessary mechanical properties. Experience, combined with the quality-assurance equipment and personnel of the metalcasting facility, indicates if the metalcasting facility can achieve high-quality cast components. The ability of the metalcasting facility to hold tight dimensions regarding the location of critical features and the location of contoured mold lines is important. Special equipment, techniques, and procedures to mold, trim, heat treat, straighten, machine, and inspect are necessary to accomplish this on a production basis. The organizational structure of a metalcasting facility varies to some degree with the size of the facility. However, the organizational structure of most metalcasting facilities is basically the same:
senior management, production, quality assurance, marketing/sales, and maintenance/operations. The larger facilities usually have a greater depth of personnel in each department. Smaller metalcasting facilities are under direct day-byday control of fewer people. Sometimes, only the owner can make decisions. Large organizations tend to delegate authority, and the owner or chief executive officer is involved only in business decisions on an administrative level. Metalcasting facilities that produce quality cast parts need the technical capability to develop metallurgical and dimensional procedures and to solve metallurgical and dimensional problems as they occur. A degreed metallurgist with several years of on-site experience should be a permanent part of the organizational structure. Each cast component is produced in a manner that is specific to that part. Casting procedures must be adequately documented, and methods should be in place to monitor and record the actual casting process. The details of each operation should be maintained in a manner that permits them to be retrieved easily each time production is required. Photographs can aid considerably in the process description. Details on rigging, chilling, coring, molding and mold assembly, melting, pouring, processing, welding, heat treatment, straightening, and final inspection should be recorded during processing and should be available for review. A quality-control department should be staffed to implement the necessary inspections and testing throughout the entire metalcasting facility process (e.g., Fig. 6), from the receipt of raw material through the shipment of the finished cast component. There should be a production-control department that has an efficient system of processing each purchase order, from receipt of the order to final shipment, and that can retrieve daily input from the floor regarding the production status of each order. The reliability of a metalcasting facility in meeting customer requirements is best determined by direct experience and demonstrated reliability on previous orders. If a history has not been directly established, references and verified performance on other contracts should be provided. However, performance must be considered at two levels: a demonstrated record of performance across all orders as well as the capability and expertise to meet the performance, cost, and schedule requirements of a specific cast component in terms of the required alloy, complexity, molding technique, and quality control. Purchasing Tasks, Responsibilities, and Schedule. The final step in developing the purchasing plan is to lay out the specific tasks, responsibilities, and schedules that must be achieved to meet the required production and delivery times. Schedule layout is an essential element to integrating and coordinating the various tasks that must be accomplished in purchasing metal castings. One of the largest frustrations felt by a casting user is when a metalcasting facility does not meet expectations. It is important to
accurately and completely specify requirements, because inferior specifications will lead to poor performance or unnecessary costs due to overspecification (Table 1). Ideally, cast components are ordered after the metalcasting facility quotes on the job and all specification requirements are given in the purchase order (PO) for accurate pricing. If the metalcasting facility has taken exception to any of the specifications, these must be noted in the PO.
Fig. 5
Questionnaire for foundries—short form
Fig. 6
Gas testing for quality assurance
Metalcasting Technology and the Purchasing Process / 21 Table 1 Sample cost differences attributed to overspecification Overspecified procedure
Additional costs, $/lb
Melting Heat treatment for riser removal Weld repair, preheat Increased cracking susceptibility
0.05 0.08 0.04 0.04
Total additional cost
0.21
Guidelines on Specifications for Accurate Pricing. The following is a guide for cast component purchasers to navigate their way through the pitfalls of over- or underspecification and toward accurate pricing to receive the most from their cast parts. For new jobs (first-timeordered castings), the following information must be included in the RFQ and PO. Parent number and engineering drawing with the correct revision for both casting and machining drawings must be supplied. The drawings must be legible, with tolerances clearly indicated. The referenced drawing number and revision in the PO must coincide with the actual number and revision of the drawing. Quantity, price, and delivery must be clearly defined, with no misunderstandings. For example, an order for 5000 ductile iron automotive connecting rods can read, “5000 castings— 1000 pieces by October 1999, 500/month thereafter.” Castings can be supplied in various conditions, such as blasted (shot or grit), ground, machined (rough or finished), heat treated, pickled, and/or painted. They can be shipped in open or closed boxes, crates and metal tubs, or individually packaged for protection. For every metalcasting specification, there is a particular grade or class of alloy. Examples of terminology in material specifications are as follows: In low-alloy steel, the grade “low” alloy
refers to the minimum mechanical properties required. For example, ASTM A148 grade 105-85 means that the test bars must meet or exceed 105,000 psi tensile strength and 85,000 psi yield strength. Although not directly noted by the grade, there also are minimum values for ductility, as measured by the percent elongation and reduction of area from a tensile test. In ductile iron castings, the tensile yield and elongation minimum requirements are specified in the grade. For example, the ductile iron specification ASTM A 536 class 7050-05 means that the minimum requirements are 70,000 psi tensile strength, 50,000 psi yield strength, and 5% elongation. In gray iron castings, the class refers to the minimum tensile strength and the size test coupon from which properties are determined. For example, ASTM A48 class 3DB requires a 30,000 psi tensile strength determined from a 30.5 mm (1.2 in.) diameter test bar. Class A requires properties to be determined from a 22.4 mm (0.88 in.) test bar.
Aluminum cast components, similar to cast
iron, can be supplied in the as-cast or heat treated condition. The most critical aspect of ordering cast components is quality requirements. How does the casting purchaser assure that the quality requirement is met? It is difficult to assign quality requirements to the casting, especially nondestructive testing (NDT), but without sufficient knowledge of existing NDT methods and specifications, the result is either an overly conservative approach or one that is too stringent. Therefore, some knowledge of the test methods and quality levels is an asset for casting purchasers to ensure that they receive what they need. In addition to the standard requirements, the purchaser often has supplementary requirements that are not covered under the designated specification. For example, a steel casting may require good weldability, and as such, the composition may be adjusted to ensure the carbon equivalent is below 0.5. On the other hand, an application in which the casting is subjected to impact loading at colder temperatures may require that Charpy V-notch tests be performed at the metalcasting facility. A critical aluminum alloy casting may require traceability to the heat from which it was produced. This can be accomplished by identification of the casting by serial number. Stainless steel castings, which require superior corrosion properties, can have a maximum allowed carbon content. Certifications are required only when they are critical to the cast component performance. Common specifications are chemistry (usually obtained from an analysis of the heat but also from test bars). Tensile properties also are commonly specified from a test bar but less commonly from a test block of equivalent section thickness as the casting. Other specifications may include liquid penetrant, radiography, ultrasonic, and magnetic particle testing. It is sometimes necessary to record and certify heat treatment data. In more critical applications, the actual heat treatment chart may be required. Expectations for the sample casting must be clearly defined. Are they required to be liquid penetrant inspected? Does each component require an x-ray? Is 100% layout required? Is written sample approval required prior to production? Are castings subject to third-party inspection? If sample requirements are not listed, the metalcasting facility must assume that sample approval is not required prior to production. For repeat orders with no changes, it may be sufficient to reference the requirements from the original PO. For repeat orders with an engineering drawing revision, this must be noted clearly on the PO, with the new drawing attached.
Requesting and Evaluating Quotations/Bids The second step of the purchasing process is requesting and evaluating quotations/bids from
metalcasting facilities. This section describes how to locate and identify potential metalcasting facilities, request and receive quotations, and then evaluate and rank the offer for best value. Locate and Identify Suitable Metalcasting Facilities. There are many ways to locate suitable metalcasting facility suppliers: current casting suppliers, names selected from a buyer’s guide, advertising material and magazines, and recommendations from other buyers and engineering staff. In all cases, the buyer must actively seek reliable information about sources of supply and try to match metalcasting facility capabilities and capacity to purchasing needs. Replies should be studied carefully because they will provide the guidelines for future purchasing decisions. For example, the buyer will know whether the supplier can produce the grade of material required in sufficient quantity. Attitudes toward quality can be assessed from the percentage of total production and qualitycontrol personnel directly involved in the quality-control program. The buyer should recognize that the strong emphasis of a metalcasting facility on quality control indicates it is probably producing high-grade cast components. The price structure will reflect this commitment to quality. If the component is very complex, look for a supplier who specializes in this type of cast component. Discuss capacity, to avoid disappointment after the order is placed. Buyers would do well to make buying decisions on the realities of the situation: price, quality, and inspection standards required for any given set of physical and mechanical properties. Perhaps the most important point, when all other conditions have been satisfied, is to look at the current pattern of production at the metalcasting facility. Important factors can affect prices and quality of the cast part, such as whether or not the metalcasting facility is high volume, casts small components, is a jobbing shop, and so on. An initial survey of casting sources will provide a list of candidate suppliers, based on capabilities and expertise. With a source list, the next purchasing step is to request bids. Maintaining a Good Relationship. The buyer-supplier relationship does not end when the appropriate metalcaster is found. A close, mutually beneficial relationship must be established between the casting user and the metalcasting facility to produce acceptable, cost-effective cast components. The metalcasting facility must be encouraged to learn more about the casting application, including high-stress or critical sections. With this knowledge, it can suggest additional safeguards for good performance. The metalcasting facility also can suggest various quality requirements to assure that purchasers receive what they want. In some situations, a redesigned casting can have greater integrity at the critical sections. The casting purchaser (and user) also is encouraged to learn more about castings and processes, including newer technologies for producing high-quality
22 / Applications and Specification Guidelines cast parts. Cleaner casting practices allow weight savings in many types of cast components without sacrificing quality. Send Out the Bid Package and Review Quotations. The bid package consists of an RFQ form along with supporting technical drawings/information. The bid package should include contact information for the cast component buyer and the production designer as well as the closing dates for quotations. Requests for quotations may be distributed by both paper and electronic means, such as fax and e-mail. During the quotation process, metalcasting facility engineers often call the buyer and/or the designer for additional information on technical, quality, quantity, cost, and schedule and service issues. The buyer and the design engineer should be prepared to discuss these issues with the responding metalcasting facilities. When quotations are received, they should be reviewed for completeness and responsiveness. Incomplete quotations should be returned to the sender for review and resubmission. Grade and Rank the Bids/Quotations. When all the quotations have been received, they must be graded against the selection criteria defined in the purchasing plan. This may be an informal process or a detailed review and evaluation of the bids, based on the size, value, complexity, and risk involved in the component purchases. Evaluation of the supplied quotations consists of applying the selection/criteria defined in the original purchasing plan: Assess the quote against the specific quality,
quantity, schedule, technical service, and cost targets for the casting. Determine the overall reliability and value of the individual bids against the supplychain management strategy. Rank/classify the quotations in terms of the best value.
Select a Supplier and Negotiate Contract Terms After the initial bids have been received, evaluated, and ranked, the third step is selecting the qualifying and best-value bids. If the component is well defined in terms of simple shape/complexity and of relatively low risk, the quotation may be sufficient for closing the contract and moving into production. However, for the more complex, high-value cast components, additional engineering design and contract negotiation may be required to reach final agreement. Specifically, it may be necessary to: Perform a final review and refinement of the
design for manufacturability with the casting engineer Finalize technical, quality, and schedule requirements Review and resolve nontechnical contractual issues
In contract negotiations, all parties should consider the following factors:
Find a better way to phrase existing clauses Improve the agreement by adding clarifying
clauses
Level and scope of technical service, includ
ing design consultation, production, quality control, and product improvement Detailed instructions and guidance on specifications and quality-assurance requirements Assigned responsibilities for and integration of supporting engineering activities among the various production activities. These include the patternmaker, metalcasting facility, machine shop, and heat treater. Cost of quality assessment (rejects, delays, rework, etc.) Scheduling and lead-time planning Product liability
After these factors are considered, the next steps are to review, negotiate, and resolve cost details, performance clauses, and conditions of sale and then finalize and sign the contract. Figure 7 is a sample template on conditions of sale that are considered in casting purchase contracts. Terms and Conditions for Sale Agreements in the Metalcasting Industry. Over the years, there have been constant inquiries about what constitutes typical terms and conditions of sale in the metalcasting industry. There is no association-endorsed standard agreement. However, the North American Die Casting Association compiled terms and conditions of sale information from dozens of metalcasting operations around the country in 2002. Each agreement was separated into various clauses to constitute a comprehensive agreement between the seller and the buyer. If there was a significant redundancy from one agreement to the next, then the exact text for a particular passage was not repeated. One section identifies the common points for each clause that creates the agreement. In addition, the percent of agreements that contain each common point is reported. Finally, a list of atypical clauses is discussed in the section about unique or special terms. This section should be given special attention, since several adept clauses may protect the interest of a foundry. The Report on Terms and Condition of Sale Agreements in the Die Casting Industry—2002 is meant to provide the foundry with information on various terms and conditions currently being supported by the metalcasting industry. (American Foundry Society (AFS), American Metalcasting Association (AMC), Advanced Technology Institute (ATI), Defense Logistics Agency (DLA), North American Die Casting Association (NADCA), Non-Ferrous Founders Society (NFFS), and Steel Founders’ Society of America (SFSA) neither endorse nor recommend the adaptation of any of the terms or conditions documented in this report, which is available through NADCA at http://www.diecasting.org/publications/default.htm.) These terms are available electronically for cutting and pasting to help those purchasing castings to:
The principal sections include:
Alloy and tolerances Cancellation Clerical errors Credit in returns Delivery Design changes Dispute resolution Engineering disclaimer Errors in quantity Inserts Overage allowances Packaging Patents Patterns, foundry, tooling, and dies Payment and credit terms Preliminary agreement Pricing Samples Taxes Unique or special terms Warranty
Contract Fulfillment and Continuous Improvement A signed contract is not the completion of the purchasing process. The contract must be completed by two actions: on-time delivery of a fully qualified product at the agreed cost by the metalcasting facility, and prompt and timely payment by the buyer. During the life of the contract, all parties have a valuable stake in continuous improvement in quality, delivery, and best value. Specific issues that must be considered during the production cycle include: Open and timely communication (by visits,
phone calls, e-mail, etc.) between customer and supplier on project status, shortfalls, potential problems, forecasted changes, and opportunities for value improvement Monitoring of quality, delivery, cost, and performance by both the supplier and the customer Identification of opportunities to improve quality and increase value to the final customer by the customer and supplier
Considerations in Purchasing Metal Castings The metalcasting process requires designers to consider a specific list or requirements prior to purchasing that they may not otherwise be familiar with. The following is a list of the most significant considerations when attempting to determine the overall cost and design requirements of a metal casting.
Metalcasting Technology and the Purchasing Process / 23
Fig. 7
Conditions of sale—sample contract
24 / Applications and Specification Guidelines Production Cost Factors. Efficient and effective purchasing of cast components requires the use of a considerable amount of specialized information. With the increased emphasis on value-analysis techniques, purchasing agents must evaluate the many important factors that directly influence the total cost of a component part. The total cost factors are usually grouped under the general headings of quality, service, price, and delivery. Each buyer will determine the relative importance of these factors, based on particular circumstances. Since a universal formula is not possible, the following information will assist purchasing managers by covering the important criteria that affect the cost of a cast component. It is important that the casting supplier be included during the early stages of the design and purchase planning. Regardless of the type of metalcasting facility, dollar costs will break down into three main categories: materials, direct labor, and expenses. Materials may be subdivided into raw materials (primary metal and alloying constituents) and process materials (coke, sand, molding materials, bonding agents, shot, paint, etc.). Automation has markedly improved productivity in the metalcasting facility industry, but labor remains a high percentage of costs. Expenses are extremely variable, depending on the metalcasting facility facilities and the cost-accounting procedures. Although there is no magic formula that can be universally applied to metalcasting facility costs, it is possible for the buyer to develop a basic understanding of the cost factors for any metalcasting facility or particular cast component. Numerous factors influence the total cost of a finished cast component. These are considered in two groups: factors directly affecting the cost of a cast component and indirect cost factors for a finished casting. Dimensional tolerances and allowances are also discussed, since they have an important bearing on casting costs. Size, weight, and complexity of a metal casting are generally the most important cost factors, because they directly influence the materials and labor required for its production. Weight of the metal in a cast part also is important, but it no longer dominates the cost, as was true at one time. Overall size of a casting also is important, because several tons of mold material and other expendables are involved in the casting process for each ton of cast components produced. A large, thin casting that must be made in a large mold can be much more expensive than a heavier, simpler part made in a small mold. When designing a cast component, complexity is also an important factor. In general, the more complex a design, the more difficult it will be to manufacture. This may lead to higher costs. However, this is not an ironclad rule. With careful design, complexity can be molded into the metal casting without increased production cost (Fig. 8). To minimize the cost of complex designs, the engineer should call on the metalcasting facility to make manufacturing
suggestions at the early design stage. The metalcasting facility can help to minimize costs and problems with such things as draft, fillet radius, parting line locations, coreprints, wall thickness, and surface finish. The type of pattern equipment provided is an important factor in the total cost of producing a metal casting. The term pattern equipment includes coreboxes, if they are necessary, and any jigs or fixtures used for checking or assembling cores or for checking the dimensions of the part. Note that part quality is only as good as the pattern quality. Geometric complexity of a design raises costs because the mold must be much larger in order to contain external coring to form complex outer configurations. Cored castings are more costly. Cores must be made ahead of time, transported to the molding department, and placed carefully in the mold. However, equally important is the extra cleaning and trimming that is usually necessary on cored cast components. At the same time, the cores make the metalcasting process unique since they allow for the creation of internal cavities while the external surfaces of the part are produced. Wall thicknesses are an important design feature in metal castings, with the part configuration and alloy selection playing a major role in determining what wall thicknesses can be cast. As an example, in aluminum castings, the normal minimum wall thicknesses will range from 3.0 to 6.4 mm (0.12 to 0.250 in.). Depending on the part size and the area involved, thinner wall sections, down to 2.0 mm (0.080 in.), are possible over limited areas (65 cm2, or 10 in.2). Type of Metal. The cost of a casting will vary in three ways, based on the alloy used: Inherent cost of the alloy: Cost depends on
both the baseline cost of the metal along with the alloying elements in the alloy. Purity of the alloy: Premium prices must be paid by the metalcasting facilities to obtain alloys with fewer impurities. The lower the impurity level requirements of the alloy, the fewer the gating and risering returns allowed in melt charges. Castability of the alloy. In steel and aluminum alloys, various alloy compositions have differing levels of fluidity, weldability, and castability. With reduced castability, additional metal must be poured into the mold in the form of flow channels and feed heads to assure a sound casting. For such parts, the weight of the actual metal in the finished cast component may be less than half of the total weight poured. This total-metalpoured factor is especially important when considering alloys with costly alloying elements, since the alloying elements must be added to all of the metal poured into a mold. Of these three considerations, castability of the alloy has the most effect on the cost of a cast component.
Fig. 8
Comparison of (a) fabricated agricultural component and (b) redesign to a one-piece casting. Metalcasting allows production of geometrically complex shapes, consolidating multiple welded components into one casting. This casting provided a cost-savings while improving functionality.
Dimensional Allowances and Tolerances. Because dimensional allowances and tolerances affect both the cost and delivery of metal castings, it is important that part buyers have a working knowledge of the major factors influencing them. Realistic tolerance requirements are of prime importance in minimizing overall costs. Tolerances that are more rigid than necessary increase metalcasting facility production costs and lengthen delivery schedules. Those that are too loose result in castings that require more expensive machining and usually have a heavier section size than required. The following is a brief review of the more important dimensional allowances and tolerances that should be considered. Shrinkage Allowance. When metal is poured into and takes shape in the mold, the metal is still at or above its liquidus temperature. As the metal solidifies and cools to room temperature, the metal contracts. Thus, to obtain a casting of desired size, the pattern is made slightly oversized. This addition to the dimensions is called a patternmaker’s shrinkage allowance. Another type of shrinkage that occurs while the liquid metal solidifies in the mold is called solidification shrinkage. This aspect of metal solidification is corrected by feed heads, risers, or shrink-bobs, which are attached to the casting when they are poured. This is a separate phenomenon and should not be confused with pattern shrinkage allowance. Draft Allowance. This is the term describing taper on the vertical faces of a pattern that allows the pattern to be removed without binding or tearing the mold. Draft will normally vary between 1 and 3 of taper. Small or machinedrawn patterns require minimum draft. Large patterns, inside pockets on patterns, and handdrawn patterns generally require a greater draft. Under ideal conditions, some patterns as deep as 150 mm (6 in or more may be drawn with almost no draft, using special techniques. Maximum draft should be provided where it does not interfere with desired casting geometry.
Metalcasting Technology and the Purchasing Process / 25 The metalcasting facility should be consulted so that draft allowance may be minimized or even eliminated entirely on surfaces where it must be machined off. For a no-draft condition, a loose piece in the pattern or corebox can be used. This will increase the cost of the part due to an expensive pattern and higher labor costs in molding and coremaking. Alternatively, the buyer should consider the lost foam process. Cast tooling or datum points should be fixed on designated no-draft surfaces. Machining Allowances. Additional stock that is added to the surface of a metal casting so that it can be finished to size by machining is called machine-finish allowance. Thickness of this allowance will depend on size and type of the cast component, the type of surface, how it is to be machined, and how accurately the casting can be made. Large castings and those made on a low-volume basis will generally have larger machining allowances than smaller castings or components produced in large quantities. With newly designed cast parts, a full machining allowance is usually added. After some experience is gained with the pattern, machining allowances may be reduced to the minimum necessary. To avoid local surface hard spots, a ground finish can be applied directly to a cast surface, with just enough stock so that the surface will clean up well. A general guide to machining allowances is given in Table 2. Distortion Allowance. When one part of a casting cools more rapidly than another, the part may become distorted. The first line of defense is to pay attention to design details, such as avoiding sharp changes in adjacent sections. When this is not possible, distortion often can be minimized by special metalcasting facility practice to control heat extraction or by followup heat treatment. In some cases, the pattern may be intentionally distorted so that the casting will be true. This is called distortion allowance. Tolerances. Because castings are individually designed and process details are engineered for the characteristics of each pattern, casting requirements such as surface smoothness, dimensional tolerances, or machining allowances cannot be stated in a general specification. These requirements usually are established through mutual agreement between the casting user and the metalcasting facility. Ordinarily, variation in casting dimensions can be held to within half of the patternmaker’s Table 2
Machining allowance guide On bore
mm
in.
mm
Up to 305 330–610 635–1067 1092–1524 1549–2032 2057–3048 Over 3048
Up to 12 13–24 25–42 43–60 61–80 81–120 Over 120
3.2 4.8 6.3 7.9 4.8 11.1
in.
Outside surface mm
1=8 2.4 3=16 3.2 1=4 4.8 5=16 6.3 3=16 7.9 7=16 9.5 Special instructions
Use the proper shrink factor in pattern
Guide to pattern machine-finish allowance Pattern size
shrinkage allowance. On a simple 305 mm (12 in.) casting dimension, this would be a tolerance of þ2.1 mm (þ 0.0825 in.). Closer tolerances often are obtained on repetitive production where corrective adjustments can be made. For such cast parts, a tolerance of 1.3 mm in a 203 to 305 mm (þ0.05 in. in an 8 to 12 in.) dimension may be con sidered normal. Special procedures have been developed to produce some cast components with dimensions as close as 0.13 mm (0.005 in.). However, note that tolerances do not vary proportionally with the component dimensions. Dimensions within each half of the mold are generally more accurate than those across the parting plane or those related to a cored surface. The type of pattern equipment used directly affects dimensional tolerances. This is particularly true when machine molding or coremaking methods are employed. Metal patterns and coreboxes that are machined to size often will provide closer tolerances than will cast-to-size metal or wooden patterns. High-strength cast components frequently demand tight dimensional tolerances or thin-wall sections. In either, higher production costs are incurred if a systems approach is not used by the OEM to work with the metalcasting facility to recognize the true requirements of the component. Frequently, components are overspecified. In other cases, the design can be altered to take advantage of the capabilities of the many casting processes that are commercially available to meet true design requirements. The factors that most affect dimensional tolerances are:
in.
3=32 1=8 3=16 1=4 5=16 3=8
design. This is less of a problem on smaller castings. For larger castings, the metalcasting facility must have experience with the type of part required so it can make proper judgments about what shrinkage factors to use. Variability in the pattern equipment and coreboxes from nominal blueprint dimensions. To minimize this, quality tooling is required. Variability in the molds and cores themselves due to the people, equipment, and materials used in their production. Quality metalcasting facilities are aware of this and have controls and gages in place to minimize the effect. Variability in cleaning and finishing operations performed on the part. The metalcasting facility will perform inspections after this process to assure that finishing operations are done properly. Distortions due to the heat treatment of the castings. The metalcasting facility should plan to use heat treatment fixtures and special quenchants where the potential for this problem exists.
The quality of both the tooling and the production equipment is of great importance in dimensional tolerances. Even if only a small quantity of castings is required, the tooling
must be of prime quality if the dimensional tolerances are to be satisfied. Designers must be aware that there are additional costs incurred when tight tolerances are specified. Linear tolerances often must be increased for larger casting sizes. Critical dimensions should he identified to the metalcasting facility so that proper tooling will be built. Cast Surface Finish. There is no standard finish specification for cast components. Although root mean square values for cast surfaces are sometimes given on drawings, they are of questionable value because there is no dependable method of quantitative surface roughness measurement for unmachined cast surfaces. However, sample castings of the type required provide a satisfactory way of designating desired surface finish and observing what can be produced. In some casting applications, the degree of roughness of the cast surface is critical to successful use. Dimensions that are held within close tolerances may be ineffective if the surface is not smooth enough for accurate measurement or proper positioning in a fixture. While a textured surface is desirable or even necessary for some finishing or coating procedures, a surface that is too rough can be detrimental to finish quality and may require extra finishing, with increased costs. Although a smooth surface on a cast part often is considered an aspect of its quality, this is not an accurate indicator of the overall quality of the casting. Surface quality can involve several characteristics that may not be evident or even measurable. Superficial roughness on the casting can be caused by duplicating the rough surface of a pattern in metal or by an improperly made mold. However, most of the surface roughness of iron castings is the direct result of the highenergy abrasive cleaning methods in common use. Accurately evaluating a cast component surface roughness is difficult. Instruments and methods that are used successfully for measuring the roughness of machined surfaces are unsuitable for cast component surfaces. A sand cast surface is not composed of regularly spaced surface features, as found on a machined surface. The size distribution of the sand mix produces a much more complex and richer pattern of surface features. A finish that varies from location to location on a casting (due to inherent variations in sand compaction around a complex engineered geometry) further complicates the problem of measurement. The surface texture or roughness is generally specified by the term root mean square (rms) or arithmetic average (AA). The rms yields approximately 11% higher values than AA, but the difference is not significant. The surface finish may be specified as 300 rms maximum, 115 AA maximum, or C50 maximum on a visual comparator. Finer finish requirements will require special molding media or secondary operations,
26 / Applications and Specification Guidelines such as grinding or machining. This will add to the cost. Measurement should be made by surface comparator. The profilometer probe diameter acts as a filter for the input data. Fine surface areas should be identified and specified on drawings. Order Quantity. An important factor in selecting appropriate tooling is order quantity. Even with a given pattern, the order quantity may have an important influence on casting cost. One reason for this is that there are a number of standard operations in the metalcasting facility that must be performed for each part order, regardless of the quantity involved. Some of these are: Quoting, acknowledgment, and recording of
the order
Scheduling production and issuing shop
work orders
Retrieving pattern equipment from storage
and inspecting and transporting it to production departments Setting up tooling for production (i.e., laying out or attaching to machines and lining up proper flasks) If cores are required, delivering core equipment to coreroom, making cores, and transporting them to the molding line Shipping and billing Efficiency of production operations is also greatly influenced by quantity. When an order is issued subject to release, the total quantity is important in the selection of metalcasting facility tooling, but each release is normally considered as a separate order by the metalcasting facility, since each production order requires separate processing. The purchasing manager can determine the most economical lot size for cast parts when he knows what his casting requirements will be for some time in advance. Due to usual billing practices, less than a one-month supply is seldom practical. Savings realized by larger releases should be weighed against inventory costs, which include investment, space, and records. Possible changes in product demand and casting design or prices may not be equated directly in numbers but are incorporated as a matter of judgment by experienced buyers. The production quantity and the production rate are major factors in selecting the casting process and the type of mold pattern/tooling for the casting run. Various metalcasting processes have different production rate capabilities, different tooling costs, and different design change flexibility. Sand cast components often are preferred, because design changes can be made with relative ease. Two factors play important roles in tool materials selection. The first factor is the total quantity of cast parts needed. The second factor depends on the delivery period involved. Quality wood tools will generally produce up to 100 parts within approximately a year. Over a longer period of time, the tool may distort and
produce castings with dimensional discrepancies. Wood and plastic tooling should be usable for up to 500 parts in a moderate time frame (one to five years) with periodic refurbishment. Whenever small lot quantities are required for more than three years, metal or plastic tooling should be used to assure only minor refurbishment costs. Metal tooling is recommended when more than 2000 castings are required annually. Additional Operations. Often, the metalcasting facility can perform additional production operations that may increase the direct cost of the casting but will reduce the total cost of the finished product. Heat treating, painting or coating, and inspections that are more sophisticated and testing procedures are typical of additional operations that may be efficiently performed by the metalcasting facility. For quantity production runs, castings may be justified or targeted in a gaging fixture to provide accurate locating points for subsequent machining operations. This assures that the correct amount of stock is available to be removed on each critical finish surface. Technical Services. Close cooperation through concurrent engineering between the metalcasting facility and its customer is the approach required to yield the greatest benefit for buyer and supplier. However, for the more complex engineered cast components, an additional amount of service or technical assistance may be required from the metalcasting facility. This commonly occurs in development of new products. Assistance with component design, information on properties and metallurgy, troubleshooting, pattern assistance, or pattern models and experimental castings may be provided by the metalcasting facility. These services may be billed to the customer at cost or included in the price of the parts. However, in either case, quotations from a metalcasting facility that stands ready to provide its customers with this assistance cannot be compared directly with quotations from a metalcasting facility that does not. The latter type of metalcasting facility should not be ignored, because it can be a good, economical source for castings. This is particularly true where special problems or requirements do not exist and where additional services are not necessary. When a complex casting will be produced in large quantities, a model casting may be made to assist in planning production equipment. The model will assist in establishing the most efficient metalcasting method by providing an opportunity to establish critical factors such as the parting line and core prints on the actual shape. This also can be of assistance in planning for subsequent processes, such as machining. Types of Price Quotations. At one time, metal was expensive compared with labor, and castings were commonly sold by the pound. This was a simple buying method and worked well because the weight of a cast part was a more important part of total cost than was the
workman’s time to make it. This is no longer true. With the tremendous increase in cost of labor and mechanized equipment in a modern metalcasting facility, the value of the metal in most castings has become a secondary factor. Accurate cost accounting must include cost of the worker’s time, value of materials used, and overhead costs. This provides price quotations that are based on cost per casting and a ready reference for the purchaser to arrive at unit cost. The cost-per-piece can easily be converted to cost-per-pound, but such a figure cannot be justly compared to other castings that may be ordered in different quantities or require more or less man-hours to be produced. Some metalcasting facilities, however, still retain the price-per-pound method of quoting for specific types of parts. Thus, equivalent quotations based on price-per-pound and price-per-piece may show variations when compared for individual castings. This is particularly true when the quotation is based on an estimated casting weight. Variations of actual from estimated weight still affect the cost more on a price-per-pound basis than on a price-perpiece basis. Lead Time and First-Article Delivery. Essentially, the metalcasting process involves pouring molten metal into a cavity close to the final dimensions of a desired component. It is the most direct and simplest forming method available for metals and is often an essential production factor in rapidly manufacturing finished metal parts. Lead time is commonly defined as the amount of time between contract agreement and delivery of the first article, either prototype or initial production. Lead time is reduced with metalcasting compared to other manufacturing methods by eliminating and/or shortening the time required for tool production, parts ordering and delivery, assembly, finishing, and machining. In the metalcasting arena, new rapid prototyping (RP) tools (Fig. 9) are being developed to produce patterns and tools for castings. The RP tools can dramatically reduce lead times and production costs required for patterns and tools, particularly for designs that have complex detail that would be difficult to produce by conventional machining. An example of an investment casting produced with rapid tooling is shown in Fig. 9. Quality Assurance and Specifications. Quality assurance has benefits for both the buyer and the seller of metal castings. The buyer is assured of a cast component that meets performance requirements and the designated specifications. The metalcasting facility has confidence in its production process, the documentation of the process methods, and the application of quality standards in its day-to-day operations. All parties benefit because a high-quality product is delivered on schedule without cost overruns. Standard specifications for various types and grades of alloys and castings are promulgated by several national organizations and
Metalcasting Technology and the Purchasing Process / 27
Fig. 10
Fig. 9
Example of rapid prototyping process with investment casting. (a) Valve body prototype prepared for investment casting manufacturing without hard tooling. (b) Soft tooling for production investment casting mold. (c) Final casting. With rapid tooling, metal casting has developed to the point of producing castings in days, with no tooling.
government agencies. The most commonly used specifications for alloys and metal castings in the United States and Canada are those of the ASTM International, which is composed of many volunteer committees that write and maintain the standard specifications. Each committee is assigned jurisdiction over standards for one type of material, application, or testing methodology. The Society of Automotive Engineers (SAE) also publishes commonly used specifications for cast components. Other organizations, such as the American Society of Mechanical Engineers and the American Petroleum Institute, have specifications concerning castings for specific applications, but the material portion of these specifications typically refers to an appropriate ASTM International specification. The engineering departments of some companies that design and use a large number of castings write their own specifications. Some casting applications require an unusual combination of
Indentation profiles of three types of hardness testing methods
properties that would not be provided by cast parts made to standard specifications. Types of Specifications. There are several methods of specifying cast components. Specifications can be written to designate the properties of the metal that are specifically needed (strength, hardness, finish, accuracy, or special characteristics, such as resistance to heat or corrosion). In addition, there are many casting applications in which the ordinary properties of the designated type of alloy are quite adequate. In purchasing metal castings for noncritical applications, there may be an advantage in economy and availability to ignore standard metal specifications. The purchase order simply could designate the generic type of alloy by name (gray iron or malleable iron). Other requirements, such as for dimensional tolerances, may be added. When the mechanical properties of the metal are important, there are two commonly used methods to designate this requirement. The more frequently used procedure for engineering metal castings uses a standard test bar that is poured along with the specified lot of parts. The required minimum tensile properties should be obtained in testing a sample from this test bar. Most ASTM International specifications apply this method to qualify the metal used to pour the casting. The actual properties of the metal in the part will depend on its characteristics and the metal cooling rate in the various sections of the casting. The relation between properties in the casting and the test bar will depend on their relative cooling rates. For example, the properties of gray iron are especially sensitive to the section thickness in which they are cast, so a series of standard test bars of increasing size have been established specifically for gray iron and related to various casting thicknesses. One of these standard test bars, or, if necessary, a larger- or smaller-sized one, should be selected to approximate the cooling rate in the critical sections of the cast component. Malleable iron is not poured into heavy sections, and since all malleable castings are heat treated, a single-sized test bar is used. Ductile iron has some variation in properties with metal section when not heat treated.
However, a single-sized test bar is satisfactory, except for very large castings. In another approach, a commonly used method of specifying iron for castings produced in larger quantities is based on the microstructure and Brinell hardness of the metal in the part itself at a designated location. This method was originally developed for automotive and agricultural implement castings but has proven satisfactory for many other types of iron castings produced in quantity. When the microstructure of the iron is normal, there is a good general relationship between Brinell hardness and tensile properties; however, this relationship is not part of the specification, nor are the tensile properties of the iron stipulated. The suitability of a casting for its intended use as produced in a particular type of iron at a specified hardness level must be established for each application. This is typically done by trial use or by simulated load tests. Some castings purchased under SAE specifications for critical applications are tested for a specified Brinell hardness (Fig. 10). Standard Specifications. Many casting purchases are referenced to standard specifications, primarily to those of ASTM International and SAE. In addition, there are also cast parts purchased without metal specifications other than simply stating the generic type of alloy wanted. This purchasing procedure is satisfactory when the requirements of the casting are quite ordinary and typical for that type of alloy. Some metalcasting facilities will not accept orders that require metal property tests but specialize in providing castings with other specified requirements. Where metal properties are to be specified, the use of a standard specification provides the buyer with a maximum number of sources and possible economic advantages. The use of a familiar specification saves time and reduces the possibility of mistakes. Even when purchasing cast parts with unusual requirements, it is preferable to use an appropriate standard specification as a base and then add additional requirements. Of course, the additional requirements should be clearly stated and given with the reference number of the standard specification on the purchase order.
28 / Applications and Specification Guidelines Quality-Management Systems. A buyer may ask a caster, “How good is your quality, and how do you control it?” The caster has two possible replies: (1) The caster can provide a detailed description of the quality-management system and wait for the buyer to review and qualify the methods and procedures, or (2) the reply can be, “My quality system has been reviewed and certified against a recognized quality-management standard. Here is the certification.” A certified quality-management program has two basic features:
Definition and documentation of the policies
and procedures required that meet the specified quality standards Application and practice of the quality procedures on a day-to-day basis, supported by traceable records There are several quality-management standards and certifications that are widely accepted in the manufacturing community and are applicable to the metalcasting market: ISO 9000, QS 9000, AS 9000, and NQS 9000.
SELECTED REFERENCES Cost-Effective Casting Design: A Designer’s
and Purchaser’s Reference, American Foundry Society, Inc. Engineered Casting Solutions, American Foundry Society, Inc., www.castsolutions. com Principles of Purchasing Castings, American Foundry Society, Inc. Website for casting buyers and designers: www.metalcastingdesign.com
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 31-40 DOI: 10.1361/asmhba0005189
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Chemical Thermodynamics and Kinetics Revised by Seymour Katz
TWO SUBJECTS that underlie all foundry processes are chemical thermodynamics and chemical kinetics. Relationships derived from these two disciplines of physical chemistry are extremely useful in providing a quantitative framework for explaining and predicting chemical behavior. This includes such presolidification chemical processes as preparing liquid metal for casting (going back to the winning of ores, if desired), melt purification from scrap recycling, and alloy additions. Chemical processes are also involved in making molds and cores, the casting and solidification processes, the microstructural development of alloys, and the phase changes that occur after casting (for example, by heat treatment). This article introduces the fundamental concepts of chemical thermodynamics and chemical kinetics in describing presolidification phenomena. Thermodynamics defines the most stable chemical system (equilibrium system) that can be produced, given a set of starting conditions (compositions, temperature, pressure, and so on). An example is the phase diagram. It gives the stable phases that coexist at equilibrium, given the composition, temperature, and pressure. Chemical thermodynamics is only concerned with the initial and final states of a system, while chemical kinetics is concerned with how and at what rate the system is transformed from the initial to the final state. The practical importance of chemical kinetics implies that systems often fail to reach the most stable condition and that the path taken plays a major role in determining the speed of the reaction. In general, kinetics assumes its greatest importance when atomic movement is restricted, such as at low temperatures or in solids. Thermodynamics most often applies when atoms have high mobility, as in gases and liquids and at high temperatures. Another factor that determines the relative applicability of thermodynamics and kinetics is time. The less time available for a process, the more likely that kinetics will govern the chemical state of the system. The thermodynamics and kinetics are fundamental in the composition control of liquid metal. Composition control consists of two areas: alloy addition and melt purification. Both are of vital importance in producing quality
castings. The chemical and physical processes involved in alloy addition do not vary much with base metal; therefore, this subject can be treated generically in this article. On the other hand, melt purification processes tend to be material-specific and vary according to base metal. For example, metal purification may involve processing of scrap, which is a common source material used for iron and aluminum castings. By its nature, scrap is of uncertain composition, both with respect to concentrations of desirable alloy elements and the presence of deleterious elements. Thus, despite all precautions in melting, adjustments in melt composition may be required. Following the introduction to chemical thermodynamics and kinetics in this article, the next two articles provide pertinent thermochemical data for ferrous and nonferrous (aluminum and copper) casting alloys. Casting defects with origins in the liquid state (gas defects and inclusions) are covered in the articles “Gases in Metals” and “Inclusion-Forming Reactions.” Nomenclature of symbols used in physical chemistry is given in Table 1.
mass. Common units of H are joules/mole, calories/gram, or Btu/pound. Thus, the total enthalpy of ni moles of material i is: ðHi Þtotal ¼ ni Hi
(Eq 1)
For a mixture of materials, the total enthalpy is the sum of the individual enthalpies: Htotal ¼
X
ni Hi
(Eq 2)
i
The heat capacity, CP, given in joules/mole K, and so on is a measure of how H varies with temperature and is defined as: CP ¼
@H @T P
(Eq 3)
where the subscript P indicates that the values refer to constant pressure conditions. Enthalpy and heat capacity govern the thermal state of a system and the heat generated or absorbed by chemical reactions. These variables are particularly important for establishing the energy requirements for processes, for determining rates of solidification, and for heat treatment processes. Enthalpy and heat capacity data are tabulated in Ref 1.
Chemical Thermodynamics
Gibbs Free Energy
There are three thermodynamic laws that provide the basis for the relationships that describe chemical systems. These laws hold that perpetual motion and absolute zero temperature can never be achieved. The relationship between the thermodynamic laws and, for example, phase diagrams is not immediately apparent. However, the connection exists and is testimony to a great deal of brilliant theoretical work carried out over the last 100 years. For metallurgical systems, the most important thermodynamic variables are enthalpy and Gibbs free energy. The former governs the disposition of heat, and the latter governs chemical equilibrium.
Conditions for Equilibrium. Chemical equilibrium is a condition in which the chemical components of a system have no tendency to change. In mechanical systems, equilibrium is understood as the condition in which the potential energy is at a minimum. For chemical systems at constant temperature and pressure, an analogous statement for chemical equilibrium is that the Gibbs free energy is at a minimum, or equivalently from calculus:
Enthalpy and Heat Capacity The heat content of a material is termed its enthalpy, H. Because the value of the enthalpy is proportional to the amount of material present, H is usually tabulated as enthalpy per unit
dGðP; T; ni . . .Þ ¼ 0
(Eq 4)
For systems in which chemical change can occur, dG can be expressed in partial differential form:
@G @G dT þ @T P;ni ... @P T;nj @G dP þ dni þ @ni T;P;nj
dG ¼
(Eq 5)
where ni stands for moles of component i. The term ð@G=@ni ÞT;P;nj ; is called the partial
32 / Principles of Liquid Metal Processing Table 1 Physical chemistry nomenclature ai aoi A A0 Ci Cp Cs CE dp D DR eii E fi F* F Gi DG Gi Goi xs G Goi Goi ð%iÞ Goi ðXi Þ h hi Hi Hio H i
Activity of component i relative to pure material standard state Activity of component i in the pure material standard state Interfacial surface area Surface area of an alloy particle Concentration of component i Heat capacity Sulfide capacity Carbon equivalent Particle diameter Diffusion coefficient Desulfurization ratio Interaction coefficient for 1 wt% standard state Cell potential Activity coefficient for 1 wt% standard state Integral free energy of solution Faraday’s constant Gibbs free energy of component i Integral molar free energy Partial molar free energy of component i Molar free energy of component i Excess partial molar free energy of component i Free energy of component i in solution at the 1 wt% standard state Free energy of solution from pure material to 1 wt% standard state Free energy of solution from the pure material to the hypothetical pure material standard state Convective heat transfer coefficient Activity of component i relative to the 1 wt% standard state Molar enthalpy of component i Standard enthalpy of formation Partial molar enthalpy of component i
molar free energy, Gi , or the chemical potential. It is clear from Eq 4 and 5 that, at equilibrium and at constant temperature and pressure, the following relationship exists: 0 ¼ Gi dni þ Gj dnj þ . . .
(Eq 6)
When the change in composition is dictated by a chemical reaction, stoichiometry provides a relationship between dni, dnj, and so on, and Eq 6 can suitably be transformed to: 0 ¼ ni Gi þ nj Gj þ . . .
(Eq 7)
where ni, nj, and so on are the stoichiometric coefficients in the chemical reaction equation. For products and reactants, the stoichiometric coefficients are positive and negative numbers, respectively. As an example, for the chemical reaction: 4Al þ 3O2 ¼ 2Al2 O3
(Eq 8)
Equation 7 becomes: 0 ¼ 2GAl2 O3 4GAl 3GO2
(Eq 9)
Tabulation of Thermodynamic Data. To make thermodynamic calculations convenient, a system was developed for tabulating thermochemical data so that relatively small amounts of data could describe a wide range of chemical conditions. This was done by defining special
xs
H i J k, k0 , kB KA kH km KmN K m,m0 Mi ni Ni pi poi P P S, P M
Q Q r, r0 R S Sc Sr S i xs S i Sio
Excess partial molar enthalpy of component i Mass flux Rate constants Constant related to the absorption of an element on a surface Constant related to hydrogen solubility in iron Mass transfer coefficient Liquid-phase mass transfer coefficient Equilibrium constant Solubility factors giving change in carbon solubility in Fe-C-X from unit mass addition of a third element Molecular weight of component i Moles of component i Mole fraction of component i Pressure of component i in the gas phase Pressure of component i in its standard state Constant pressure conditions Partition coefficient. Ratio alloy element concentration between liquid and solid phases in equilibrium. The superscripts S and M refer to stable and metastable products of Fe-C solidification, respectively. Heat flux Gas flow rate Radius of a rod Universal gas constant Ratio of surface area to volume Ratio carbon concentration: iron alloy melt to eutectic Rectified saturation degree, that is, weight fraction Fe-C eutectic formed on solidification of hypoeutectic iron Partial molar entropy of component i Excess partial molar entropy of component i Molar entropy of component i in the pure material standard state
conditions called standard states. Thus, for a chemical compound i that is an ideal gas at pressure pi,Gi is now written: Gi ¼ Goi þ RT ln
pi poi
(Eq 10)
where Goi is the free energy per mole of i in its standard state and poi is the pressure of i in its standard state. The second term on the right-hand side of Eq 10 is the free energy change to take a mole of i from its standard state poi to pi. For gases, the usual standard state is poi ¼ 1 atm. Thus, poi is usually omitted in writing, Eq 10. Both Gi and Goi are absolute values that cannot be measured. Changes in these variables can be measured from an arbitrary baseline. By convention, the baseline is taken as the elements in their standard states for which ðGoi Þelements ¼ 0 is defined at all temperatures. To indicate that the values are relative to a baseline, Eq 6, 7, and 10 are rewritten by replacing Gi and Goi with Gi and Goi , respectively. Values of Goi for many compounds (called the standard free energy of formation) and phase changes have been tabulated (Ref 1, 2). Because Goi is a function of temperature, the most common tabulations are based on another thermodynamic definition of Goi , that is: Goi ¼ Hio T Sio
(Eq 11)
t t tm T V V0 V0 w W Xi
Time Celsius temperature Melting time Kelvin temperature Gas flow rate Gas volume per tonne of metal Volume of an alloy particle Mass fraction: desulfurization to iron Weight of melt Mole fraction of component i
Greek symbols a Phase designation aij Alpha activity coefficient function for Darken’s quadratic formalism gi Activity coefficient of component i relative to pure material standard state io Henry’s law constant d Boundary layer thickness Interaction coefficient; pure material standard eji state y Fraction of surface sites covered by surface-active atoms k Empirical constant indicating iron compositions that have the same microstructure at a given cooling rate l Particle shape factor ni Stoichiometric coefficient for component i in a reaction r Density Superscripts B e, equil f ‘ s S *
Bulk property Equilibrium condition Final condition Liquidus concentration Solidus concentration Surface condition Interface condition
where Hio and Sio are the standard enthalpy and the entropy of formation, respectively. The tabulations provide Hio and Sio data. The value of Goi is calculated for any temperature using Eq 11. Equilibrium Constant. Substituting Eq 10 into Eq 7 gives the most important relationship governing reaction equilibrium: X
ni Goi ¼ RT
i
X
ln pni i
(Eq 12)
i
Here, the thermodynamic data on the left are related to actual pressures (concentrations) on the right. Equation 12 is ordinarily written in abbreviated form as: Go ¼ RT ln K
(Eq 13)
where K is the equilibrium constant. The form of K is illustrated in the following. Given the gas phase chemical reaction: 2H2 þ O2 ¼ 2H2 O
(Eq 14)
the equilibrium constant is: K14 ¼
p2H2 O p2H2 pO2
(Eq 15)
Activity and Condensed Phase Equilibrium. Metallurgical applications are primarily concerned with condensed phases (solids and liquids), which do not behave in the manner of an ideal gas. That is, for a condensed
Chemical Thermodynamics and Kinetics / 33 material, Goi does not vary with the logarithm of the concentration as may be inferred from Eq 12, where, for an ideal gas, Goi varies directly with ln pi. To preserve the useful equilibrium constant expression for nonideal, condensed phases, Eq 10 and 12 are retained in form but rewritten to enable the use of data expressing nonideality. Thus: ! ai aoj
(Eq 16)
ln ani i
(Eq 17)
Gi ¼ Goi þ RT ln X
ni Goi ¼ RT
i
X i
The difference between Eq 16 and 10 and Eq 17 and 12 is that a term ai, called activity, replaces the pressure term. The activity is a function of concentration but usually not a simple function of concentration. Analogous to the case with gases, the activity in the standard state aoi is taken as unity. The equilibrium constant is now defined in terms of activity: ln K ¼
X
ln ani i
(Eq 18)
i
Activity Coefficients—Departure from Ideality. Common expressions for activity are: ai ¼ gi Xi
(Eq 19)
hi ¼ fi ð%iÞ
(Eq 20)
where the concentration terms Xi and (%i) are the mole fraction and weight percent, respectively, of the solute i, and gi and fi are the activity coefficients that measure departure from ideality, that is, generally gi and fi 6¼ 1. The terms ai and hi are often called the Raoultian activity and the Henrian activity, respectively, to distinguish the standard state being used. The Raoultian activity is usually used in conjunction with the pure material standard state, while the Henrian activity is used with the 1 wt% standard state (to be explained). The latter standard state and Eq 20 are used to describe dilute solutions and have been extensively applied to ferrous systems, especially steel. 1.0 A
T = 1873 K
Activity, ai
0.6
aNi aFe
0.4
γ⬚
Ni
0.2 B 0 0 (Fe)
Activity coefficient, γi
Raoult’s law
0.8
Henry’s law 0.2
0.4
0.6
0.8
1.0 (Ni)
Mole fraction, XNi
Fig. 1
Activities of iron and nickel in their binary solutions at 1873 K. Source: Ref 1, 3
The pure material standard state and Eq 20 are used in most other cases. The two standard states, their activities, and their activity coefficients are illustrated with Fig. 1, which gives the activity composition relationships for the liquid iron-nickel system. The point “A” represents the pure material standard state for nickel (pure nickel in its most stable form—liquid—at the temperature of concern, 1873 K). Activity, given on the ordinate, is based on Eq 19. The ideal solution line (Raoult’s law line), where aNi = XNi, is labeled. The change in the partial molar free energy of nickel in transferring nickel to solution is given by Eq 21, which was obtained by substituting Eq 19 into Eq 16: GNi ¼ GoNi þ RT ln XNi þ RT ln gNi
(Eq 21)
In this case, Goi ¼ 0 because the standard state and the baseline conditions are identical. The term RT ln XNi is the change in GNi ideal for an ideal solution GNi . The term RT ln gNi provides the correction for the real system. As seen, gNi < 1. Thus, GNi is more negative ideal than GNi . Larger negative free energy values indicate greater stability. Thus, real solutions of nickel in iron are more stable than ideal solutions of nickel in iron. A common characteristic of the activity of a component in solution is the relatively linear portion of the curve as Xi ! 0. This limiting slope, designated goi (Fig. 1), is known as the Henry’s law constant. The 1 wt% solution standard state is marked point “B” in Fig. 1. It falls within or near the range in which the activity coefficient is a constant ðgoi Þ. By defining the standard state at point “B”, fi = 1 in the linear region and conveniently, hi = (%i). The expression that corresponds to Eq 21 for the 1 wt% standard state is: ^o þ RT lnð%NiÞ GNi ¼G Ni þ RT lnðfNi Þ
(Eq 22)
^o to indiIn Eq 22, GoNi is designated G Ni cate that it refers to the 1 wt% standard state. Clearly, for any iron-nickel composition, GNi is the same whether it is calculated by Eq 21 or 22. Therefore, the choice is one of ^o 6¼ 0. It is the convenience. In this case, G i change in free energy in taking pure liquid nickel to point “B” on the Henry’s law line. The difference in Goi between standard states A and B is: ^o ¼ Go þ RT ln aB G i i aA Ms loi ¼ Goi þ RT ln Ms þ 99Mi
(Eq 23)
where the activities for composition A and B are Raoultian, and Ms and Mi are the molecular weights, respectively, of solvent and solute. The equilibrium constant expressions (Eq 17) for the pure material and 1 wt% standard states are: X i
ni Goi ¼ RT
X
lnðgi Xi Þni
(Eq 24)
X
^o ¼ RT ni G i
X
ln½fl ð%iÞni
(Eq 25)
i
o
Very often Gi for components in the equilibrium constant are expressed in different standard states. There is no problem with this as long as appropriate values Goi or ^o are used. G i Tabulations of Activity Coefficient Data. Tabulations of Goi data were already discussed. From Eq 23 to 25, it is seen that data for goi , gi, and fi are also needed. Tabulations of goi and ^o for elements in solution in liquid iron, G i cobalt, and copper are given in Ref 2. Similar data for aluminum are included in the article “Thermodynamic Properties of AluminumBase and Copper-Base Alloys” in this Section. Data for gi and fi for iron, copper, and aluminum multicomponent alloys are available in Ref 4 to 6. These data are based on a formalism (Ref 7) that is discussed in the article “Thermodynamic Properties of Aluminum-Base and Copper-Base Alloys” in this Section. Another formalism (Ref 8, 9), found in the article “Thermodynamic Properties of Iron-Base Alloys” in this Section, is used to describe the activities of carbon and silicon in multicomponent ferrous alloys. Most of the tabulated values were obtained from studies in dilute solution (Henry’s law region). From Fig. 1, it can be seen that the shapes of the activity curves are complex. Therefore, data for the Henry’s law region have only limited applicability outside the region. To describe Gi over the entire composition range, tables giving Gi as a function of composition are used. These data are usually available only for binary alloy systems. Tables for important aluminum and copper binary systems are given in the article “Thermodynamic Properties of Aluminum-Base and CopperBase Alloys” in this Section. Comparable data for iron are available in Ref 10. Because Gi is temperature dependent, the data are commonly given as H i and S i , where Gi is obtained by: Gi ¼ H i T S i
(Eq 26)
Examples of the use of Eq 1 to 26 are found in other articles in this Section.
Phase Diagrams Phase diagrams describe the phases that coexist at equilibrium. Because chemical equilibrium is involved, phase diagrams have a thermodynamic basis. The complexity and variety of relationships demonstrated by phase diagrams place a detailed description beyond the scope of this article. A qualitative demonstration of the interrelationship between phase diagrams and thermodynamics is given. For a multiphase system, a condition of equilibrium is that the partial molar free energy or chemical potential of each component must be the same in all phases (a, b, . . .):
34 / Principles of Liquid Metal Processing a
b
Gi ¼ Gi
(Eq 27)
Thus, for example, the chemical potential of carbon in carbon-saturated iron is the same as graphite, despite there being less than 5 wt% C in saturated iron. A similar situation exists with respect to silicon/aluminum-silicon alloy equilibrium, where the chemical potential of pure 0
silicon is matched by less than 15 wt% Si in solution. Phase diagrams describe stability relationships for compounds and solutions (solid and liquid). To understand the relationships based on thermodynamics, it is necessary to compare the phase diagram (temperature-composition plot) with the equivalent free energy-composition plot. To construct the free energy-composition plot, it must be recognized that for a binary solution the integral molar free energy is:
Free energy (F *), cal
G ¼ X1 G1 þ X2 G2 −500
−1000
F ¼ G ðX1 G1 þ X2 G2 Þ
(ΔG1 – ΔG1⬚)
−1500
−2000
(Eq 28)
Substituting Eq 16 into Eq 28 and assuming ideal solutions (ai = Xi) yields F*, the free energy change in forming 1 mole of an ideal, binary solution from components in the pure standard state (taken as the pure liquid in this case): ¼ RT ðX1 ln X1 þ X2 ln X2 Þ (ΔG2 – ΔG2⬚) 0
0.75
0.50
0.25
1.0
Mole fraction (X2)
Fig. 2
Free energy of formation of an ideal solution from liquid components in the pure material standard state. Also illustrated is the graphic method of tangents to determine the partial molar free energies of solution.
(Eq 29)
A plot of F* as a function of X2 is given in Fig. 2. Clearly, F* < 0 with a minimum at X2 = 0.5. For 6 1), the curve can be higher nonideal solutions (gi ¼ or lower than the ideal line, and it can be asymmetrical. It can be shown that the intercepts of a tangent line drawn at any point on the curve will give ðG1 G1 Þ and ðG2 G2 Þ (Ref 1). These are the partial molar free energies of solution from the liquid standard state.
0
For metallic solutions, Gi ¼ 0 ; therefore, GI is obtained directly from this plot. The free energy of a phase of definite composition is represented as a point on the free energy-composition diagram. Clearly, if such a phase were in equilibrium with a solution, the equilibrium solution and the compound would have to lie along the same tangent line in order for Eq 27 to hold. This condition is illustrated in Fig. 3(a). Figure 3(b) shows the phase diagram that results from the thermodynamic constraints. In Fig. 3(a), F* < 0 for pure solid phase 1 because the liquid was chosen as the standard state, and at the temperature chosen, the liquid is less stable than the solid. The equilibrium between a liquid and a primary solid solution is illustrated in Fig. 4. From Fig. 4(a), the respective equilibrium concentrations lie along a common tangent, thus satisfying the equilibrium conditions expressed by Eq 27. The phase diagram that results from these thermodynamic conditions is given in Fig. 4(b). For small values of X2, either solution can theoretically exist. The solid phase is expected to be present, however, because it is more stable (lower F*). If, by virtue of such factors as rapid cooling, the liquid phase is retained in a region where the solid is more stable, a metastable condition would exist. This is not uncommon in systems of metallurgical interest. Common examples are cementite (Fe3C), which is metastable with respect to iron and graphite, and the glass (supercooled liquid) condition of slags, which is metastable with respect to crystalline compounds.
Solid Solution
B
Composition of solution in equilibrium with pure solid
Solid + solution
α solid solution
Free energy (F *), cal
Free energy (F *), cal
0 A
Solution
Liquid solution
(a)
Temperature (T )
(a)
Temperature (T )
This line corresponds to temperature of free energy chart above
T1 L
α+L
β+L
α
0 0
(b)
Fig. 3
1.0 Mole fraction (X2) Relationship between (a) free energy-composition and (b) temperature-composition diagrams for the equilibrium between a solution and a pure solid component. Source: Ref 11
β
α+β
(b)
Fig. 4
1 Mole fraction (X2)
Relationship between (a) free energycomposition and (b) temperaturecomposition diagrams for the equilibrium between a liquid and a primary solid solution. Source: Ref 12
Chemical Thermodynamics and Kinetics / 35
Chemical Kinetics The rates of chemical processes depend so much on local conditions that it is difficult to describe the processes in the universal terms used in thermodynamics. The rates of chemical change in solid cast metals are usually governed by solid diffusion. This subject is examined in the Section “Principles of Solidification” in this Volume. For processes involving fluids at high temperatures, the rates of chemical reactions are inherently rapid and usually not rate limiting. The processes that generally limit the rate are nucleation of the product phase and interphase mass transport (moving material between immiscible phases).
Nucleation The problem of nucleation is universal to the kinetics of solidification, and the subject is treated more extensively in the article “Nucleation Kinetics” in this Volume. However, a few general comments are in order. Nucleation of a phase in an inclusion-free metal requires supersaturation with respect to the material being precipitated. For this condition, the chemical potential of the material in the melt exceeds that in the precipitate, which is the requirement for spontaneous reaction. Initially, the diameter of the precipitated phase is in the atomic range. Thermodynamics dictates that the activity of particles increases with decreasing radius of curvature. Therefore, the activity of atomic-sized particles is very high. This in turn requires very high melt supersaturation to meet the requirement for spontaneous nucleation. This barrier to the generation of nuclei from a homogeneous melt, imposed by thermodynamics, makes it unlikely that appreciable nucleation occurs by this mechanism. Instead, nucleation is initiated on particles purposely introduced to the melt before solidification. For example, TiB2 particles are in use to nucleate aluminum melts (Ref 13), and complex oxysulfide particles nucleate graphite in gray and ductile iron (Ref 14, 15).
Interphase Mass Transport Because high-temperature reactions are rapid, the rate-limiting step in many cases of heterogeneous reaction is the diffusion of reactants to and from the reaction interface. This is particularly the case when the reactants are dilute. Because the reaction rate in these cases is rapid relative to rates of diffusion, the concentration of a reacting species at the reaction interface is either above or below the bulk concentration. The condition that exists depends on whether the activity of the reactive species in the second phase is higher or lower than the activity in the bulk. The model, frequently used in quantifying mass transfer limited kinetics, assumes that there is a thin stagnant film at the interface
across which a concentration gradient exists. The model is schematically represented in Fig. 5. Transport across the boundary layer is assumed to occur only by diffusion. Applying Fick’s first law of diffusion, the flux of species i across the boundary layer is: J¼
D C CiB ¼ km Ci CiB d i
(Eq 30)
where D is the diffusion coefficient, d is the boundary layer thickness, and Ci and CiB are concentrations of the diffusion species, respectively, at the interface and in the bulk. The term D/d is seldom used; instead, the mass transfer coefficient km is used to represent D/d. When phase 2 is a fluid, d decreases with increasing fluid velocity along the interface (that is, increased convection). Thus, km is a function of the fluid dynamic condition. To determine the rate of reaction, that is, J, it is necessary to know Ci . This is difficult to determine experimentally. To circumvent the problem, one of two assumptions is commonly made. For extremely fast reactions in which the equilibrium constant is very large, Ci can be assumed to be 0. Thus, Eq 30 reduces to: J ¼ km CiB
J ¼ km Ciequil CiB
(Eq 32)
The value of Ciequil is determined from the equilibrium constant expression, knowing the activities of all the reactants and products except Ciequil . A number of examples of this treatment are found in succeeding articles in this Section. There are cases in which the reaction rate (rate constant k) at the interface, unimpeded by diffusion, is of the same order as the mass transfer limited rate. In this case, the observed rate and the rate constant k0 are affected by both processes. This relationship is illustrated as follows. Because diffusion of i through the boundary layer and the reaction of i at the interface occur sequentially, the rates of both processes must be the same at steady state. Thus: J ¼ km Ci CiB ¼ k Ci Ciequil
1 ¼ k0
1 1 þ km k
δ
Stagnant boundary layer
Fig. 5
Boundary layer model used for mass transfer limited kinetics
It is clear from Eq 34 and 35 that for the condition k km, Eq 34 reduces to Eq 32. Applications in which mixed control is important are nitrogen absorption in iron alloy melts (Ref 17) and coke gasification at intermediate temperature (Ref 16).
Kinetics of Alloy Additions As previously noted, the chemical and physical processes involved in alloy addition do not vary much with base metal and can be treated generically. There are two fundamental mechanisms governing the kinetics of alloy additions, depending on whether the alloy is liquid or solid at the temperature of the metal bath. These processes are schematically illustrated in Fig. 6.
Dissolution of Alloy with Melting Point Lower than Bath Temperature
(Eq 33)
where the last term gives the first-order rate of reaction occurring at the interface. Eliminating Ci in the last two terms in Eq 33 yields: J ¼ k0 Ciequil CiB
C Bi
Phase 1
(Eq 31)
This assumption has been used to describe the high-temperature gasification of coke (Ref 16). The second assumption is that equilibrium between the two phases is established at the reaction interface. In this case, Eq 30 is written as:
where:
Phase 2 C *i
(Eq 34)
(Eq 35)
For the case where the alloy addition is liquid at the melt temperature (Fig. 6, top row sequence), addition of the cold alloy at time t0 is quickly followed by formation of a layer of frozen base metal on the alloy surface (t1) due to heat extraction by the alloy. At longer times (t2), the shell thickness increases, while melting occurs at the alloy surface. As the interior of the alloy heats up, the thermal demand decreases. As a result, the frozen base metal shell begins to redissolve while the thickness of the liquid
36 / Principles of Liquid Metal Processing Superheat, ⬚F
Time t1
t0
t2
t3
t4
tm
840
Alloy liquid at bath temperature
32
48
64
80
96
112
60
70
720
Melting time, s
600
Alloy solid at bath temperature
480
360 C 240
Solid alloy Liquid alloy
B 120
Solid base metal
Fig. 6
A
Two kinetic paths for melting and/or dissolving alloy additions in liquid metal baths. Source: Ref 18
0
20
30
40
50
Superheat, ⬚C
Table 2 Data for melting time calculations for a 100 mm (4 in.) diameter sphere of silicomanganese in a steel melt Condition
Fig.
DH, J/g
r0, g/cm3
V0/A0, cm
h, J/g K
TL , K
ðTBL TL Þ, K
t m, s
387 307
5.6 5.6
1.67 1.67
0.15 0.15
1803 1395
20 418
1200 45
Frozen shell No frozen shell Source: Ref 18
7
Predicted melting/dissolution times for silicomanganese spheres in quiescent steel baths versus steel superheat temperature. Particle diameters of the spheres are as follows: A, 25 mm (1 in.); B, 50 mm (2 in.); C, 100 mm (4 in.). Source: Ref 18
requirement for melting, divided by the rate at which heat is supplied from the bath (Ref 18), that is:
Table 3 Data for melting time calculations for a 10 mm (0.4 in.) diameter aluminum sphere in different liquid metal melts Base metal
DH, J/g
r0, g/cm3
V0/A0, cm
h, J/g K
TL , K
ðTBL TL Þ, K
tm, s
Steel Cast iron Copper Aluminum
576 576 576 576
2.7 2.7 2.7 2.7
0.167 0.167 0.167 0.167
0.15 0.15 0.15 0.15
1823 1723 1323 1023
890 790 290 90
1.9 2.2 6.0 19.2
annulus increases (t3). In the final stage, beginning at t4, the redissolution of the base metal shell is complete, and as a result, the liquid alloy held in the annular space is released to the melt. Any alloy that remains solid is exposed directly to the liquid metal bath and is progressively incorporated into the melt by the surface melting and convection. Dissolution is completed at tm (melting time). The length of the incubation period (t0 t4) and the length of the alloy dissolution period (t4 tm) are governed by heat transfer. Conditions favoring rapid alloy dissolution are high superheat temperatures, small alloy particle size (large surface for heat transfer), and low alloy thermal conductivity. The heat flux to the alloy particle is given by: Q ¼ h TLB TL
(Eq 36) 2
1
where Q is the heat flux (J m s ), h is the convective heat transfer coefficient (J m2 s1 K1), TLB is the temperature of
the bulk liquid (in degrees Kelvin), TL is the melting point temperature of the alloy at the solid/liquid interface (in degrees Kelvin), and TLB TL is the superheat temperature (in degrees Kelvin). During the incubation period, the superheat temperature is less than that in the alloy dissolution period because TL corresponds to the melting point of the base metal and alloy in the former and latter cases, respectively. For steel and aluminum melts, superheat temperatures in the incubation period are generally smaller (20 to 100 K) than for cast iron (200 to 300 K). In the alloy dissolution period, the superheat temperature for a given alloy addition increases with increasing melting point of the base metal. Thus, superheat temperatures, and therefore dissolution rates, increase in the following order: aluminum < copper < cast iron < steel. Melting times, tm, can be estimated by assuming that tm is equal to the total heat
tm ¼
Hr0 V0 QA0
(Eq 37)
where DH is the heat supplied to 1 g of alloy at the time it dissolves, and r0, V0, and A0 are the density, volume, and surface area of the alloy particle. The variable Q is defined by Eq 36. Equation 37 is most accurate when applied at the extremes of possible superheat temperatures. For small superheat temperatures, it can be assumed that by the time the frozen shell redissolves, the alloy is a liquid at the melting point of the base metal. At high superheat temperatures, it can be assumed that no frozen layer forms and that the alloy liquefies at its melting point. These two cases have been examined for the dissolution of a 100 mm (4 in.) particle diameter B(dp) sphere of silicomanganese in a steel melt TL ¼ 1823 K . Data for the calculations are given in Table 2. The very long melting time of 1200 s obtained by restricting the calculations to a small superheat temperature (20 K) is in good agreement with the 900 s obtained with a more sophisticated model (Fig. 7). The difference in tm for the two extreme cases considered in Table 3 underscores the importance of superheat temperature in alloy dissolution processes. This is clearly illustrated in Fig. 7, in which the role of particle diameter is also demonstrated. It is apparent from Fig. 7 that the large value of tm in Table 2 for the case
Chemical Thermodynamics and Kinetics / 37 a common foundry practice that is necessary for control of cast iron composition. The rate at which carbon dissolves in cast iron is controlled by the rate of diffusion of carbon from carbon-saturated iron at the carbon/iron interface into the bulk liquid (Ref 21–23). Thus, Eq 38 can be written:
Density ratio 40
0
0.2
0.4
0.6
0.8
1.0
Liquid iron-carbon alloy
Carbon
C sat * Carbon concentration
20
10
Total immersion time, s
6 4
d 2 Distance C
1
Fig. 9
0.6
B
0.4
A
0.2
0.1 0
1
2
3
4
5
6
7
Addition density, g • cm-3
Fig. 8
Total immersion times versus density ratio for alloys entering a stagnant steel bath at 7.67 m s1 (25.2 ft s1) (drop height: 3.0 m, or 10 ft). Particle diameters are as follows: A, 10 mm (0.4 in.); B, 50 mm (2 in.); C, 250 mm (10 in.). Source: Ref 18
J ¼ km ðCsat CLB Þ
C BL
TLB TL ¼ 20 and dp = 100 mm can be substantially B reduced with moderate changes in TL TL and dp. The range of melting times obtained with the addition of an alloy to different base metals is examined subsequently. Consider the addition of pure aluminum to melts of steel, cast iron, copper, and aluminum under conditions where no frozen shell is formed. In this case, TL is the same (933 K), but TLB varies. Pertinent data for the calculation and the calculated melting times are given in Table 3. It is clear from Table 3 that the duration of alloy melting becomes a more serious problem as the melting point of the base metal decreases. The short melting times predicted for steel and cast iron melts in Table 3 highlight another important consideration in alloy addition, namely, immersion time due to alloy buoyancy. Because iron is denser than common alloy additions such as carbon, silicon, and aluminum, buoyant forces prevent these alloys from being easily submerged for times approaching tm. This is illustrated in Fig. 8, which plots immersion times versus the density ratio, ralloy/ rbase metal. For 10 mm (0.4 in.) aluminum particles (density ratio: 0.39) that are dropped
Carbon concentration profiles for a carbon rod dissolving in an iron-carbon melt
a distance of 3 m (10 ft) to the surface of an iron bath, the immersion time is less than 1 s, while the melting time (from Table 3) is 1.9 s. Immersion times can be increased by entraining particles in strong recirculating flows or by wire and rod feeding (Ref 19, 20). These conditions have higher associated heat transfer coefficients, which further facilitate dissolution.
Dissolution of Alloy That Is Solid at Bath Temperatures The processes involved in the dissolution of this type of addition are schematically illustrated in the bottom row sequence in Fig. 6. As in the previous case, heat transfer controls the duration of the incubation period (t0 t4). However, in this case, mass transfer is the controlling factor in the alloy dissolution period (t4 tm). In general, mass transfer controlled dissolution is slower than heat transfer controlled dissolution, and the number of variables controlling the rate is greater. For mass transfer control, melt temperature and fluid dynamic considerations are important, as in the heat transfer controlled case. However, melt composition and diffusion variables are also important. The flux of alloy element to the melt is given by: J ¼ km ðCL CLB Þ
(Eq 38)
where J is the mass flux of alloy element (g cm2 s1), km is the mass transfer coefficient (cm s1), CLB is the concentration of the alloy element in the bulk melt (g cm3), and CL is the concentration of the alloy element in the melt at the solid/liquid interface (g cm3). To illustrate mass transfer limited dissolution processes, two industrially important examples are examined as follows. Carbon Dissolution in Cast Iron. Carbon is the most important alloy element in cast iron. The addition of carbon to iron-carbon melts is
(Eq 39)
where Csat indicates carbon-saturated conditions at the carbon/melt interface. The conditions in the melt are qualitatively described in Fig. 9. Referring to Fig. 9 and Eq 39, the flux of carbon, and therefore the dissolution rate, increases with decreasing bulk carbon concentration CLB , increasing carbon concentration at saturation Csat (increasing bath temperature), and decreasing boundary layer thickness, d. The boundary layer thickness is a function of the fluid velocity, and it affects the value of km. CLB ) on The effect of the difference (Csat the dissolution rate of carbon is illustrated in Fig. 10 with data for the dissolution of graphite granules in induction-stirred cast iron melts. Theoretical curves are drawn for different rate constants k and particle shape factors l, where l is defined as:
¼
Sparticle Ssphere
(Eq 40)
where Sparticle is the surface-to-volume ratio of the graphite particles, and Ssphere is the equivalent surface-to-volume ratio if the particles are spherical. The role of boundary layer thickness (fluid dynamics) on dissolution kinetics is illustrated in Fig. 11, which indicates the variation of the rate constant (dissolution rate) with the peripheral velocity (rotational speed) of a cylindrical graphite rod. The line drawn through the data has a slope of 0.7, which is expected from earlier correlations (Ref 24). Dissolution of Steel in Iron-Carbon Melts. Another well-established case of mass transfer controlled dissolution is the dissolution of steel in iron-carbon melts (Ref 25–28). This process is critical in the electric melting of steel and cast iron. For this case, the rate of steel dissolution increases with melt carbon concentration (Fig. 12), melt temperature (Ref 27), and bath velocity (Ref 25, 26, 28) (affecting the boundary layer thickness). The data can be accounted for on the basis that carbon diffuses from the bulk of the melt to the steel/liquid iron interface, where steel is isothermally converted to a liquid with a liquidus composition, CL‘ . The conditions existing in the bath are illustrated in Fig. 13. Figure 13(a) shows the carbon concentration profiles in the liquid bath in the vicinity of a dissolving steel bar. The meaning of the carbon concentration notation is indicated in Fig. 13(b). As indicated in Fig. 13(a), carbon diffuses across the liquid boundary layer, while equilibrium conditions are maintained at the solid/liquid interface. Thus, the carbon concentrations in the solid and the liquid
38 / Principles of Liquid Metal Processing 1.0
1000 15
800
λ = 2.0
10
600 λ = 2.5
0.8
Time to dissolve, s
200
3 λ = 1.4 2
100
λ = 1.6
80 60
k = 0.007 λ = 1.6
1
k = 0.007
Time to dissolve, min
5
0.6
0.4
0.2
k = 0.02
40
Fractional thickness, r /r0
400
B A
k = 0.02 20
0 k = 0.03
0
10
20
30
40
Immersion time, min 10 0
0.5
1.0
2.0
1.5
2.5
3.0
Fig. 12
Melting curves for 75 mm (3.0 in.) diameter steel rods immersed in iron-carbon baths. A, high (4.6% C) iron-carbon bath; B, low (2.1% C) ironcarbon bath. Plotted is the fractional thickness (the fraction of the rod diameter remaining undissolved) versus immersion time. Source: Ref 27
3.5
2 Weight fraction of solution (C*sat – CB L ) ⫻ 10
Fig. 10
Time of dissolution of carbon particles (10 + 14 mesh) versus (Csat CLB ). Source: Ref 23
J ¼ km ðCLB CL‘ Þ
0.1
(Eq 41)
The carbon flux can also be defined in terms of a carbon balance: B ‘ r C dr J ¼ S ‘L CSB dt rL
Mass transfer coefficient (km), cm • s−1
0.05 0.7 k = (const) V i
where rBS and r‘ L are the respective densities of solid steel and liquid iron at the liquidus composition, CSB is the bulk carbon concentration of solid steel, and dr/dt is the rate of change of the radius or thickness of the steel with time. By combining Eq 41 and 42 and integrating between the limits r = r0 at t = 0 and r = 0 at t = tm, the isothermal melting time tm is given as:
0.02
0.01
(C *sat – C BL )
tm ¼
0.8–2.2 wt % 2.2–3.7 wt % 3.7–5.0 wt %
0.005
0.002 2
5
10
20
50
100
200
400
Peripheral velocity (Vi), cm • S−1
Fig. 11
Experimental mass transfer coefficient versus peripheral velocity for the dissolution of a rotating carbon rod in an iron-carbon melt. Source: Ref 22
at the interface are the isothermal solidus Css and the liquidus CL‘ concentrations, respectively. Dissolution takes place by isothermal conversion of the steel into a liquid, that is, by a chemical melting process. It is clear from
(Eq 42)
Fig. 13(b) that as temperature increases, CL‘ decreases significantly. Thus, the driving force for diffusion increases with temperature. Based on Eq 38 and Fig. 9, the flux of carbon entering the steel is:
B ðrBS CL‘ =r‘ L Þ CS r0 ‘ B km ðCL CL Þ
(Eq 43)
In laboratory studies, Eq 43 accurately predicted steel dissolution rates in a stagnant bath (Ref 27). Research has suggested that melting times are reduced by approximately a factor of 2 in induction-stirred melts (Ref 26). Other work has indicated that tm can be reduced by an order of magnitude with strong convection (Ref 28). On the basis of Eq 43 and the mass transfer correlation for rapidly dissolving rods (Ref 29), correlations were developed relating tm to bath temperature and to carbon concentration (Ref 27). This is given in Fig. 14 for a steel thickness of 2.5 mm (0.098 in.). Because tm is directly proportional to the thickness of the steel, Fig. 14 can be used to estimate tm for a wide range of steel thickness by making the appropriate thickness correction. Checks of the
Chemical Thermodynamics and Kinetics / 39
Steel
Liquid ironcarbon alloy
4. 5.
C BL
Carbon concentration
6.
7.
C *Ll
8. 9.
C *Ss δ
10. C BS
11. Distance (a)
12. 13.
1600 δ+L
2600
C BL
15. 2400
1300 γ+L
1200
2200
1100
Temperature, ⬚F
C *Ll
C *Ss
C BS
1400 Temperature, ⬚C
14.
2800
Liquid alloy
δ
1500
2000 γ + Fe3C
1000
16. 17.
1800
900 0
0.5
Fig. 13
1
2
3
4
5
18.
Carbon, wt%
(b)
(a) Carbon concentration profiles for a steel rod dissolving in an iron-carbon melt. (b) Iron-carbon phase diagram defining the carbon concentrations noted in (a). Source: Ref 27
predictions from Fig. 14 in production operations have indicated that predicted tm values are realistic (Ref 27). ACKNOWLEDGMENT Revised from the articles “Principles of Physical Chemistry” and “Composition Control” by Seymour Katz in Casting, Vol 15, Metals Handbook, 9th ed., ASM International, 1988, p 50–55 and 71–81
REFERENCES 1. Y.K. Rao, Stoichiometry and Thermodynamics of Metallurgical Processes, Cambridge University Press, 1985, p 285–287, 298, 383–394, 880–891 2. E.T. Turkdogan, Physical Chemistry of High Temperature Technology, Academic Press, 1980, p 5–26, 81, 82 3. G.R. Belton and R.J. Fruehan, The Determination of Activities by Mass Spectrometry. I. The Liquid Metallic Systems Iron-
19. 20.
21.
22.
Nickel and Iron-Cobalt, J. Phys. Chem., Vol 71, 1967, p 1403–1409 G.K. Sigworth and J.F. Elliott, The Thermodynamics of Liquid Dilute Iron Alloys, Met. Sci., Vol 8, 1974, p 298–310 G.K. Sigworth and J.F. Elliott, The Thermodynamics of Dilute Liquid Copper Alloys, Can. Metall. Q., Vol 13, 1974, p 455–461 G.K. Sigworth and T.A. Engh, Refining of Liquid Aluminum—A Review of Important Chemical Factors, Scand. J. Metall., Vol 11, 1982, p 143–149 C. Wagner, Thermodynamics of Alloys, Addison-Wesley, 1962, p 51 L.S. Darken, Thermodynamics of Binary Metallic Solutions, Trans. Metall. Soc. AIME, Vol 239, 1967, p 80–89 L.S. Darken, Thermodynamics of Ternary Metallic Solutions, Trans. Metall. Soc. AIME, Vol 239, 1967, p 90 R. Hultgren et al., Selected Values of the Thermodynamic Properties of Binary Alloys, American Society for Metals, 1973 L.S. Darken, Thermodynamics and Physical Metallurgy, American Society for Metals, 1950 L.S. Darken and R.W. Gurry, Physical Chemistry of Metals, McGraw-Hill, 1953, p 235 Aluminum Casting Technology, American Foundrymen’s Society, 1986, p 21 M.J. Lalich and J.R. Hitchings, Characterization of Inclusions as Nuclei for Spheroidal Graphite in Ductile Cast Irons, Trans. AFS, Vol 84, 1976, p 653–664 B. Francis, Heterogeneous Nuclei and Graphite Chemistry in Flake and Nodular Cast Irons, Metall. Trans. A, Vol 10, 1979, p 21–31 S. Katz, The Properties of Coke Affecting Cupola Performance, Trans. AFS, Vol 90, 1982, p 825–833 R.J. Fruehan, B. Lally, and P.C. Glaws, A Model for Nitrogen Absorption in Iron Alloy Melts, Fifth International Iron and Steel Congress Process Technology Proceedings, Vol 6, The Iron and Steel Society of AIME, 1986, p 339–346 R.I.L. Guthrie, Addition Kinetics in Steelmaking, Proceedings of the 35th Electric Furnace Conference, Iron and Steel Society of AIME, 1978, p 30–41 F.A. Mucciardi and R.I.L. Guthrie, Aluminum Wire Feeding in Steelmaking, Trans. ISS, Vol 3, 1983, p 53–59 J.W. Robison, Jr., “Ladle Treatment with Steel-Clad Metallic Calcium Wire,” Paper 35, presented at Scaninject III, Part II, MEFOS, Lulea, Sweden, 1983 L. Kalvelage, J. Markert, and J. Po¨tschke, Measurement of the Dissolution of Graphite in Liquid Iron by Following the Buoyancy, Arch. Eisenhu¨ttenwes., Vol 50, 1979, p 107–110 R.G. Olsson, V. Koump, and T.F. Perzak, Rate of Dissolution of Carbon in Molten
40 / Principles of Liquid Metal Processing Bath temperature, ⬚F 2400
2300 3.5
2500
3.0
2600
2700
2% C bath 4% C bath
1.5% C bath
3% C bath
Melting time, min
2.5
2.0
1.5
1.0
5% C bath 0.5
0 1260
1315
1370
1425
1480
Bath temperature, ⬚C
Fig. 14
Predicted melting times for vertical steel plates and cylinders, 2.5 mm (0.098 in.) scrap thickness and 0.1% C composition, [(%S)i = 0.03] immersed in stagnant iron-carbon baths at different bath temperatures. Source: Ref 27
23. 24.
25.
26.
Fe-C Alloys, Trans. Met. Soc. AIME, Vol 236, 1966, p 426–429 O. Angeles, G.H. Geiger, and C.R. Loper, Jr., Factors Influencing Carbon Pickup in Cast Iron, Trans. AFS, Vol 74, 1968, p 3–11 M. Eisenberg, C.W. Tobias, and C.R. Wilke, Mass Transfer at Rotation Cylinders, Chem. Eng. Prog. Symp. Series, Vol 51, 1955, p 1–16 R.G. Olsson, V. Koump, and T.F. Perzak, Rate of Dissolution of Carbon Steel in Molten Iron-Carbon Alloys, Trans. Met. Soc. AIME, Vol 233, 1965, p 1654–1657 R.D. Pehlke, P.D. Goodell, and R.W. Dunlap, Kinetics of Steel Dissolution in Molten
Pig Iron, Trans. Met. Soc. AIME, Vol 233, 1965, p 1420–1427 27. R.I.L. Guthrie and P. Stubbs, Kinetics of Scrap Melting in Baths of Molten Pig Iron, Can. Metall. Q., Vol 12, 1973, p 465–473 28. K. Mori and T. Sakuraya, Rate of Dissolution of Solid Iron in Carbon-Saturated Liquid Iron Alloys with Evolution of CO, J. Iron Steel Inst. Jpn., Vol 22, 1982, p 964–990 29. P.T.L. Brian and H.B. Hales, Effects of Transpiration and Changing Diameter on Heat and Mass Transfer to Spheres, AICHE J., Vol 15, 1969, p 419–425
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 41-55 DOI: 10.1361/asmhba0005190
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Thermodynamic Properties of Iron-Base Alloys Doru M. Stefanescu, Ohio State University Seymour Katz, S. Katz Associates
THE STRUCTURE AND PROPERTIES of cast metals are sensitive to numerous impurities. Because purification of melts generally adds considerable cost to castings, the lowest cost and surest defense against contamination is careful selection of scrap. Purification is generally reserved for elements that are so pervasive that avoidance is impossible. This is exemplified by sulfur and oxygen removal from cast iron and steel, respectively, and removal of alkali and alkaline earth elements from aluminum. Because of their immediate importance, aspects of the physical chemistry of these processes are reviewed. The structures and properties of steel and cast iron also depend on the control of the solidification structure in castings. To accomplish this, it is necessary to understand and control the thermodynamics of the liquid and solid phases and the kinetics of solidification (nucleation and growth of various phases). The information linked to the practical interests that thermodynamics can provide when considering iron-base alloys encompasses a rather wide range. Only two issues are addressed in sections on thermodynamics in this article. The first is the calculation of solubility lines, relevant to the construction of phase diagrams, and the second is the calculation of the activity of various components, which allows for the determination of probability of formation and relative stability of various phases. Alloying elements are then discussed in terms of their influence on the activity of carbon, which provides information on the stability of the main carbonrich phases of iron-carbon alloys, that is, graphite and cementite. Information on the iron-oxygen system can be found in the article “InclusionForming Reactions” in this Volume.
the level of inclusions, leading to stronger and more fatigue-resistant steels. For cast iron, desulfurization is practiced in the manufacture of ductile iron castings in order to develop spherical graphite morphology. Ductile iron is used in applications where high fracture toughness is needed. Sulfur is removed from iron and steel when the metals are liquid. Although a variety of reagents are employed to remove sulfur, namely, calcium, magnesium, sodium, and rare earths, the most important is calcium. Common forms of calcium include the metal; alloys such as calcium silicon (CaSi); the oxide, calcium oxide (CaO); and the carbide, calcium carbide (CaC2). Despite the use of various forms of calcium, the governing chemical reaction in all cases appears to be the same (Ref 1, 2): CaO þ S ¼ CaS þ O
(Eq 1)
The equilibrium constant for the reaction (Eq 1) is: aCaS hO K9 ¼ aCaO hS
(Eq 2)
For both cast iron and steel, target sulfur concentrations after desulfurization are in the range of 0.006 to 0.010% S. In electric-melted cast iron and steel, the sulfur levels before desulfurization are 0.02 to 0.03% S, while input sulfur levels to the cupola are generally much higher—0.1 to 0.2% S. Requirements for Desulfurization. For reasons related to reaction kinetics and thermodynamics, the final sulfur (%S)f concentration achieved in desulfurization processes depends on three factors:
Purification of Ferrous Melts
Initial sulfur concentrations (%S)i Amount of desulfurizer used, usually expressed
One of the most important processes involved in cast iron and steel production is desulfurization. For steels, desulfurization is necessary to reduce
as the weight fraction of desulfurizer to metal, W Extraction efficiency of the desulfurizer, which is measured by the desulfurization
ratio (DR), that is, the ratio of sulfur concentrations: desulfurizer to metal The three variables can be related as follows through a mass balance on sulfur: ð%SÞf ¼
ð%SÞi 1 þ W ðDRÞ
(Eq 3)
For liquid-desulfurizing slags that are not saturated with respect to calcium sulfide (CaS), the maximum value of DR, that is, the equilibrium value, can be predicted from thermodynamic considerations: ðDRÞmax ¼
ð%SÞslag ð%SÞT
¼ max
CS fS K14 K15 hO
(Eq 4)
where CS is the slag sulfide capacity, defined as (Ref 3): CS ¼ ð%SÞslag
1=2 pO2 pS2
(Eq 5)
and K14 and K15 are the respective equilibrium constants for the reactions: 1 O2 ¼ O 2
(Eq 6)
1 S2 ¼ S 2
(Eq 7)
and fS is the activity coefficient for 1 wt% S in the standard state. Equation 4 can be derived from Eq 2. Using known input and desired output sulfur values, Eq 3 gives the required desulfurization ratios as a function of weight fraction desulfurizer. These data are plotted in Fig. 1 for two cases. In the electric-melting case, initial and final sulfur concentrations were assumed to be 0.03 and 0.008% S, respectively. In the cupola-melting
42 / Principles of Liquid Metal Processing
Weight ratio, Wdesulfurizer/Wiron
1
Table 1 Theoretical weight ratios: desulfurizer to iron to achieve 0.008% S in various systems T, K
CS ( 104)(a)
fS
ho ( 104)(b)
44 CaO-15MgO-5Al2O3-36SiO2 CaOsat-Al2O3
1773 1773
2.7 59
5 5
1.3 5.0
90 417
CaOsat-Al2O3 CaOsat-Al2O3
1773 1873
59 316
5 1
5.0 0.45
417 5280
Type of melt
A. cast iron—cupola B. cast iron—cupola (ladle desulfurized) C. cast iron—electric D. steel—electric
0.1 B A
100
(DR)max(c)
W
0.13 0.027 0.0065 0.00052
(a) CS is obtained from optical basicity correlations (Ref 3). fs is based on iron composition (Ref 4). (b) Case A: ho is governed by Si/SiO2 equilibrium based on respective concentration in iron and slag (Ref 5). Cases B and C: ho is governed by Si/SiO2 equilibrium with aSiO2 ¼ 1 due to ladle exposure to air (Ref 6). Case D: ho is governed by Al/Al2O3 equilibrium with %Al = 0.05 (Ref 1), assumed no contact with air. (c) K14 and K15 data from Ref 7
0.01
10-3 10
Slag composition
103
104
104
Desulfurization ratio (DR), (% S)slag/(%S)iron
Fig. 1
The best desulfurizing cupola slags have lower
sulfide capacity and desulfurization ratio than slags used in ladle desulfurization; the consequence is the need for much larger quantities of slag*. Compared to steel, cast iron desulfurization benefits from higher fS because of the presence of relatively high concentrations of carbon and silicon in cast iron. Ladle desulfurization systems that are exposed to air suffer higher hO and, as a result, poorer desulfurization than may otherwise be anticipated. For the cupola slag case in Table 1, the thermodynamically predicted value of (DR)max = 90 is in good agreement with measured data (Fig. 2). This indicates that the cupola desulfurization process operates close to equilibrium levels. Further evidence for this is given Fig. 3, which plots desulfurization data for a cupola operated with varying amounts of municipal ferrous refuse in the charge. The upper portion of Fig. 3 plots CS and oxygen activity. The latter is expressed as the partial pressure of oxygen. *The actual differences are less than those indicated. Calcium sulfide has only limited solubility in CaO-Al2O3 slags. As a result, a minimum W = 0.01 is needed to maintain the CaS-unsaturated condition.
103 Desulfurization ratio, aCaS/as
case, the equivalent concentrations were 0.10 and 0.008% S. The cupola line applies for both cupola iron and ladle-desulfurized cupola iron. The relationships illustrated in Fig. 1 are useful in defining systems that will provide the necessary desulfurization conditions. Four systems are compared in Table 1. These cover cupola- and electric-melted cast iron and electric-melted steel. Also examined are two liquid slag systems: a basic cupola slag (dicalcium silicate saturated) and a CaO-saturated CaOAl2O3 slag. Table 1 shows that:
100
Desulfurization ratio, (%S)slag /(%S)iron
Weight fraction desulfurizer required to achieve given desulfurization ratios. Plots for electricmelted iron and steel (curve A) assume initial sulfur concentrations [%(S)i = 0.03] different from those of cupola-melted iron [(%S)i = 0.10] (curve B).
100
10
10 0.7
0.9
1.1
1.3
1.7
1.5
1.9
2.1
2.3
Slag basicity
Fig. 2
Desulfurization ratio-basicity comparisons for cupola (closed circles) and laboratory data (open circles). Equilibrium values are indicated by the angled line. The vertical line indicates the slag basicity above which slags are saturated at 1500 C (2730 F), with respect to dicalcium silicate. This is the point at which the observed DR should equal (DR)max. Source: Ref 2
The lower portion of Fig. 3 is a comparison of (DR), measured and calculated. The good agreement found is evidence that near-equilibrium conditions existed. For the cases discussed previously, the oxygen activity in cupola iron has been found to be governed by the reaction (Ref 5, 8):
Mn þ O ¼ MnO
(Eq 8)
Therefore, the overall cupola desulfurization reaction is: CaO þ S þ Mn ¼ CaS þ MnO
(Eq 9)
Thermodynamic Properties of Iron-Base Alloys / 43
106
0.01
%C ⬚C 3.7 1550 4.0 3.7 1500 4.0
10−3
10−5 100 Desulfurization ratio
Inlet S 0.015 – 0.03% Inlet S 0.06 – 0.08%
(6 ⫻ 10−4)
Measured Calculated
50
0
0.5
1.0
1.5
3.7 4.0
1450
3.7 4.0
1400
2.0
2.5
3.0
Silicon concentration in iron, wt %
20 10
Fig. 4
Equilibrium sulfur concentrations for CaO desulfurization calculated with Eq 11 and 12, assuming aCaO ¼ aCaS ¼ aSiO2 ¼ 1. Source: Ref 2
5
Measured sulfur concentration, wt %
10−4
Sulfur concentration in iron, wt %
107
10
(po2)−1/2
Slag sulfide capacity
Cs (p o2)−1/2 Mn −3
0.1
(0.02)
108
0.01
0.01
2 1
0
20
40
60
80
100
Municipal refuse in charge, %
Fig. 3
Slag sulfide capacities, oxygen activities, and desulfurization ratios, measured and calculated, for a cupola operated with municipal ferrous refuse as a charge material. Source: Ref 8
Considerably lower sulfur could be achieved (Ref 2) if hO were governed by: C þ O ¼ CO
(Eq 10)
However, equilibrium for this reaction has not been observed. This discussion has concerned CaS-unsaturated slags. However, a desulfurizing slag, saturated with CaS, can in many cases continue to desulfurize as long CaO is present. In this case, the ultimate sulfur levels are not as low as those for CaS-unsaturated slags. Nevertheless, CaSsaturated slags can produce sulfur levels that are more than sufficient for cast iron and steel applications. In fact, CaS-saturated conditions are often employed in industrial applications because much less desulfurizer is required. In addition to the CaS-saturated liquid slags, totally solid slags such as CaO and CaC2 are in the same category, because CaS does not form solid solutions with these materials. Also in this category are desulfurizers containing small amounts of liquid phase. These materials possess the desirable properties of liquid and solid slags. That is, they combine the fast reaction rates of liquid slags with the large desulfurizing capacities (high CaO concentration) of solid slags. The final sulfur concentrations attainable under CaS-saturated conditions can be obtained from Eq 2 by setting aCaS = 1. For ladle desulfurization, the most important industrial desulfurizers are also CaO-saturated, that is, aCaO = 1.
Applying the condition of double saturation to Eq 2 yields:
1 10−3
ð%SÞr ¼
hO K9 fS
(Eq 11)
In the ladle desulfurization of cast iron, hO is governed by the reaction (Ref 2): 1 1 Si þ O ¼ SiO2 2 2
1
(Eq 12)
where aSiO2 ¼ 1. This is attributed to the exposure of the iron to air (Ref 2). Sulfur concentrations obtained with Eq 11, assuming siliconsilicon dioxide (Si-SiO2) equilibrium and aSiO2 ¼ 1, are given in Fig. 4. In continuous ladle-desulfurization processes, equilibrium sulfur levels are achieved when input sulfur levels are low but are only approached when input levels are high. This is illustrated in Fig. 5 with desulfurization data for CaC2 (Ref 2). Other desulfurizers such as CaO-CaF2 behave similarly. High desulfurization rates are needed for effective desulfurization in the short times required. For CaO-base desulfurizers, the presence of relatively small amounts of liquid phase (<25 vol%) significantly increases the rate of desulfurization. This is illustrated in Fig. 6, in which the rates of desulfurization by CaO with varying amounts of calcium fluoride (CaF2) are compared. The faster rates obtained with CaF2 additions were due to the formation of a CaOCaF2 liquid phase that provided a path for CaS diffusion from the reaction interface into the porous CaO particle. This prevented the normal rapid development of an impervious CaS coating on the CaO surface (Ref 10, 11). A similar explanation, involving liquid-phase formation, was used to account for the higher
(6 ⫻ 10−4) 0.0006
0.001
0.002
0.004
Theoretical sulfur concentration, wt%
Fig. 5
Comparisons of the sulfur concentrations in production-continuous desulfurization using CaC2 with the sulfur concentrations from Eq 11 and 12. Open circles indicate electric-melted iron. Closed circles indicate cupola iron. Source: Ref 2
CaO desulfurization rates of steel when aluminum was concurrently added (Ref 12, 13). Another rate-controlling variable in ladle desulfurization is bath agitation. For gas-stirred melts, the desulfurization rate constant k is _ that is, k / Q_ n . a function of gas flow rate Q, For well-dispersed solid-liquid mixtures, the rate of diffusion-controlled interphase reaction _ that is, n is a relatively weak function of Q, 0.2 to 0.4 (Ref 14–16). For a poorly dispersed system, such as a ladle of iron with a cover slag of desulfurizer, agitation has a much greater influence on the rate constant, with n typically in the range of 1.0 to 2.5 (Ref 14–16). This is due to the entrainment of increasing amounts of top slag into the liquid metal with increasing _ This effect is illustrated in Fig. 7, which Q. plots the apparent mass transfer coefficient for desulfurization versus the gas flow rate. Deoxidation. In the manufacture of steel, FeO-saturated conditions (0.2%O) are approached as a result of oxygen blowing to remove carbon and silicon as the respective oxides. To make a steel product, the oxygen levels are subsequently lowered to 0.005 to 0.02% O by reactive alloy additions of
44 / Principles of Liquid Metal Processing 1
0% CaF2
0.5
2%
0.2
%S/ %So
8%
0.1
5% CaF2
0.05
20%
10% 0.02
0.01 0
10
20
30
40
50
60
Time, min
Fig. 6
Desulfurization rates of carbon-saturated iron, containing 0.4% Si, with CaO and varying amounts of CaF2 at 1450 C (2640 F). Source: Ref 9
For single-element deoxidation, expressed by:
Desulfurization rate constant (k ), s -1 ⫻ 10-3
10
xM þ yO ¼ Mx Oy
5
(Eq 13)
the equilibrium oxygen concentration in the liquid steel, obtained by rearranging the equilibrium constant expression, is:
2 1
•
k a Q0.25
•
k a Q 2.1
0.5
ð%OÞ ¼
0.2 0.05 0.1
0.2
0.5
1
2
5
10
Gas flow rate (Q ) , m3 • kg-1 • s-1 ⫻ 10-6
Fig. 7
Dependence of CaO desulfurization rate constant on the rate of gas flow through a reactor. Source: Ref 16
electropositive elements. The low oxygen concentrations are necessary to maintain the number of oxide inclusions at a suitable low level, to prevent the formation of CO blowholes, and to ensure effective desulfurization.
a
Mx Oy x y K22 ð%MÞx fM fO
y1 (Eq 14)
The calculated equilibrium values of oxygen solubility in liquid iron at 1600 C (2910 F), based on Eq 14, are given in Fig. 8 for several elements as a function of the concentration of the element, assuming aMx Oy ¼1 . As seen among the common deoxidizers, aluminum produces the lowest oxygen. Although not shown, rare earth metals produce even lower oxygen (Ref 17). Figure 8 plots data for oxygen concentration and oxygen activity. In all cases, oxygen activity decreases with increasing levels of deoxidizer, but the oxygen concentration can go through a minimum. This occurs in cases where the
interaction coefficient eM O is a large negative number, and as a result, fO decreases significantly even at relatively low concentrations of M. To obtain exact values for [%O] or hO, equations for K22 can be found in Ref 1 and 7. As indicated by Eq 14, the oxygen concentration can be reduced if a complex deoxidation product is formed so that aMx Oy <1. Common deoxidation systems of this type are Si-Mn, Al-Si-Mn, or Al-CaO. An advantage, in addition to lower oxygen, is that less deoxidizer is required in solution to achieve a given level of oxygen. This is illustrated in Fig. 9, which plots the oxygen activity of iron as a function of the concentration of CaO in the calcium aluminate inclusion. Separate lines are given for aluminum concentrations ranging from 0.001 to 0.05% Al. Also indicated is the oxygen activity when pure aluminum oxide (Al2O3) is the reaction product. The data in Fig. 9 show that an oxygen activity of 4 ppm can be produced at three aluminum levels—0.002, 0.005, or 0.01%—depending on whether the respective deoxidation product was CaO-saturated calcium aluminate, CaAl2O4-saturated calcium aluminate, or Al2O3. The levels of total oxygen measured in steel melts considerably exceed equilibrium levels (Ref 7). Two possible explanations are the kinetic limitations in the deoxidation reaction and the ineffective separation of deoxidation products from the melt. Consideration of the first possibility suggests that the most important kinetic limitation would be oxygen and/or deoxidant solute diffusion to inclusions in the melt (Ref 7). Considering the presence of 105 to 107 inclusion particles in a cubic centimeter of melt, deoxidation to equilibrium levels of 10 to 20 ppm would require a few minutes. Appreciably longer times would be needed to reduce oxygen to less than 1 ppm, as with rare earth additions. For most cases, oxide formation appears faster than oxide particle separation from the bath. Thus, for applications that depend on dissolved oxygen, such as desulfurization or CO blowhole formation, equilibrium oxygen values can be assumed to exist after relatively short reaction periods. For considerations of the inclusion content of steel, the kinetics of particle separation need to be considered. The limiting case of particle flotation at velocities obtained from Stokes’ law calculations is shown in Fig. 10. Inclusion size versus time of flotation is plotted. The particles produced during deoxidation are generally small (<5 mm, or 200 min.); accordingly, separation times are long. The times indicated in Fig. 10 are representative of conditions in a quiescent bath. Times are considerably shorter for stirred melts because articles grow by collision and coalescence to sizes of 30 to 100 mm (Ref 7). The efficiency of coalescence of particles is material dependent. The energy of adhesion of particles is determined by the wetting characteristics of the particle, measured by wetting angle. For example, the force of adhesion of Al2O3
Thermodynamic Properties of Iron-Base Alloys / 45 20
1
0.002 % Al 0.001
Oxygen concentration Oxygen activity ao ∫ wt% O for %X Æ 0 Iron oxide saturation
Oxygen activity of iron, ppm 0
V2O4
0.1
MnO Cr2O4
SiO2 Oxygen is solution, wt %
16
B2O3
CO
Cr2O3
B2O3
V 2O 3
0.01
Ti2O3
10-3
0.005
0.01
0 30
0.01
CaO saturated
0.05 40
50
60
70
CaO concentration, mass%
ZrO2
Fig. 8
8
0.05
Al2O3
Fig. 9
Ti2O3 0.01
0.002 0.005
4 CaAl2O4 saturated
Ti3O5
10-4 -3 10
12
0.1 Deoxidizer in solution, wt %
1
10
Concentration of oxygen and aluminum in liquid steel in equilibrium with calcium aluminate at 1600 C (2910 F); arrows indicate oxygen values when the reaction product is pure Al2O3. Source: Ref 18
Deoxidation equilibria in liquid iron alloys at 1600 C (2910 F). Source: Ref 17
107
106
10
15 105
20 30 40 0.1
A
B
1
C
10
Z, number of nuclei • cm−3
Inclusion size for 0.05%O removal, µm
5
100
Flotation time, min
Fig. 10
Calculated time of flotation of inclusions in stagnant melts as a function of inclusion size. Melt depth: A, 50 mm (2 in.); B, 500 mm (20 in.); C, 2000 mm (80 in.). Source: Ref 7
particles with a wetting angle of 140 is twice that for SiO2 with a wetting angle of 115 . On this basis, Al2O3 inclusions are expected to cluster more easily than SiO2 inclusions, a phenomenon that has been observed in practice (Ref 7).
Thermodynamics of Ferrous Systems For binary systems, only the iron-carbon and iron-silicon systems are discussed in this article. Information on the iron-oxygen system is available in the article “Inclusion-Forming Reactions” in this Volume.
Fig. 11
Phase diagram for the binary iron-carbon system. Source: Ref 19
The Iron-Carbon System The phase diagram of the binary iron-carbon system includes the stable (iron-graphite) and
metastable (Fe-Fe3C) equilibria. The most recent diagram is shown in Fig. 11. The solubility of carbon in iron is described in Eq 15–20, Eq 21. (Ref 20, 21).
46 / Principles of Liquid Metal Processing For the stable system, that is, with graphite as the equilibrium high-carbon phase, Eq 15–17 are used. For the liquid:
The general equation used to describe the activity of a component 2 in a component 1 according to the Darken formalism (Ref 22) is:
wt%Cmax ¼ 1:3 þ 2:57 103 t
log
(Eq 15)
in the interval 1152 to 2000 C (2106 to 3632 F), with t being the temperature in degrees Celsius. Using thermodynamic quantities, Eq 15 can be written as: log XCmax
12:728 ¼ T þ 0:727 log T 3:049
(Eq 16)
where XCmax is the maximum solubility of graphite in liquid iron in mole fraction, and T is the temperature in degrees Kelvin. Equation 16 is valid for the interval 1425 to 2300 K. For the austenite: wt%Cmax ¼ 0:435 þ 0:355 103 t þ 1:61 106 t2
wt%Cmax
(Eq 18)
(Eq 19)
For the austenite:
(Eq 23)
1180 0:87 T
(Eq 24)
Using the natural logarithm (base e) instead of the common logarithm (base 10), Eq 22 can be written as follows for the iron-carbon systems: 2714 ln gC ¼ 2 T 2920 þ þ 4:01 2XC XC2 T
(Eq 25)
153 aFe-Si ¼ þ 3:02 T
log goSi ¼ 0:914
The Iron-Silicon System
6863 T
(Eq 26)
(Eq 27)
As shown in Fig. 13, calculations with Eq 26 and 27 are in good agreement with experimental data (note that the equation in Fig. 13 results from introducing Eq 26 and 27 into Eq 22 and using a natural logarithm rather than a common logarithm).
Ternary Iron-Base Alloys Using a Taylor series expansion for the excess partial molar free energy, one researcher derived an expression for the activity coefficient of component 2 in a multicomponent solution, which for dilute solutions can be written as (Ref 24): ln f2 ¼ ln
g2 X2 e22 þ X3 e23 þ X4 e24 ; . . . go2
(Eq 28)
where g2 is the activity coefficient of component 2 dissolved in the three-component system, go2 is the activity coefficient of component 2 in the range of infinitely dilute solution, and e22, e23, e24, . . . are coefficients defined as: @ ln f2 @X2 limit @ ln f2 ¼ X1 !1 @X3 limit
e22 ¼ X1 !1 e23
(Eq 20)
For the ferrite: 10; 908 wt%Cmax ¼ 1:8 103 exp T
1270 aFe-C ¼ þ 1:74 T
Equation 25 is in good agreement with experimental data, as shown in Fig. 12.
wt%Cmax ¼ 0:628 þ 1:222 103 t þ 1:045 106 t2
where g2 is the activity coefficient of component 2 = a2/X2 (a2 being the activity of component 2 and X2 the mole fraction of component 2); go2 is the activity coefficient of component 2 in the range of infinitely dilute solutions, that is, when %component 2 ffi 0; it is only a function of temperature (g2 ! go2 as X2 ! 0); and a12 is the function of temperature, usually of the shape A/T + B, where A and B are constants. For the activity of carbon in an iron melt, Eq 22 is valid for component 1 (iron) and component 2 (carbon) with (Ref 23):
log goC ¼
For the metastable system, that is, with Fe3C as the equilibrium high-carbon phase, Eq 19–21 apply. For the liquid: wt%Cmax ¼ 4:34 þ 0:1874ðt 1150Þ t 200 ln 1150
(Eq 22)
(Eq 17)
For the ferrite: 11; 460 ¼ 2:46 103 exp T
g2 ¼ a12 ðX22 2X2 Þ go2
with the following remarks (Ref 23): XSi 0.2; component 1 is iron; component 2 is silicon:
(Eq 29)
Equation 22 can be used to calculate the activity of silicon in the iron-silicon system,
For infinitely dilute solutions, that is, X2 ! 0, X3 ! 0: e23 ¼ e32
Fig. 12
Fig. 13
(Eq 21)
Equation 15–21 can be used to calculate the maximum solubility of carbon in various phases (liquid, austenite, ferrite) as a function of temperature. For example, to calculate the carbon content of the stable eutectic, Eq 15 is used with t = 1152 C (2106 F) to obtain %Cmax = 4.26%. Prediction of phase stability relies heavily on knowledge of the activity of various elements. In the specific case of iron-carbon alloys, an increase in the activity of carbon in the liquid parallels a greater ability of carbon to separate as graphite; that is, the activity of carbon is a measure of an increased tendency of the alloy to solidify as the stable system. On the other hand, a decreased activity of carbon reflects the carbide-promoting behavior of the system, that is, the metastable solidification tendency.
Dependence of the activity coefficient of carbon on the mole fraction of carbon in liquid iron at 1550 C (2820 F). Source: Ref 23
Dependence of the activity coefficient of silicon on the mole fraction of silicon in liquid iron at 1600 C (2910 F). Source: Ref 23
Thermodynamic Properties of Iron-Base Alloys / 47 Simplification of the Darken formalism for infinite dilutions gives the interrelationship between e and a, as follows (Ref 22): e22 ¼ 2a12 2:303 e33 ¼ 2a13 2:303 e23 e32 ¼ ¼ a23 a12 a13 2:303 2:303
ðEq 30Þ
The solute concentration is often given in weight percent. For a ternary system, Eq 28 is written: log fi ¼ eii ½%i þ eij ½%j
(Eq 31)
where the interaction coefficient e is related to E by: eii ¼ 230ðMi =M1 Þeii þ ½ðM1 Mi Þ=M1 eij ¼ 230ðMj =M1 Þeij þ ½ðM1 Mj Þ=M1
log ðEq 32Þ
Data on interaction coefficients in liquid iron for carbon, hydrogen, nitrogen, oxygen, and sulfur are given in Table 2. For metallurgical reactions involving dilute solutions, it is often necessary to change the standard state from pure component to that of 1 wt% in solution. The free energy change is given by: Gs ¼ RT ln
0:5585goi Mi
available in Ref 25. Only the Fe-C-Si, Fe-C-P, and Fe-C-Mn systems are discussed in this article. The Fe-C-Si System. The addition of silicon to a binary iron-carbon alloy decreases the stability of Fe3C, which is already metastable, and increases the stability of ferrite (the a field is enlarged, and the g field is constricted). The equilibrium diagrams in Fig. 14 show that as the silicon content in the Fe-C-Si system increases, the carbon contents of the eutectic and eutectoid decrease (%Cmax decreases), while the eutectic and eutectoid temperatures increase (Ref 26). The activity coefficient, g2, of component 2, dissolved in a three-component system as a function of the concentration X2 and X3 of dissolved components 2 and 3, is described by Eq 34 (Ref 22):
(Eq 33)
Free energies of solution DGs of selected elements in liquid iron are given in Table 3.
g2 ¼ 2a12 X2 þ ða23 a12 a13 ÞX3 go2 þ a12 X22 þ a13 X32 þ ða12 þ a13 a23 ÞX2 X3
ðEq 34Þ
where a12 and a13 are the modified interaction parameters of the two-component systems, a23 is the modified interaction parameter that describes the activity of both dissolved components, and go2 is the activity coefficient of component 2 in the range of infinitely dilute solution. Starting from Eq 34, the activity of carbon and silicon in an Fe-C-Si melt can be calculated as a function of temperature and concentration with the following relationships (Ref 23): log aC ¼ð1180=T Þ 0:87 þ ½ð2540=T Þ þ 3:48 XC
Ternary Fe-C-X Systems
½ð1270=T Þ þ 1:74 XC2 þ log XC
Detailed equilibrium diagrams of the Fe-C-B, Fe-C-Cr, Fe-C-Cu, Fe-C-Mn, Fe-C-Mo, Fe-CN, Fe-C-Ni, Fe-C-Si, and Fe-C-W systems are
þ ½ð1423=T Þ þ 4:75 XSi 2 ½ð153=T Þ þ 3:02 XSi
½ð1423=T Þ þ 4:76 XC XSi
ðEq 35Þ
log aSi ¼ ð6863=T Þ þ 0:914 þ ½ð306=T Þ þ 6:04 XSi 2 ½ð153=T Þ þ 3:02 XSi þ log XSi
þ ½ð1423=T Þ þ 4:76 XC ½ð1270=T Þ þ 1:74 XC2 ½ð1423=T Þ þ 4:76 XC XSi
ðEq 36Þ
The Fe-C-P System. Because the partition coefficient of phosphorus between solid and liquid iron is low, phosphorus segregates into the liquid during solidification, and the melt becomes supersaturated in phosphorus, which results in the formation of Fe3P. Therefore, the four possible solid phases of the stable and metastable systems are austenite (g), Fe3P, Fe3C, and graphite. The calculated liquidus surfaces for the four phases and the eutectic lines formed by their interactions are shown in Fig. 15, which is a superimposition of the g-Fe3P-Fe3C and the g-Fe3P-graphite (Gr) diagrams. Two ternary eutectic points are shown in Fig. 15. To the left is L ! g + Fe3P + Gr at 954 C (1749 F) and yC = 0.099, yP = 0.123. To the right is L ! g + Fe3P + Fe3C at 948 C (1738 F) and yC = 0.106, yP = 0.123. Here, yP = XP/(1 XC), and yC = XC/(1 XC), where XP is the weight fraction of phosphorus and XC is the weight fraction of carbon. In general, it is agreed that phosphorus does not have a strong graphiteor carbide-stabilizing influence. The Fe-C-Mn System. The influence of increasing manganese contents in the Fe-C-Mn system over the equilibrium phase diagram is shown in Fig. 16. It is evident that manganese: Decreases the eutectoid temperature Increases the eutectoid interval, that is, the
range of temperature and carbon contents over which a, g, and carbide can coexist Decreases the carbon content in the eutectoid and in the eutectic Increases the eutectic temperature (by approximately 3 C, or 5 F, for each 1% Mn)
Table 2 Interaction coefficients in ternary iron-base alloys for carbon, hydrogen, nitrogen, oxygen, and sulfur at 1600 C (2910 F) Element j
Aluminum Boron Carbon Cobalt Chromium Copper Manganese Molybdenum Nitrogen Niobium Nickel Oxygen Phosphorus Sulfur Silicon Tin Titanium Vanadium Tungsten Zirconium Source: Ref 4, 7
eCj
eHj
eNj
eOj
eSj
0.043 0.24 0.14 0.008 0.024 0.016 0.012 0.008 0.11 0.06 0.012 0.34 0.051 0.046 0.08 0.041 ... 0.077 0.006 ...
0.013 0.05 0.06 0.002 0.002 <0.001 0.001 0.002 ... 0.002 0 0.19 0.011 0.008 0.027 0.005 0.019 0.007 0.005 ...
0.028 0.094 0.13 0.011 0.047 0.009 0.02 0.011 0 0.06 0.01 0.05 0.045 0.007 0.047 0.007 0.53 0.093 0.002 0.63
3.9 2.6 0.13 0.008 0.04 0.013 0.021 0.004 0.057 0.14 0.006 0.20 0.07 0.133 0.131 0.011 0.31 0.14 0.009 (3.0)
0.035 0.13 0.11 0.003 0.011 0.008 0.026 0.003 0.01 0.013 0 0.27 0.29 0.028 0.063 0.004 0.072 0.016 0.01 0.052
Table 3 Free energies of solution in liquid iron at 1 wt% Mi, liquid iron at 1600 C (2910 F) Element i
Aluminum (l) Carbon (g) Cobalt (l) Chromium (s) Copper (l) ½ H2(g) Manganese (l) ½N2 (g) Nickel (l) ½O2 (g) ½P2(g) ½ S2(g) Silicon (l) Titanium (s) Vanadium (s) Tungsten (s) Zirconium (s)
goi
D Gs (cal g atom1)
0.029 0.57 1.07 1.14 8.6 ... 1.3 ... 0.66 ... ... ... 0.0013 0.038 0.1 1.2 0.043
15,100 6.67T 5400 10.10T 240 9.26T 4600 11.20T 8000 9.41T 8720 + 7.28T 976 9.12T 860 + 5.71T 5000 7.42T 28,000 0.69T 29,200 4.60T 32,280 + 5.60T 31,430 4.12T 7440 10.75T 4950 10.90T 7500 15.20T 8300 11.95T
48 / Principles of Liquid Metal Processing
Fig. 14
Influence of silicon content on the solubility lines and equilibrium temperatures of the iron-carbon system. (a) to (c) Source: Ref 26, (d) to (g) Source: Ref 25
The carbide M3C, which is (FeMn)3C, is stable over a wide range of manganese and iron contents, extending at 1000 C (1832 F) from Mn3C to nearly pure Fe3C. At high manganese contents (for example, above 40% Mn), other manganese carbides such as M5C2, M7C3, M23C6, and M4C1+x(e) can form (Ref 25). Influence of a Third Element on the Equilibrium Temperatures. Various elements, dissolving in liquid and solid iron phases, change the equilibrium temperatures on the iron-carbon diagram, as indicated in Table 4. This results in increased or decreased a and g fields.
Of particular interest for the case of cast iron is the influence of third elements on the stablemetastable eutectic interval (Tst Tmet). In general, elements that increase the Tst Tmet interval promote graphite formation, while those that decrease the interval promote carbide formation. Based on their specific influence on the Tst Tmet interval, third elements can be classified into the following four groups (Fig. 17): Strong graphitizers that increase Tst and
decrease Tmet, such as silicon, aluminum, nickel, and copper (Fig. 17a)
Weak graphitizers that decrease both Tst and
Tmet but increase the Tst Tmet interval overall, such as phosphorus and arsenic (Fig. 17b) Strong carbide stabilizers that decrease Tst but increase Tmet, such as chromium, vanadium, and manganese (Fig. 17c) Weak carbide stabilizers that decrease both Tst and Tmet, such as molybdenum and tungsten (Fig. 17d) st ¼ Tst Tmet interval Real values for the Tmet for a number of elements can be calculated with
Thermodynamic Properties of Iron-Base Alloys / 49 the data given in Table 4, which also includes data on the influence of some elements on the eutectoid temperature and on the maximum solubility of carbon in austenite. Two examples of the use of data in Table 4 are provided as follows.
having 2% Si, 0.5% Mn, and 1% Cu by weight: Tst ¼ 1154 C þ 4% Si 2% Mn
Example 1: Calculation of the value of the Tst Tmet eutectic interval for a cast iron
TFe-C-Mn-Cu ¼ 1166 1117:2 ¼ 48:8 C
þ 5% Cu ¼ 1166 C Tmet ¼ 1148 C 15% Si þ 3% Mn 2:3% Cu ¼ 117:2 C TFe- C 1154 1148 ¼ 6 C
Fig. 15
Calculated liquidus surfaces in the Fe-C-P phase diagram. Source: Ref 27
Fig. 16
Vertical sections through the Fe-C-Mn ternary phase diagram at (a) 4.92, (b) 12.8, and (c) 19.7 wt% Mn
Example 2: Calculation of the temperatures of the point of maximum carbon solubility in austenite, TE and TE0 , for the aforementioned iron: TE ¼ 1154 C 10%Si þ 3:2% Mn 2%Cu ¼ 1133:5 C TE0 ¼ 1148 C þ 2:5% Si 2% Mn þ 5:2%Cu ¼ 1157:2 C Some typical experimental results for silicon, chromium, and vanadium are shown in Fig. 18. These results are in agreement with theoretical predictions. Partition of a Third Element Between Various Phases in the Fe-C-X System. It has been established from the thermodynamic stability of phases involved in the eutectic solidification of Fe-C-X alloy that the differences in the equilibrium partition coefficients DP (= PM PS, where PM and PS are the partition coefficients of element X between liquid and eutectic metastable or stable solid), the partition of element X between austenite g=L and liquid, PX , and the partition of element X Fe C=g between cementite and austenite, PX 3 , are related to one another and have similar effects on graphitization during eutectic solidification (Ref 28). Calculated and experimental data for partition coefficients of various elements in phases of iron-carbon alloys are given in Table 5. Figure 19 illustrates the correlation between the graphitizing influence of a third element (expressed by the normalized addition of each element required to increase or decrease the chill depth by a given amount) and the DP value. It is quite obvious that a definite correlation exists, although silicon, aluminum, and titanium deviate from the line. These three elements have a strong affinity for nitrogen; this may explain their higher-than-expected graphitizing influence. Indeed, it is accepted that the nitrides may act as
50 / Principles of Liquid Metal Processing nuclei for the stable eutectic, promoting gray solidification. This is why experiments show that titanium, for example, can act either as a carbide or graphite promoter, depending on the nitrogen content. Influence of a Third Element on Carbon Solubility in the Fe-C-X System. The influence of a third element on the solubility of carbon can be expressed by (Ref 20): XCX
¼
XCXmax
XCmax ðmole fractionÞ
ðEq 37aÞ
concentrations. In the region of low concentration of the third element (%X = 0 to 5%), the change in the solubility of carbon can be represented by the temperature-independent linear equation: XCX ¼ m XX
ðEq 38aÞ
or
%CEX ¼ m0 g %X 0
%C ¼ m %X X
ðEq 38bÞ
%C ¼
X %Cmax
%Cmax %Cmax ðwt%Þ
ðEq 37bÞ
where XCX and D%CX represent the increase or decrease of carbon solubility in mole fraction or weight percent, respectively; XCmax and % Cmax represent the saturation concentration in the iron-carbon system calculated from Eq 15 or 16; and XCXmax and %CX max are the saturation concentration in the Fe-C-X system. The changes in carbon solubility resulting from additions of third elements, X, are shown in Fig. 20. These data are valid in the range from 1200 to 1700 C (2190 to 3090 F), with the exception of silicon and sulfur at high
where m and m are solubility factors. In Eq 38a and 38b, m and m0 as well as XCX and D% CX are positive for an increase in the solubility of carbon and negative in the other case. It is also generally accepted that carbide-promoting elements increase the solubility of carbon (activity coefficient of carbon in solution is decreased), while graphite-promoting elements decrease it (activity coefficient of carbon in solution is increased). Therefore, third elements that have a negative solubility factor promote graphitization, while elements that have a positive solubility factor promote carbide formation. The value of the solubility factor is proportional to their effect. It must be noted, however, that the accuracy of the prediction of the behavior of elements based on solubility factor is questioned in
Table 4 Influence of a third element on the temperature change of critical points on the iron-carbon diagram +, increase; , decrease. Maximum solubility of carbon in austenite, C/wt%X Element
Metastable (E)
Silicon Copper Aluminum Nickel Cobalt Chromium Manganese Molybdenum Tungsten Boron Nickel Titanium Vanadium Phosphorus Sulfur
10 to 15 2 14 4.8 ... +7.3 +3.2 ... ... ... +6–8 180 ...
0
Stable (E )
+2.5 +5.2 +8 +4 ... 2 ... ... ... ... 180 ...
ðEq 39Þ
Data for m0 g are also given in Table 6.
0
X
Ref 28 for elements that have a strong influence on the austenite liquidus. As indicated in Table 6, experimental values are close to the theoretical values for many elements. Equations similar to Eq 38a and 38b can be used to calculate the change in the maximum solubility of carbon in austenite upon the addition of a third element:
Multicomponent Iron-Carbon Systems Carbon Solubility in Multicomponent Systems. Assuming that the values in Table 6 for m and m0 , which were determined for the ternary Fe-C-X system, can be extended to multicomponent systems (in this case, the effect of alloying elements on the solubility of carbon can be considered to be additive, as shown in Eq 41, at least in the region of low concentrations), the saturation concentration of carbon in multicomponent systems can be calculated as (Ref 20): %CSi;Mn;P;S;... ¼ %Cmax þ %CSi;Mn;P;S;... max
ðEq 40Þ
which is Eq 37b rewritten for multicomponent systems. In turn: %CSi;Mn;P;S;... ¼ %CSi þ %CMn þ %CP max
Eutectoid, C/wt%X
Eutectic, C/wt%X 0
Metastable (S) Stable (S ) Metastable C Stable C
+8 ... +10 20 ... +15 9.5 + + ... ... ... +15 + ...
0–30 10 +10 30 ... +8 3.5 + + ... ... ... + +6 ...
10 to 20 2.3 15 6 ... +7 +3 ... ... ... 6–8 37 ...
+4 +5 +8 +4 ... 2 ... ... ... 30 ...
þ %CS þ . . .
Eutectic, C/at. %X(a) 0
Metastable C Stable C
3.25 3.32 7.74 1.33 2.23 4.97 2.33 8.90 12.13 15.47 19.67 16.91 ... 16.05 .. . .
ðEq 41Þ
0
10.91 8.41 0.91 6.50 1.48 10.23 7.15 12.36 15.55 18.53 16.87 18.91 ... 16.98 18.53
(a )Calculated results. Source: Ref 21, 28
Then, using Eq 15 and data from Table 6, Eq 40 becomes: %CSi;Mn;P;S;... ¼ 1:3 þ 2:57 103 t 0:31% Si max þ 0:027% Mn 0:33% P 0:4% S . . .
ðEq 42Þ
Saturation Degree and Carbon Equivalent. The iron-carbon diagram with letter notations for critical points (Fig. 21) is used in this section. It is well known that mechanical properties of cast iron strongly depend on the amount and shape of graphite. In turn, for hypoeutectic gray irons, the amount of graphite depends on the amount of eutectic. The amount of eutectic can be calculated with the lever rule. For hypoeutectic irons, one can write: Amount of eutectic Amount of eutectic þ amount of austenite %Canal %C0E ¼ %Canal %C0E þ %C0C %Canal %Canal %C0E ¼ ðEq 43Þ %C0C %C0E
Sr ¼
Fig. 17
Classification of the influence of a third element in the Fe-C-X system on the graphite- or carbide-promoting tendency in cast iron, based on their influence on the eutectic stable and metastable temperatures. (a) Strong graphitizers. (b) Weak graphitizers. (c) Strong carbide stabilizers. (d) Weak carbide stabilizers
where Sr is the rectified saturation degree, % Canal is the analyzed carbon content of cast iron, and %C0 C and %C0 E represent the carbon content
Thermodynamic Properties of Iron-Base Alloys / 51
Fig. 18
Influence of (a) chromium, (b) silicon, and (c)
Table 5
Equilibrium partition coefficients of a third element X in the Fe-C-X system g=L
Fe C=L
PX (a)
PX 3
(a)
Element X
calc
exp
calc
exp
PS
PM
DP = PM PS
Silicon Copper Aluminum Nickel Cobalt Chromium Manganese Molybdenum Tungsten Boron Nitrogen Titanium Phosphorus Sulfur
1.71 1.57 ... 1.46 1.18 0.53 0.70 0.41 0.23 0.06 2.04 0.04 0.15 0.06
1.72 1.62 1.15 1.61 1.13 0.55 0.75 0.38 0.42 ... 2.04 ... ... ...
0 0.12 ... 0.43 0.59 1.96 1.03 0.60 0.42 0.22 2.12 0.09 0.08 ...
0.05 0.08 0.03 0.32 0.60 1.95 1.21 0.84 0.88 ... ... 0.27 0.09 ...
1.55 1.43 1.05 1.33 1.07 0.48 0.64 0.37 0.21 0.06 1.86 0.04 0.14 0.06
0.78 0.78 0.55 0.90 0.85 1.32 0.90 0.52 0.33 0.15 2.09 0.07 0.11 ...
0.77 0.65 0.50 0.43 0.21 0.84 0.26 0.15 0.12 0.09 0.23 0.03 0.03 ...
(a) calc, calculated; exp, experimental. Source: Ref 28
temperature of the multicomponent system. Further simplification gives: SC ¼
Fig. 19
Correlation between the partition coefficient, DP, and the graphitizing influence of a third element in the Fe-C-X system. Source: Ref 28
of the eutectic and of the austenite, respectively, at the eutectic temperature in the multicomponent system (Fig. 21). Assuming the solubility factors can be extrapolated to the hypoeutectic region, %C0 C can be calculated with Eq. 42. A similar approach can be used to calculate %C0 E using data in Table 6 for m0 g: %C0E ¼ 2:11 0:11% Si 0:35%P þ 0:006% Mn 0:08% S
ðEq 44Þ
In foundry practice, the following simplified form of Eq 43 is used: SC ¼ ¼
%Canal %C0C %Canal 4:26 0:31% Si 0:33% P þ . . .
ðEq 45Þ
For more accuracy, %Cmax as a function of temperature rather than %Cmax = 4.26 can be used, where the temperature is the eutectic
%Canal 4:25 0:3ðSi þ PÞ
ðEq 46Þ
where SC < 1 for hypoeutectic irons, SC = 1 for eutectic irons, and SC > 1 for hypereutectic irons. Saturation degree is used in most of the European literature. In the Anglo-Saxon literature and foundry practice, carbon equivalent rather than saturation degree is used. Carbon equivalent (CE) can be calculated as: CE ¼ %Canal %CSi %CMn %CP %CS . . . ¼ %Canal þ 0:31%Si þ 0:33% P 0:027% Mn þ 0:4% S
ðEq 47Þ
A eutectic iron has CE = 4.26%. It must be noted that although Sr gives directly the amount of eutectic in the structure (for example, Sr = 0.9 means 90% eutectic), SC and CE, although easier to calculate, do not allow for exact estimation of the amount of eutectic.
Structural Diagrams For the practicing metallurgist involved in cast iron production, one of the main applications of the thermodynamics of the iron-carbon system is the calculation of structure-composition
correlations. The so-called structural diagrams are used for this purpose. The first and simplest structural diagram is the Maurer diagram (Fig. 22a), in which the as-cast structure is considered to be solely the product of the carbon and silicon contents of cast iron. Because it was recognized long ago that solidification rate, as well as chemical composition, plays a significant role in the formation of the as-cast structure of casting, especially for cast iron, Laplanche (Ref 29) further developed the Maurer diagram to include the influence of cooling rate through the section size (or bar diameter) of the casting (Fig. 22b). The k lines on the diagram are lines of same structure (same degree of graphitization) at a given cooling rate. The value k is an empirical function of composition, as follows: k ¼ 4=3 Si 1
5 3C þ Si
ðEq 48Þ
The correlation between k and cooling rate is given in Table 7. Thus, if one wants to produce a pearlitic iron in casting with equivalent section sizes of 10 to 30 mm (0.4 to 1.2 in.) bar diameters, Table 7 would suggest a k of 0.85 to 2.35. For example, k = 1.8 is then chosen. A saturation degree, SC, is also selected depending on the required mechanical properties (for example, SC = 0.8), and the intersection of the k = 1.8 with SC = 0.8 will then give the carbon and silicon required for the iron.
52 / Principles of Liquid Metal Processing A further contribution to structural diagrams was made by Patterson and Doepp (Ref 30). This diagram (Fig. 23) encompasses not only structure, composition, and cooling rate but also a broad range of mechanical properties. Part of this diagram is based on Laplanche’s k lines. As an example of the way this diagram is used, assume it is desired to pour a 50 mm (2 in.) diameter cast iron bar (or equivalent section size) with a pearlitic-ferritic structure. A heavy horizontal line is drawn from the y-axis at the 50 mm (2 in.) diameter bar location to the middle of zone IIb. This corresponds to a determination of k = 2.2. This k line can be followed to the lower diagram until it intersects a preselected saturation degree—in this case, SC = 0.6. The composition is now read as C = 2.6% and Si = 3%, together with the equilibrium tem! perature for the ðSiO2 Þ þ 2½C ½Si þ fCOg reaction, which is 1460 C (2660 F). An additional explanation of the significance of this reaction can be found in the article “Solidification of Eutectic Alloys: Cast Iron” in this Volume. It must be pointed out that foundry variables such as superheating and holding during melting, charge composition (raw materials), and inoculation can cause significant deviation from these calculations.
ACKNOWLEDGMENTS Reviewed and adapted from Seymour Katz, Composition Control, Casting, Vol 15, Metals Handbook, 9th Edition, ASM International, 1988, p71–82; and D.M. Stefanescu, Thermodynamic Properties of Iron-Base Alloys, Casting, Vol 15, Metals Handbook, 9th Edition, ASM International, 1988, p 61–70 REFERENCES 1. R.J. Fruehan, Ladle Metallurgy: Principles and Practices, Iron and Steel Society of AIME, 1985, p 7–8 2. S. Katz and C.F. Landefeld, Cupola Desulfurization, Cupola Handbook, American Foundrymen’s Society, 1984, p 351–363 3. I.D. Sommerville, The Capacities and Refining Capabilities of Metallurgical Slags, Foundry Processes: Their Chemistry and Physics, S. Katz and C.F. Landefeld, Ed., Plenum Press, 1988, p 101–133 4. G.K. Sigworth and J.F. Elliott, The Thermodynamics of Liquid Dilute Iron Alloys, Met. Sci., Vol 8, 1974, p 298–310 5. S. Katz and H.C. Rezeau, The Cupola Desulfurization Process, Trans. AFS, Vol 87, 1979, p 367–376 6. S. Katz, D.E. McInnis, D.L. Brink, and G.A. Wilkinson, The Determination of Aluminum in Malleable Iron from Measured Oxygen, Trans. AFS, Vol 88, 1980, p 835–844 7. E.T. Turkdogan, Physical Chemistry of High Temperature Technology, Academic Press, 1980 Table 6
Solubility factors of various third elements for carbon-saturated Fe-C-X melts
Influence of third elements, X, on the solubility of carbon in molten iron. (a) In mole fraction. (b) In weight percent. Source: Ref 20
Fig. 20
5 6 13 14 15 16 22 23 24 25 26 27 28 29 32 33 41 42 44 50 51 52 73 74 76 78 79 92
Symbol
B C Al Si P S Ti V Cr Mn Fe Co Ni Cu Ge As Nb Mo Ru Sn Sb Te Ta W Os Pt Au U
D %CX = m0 %X
DXX C ¼ m XX (a)
Third element X Atomic No.
8. S. Katz and V.R. Spironello, Effect of Charged Aluminum on Iron Temperature, Silicon Recovery and Desulfurization in an Iron-Producing Cupola, Trans. AFS, Vol 92, 1984, p 161–172 9. C.F. Landefeld and S. Katz, Kinetics of Iron Desulfurization by CaO-CaF2, Proceedings of the Fifth International Iron and Steel Congress, Vol 6, Iron and Steel Society of AIME, 1986, p 429–439 10. S. Katz and C.F. Landefeld, Plant Studies of Continuous Desulfurization with CaOCaF2-C, Trans. AFS, Vol 93, 1985, p 215–228 11. S. Katz and B.L. Tiwari, A Critical Overview of Liquid Metal Processing in the Foundry, Foundry Processes: Their Chemistry and Physics, S. Katz and C.F. Landefeld, Ed., Plenum Press, 1988, p 1–52 12. J. Niederinghaus and R.J. Fruehan, Desulfurization Mechanisms for CaO-Al and CaO-CaS in Carbon Saturated Iron, Metall. Trans. B, to be published 13. E.T. Turkdogan, Physiochemical Phenomena of Mechanisms and Rates of Reaction in Melting, Refining and Casting of Foundry Irons, Foundry Processes: Their Chemistry and Physics, S. Katz and C.F. Landefeld, Ed., Plenum Press, 1988, p 53–100 14. S. Asai and I. Muchi, Fluid Flow and Mass Transfer in Gas Stirred Ladles, Foundry Processes: Their Chemistry and Physics, S. Katz and C.F. Landefeld, Ed., Plenum Press, 1988, p 261–292 15. S.-H. Kim and R.J. Fruehan, Physical Modelling of Liquid/Liquid Mass Transfer in Gas Stirred Ladles, Metall. Trans. B, Vol 18, 1987, p 381–390
0 D%CX C ¼ mg %X
m exp
m calc
Validity
m exp
m calc
Validity
m0 g
Validity
0.575 0.67 0.52 0.71 0.81 1.00 +0.508 +0.359 +0.215 +0.105 0.0 0.10 0.20 0.313 0.677 0.734 +0.51 +0.24 0.0 0.70 0.79 0.85 +0.49 +0.256 0.0 0.224 0.333 +0.53
0.51 0.68 0.51 0.68 0.85 1.02 +0.44 +0.33 +0.22 +0.11 0.0 0.11 0.22 0.33 0.66 0.77 +0.33 +0.22 0.0 0.66 0.77 0.88 +0.33 +0.22 0.0 0.22 0.33 +0.55
<0.04 ... <0.05 <0.007 <0.048 <0.004 <0.013 <0.1 <0.13 <0.30 ... <0.20 <0.10 <0.02 <0.1 <0.07 <0.045 <0.14 <0.008 <0.02 <0.05 <0.005 <0.03 <0.05 <0.013 <0.015 <0.024 <0.01
0.539 0.610 0.220 0.310 0.331 0.405 +0.159 +0.105 +0.064 +0.028 0.0 0.027 0.051 0.076 0.144 0.151 +0.058 +0.014 0.020 0.110 0.119 0.122 +0.004 0.015 0.035 0.052 0.060 0.006
0.462 0.622 0.215 0.294 0.349 0.414 +0.138 +0.097 +0.062 +0.029 0.0 0.029 0.055 0.080 0.141 0.159 +0.030 +0.012 0.023 0.104 0.117 0.126 0.009 0.01 0.035 0.052 0.060 0.005
<1.0 ... <2.0 <5.5 <3.0 <0.4 <1.6 <11.0 <10.0 <25.0 ... <40.0 <8.0 <3.8 <14.0 <10.0 <9.0 <2.0 <1.5 <4.5 <15.0 <1.5 <11.0 <18.0 <4.0 <6.0 <10.0 <5.0
... ... +0.04 0.11 0.35 0.08 2.08 +0.18 0.07 +0.006 ... 0.017 0.07 +0.014 ... ... ... 0.3 ... ... ... ... ... 0.12 ... ... ... ...
... ... <2 <6 <0.4 <5 <0.8 <1 <20 <60 ... <20 <6 <5 ... ... ... <4 ... ... ... ... ... <12 ... ... ... ...
(a) exp, experimental; calc, calculated. Source: Ref 20
0
Thermodynamic Properties of Iron-Base Alloys / 53 16. J. Ishida, K. Yamaguchi, S. Sugiura, N. Demukai, and A. Notoh, Denki Seiko, Vol 52, 1981, p 2–8 17. E.T. Turkdogan, Ladle Deoxidation, Desulfurization and Inclusions in Steel—Part I: Fundamentals, Arch. Eisenhu¨ttenwes., Vol 54, 1983, p 1–10 18. E.T. Turkdogan, Slags and Fluxes for Ferrous Ladle Metallurgy, Ironmaking Steelmaking, Vol 12, 1985, p 64–78 19. T.B. Massalski et al., Ed., Binary Alloy Phase Diagrams, Vol 1, American Society for Metals, 1986 20. F. Neumann, The Influence of Additional Elements on the Physico-Chemical Behavior of Carbon in Carbon Saturated Molten Iron, Recent Research on Cast Iron, H.D. Merchant, Ed., Gordon and Breach, 1968, p 659
Fig. 21
21. N.G. Girsovitch, Ed., Spravotchnik po tchugunomu litja (Cast Iron Handbook), Mashinostrojenie, 1978 22. L.S. Darken, The Thermodynamics of Ternary Metallic Solutions, Trans. TMS, Vol 239, 1967, p 80, 90 23. F. Neumann and E. Do¨tsch, Thermodynamics of Fe-C-Si Melts with Particular Emphasis on the Oxidation Behavior of Carbon and Silicon, The Metallurgy of Cast Iron, B. Lux et al., Ed., Georgi Publishing, 1975, p 31 24. C. Wagner, Thermodynamics of Alloys, Addison-Wesley, 1952 25. Metallography, Structures and Phase Diagrams, Vol 8, Metals Handbook, 8th ed., American Society for Metals, 1973, p 400–416 26. E. Piwowarsky, Hochwertiger Gusseisen, 2nd ed., Springer-Verlag, 1958
The iron-carbon diagram with letter notations. S.S., soild solution
27. M. Hillert and P.O. So¨derholm, White and Gray Solidification of the Fe-C-P Eutectic, The Metallurgy of Cast Iron, B. Lux et al., Ed., Georgi Publishing, 1975, p 197 28. A. Kagawa and T. Okamoto, Partition of Alloying Elements on Eutectic Solidification of Cast Iron, The Physical Metallurgy of Cast Iron, H. Fredriksson and M. Hillert, Ed., North-Holland, 1985, p 201 29. H. Laplanche, The Maurer Diagram and Its Evolution and a New Structural Diagram for Cast Iron, Foundry Trade J., No. 1669–1671, 1948, p 191, 225, 249 30. W. Patterson and R. Doepp, Giessereiforschung, Vol 21 (No. 2), 1969, p 91
54 / Principles of Liquid Metal Processing
Fig. 22
Structural diagrams for cast iron. (a) Maurer diagram. (b) Laplanche diagram. See also Fig. 23. Source: Ref 29
Table 7
k lines for various structural ranges
Bar diameter mm
in.
k for mottled structure
k for pearlitic structure
k for pearlitic-ferritic structure
30 20 10
1.2 0.8 0.4
0.65–0.85 0.75–1.10 1.05–1.50
0.85–2.05 1.10–2.25 1.50–2.35
2.05–3.10 2.25–3.40 2.235–3.50
Thermodynamic Properties of Iron-Base Alloys / 55
Fig. 23
Patterson and Doepp structural diagram for cast iron. See the corresponding text for details. Source: Ref 30
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 56-63 DOI: 10.1361/asmhba0005191
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Thermodynamic Properties of Aluminum-Base and Copper-Base Alloys* THIS ARTICLE provides accessible information on the thermodynamic properties of liquid aluminum-base and copper-base alloys. Three alternative means have been used to report the thermodynamic data: Activities and activity coefficients Partial and integral molar thermal properties Interaction coefficients
symbol Gxsi . It represents the contribution to Gi resulting from the departure of solution from ideal behavior. Thus, one may write: xs
Gi ¼ RT ln gi xs Gi
can also be
Gi ¼ H i T S i ¼ RT ln gi
(Eq 4)
Corresponding to Eq 1, expressed as: xs
In addition, the utility of phase diagrams is briefly discussed. Knowledge of activities and activity coefficients is necessary in describing solution behavior and in solving problems that involve chemical equilibria. The thermal properties are useful in understanding the liquid state and in correlating data on solution behavior. The interaction coefficients provide a simple means of calculating activity coefficients in dilute solutions and are also used to correlate experimental data on dilute solutions.
Activities and Thermal Properties The activity a of constituent i in a solution is given by the relationship: Gi ¼ H i T S i ¼ RT ln ai
(Eq 1)
where Gi , H i , and Si are the partial molar free energy, enthalpy, and entropy, respectively, of constituent i in solution relative to the pure constituent i at the solution temperature as the standard state; T is the temperature (in degrees Kelvin); and R is the universal gas constant. Thus, from the knowledge of H i and Si as functions of temperature and solution composition, one may calculate Gi , ai, and gi (activity coefficient of component i relative to pure material standard state) for all conditions of solution. It is important to realize, however, that thermal property values vary with solution composition, unlike the properties of pure substances. Replacing ai in Eq 1 with Xi gi yields: Gi ¼ RT ln Xi þ RT ln gi
(Eq 2)
where Xi is the mole fraction of component i. In Eq 2, the term RT ln gi is called partial molar excess free energy and is expressed by the
(Eq 3)
xs
xs
However, because H i for the ideal solution is 0 by definition, H xsi and are H i are identical. Therefore, Eq 4 can be rewritten as: xs
xs
Gi ¼ H i T S i ¼ RT ln gi
(Eq 5)
Equations 1 and 5 are commonly used to calculate activities and activity coefficients from the thermal properties of the solution. Although the partial molar thermal properties of individual constituents in a solution are most useful in solving chemical process problems, the integral properties are valuable in the discussion of the nature of solutions and also serve in some cases as the source of data on the partial properties. The integral molar free energy is related to the partial molar free energy by the following relationship: Gi ¼ Xi Gi þ X j Gj þ . . . Xn Gn
(Eq 6)
Similar equations can also be written to express other integral thermal properties. The following tables compile, thermodynamic, data in the form of activities, activity coefficients, partial molar thermal properties, and integral molar properties for selected aluminum-base and copperbase alloys. The alloy systems have been chosen based on their commercial importance and on the concentration of the major alloying elements. In general, the concentrations of the alloying elements in the commercial alloys are below 10 at.%, and under this condition the interaction coefficient method described as follows provides a convenient means of calculating the activities of the constituents in solution. When the concentrations are close to or greater than 10 at.%, the interaction coefficient method does not yield accurate values. With this factor in mind, thermodynamic data have
been selected for the following alloy systems: aluminum-magnesium, aluminum-silicon, copperaluminum, copper-nickel, copper-tin, and copper-zinc. Aluminum-magnesium thermodynamic data:
Tables 1 and 2
Aluminum-silicon
thermodynamic data: Tables 3 and 4 Copper-aluminum thermodynamic data: Tables 5 and 6 Copper-nickel thermodynamic data: Tables 7 and 8 Copper-tin thermodynamic data: Tables 9 to 17 Copper-zinc thermodynamic data: Tables 18 and 19
Interaction Coefficients The use of interaction coefficients, first suggested by Wagner (Ref 4), provides a convenient means of organizing thermodynamic data on dilute solutions. Wagner proposed a Taylor series expansion for the excess partial molar free energy in order to express the logarithm of the activity coefficient of a dilute constituent in a multicomponent solution. The values of the activity coefficients calculated from this method are as accurate as the original data from which the interaction coefficients are determined, provided it is applied to solutions that are quite dilute (Xi < 0.1). Consider a dilute solution of i, j, and k dissolved in a common solvent, s. If all but the first-order terms of the Taylor series expansion are neglected, the activity coefficient of solute i is expressed by the relationship: @ ln gi @Xi @ ln gi @ ln gi þ Xk þ Xj @Xj @Xk
ln gi ¼ ln goi ¼ Xi
(Eq 7)
In Eq 7, the partial derivatives are called interaction coefficients and are expressed by the symbol eji , where the superscript denotes the constituent that affects the activity coefficient of the subscript constituent. Thus:
*Updated from B.L. Tiwari, Thermodynamic Properties of Aluminum-Base and Copper-Base Alloys, Casting, Vol 15, ASM Handbook, ASM International, 1988, p 55–60.
Thermodynamic Properties of Aluminum-Base and Copper-Base Alloys / 57 Table 2 Integral thermal properties for liquid aluminum-magnesium alloys at 1073 K
Table 1 Activities, activity coefficients, and partial molar thermal properties for liquid aluminum-magnesium alloys at 1073 K Al component Alð‘Þ ¼ Alðin alloyÞð‘Þ XAl
aAl
0.9997 0.9992 0.9989 0.9945 0.9886 0.9785 0.9553 0.8870 0.8095 0.6600 0.4175 0.2510 0.1792 0.0450
gAl
DGAl , cal/g-atom
DSAl , cal/K-g-atom
0.9 1.7 2.3 12.8 26.6 45.0 93.5 251.6 473.7 934.7 1893.4 2874.9 3839.5 7988.8
0.0091 0.0016 0.0105 0.0203 0.0417 0.0420 0.0958 0.2157 0.3790 0.7806 1.1548 1.2240 0.2920 3.8840
DGAl , cal/g-atom
0.2 0.0 0.0 0.9 1.9 1.3 4.3 4.3 22.7 48.5 30.3 74.0 173.4 1373.6
0.9996 0.9992 0.9989 0.9940 0.9876 0.9791 0.9571 0.8887 0.8008 0.6451 0.4115 0.2597 0.1652 0.236
0.9999 1.0000 1.0000 0.9996 0.9991 1.0006 1.0020 1.0020 0.9894 0.9775 0.9859 1.0353 0.9219 0.5251
9 0 9 9 18 0 9 20 67 97 654 1561 4153 12157
aMg
gMg
DGMg, cal/g-atom
DSMg, cal/K-g-atom
DHMg , cal/g-atom
DGxs Mg, cal/g-atom
0.00022 0.00063 0.00081 0.0048 0.0107 0.0183 0.0366 0.0929 0.1679 0.3097 0.5176 0.6367 0.7184 0.9198
0.8060 0.7974 0.7814 0.8780 0.9353 0.8535 0.8188 0.8221 0.8813 0.9109 0.8885 0.8500 0.8752 0.9631
18,017 15,696 15,171 11,358 9683 8530 7051 5069 3806 2500 1405 963 705 178
15.4180 13.9463 12.6176 9.6420 7.4045 6.9708 5.4069 3.6123 2.7957 1.6331 1.2733 1.3102 1.7577 1.1672
1470.8 729.8 1630.2 1010.9 1736.7 1049.0 1248.7 1192.1 805.9 747.6 38.2 443.2 1180.8 1074.3
459.73 482.71 525.87 227.42 142.53 337.68 424.12 418.20 269.73 198.80 251.88 346.49 284.20 79.98
Mg component Mgð‘Þ ¼ Mg ðin alloyÞð‘Þ XMg 0.00027 0.00079 0.00104 0.0055 0.0114 0.0215 0.0447 0.1130 0.1905 0.3400 0.5825 0.7490 0.8208 0.9550
XMg
DG, cal/g-atom
DS, cal/K-g-atom
DH, cal/g-atom
DGxs, cal/g-atom
0.00027 0.00079 0.00104 0.0055 0.0114 0.0215 0.0447 0.1130 0.1905 0.3400 0.5825 0.7490 0.8208 0.9550
5.7 14.2 18.1 75.7 136.9 227.3 404.8 796.1 1108.7 1466.9 1608.9 1442.6 1266.9 529.7
0.01320 0.01274 0.02365 0.07366 0.12577 0.19086 0.33346 0.59962 0.83951 1.07043 1.22383 1.28856 1.39046 0.93986
8.61 0.58 7.29 3.34 2.04 22.54 47.28 152.48 207.80 318.19 295.28 59.77 225.25 478.87
0.34 0.39 0.55 2.39 3.53 6.00 14.91 43.49 69.79 99.61 159.37 240.99 264.35 138.20
Source: Ref 1
Table 4 Integral thermal properties for liquid aluminum-silicon alloys at 1100 K ð1 xÞAlð‘Þ þ xSið‘Þ ¼ Alð1xÞ Sixð‘Þ XSi
0.1 0.2 0.3 0.31(a)
DG, cal/g-atom
DGxs, cal/g-atom
1210 1931 2403 2430 (þ100)
499 837 1067 1076 (þ100)
(a) Phase boundary. Source: Ref 2
Source: Ref 1
If the solution of i in s obeys Henry’s law over a small range of concentration, then gi is a constant over this range, and eii ¼ 0. It can be shown by using the Gibbs-Duhem equation that:
Table 3 Activities, activity coefficients, and partial molar thermal properties for liquid aluminum-silicon alloys at 1100 K Al component Alð‘Þ ¼ Alðin alloyÞð‘Þ XAl
aAl
1.00 0.90 0.80 0.70 0.69(a)
ð1 xÞAlð‘Þ þ xMgð‘Þ ¼ Alð1xÞ Mgxð‘Þ
xs
DHAl , cal/g-atom
gAl
1.000 0.859 0.698 0.503 0.473 (þ0.01)
xs
DGAl , cal/g-atom
DGAl , cal/g-atom
0 333 786 1500 1638 (þ50)
0 102 298 720 826 (þ50)
1.000 0.954 0.872 0.719 0.685 (þ0.02)
Si component Sið‘Þ ¼ Siðin alloyÞð‘Þ DGSi, cal/g-atom
XSi
aSi
gSi
0.00 0.10 0.20 0.30 0.31(a)
0.000 0.016 0.051 0.127 0.147 (þ 0.02)
0.040 0.155 0.255 0.424 0.473 (þ 0.60)
1 9106 6510 4509 4194 (þ 300)
xs DGSi , cal/g-atom
7060 4073 2991 1877 1634 (þ 300)
DHSi, cal/g-atom 2502 ... ... ... ... (þ 50)
DSSi, cal/ K-g-atom 1 ... ... ... ...
xs DSSi , cal/ K-g-atom
4.144 ... ... ... ...
(a) Phase boundary. Source: Ref 2
@ ln gi Xk ¼ 0; Xi ! 0 @Xi Xj @ ln gi ¼ 0; Xi ; Xj ! 0 eji ¼ @Xj Xk
In Eq 8, goi is the limiting value of gi at infinite dilution:
eji ¼
ln gi ¼ ln goi þ Xj eii þ
Xj eji
þ Xk eki
goi ¼ lim gi (Eq 8)
Xi ! 0
(Eq 9)
eji ¼ eij
(Eq 10)
and this reciprocal relation is most useful in determining values of e from activity data. Values of eii are determined from experimental data on the binary system i-s, and values of eji are determined from experimental data on the ternary system i-j-s. Values of go and e for aluminum-base and copper-base alloys are given in Tables 20 to 24. The standard state for and e and go described previously is the pure material at the temperature of the solution, and the concentration is expressed in mole fraction. In dealing with dilute solutions, however, it is common to use a hypothetical 1 wt% solution as the standard state. Under this condition, the Taylor series expansion corresponding to Eq 8 is: log fi ¼ eii ð%iÞ þ eji ð%jÞ eki ð%kÞ
(Eq 11)
where fi, which is the activity coefficient for 1 wt% standard state, equals ai/%i and: eji ¼
@ log fi @%j %k¼0; %i; %j!0
(Eq 12)
58 / Principles of Liquid Metal Processing Table 5 Activities, activity coefficients, and partial molar thermal properties for liquid copper-aluminum alloys at 1373 K
Table 6 Integral thermal properties for liquid copper-aluminum alloys at 1373 K
Al component Alð‘Þ ¼ Alðin alloyÞð‘Þ XAl
aAl
1.0 0.9 0.8 0.7 0.6 0.5
gAl
1.000 0.889 0.759 0.609 0.441 0.266 (þ0.02) 0.116 0.028 0.006 0.001 0.000
0.4 0.3 0.2 0.1 0.0
1.000 0.988 0.949 0.870 0.735 0.532 (þ0.04) 0.290 0.095 0.029 0.008 0.002
(1x)Al‘ þ xCu‘ ¼ Al(1
DGAl, cal/g-atom
xs DGAl , cal/g-atom
DHAl, cal/g-atom
0 320 753 1354 2235 3611 (þ150) 5873 9709 14,062 19,307 1
0 32 144 381 842 1720 (þ150) 3373 6424 9670 13,025 16,640
0 29 44 60 391 1055 (þ300) 2878 5012 5864 7415 8625
xs DSAl , cal/K-g-atom
DSAl, cal/K-g-atom
0.000 0.254 0.580 0.942 1.343 1.862 (þ0.25) 2.181 3.421 5.971 8.661 1
0.000 0.044 0.137 0.234 0.328 0.484 (þ0.25) 0.361 1.028 2.772 4.086 5.838
XCu
0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9
x)Cux(‘)
DG, cal/ g-atom
DH, cal/ g-atom
DS, cal/ K-g-atom
DGxs, cal/ g-atom
DSxs, cal/ K-g-atom
1723 2969 3955 4700 5170 (þ 150) 5289 4906 3927 2361
459 960 1445 1863 2163 (þ 150) 2254 1979 1380 800
0.921 1.463 1.828 2.066 2.190 (þ 0.1) 2.210 2.132 1.855 1.137
836 1604 2289 2863 3278 (þ 150) 3453 3240 2562 1474
0.275 0.469 0.615 0.728 0.812 (þ 0.1) 0.873 0.918 0.861 0.491
Source: Ref 2
Cu component Cuð‘Þ ¼ Cuðin alloyÞð‘Þ XCu
aCu
gCu
0.0 0.1 0.2 0.3 0.4 0.5
0.000 0.005 0.013 0.025 0.046 0.085 (þ0.005) 0.166 0.352 0.600 0.839 1.000
0.042 0.052 0.065 0.084 0.115 0.170 (þ0.01) 0.277 0.503 0.750 0.932 1.000
0.6 0.7 0.8 0.9 1.0
DGCu, cal/g-atom
xs DGCu , cal/g-atom
DHCu, cal/g-atom
1 14349 11833 10025 8396 6728 (þ150) 4900 2848 1394 478 0
8671 8067 7442 6741 5896 4837 (þ150) 3506 1875 785 191 0
4225 4849 4975 4675 4071 3271 (þ300) 1838 679 259 65 0
DSCu, cal/K-g-atom
xs DSCu , cal/K-g-atom
1 6.919 4.995 3.897 3.150 2.518 (þ0.25) 2.230 1.580 0.827 0.301 0.000
3.238 2.344 1.797 1.505 1.329 1.141 (þ0.25) 1.215 0.871 0.383 0.092 0.000
Source: Ref 2
Table 7 Activities, activity coefficients, and partial molar thermal properties for liquid copper-nickel alloys at 1823 K Cu component Cuð‘Þ ¼ Cuðin alloyÞð‘Þ XNi
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
aCu
1.000 0.902 0.814 0.740 0.677 0.611 (þ 0.02) 0.534 0.444 0.334 0.191 0.000
gCu
1.000 1.002 1.017 1.058 1.128 1.222 (þ 0.03) 1.334 1.480 1.669 1.912 2.227
DGCu, cal/g-atom
xs DGCu ,
0 374 747 1088 1414 1786 (þ 100) 2275 2941 3974 5992 1
cal/g-atom
0 8 61 204 436 725 (þ 100) 1044 1421 1856 2349 2900
Table 8 Integral thermal properties for liquid copper-nickel alloys at 1823 K (1 x)Cu(‘) þ xNi(‘) ¼ Cu(1 XNi
0.1 0.15 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9
DG, cal/ g-atom
948 1195 1377 1617 1745 1786 (þ100) 1742 1604 1349 917
x)Nix(‘)
DH, cal/ g-atom
DS, cal/ K-g-atom
DGxs, cal/ g-atom
DSxs, cal/ K-g-atom
240 341
0.652 0.842
... ... ...
... ... ...
0.006 0.003 ... ... ... ...
... ... ... ...
... ... ... ...
229 336 435 596 693 725 (þ100) 696 609 464 261
... ... ... ...
Source: Ref 2
Ni component Ni(‘) ¼ Ni (in alloy)(‘) aNi
0.000 0.185 0.341 0.455 0.539 0.611 (þ 0.02) 0.682 0.752 0.826 0.907 1.000 Source: Ref 2
gNi
1.906 1.864 1.704 1.517 1.347 1.222 (þ 0.03) 1.136 1.075 1.032 1.008 1.000
DGNi, cal/g-atom
1 6120 3899 2851 2241 1786 (þ 100) 1337 1031 692 353 0
xs DGNi , cal/g-atom
2336 2222 1931 1510 1079 725 (þ 100) 464 261 116 29 0
In Eq 11, fi is the activity coefficient, and the zeroeth-order term, log fio , disappears because the activity coefficient at infinite dilution fio is equal to 1. The reciprocal relationship, corresponding to Eq 10, is: eji ¼
Mi i e Mj j
(Eq 13)
where Mi is the atomic weight of i. The relationship between the two types of the interaction coefficients is: eji ¼
Ms ej ð2:303Þð100ÞMj i
(Eq 14)
where Ms is the atomic weight of the solvent (aluminum, copper).
Thermal Properties for Hypothetical Standard State It is a common practice in solving problems on chemical equilibria to use the hypothetical pure component (Xi) or hypothetical weight percent (%i) as the standard state. For this purpose, thermodynamic properties for the hypothetical standard states are needed. The values of go can be used to determine Gibbs free energies of mixing for elements in liquid base metal, using Eq 15 and 16: Goi ðXi Þ ¼ RT ln goi
(Eq 15)
and Goi ð%iÞ ¼ RT ln
goi Ms 100Mi
(Eq 16)
The values of Goi ðXi Þ and Goi ð%iÞ thus calculated for aluminum-base and copper-base alloys are also given in Tables 20 and 22. The principal advantage of using the hypothetical standard states is that at low concentrations, the activity of i can be set equal to its atom fraction or weight percent, depending on the composition scale used to express the standard state.
Phase Diagrams A phase diagram is essentially a map that shows the relative stabilities of various phases present at equilibrium under the varying conditions of temperature and composition. Therefore, for a given composition of an alloy, a phase diagram can be used to determine the phases, which will be present at equilibrium as the melt
Thermodynamic Properties of Aluminum-Base and Copper-Base Alloys / 59 Table 9 Activities, activity coefficients, and partial molar thermal properties for liquid copper-tin alloys at 1400 K
Table 10 Partial molar thermal properties for liquid copper-tin alloys at 633 K
Cu component Cuð‘Þ ¼ Cuðin alloyÞð‘Þ XCu
1.0 0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0
aCu
gCu
1.000 0.802 0.539 0.389 0.284 0.220 (þ0.025) 0.169 0.125 0.082 0.038 0.000
Sn(‘) ¼ Sn (in alloy)(‘)
DGCu, cal/g-atom
xs DGCu , cal/g-atom
DHCu, cal/g-atom
0 613 1717 2626 3497 4215 (þ300) 4949 5781 6971 9104 1
0 320 1096 1633 2076 2287 (þ300) 2400 2432 2493 2698 3197
0 159 749 1443 1662 1631 (þ250) 1418 1049 640 81 1050
1.000 0.891 0.674 0.556 0.474 0.440 (þ0.05) 0.422 0.417 0.408 0.379 0.317
DSCu, cal/K-g-atom
xs DSCu , cal/K-g-atom
0.000 0.325 0.691 0.845 1.311 1.845 (þ0.3) 2.522 3.380 4.522 6.561 1
0.000 0.115 0.248 0.136 0.295 0.468 (þ0.3) 0.701 0.988 1.324 1.985 3.034
XSn
0.98 0.99
aSn
gSn
0.965 0.985 (þ0.005)
0.985 0.995 (þ0.005)
DGSn, cal/g-atom
47 13 (þ5)
DGxs Sn, cal/g-atom
20 7 (þ5)
Source: Ref 2
Sn component Sn(‘) ¼ Sn (in alloy)(‘) XSn
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
aSn
0.000 0.007 0.072 0.197 0.340 0.467 (þ 0.05) 0.580 0.681 0.784 0.892 1.000
gSn
DGSn, cal/g-atom
DGxs Sn, cal/g-atom
DHSn, cal/g-atom
DSSn, cal/K-g-atom
DSxs Sn, cal/K-g-atom
1 13,706 7304 4523 3003 2119 (þ 300) 1516 1069 678 317 0
13,609 7301 2827 1173 454 190 (þ 300) 95 77 57 24 0
8000 5233 1901 252 706 681 (þ 250) 506 311 176 49 0
1 6.053 3.860 3.411 2.649 2.000 (þ 0.3) 1.444 0.986 0.610 0.261 0.000
4.006 1.477 0.661 1.018 0.828 0.622 (þ 0.3) 0.429 0.277 0.167 0.052 0.000
0.007 0.072 0.362 0.656 0.849 0.934 (þ 0.1) 0.966 0.973 0.979 0.991 1.000
Source: Ref 2
Table 11
Heats of formation of solid and liquid copper-tin alloys at 723 K ð1 xÞCuðsÞ þ xSnðsÞ ¼ Cuð1xÞ SnxðsÞ ð1 xÞCuð‘Þ þ xSnð‘Þ ¼ Cuð1xÞ Snxð‘Þ Cuð‘Þ ¼ Cuðin alloyÞð‘Þ
XSn
Phase
0.091(b)
(Cu)
DH, cal/g-atom
xSn
Phase
DCp(a), cal/K-g-atom
DH Cu , cal/g-atom
280
0.825(b)
‘
1105
1.7
...
...
... ... ...
990 760 530 (þ 50) 0
... 1.2 ...
... ... ...
... ... ...
260
... ...
... ...
1.9 (þ2) ... ...
0.204(b) 0.209
d d
1300 1260
0.850 0.900 0.950
0.244(b)
e
1686
1.000
e
1800 1920 (þ100)
... ...
0.250 0.255(b)
‘
DH, cal/g-atom
... ...
(a) Cp ¼ heat capacity. (b) Phase boundary. Source: Ref 2
Table 12
Integral thermal properties for liquid copper-tin alloys at 1400 K ð1 xÞCuð‘Þ þ xSnð‘Þ ¼ Cuð1xÞ Snxð‘Þ
XSn
0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 Source: Ref 2
DG, cal/g-atom
1922 2834 3195 3300 3167 (þ 300) 2889 2483 1937 1196
DH, cal/g-atom
666 979 934 715 475 (þ 150) 264 97 13 52
DS, cal/K-g-atom
0.897 1.325 1.614 1.846 1.923 (þ 0.24) 1.875 1.705 1.392 0.891
DGxs, cal/g-atom
1018 1442 1495 1427 1238 (þ 300) 1017 784 545 291
DSxs, cal/K-g-atom
0.251 0.331 0.400 0.509 0.545 (þ 0.24) 0.538 0.491 0.398 0.245
... ...
DC pCuð‘Þ , cal/K-g-atom
60 / Principles of Liquid Metal Processing Table 13 Measured activities of tin in liquid copper-tin alloys Temperature, K
Table 15 Recommended partial molar thermodynamic quantities for liquid alloys at 1400 K Cu component Cu‘ ¼ Cu(in alloy)‘
Activity
Composition, mole fraction of Sn
Sn
Cu
XCu
aCu
gCu
0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 0.1 0.2 0.3 0.4 0.5 0.65 0.8 0.9 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9
0.02 0.082 0.22 0.373 0.499 0.614 0.712 0.811 0.905 0.1 0.055 0.198 0.307 0.466 0.652 0.808 0.919 0.007 0.055 0.214 0.353 0.474 0.575 0.683 0.805 0.903
0.856 0.678 0.498 0.373 0.297 0.227 0.17 0.113 0.064 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ...
1.0 0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0
1.000 0.822 0.611 0.448 0.336 0.258 0.197 0.143 0.092 0.045 0.000
XSn
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
1573
1373
1073
DGCu, J/mol
Gex Cu , J/mol
1.000 0.913 0.763 0.640 0.560 0.516 0.492 0.475 0.460 0.450 0.468
0 2282 5744 9356 12,693 15,777 18,929 22,674 27,780 36,100 1
0 1055 3146 5204 6747 7708 8262 8658 9044 9294 8842
aSn
gSn
DGSn, J/mol
Gex Sn , J/mol
0.000 0.013 0.076 0.196 0.336 0.465 0.581 0.690 0.798 0.901 1.000
0.021 0.132 0.378 0.652 0.839 0.930 0.968 0.986 0.997 1.001 1.0000
1 50,370 30,058 19,002 12,716 8913 6326 4316 2633 1213 0
45,239 23,564 11,322 4986 2049 844 380 164 35 13 0
Source: Ref 3
Table 14 Recommended integral thermodynamic quantities for liquid alloys at 1400 K
Table 16 Recommended enthalpies of formation of solid and liquid alloys at 723 K
DG
0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9
7091 10,607 12,250 12,702 12,345 11,368 9824 7662 4702
DH
3360 4374 4010 2996 1838 857 210 79 94
Gex
DS
2.655 4.452 5.886 6.933 7.505 7.508 6.867 5.529 3.426
3307 4782 5139 4868 4277 3534 2713 1837 918
Note: Values in J/mol or J/mol K. Source: Ref 3
solidifies. In addition, a phase diagram can also be used to estimate the activity of the components in liquid solution and to understand the behavior of the liquid solution (Ref 9). For example, if the diagram is for a simple eutectic system, it is likely that no strong intermetallic compound exists. If, however, a strong compound in the solid state appears in the diagram, then association of dissimilar atoms is probably occurring in the liquid. Therefore, one can predict whether a liquid solution will exhibit strong negative or positive deviations from the ideal behavior from the shapes of the liquids and azeotropes. The phase diagrams of the aluminum-silicon and copper-zinc systems, which are chosen based on their commercial importance, are shown in Fig. 1 and 2. The aluminum-silicon diagram (Fig. 1) represents a simple eutectic system, and the alloys containing 8.5 to 13% Si are widely used
Sex Cu , J/mol K
0.000 0.661 1.243 2.123 3.628 5.999 9.360 13.736 19.152 26.317 1
0.000 0.215 0.612 0.843 0.619 0.235 1.742 3.725 5.770 7.172 6.891
Phase
0.090(a)
DH
XSn
(Cu)
1165
0.825(a)
0.205
d
5350
0.85
0.240
e
7650
0.90
0.435
Z
7030(b)
0.95
Phase ‘ ‘ ‘ ‘
49,228 21,394 5854 1526 3930 3704 2455 1176 341 22 0.000
DSSn, J/mol K
Sex Sn , J/mol K
1 20.697 17.289 14.663 11.890 9.012 6.272 3.923 2.124 0.882 0.000
2.849 1.552 3.907 4.652 4.272 3.249 2.025 0.957 0.269 0.006
XSn
Phase
GSn, J/mol
671
0.008 0.012 0.021 0.043 0.069
(Cu) (Cu) (Cu) (Cu) (Cu)
54,350 50,354 45,606 38,618 33,600
515
Data given relative to pure liquid Sn. Source: Ref 3
(1 X) Cu‘ þ XSn‘ ¼ Cu(1x)Snx‘ XSn
DHSn, J/mol
Table 17 Recommended partial Gibbs energy of tin in solid alloys at 1000 K
(1 X) Cu(s) þ XSn(s) ¼ Cu(1x)Snx(s)
Sex
0.038 0.291 0.807 1.337 1.742 1.912 1.788 1.368 0.723
0 1356 4004 6384 7614 7379 5825 3444 967 744 806
DSCu, J/mol K
Sn component Sn‘ ¼ Sn(in alloy)‘
Source: Ref 3
XSn
DHCu, J/mol
DH
263 104
(a) Phase boundary. (b) DH at 237 K. Source: Ref 3
to produce automotive parts. Generally, the higher the silicon content, up to the eutectic composition (12.6% Si), the greater the fluidity and, consequently, the easier an alloy is to cast. The copperzinc diagram (Fig. 2) however, shows a high solubility of zinc in solid copper-zinc solution and the formation of several other phases over a wide range of zinc concentrations. Copper alloys containing up to 35% Zn and significant amounts of other alloying elements are quite common. The copper-zinc phase diagram is used to adjust the compositions in order to produce alloys with the desired properties. REFERENCES 1. B.L. Tiwari, Metall. Trans. A, Vol 18, 1987, p 1645–1651
2. R. Hultgren et al., Selected Values of the Thermodynamic Properties of Binary Alloys, American Society for Metals, 1973 3. P.R. Subramanian, D.J. Chakrabarti, and D. E. Laughlin, Ed., Phase Diagrams of Binary Copper Alloys, ASM International, 1994 4. C. Wagner, Thermodynamics of Alloys, Addison-Wesley, 1952 5. G.K. Sigworth and T.A. Engh, Scand. J. Metall., Vol 11, 1982, p 143–149 6. J.M. Dealy and R.D. Pehlke, Trans. Met. Soc. AIME, Vol 227, Feb 1963, p 88–94 7. G.K. Sigworth and J.F. Elliott, Can. Met. Q., Vol 13, 1974, p 455–461 8. J.M. Dealy and R.D. Pehlke, Trans. Met. Soc. AIME, Vol 227, Aug 1963, p 1030–1032 9. L.S. Darken and R.W. Gurry, Physical Chemistry of Metals, McGraw-Hill, 1953 10. T.B. Massalski et al., Binary Alloy Phase Diagrams, Vol 1 and 2, American Society for Metals, 1986
Thermodynamic Properties of Aluminum-Base and Copper-Base Alloys / 61 Table 18 Activities, activity coefficients, and partial molar thermal properties for liquid copper-zinc alloys at 1200 K Cu component Cu(‘) ¼ Cu (in alloy)(‘) XZn
aCu
0.334(a) 0.4 0.5
gCu
0.432 0.334 0.219 (þ0.05) 0.139 0.085 0.049 0.023 0.000
0.6 0.7 0.8 0.9 1.0
xs DGCu ,
DGCu , cal/g-atom
2003 2614 3621 (þ400) 4705 5869 7187 9010 1
0.648 0.557 0.438 (þ0.1) 0.348 0.284 0.245 0.229 0.235
cal/g-atom
1034 1396 1968 (þ400) 2520 2998 3349 3519 3454
Zn component Zn(‘) ¼ Zn (in alloy)(‘)
aZn 0.132 0.207 0.347 (þ 10.05) 0.505 0.657 0.789 0.900 1.000
xs
gZn
DGZn , cal/g-atom
DGZn , cal/g-atom
0.398 0.517 0.695 (þ 0.1) 0.841 0.939 0.986 1.000 1.000
4814 3757 2521 (þ 400) 1630 1002 564 250 0
2199 1572 868 (þ 400) 412 151 32 1 0
(a) Phase boundary. Source: Ref 2
Table 19
Integral thermal properties for liquid copper-zinc alloys at 1200 K ð1 xÞCuð‘Þ þ xZnð‘Þ ¼ Cuð1xÞ Znxð‘Þ
XZn
DG
0.334(a) 0.4 0.5
2942 3071 3071 (þ 400)
DGxs, cal/g-atom
1423 1466 1418 (þ 400)
xZn
0.6 0.7 0.8 0.9
DG, cal/g-atom
2860 2462 1889 1126
DGxs, cal/g-atom
1255 1005 695 351
(a) Phase boundary. Source: Ref 2
Table 20
Standard Gibbs free energies for solution of elements in liquid aluminum
Solution reaction
Be(s)¼Be Bi(l)¼Bi Ca(l)¼Ca Cd(l)¼Cd Cu(s)¼Cu Fe(s)¼Fe Ga(l)¼Ga Ge(l)¼Ge ½H2 ¼ H In(l)¼In Li(l)¼Li Mg(l)¼Mg Na(l)¼Na Ni(s)¼Ni Pb(l)¼Pb Sb(l)¼Sb Si(s)¼Si Sn(l)¼Sn Zn(l)¼Zn Source: Ref 5
g0, 1100 K
19.6 24.5 0.0086 19.9 0.037 1.6 104 1.1 0.16 ... 12.3 0.40 0.18 293 0.7 106 115 3.4 0.27 4.68 1.92
DGoi for i (X), cal/g-atom
9404 2.639T 5309 þ 1.524T 10,400 7100 0.518T 1135 5.538T 27,000 þ 7.18T 832 0.52T 2761 1.176T ... 6800 1.201T 5800 þ 3.435T 3478 0.30T 8230 þ 3.798T 28,280 2.442T 9970 þ 0.363T 13,100 9.45T 9598 11.291T 5845 2.245T 2538 1.007T
DGoi for i(%), cal/g-atom
9404 9.607T 5309 11.688T 10,400 9.93T 7100 12.498T 1135 16.385T 27,000 3.411T 832 11.551T 2761 12.288T 11,664 þ 6.523T 6800 13.223T 5800 3.014T 3478 9.24T 8230 5.03T 28,280 13.132T 9970 12.832T 13,100 21.589T 9598 20.517T 5845 14.333T 2538 11.911T
Table 21 Interaction coefficients for elements in liquid aluminum i
j
«ji
Temperature, K
Ag Cu H H H H H H H H H Si Cd Cu Ga Ge In Mg Si Sn Zn Zn Li Na Mg Sn H H H H H H H H H H H
Ag H Cu Cu Cu Cu H Si Si Si Si H Cd Cu Ga Ge In Mg Si Sn Zn Si Sn Si Si Pb Ce Cu Cr Fe Mg Mn Ni Th Ti Si Sn
3.1 see eCu H 39.0 16.6 20.1 4.3 0 11.5 6.2 4.2 1.8 see eSiH 5.0 2.2 0.3 þ3.0 4.5 3.0 16.0 6.0 0.9 2.2 16.0 12 9 1.5 100 15 1 1 2 27 19 20 42 7.1 0.6
1273 973–1273 973 1073 1173 1273 973–1273 973 1073 1173 1273 973–1273 1373 1373 1023 1200 1173 1073 1100 1100 1000 963–1053 949 973 973–1073 973–1073 800 700–800 800–900 800–900 700–800 800 700–800 800–900 800–900 700–800 800–900
Source: Ref 5, 6
62 / Principles of Liquid Metal Processing Table 22
Standard Gibbs free energies for solution of elements in liquid copper
Table 24 Activity coefficients at infinite dilution in liquid metals
Temperature Element(a), i
goi , 1200 C (2190 F)
Ag(l) Al(l) As(v) Au(l) Bi(l) C(graphite) Ca(l) Cd(v) Cd(l) Co(s) Cr(s) Fe(s) Fe(l) Ga(l) Ge(l)
3.23 0.0028 4.8 104 0.14 1.25 1.4 105 5.1 104 15.6 0.53 15.4 43 24.1 19.5 0.034 0.009 ... 0.41 0.08 0.044 0.51 0.53 2.22 2.66 ... 5.27 1.3 0.05 ... 0.014 0.002 0.006 0.01 0.048 0.0328 8.5 130 0.146
1=2 H2 ðgÞ
In(l) Mg(v) Mg(l) Mn(l) Mn(s) Ni(l) Ni(s) 1 2 O2 ðgÞ
Pb(l) Pd(s) Pt(s) 1=2 Ss ðgÞ
Sb(l) Se(v) Si(l) Si(s) Sn(l) Te(v) Ti(l) V(s) Zn(l)
DGoi ðXÞ, cal/g-atom
DGoi (wt%), cal/g-atom
3900 0.32T 8630 5.84T 22,350 4630 0.73T 5960 3.6T 8550 þ 17.8T 22,200 25,700 þ 22.9T 1860 8000 11,000 12,970 2.48T 9300 0.41T 10,800 þ 0.61T 16,000 þ 1.52T 10,400 þ 8.4T 9550 þ 4.71T 40,200 þ 22.3T 8670 0.31T 1950 1550 2.31T 2340 6500 2.5T 20,400 þ 10.8T 8620 2.55T 800 10,200 þ 0.87T 28,600 þ 13.79T 12,500 18,200 15,000 2900 7.18T 8900 10,000 6730 0.31T 28,100 9.4T 5640
3900 10.52T 8630 13.84T 22,350 9.44T 4630 12.09T 5960 15.1T 8550 þ 12.0T 22,200 25,700 þ 12.7T 1860 9.0T 8000 9.0T 11,000 8.72T 12,970 11.34T 9300 9.27T 10,800 8.68T 16,000 7.5T 10,400 7.5T 9550 5.58T 40,200 15.1T 8670 7.53T 1950 8.83T 1550 11.14T 2340 9.0T 6550 11.5T 20,400 4.43T 8620 14.01T 800 10.1T 10,200 10.47T 28,600 6.03T 12,500 10.4T 18,200 9.5T 15,000 7.5T 2900 14.68T 8900 10.4T 10,000 10.53T 6730 11.74T 28,100 18.1T 5640 9.1T
C
1100–1200 1100 1000 1175–1325 1100–1300 1100–1300 800–925 ... ... ... ... 1460–1580 1460–1580 1100–1280 1255–1545 1100–1300 700–1000 650–927 650–927 1244 1244 ... ... 1100–1300 1000–1300 1500–1600 ... 1050–1250 1000–1200 1200 1550 1550 1100–1300 1200 1000–1300 ... 1150
(a) l, liquid; v, vapor; s, solid; g, gas. Source: Ref 7
Table 23 Interaction coefficients for elements in liquid copper alloys Temperature i
j
H H H H H H H H H H H H H H H H H O O O O
Ag Al Au Co Cr Fe Mn Ni P Pb Pt S Sb Si Sn Te Zn Ag Au Co Fe
O O O O O O
Ni P Pb Pt S Si
«ij
0.5 6.2 1.9 3.1 1.6 2.9 1.1 5.5 10.0 21.0 8.0 9.0 13.0 4.8 6.0 6.6 6.8 0.7 8.6 68 4.04 106/T 2183 36,000/T þ 17 700,000/T þ 385 7.4 38 19 6300
Source: Ref 5, 6
C
Temperature
F
1225 1225 1225 1150 1550 1150–1550 1150 1150–1240 1150 1100 1225 1150 1150 1150 1100–1300 1150 1150 1100–1200 1200–1550 1200 1200–1350
2235 2235 2235 2100 2820 2100–2820 2100 2100–2260 2100 2010 2235 2100 2100 2100 2010–2370 2100 2100 2010–2190 2190–2820 2190 2190–2460
1200–1300 1150–1300 1100 1200 1206 1250
2190–2370 2100–2370 2010 2190 2203 2280
i
O S S S S S S Ag Al Au Bi Ca Fe Ga Ge H Mg Mn O Pb S Sb Sn Tl Zn Zn Zn Zn Zn
j
Sn Au Co Fe Ni Pt Si Ag Al Au Bi Ca Fe Ga Ge H Mg Mn O Pb S Sb Sn Tl An Zn Zn Zn Zn
«ij
4.6 6.7 4.8 25,400/T þ 8.7 29,800/T þ 13 11.5 6.9 2.5 14 3.7 6800/T þ 1.65 20 5.7 7 13.4 1.0 9.8 6 24,000/T þ 7.8 2.7 20,800/T 15 10 4.8 4 0.38 0.72 1.185 1.40
C
1100 1115–1200 1300–1500 1300–1500 1300–1500 1200–1500 1200 1150 1100 1277 1000–1200 877 1550 1280 1255 1123 927 1244 1100–1300 1200 1050–1250 1000–1200 1300–1320 1300 1150 902 727 653 604
Temperature
F
2010–2190 2010 1830 2150–2415 2010–2370 2010–2370 1470–1700 ... ... ... ... 2660–2880 2660–2880 2010–2340 2290–2815 2010–2370 1290–1830 1200–1700 1200–1700 2271 2271 ... ... 2010–2370 1830–2370 2730–2910 ... 1920–2280 1830–2190 2190 2820 2820 2010–2370 2190 1830–2370 ... 2100
F
2010 2040–2190 2370–2730 2370–2730 2370–2730 2190–2730 2190 2100 2010 2331 1830–2190 1611 2820 2335 2290 2053 1701 2271 2010–2370 2190 1920–2280 1830–2190 2370–2410 2370 2100 1656 1341 1207 1119
Solvent
Solute
Aluminum
Silver
Copper
Magnesium Zinc
Source: Ref 2, 8
g
0.38 0.47 0.53 0.88 0.14 0.17 0.21 0.24
C
700 900 1000 800 604 653 727 802
F
1290 1650 1830 1470 1119 1207 1341 1476
Thermodynamic Properties of Aluminum-Base and Copper-Base Alloys / 63
Silicon, at.% 1500
10
0
20
30
40
50
60
70
80
90
100 2700 1414 ⬚C
2400
1300
1100
2000
900
1700
700 660.452 ⬚C
1300 577 ± 1 ⬚C
12.6
300
(Si)
(Al)
500
0
Al
10
20
30
40
50 Silicon, wt%
Fig. 1
The aluminum-silicon phase diagram. Source: Ref 9
Fig. 2
The copper-zinc phase diagram. Source: Ref 10
60
70
80
90
900
600 100 Si
Temperature, ⬚F
Temperature, ⬚C
L
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 64-73 DOI: 10.1361/asmhba0005192
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Gases in Metals Richard J. Fruehan, Carnegie Mellon University Prince N. Anyalebechi, Grand Valley Sate University
(Eq 1)
where the underlining of aluminum (Al) indicates that it is dissolved in the metal. The free energy for reaction in Eq 1 is highly negative even for very low concentrations of aluminum, indicating that the thermodynamics for the reaction are highly favorable. Similar reactions can be written for silicon, carbon, and iron. For carbon, carbon monoxide is also evolved, according to the reaction: H2 O þ C ! H2 þ CO
(Eq 2)
For the casting of aluminum and its alloys, the reaction is even more favorable. Whether or not the reaction will actually occur depends on many complex factors, including liquid-phase mass transfer and surface tension; therefore, its occurrence is difficult to predict. However, if there is excessive moisture in the sand, the reaction is highly likely to occur. The other major cause of gas porosity is the evolution of dissolved gases during casting. For example, liquid cast iron may have dissolved hydrogen and nitrogen. The solubility of these gases in the solid may be less than that in the liquid, and the gases may therefore be evolved during solidification. Whether or not this will occur depends on the amount of hydrogen and nitrogen present, the alloy being cast, chemical kinetics, and the surface tension of the alloy. Similarly, hydrogen can dissolve in liquid aluminum, but its solubility is much lower in solid aluminum
Cast Iron (Ref 1) Solubility of Hydrogen and Nitrogen. The gases in cast iron that can cause porosity are hydrogen and nitrogen. Castings from iron produced in a cupola are nearly saturated with nitrogen; therefore, the presence of nitrogen is a major concern (Ref 2). Hydrogen and nitrogen dissolve in liquid iron alloys and other metals as atomic species according to: 1
/2 H2 ðgasÞ ! H
(Eq 3)
/ N2 ðgasÞ ! N
(Eq 4)
12
The underlined symbols H and N denote hydrogen dissolved in liquid aluminum and its alloys. The solubilities of both elements in liquid and solid iron alloys have been extensively measured, and the results have been adequately summarized (Ref 3). The solubility of nitrogen and hydrogen in liquid iron is decreased by carbon. The nitrogen and hydrogen contents in iron-carbon alloys in equilibrium with 1 atm pressure of the respective gases are shown in Fig. 1 as a function of carbon content. The solubility of nitrogen at 1823 K decreases from 450 ppm for pure iron to approximately 80 ppm for an alloy containing 4.5 wt% C. The solubilities of hydrogen for the same metals are 24 and 10 ppm, respectively. The hydrogen solubility decreases with temperature and at
eutectic temperature is 6.5 ppm; temperature has only a small effect on the solubility of nitrogen. Silicon decreases the solubility of nitrogen even further. For a Fe-4wt%C-1wt% Si melt, the solubility of nitrogen is approximately 75 ppm. When cast iron solidifies, it forms austenite and graphite or cementite. The solubility of nitrogen in austenite (Ref 3) is also shown in Fig. 1. For austenite containing approximately 1.8 wt% C, the solubility is approximately 150 ppm. Therefore, when the liquid cast iron freezes, the solubility in the solid is higher than that in the liquid; this interesting phenomenon is discussed later in detail. There is some solubility of nitrogen in cementite (Ref 4). The exact amount is not known but is approximately the same as that in austenite. The solubility of hydrogen in austenite is approximately 7 ppm; therefore, the solubility of hydrogen in the liquid and solid are about equal, and there is little segregation of hydrogen during solidification (Ref 3). Kinetics of Gas-Liquid Reactions. The absorption of hydrogen and nitrogen is controlled by one of the following steps:
500
50
400
40 Hydrogen in liquid
300
30 Nitrogen in liquid
200
20 Nitrogen in austenite
100
0
Fig. 1
Hydrogen, ppm
2Al þ 3H2 O ! Al2 O3 þ 3H2
and can cause porosity. For copper, the evolution of hydrogen, water vapor, or carbon monoxide could cause gas porosity. In this article, the solubilities of the common gases present in ferrous metals, such as cast irons, and nonferrous metals, such as aluminum, copper, and magnesium, are reviewed. The kinetics of the relevant reactions, reactions during solidification, and possible methods of control or removal of the dissolved gases are discussed. The discussion is primarily focused on cast iron because more is known about the thermodynamics and kinetics of iron than other elements. However, th e same basic principles also apply to aluminum, copper, and magnesium.
Nitrogen, ppm
GAS POROSITY is one of the most serious problems in the casting of both ferrous and nonferrous metals. It is generally caused by the evolution of gases during the casting and solidification process. The gases may be the result of a reaction between the casting sand mold and the metal, or they may result from the evolution of gases dissolved in the liquid metal during solidification. An example of a chemical reaction evolving gas and causing porosity is the reaction of moisture (H2O) in a sand mold with elements in the cast iron, such as carbon, silicon, aluminum, or iron itself. For example, aluminum in cast iron often causes porosity problems, and the reaction responsible for the gas evolution is:
10 0
0
1
2 3 Carbon, %
4
5
Effect of carbon on the solubility of hydrogen and nitrogen in the iron-carbon alloys at 1823 K (i.e., liquid) and 1500 K (i.e., austenite) and 1 atm pressure
Gases in Metals / 65
(Eq 5)
where km is the mass transfer coefficient, pBN2 is the pressure of nitrogen in the bulk gas, and pSN2 is the surface pressure of nitrogen in equilibrium with the melt where: pSN2 ¼ KN2 fN2 ðwt%NÞ2
(Eq 6)
where KN is the equilibrium constant for Eq 4, and fN is the activity coefficient of nitrogen with respect to 1 wt%. However, in general, gas diffusion for these reactions is faster than the chemical reaction or liquid-phase diffusion and can be neglected in most cases. The rate of the chemical reaction is controlled by the dissociation of the nitrogen molecule on the surface, and the rate is given by: dnN2 ¼ kB Að1 yÞ pN2 peN2 dt
(Eq 7)
where kB is the rate constant for pure iron, A is the surface area, peN2 is the equilibrium nitrogen pressure given by an expression similar to Eq 6, and (1 y) is the fraction of vacant sites not occupied by surface-active elements. Certain elements are surface active on liquid iron in that they lower the surface tension of iron by covering most of the surface. For example, oxygen and sulfur are surface active on iron, and for bulk concentrations of 0.03% O or 0.03% S, over 90% of the surface sites will be covered by oxygen or sulfur. These elements therefore retard the rate of chemical reaction. At high coverage, the surface-active element (1 y) is inversely proportional to the activity, a, of the element. Therefore, the rate is given by: dnN2 kB kA A ¼ pN2 peN2 dt a
(Eq 9)
where kmN is the liquid-phase mass transfer coefficient of nitrogen, CNB is the bulk concentration of nitrogen, and CNS is the surface concentration of nitrogen. The integrated form of Eq 7 is: %N %Ne ArkmN t ¼ ln %No %Ne W
(Eq 10)
where r is the density of iron, W is the weight of the metal, %Ne is the equilibrium of the nitrogen contents, and %No is the initial nitrogen contents. The rate is often controlled by two processes in series—usually the chemical reaction and liquidphase mass transfer. The rate in this case can be obtained by equating the fluxes given by Eq 8 and 9 and solving for the surface concentration, as indicated by Pehlke and Elliot (Ref 5). 3.0 2.5
1.0
2.0 1.5
1.0 1450⬚C 1600⬚C
0.5
0.75 0.5 0.25 0
0
100
0 0
10
20
30
40
50
Fig. 2
Fig. 4
Effect of sulfur in the rate of reaction of nitrogen with Fe-CSAT-S alloys
1.0 0.75 0.5 0.25 0
Fig. 3
0
2.5
5 1/aBi
7.5
300
400
500
600
Effect of tellurium on the reaction of nitrogen with Fe-CSAT-Te alloys at 1450 C (2640 F)
1.5
1.0
0.5 as = 0.11 Present study Ref 9 0
10
Effect of bismuth on the reaction of nitrogen with Fe-CSAT-Bi alloys at 1450 C (2640 F)
200
1/[% Te]
60
1/as
(Eq 8)
where kA is related to the adsorption coefficient of the element on the surface, and the quantity kBkA/a is the overall rate constant k.
Rate constant, mol/cm2 • s • atm) ⫻ 106
JN2
km S p pBN2 ¼ RT N2
JN ¼ kmN CNB CNS
The rate of the nitrogen reaction with iron alloys containing oxygen, sulfur, chromium, and other elements has been measured by many investigators, and there is good agreement in most cases (Ref 5–10). However, the rate for carbon-saturated iron containing other elements was not investigated until recently (Ref 11). In the study, an isotope exchange technique was used to measure the rate of dissociation of the nitrogen molecule (N2). In particular, the effects of carbon, sulfur, phosphorus, lead, tin, bismuth, and tellurium on the rate were investigated. Sulfur, phosphorus, lead, bismuth, and tellurium decreased the rate, while carbon and tin had no significant effect. For example, the effect of sulfur is shown in Fig. 2 as the rate constant versus the reciprocal of the activity of sulfur (Ref 10). The activity of sulfur is in weight percent, and the activity coefficient of sulfur in carbonsaturated iron is 6.3. The rate for carbon-saturated iron with no sulfur is approximately 105 mol/ cm2 s atm at 1450 C (2640 F). For example, for 0.009 wt% S (1/as = 18), the rate is decreased to 8 107 mol/cm2 s atm. Bismuth and tellurium had an even larger effect (Ref 11), as shown in Fig. 3 and 4. The activities are relative to pure bismuth and tellurium, respectively, which have larger deviations from ideal behavior. As little as 50 ppm Te reduced the rate by 90%, and 50 ppm Bi decreased it by 80%. On the other hand, carbon is not surface active and does not affect the rate at moderate sulfur contents (0.017 wt% S) up to 4.5 wt% C (Ref 11), as shown in Fig. 5. Therefore, the
Rate constant, mol/(cm2 · s · atm) ⫻ 106
For the desorption of a gas, the same three steps are important but occur in reverse order. The process can be controlled by two or more of the steps in series. For liquid cast iron, nitrogen is the more important of the gases, and it is considered in detail. Gas diffusion is usually not rate controlling for nitrogen absorption from the atmosphere because of the pressure of N2; consequently, the driving force for diffusion is high. As a result, one of the other steps is slower and rate controlling. However, for nitrogen or hydrogen removal by an inert gas, diffusion on N2 or H2 must be considered. The flux of nitrogen,JN2 , away from the surface is given by:
Rate constant, mol/(cm2 ·s · atm) ⫻ 106
from the surface
For dilute solutions, the activity of the surfaceactive element is proportional to its concentration. Therefore, the overall rate constant is inversely proportional to the weight percent of the surface-active element in iron. Liquid-Phase Mass Transfer. If the rate is controlled by liquid-phase mass transfer, the flux of nitrogen atoms is given by:
Rate constant, mol/cm2 ·s· atm) ⫻ 106
Diffusion of the gas to the surface Chemical reaction on the surface Diffusion of the element in the liquid away
Fig. 5
0
1
2 3 4 Carbon content, wt%
5
6
Effect of carbon on the reaction of nitrogen with Fe-C-S alloys at constant sulfur activity
66 / Principles of Liquid Metal Processing pT ¼ KH2 fH2 ð%HÞ2 þKN2 fN2 ð%NÞ2
(Eq 11)
where Ki is the equilibrium constant for the gas reaction, and fi is the activity coefficient. At the end of solidification, the concentrations will be as indicated in Eq 11. Even a small amount of hydrogen may be important. For example, in the case of an Fe-3.8wt%C alloy containing only 4 ppm H, the hydrogen content in the liquid will increase slightly, because of the enrichment, to approximately 4.5 ppm just prior to the eutectic reaction. However, even for this small amount of hydrogen, its thermodynamic pressure is 0.4 atm. Therefore, if the thermodynamic pressure of nitrogen exceeds 0.6 atm, a gas pinhole may result. Calculations indicate that iron with as little as 80 ppm N and 4 ppm H may have a gas pressure exceeding 1 atm. Figure 6 shows a plot of the total pressure of hydrogen and nitrogen just prior to final solidification equal to 1 atm as a function of bulk nitrogen and hydrogen contents in an Fe-3.8C alloy. If the hydrogen and nitrogen contents are such that they are below the line, the pressure will not exceed 1 atm, and a pinhole due to gas evolution will not occur; if above the line and the pressure exceeds 1 atm, a pinhole from this source is possible. For pinholes to develop, it is also necessary to nucleate a bubble. Bubble nucleation is a complex phenomenon and is influenced by surface tension. For melts with low surface tension, bubble nucleation is favored. On the other hand, elements that reduce the surface tension also reduce the rate of N2 formation. Therefore, it is difficult to predict the net effect on the probability of bubble formation of an alloying element that reduces the surface tension of iron. Nitrogen and Hydrogen Removal by Inert Gas Flushing. The preceding discussion indicates that nitrogen and hydrogen evolution during solidification may be a cause of pinholes in castings. Therefore, it would be desirable to 120
remove a portion of these gases if possible. One method is by argon bubbling in the desulfurization reactor or ladle. The desulfurization reactor is a continuous reactor with metal flowing in and out almost continuously. However, for the present calculation, it is assumed to be a batch reactor. Because this is only an order of magnitude calculation, this assumption is not unreasonable. When argon is bubbled through the melt, nitrogen atoms combine on the bubble surface to form N2. The rate is controlled by the chemical reaction on the surface and the liquid-phase mass transfer of nitrogen to the surface in series. For the purpose of the present calculations, the following were assumed: 10 Mg (11 ton) reactor 1 m (3.3 ft) deep Argon flow rate of 0.005 m3/s (10 scfm)
through the melt
Initial nitrogen content of 100 ppm 60 mm (2.4 in.) diameter bubbles
The rate constant for iron containing a relatively small amount of sulfur is approximately 1.5 106 mol/cm2 s atm, and the mass transfer coefficient, km, is approximately 0.1 cm/s. The gas velocity, retention time, and surface area were estimated as done previously (Ref 13). Equations 8 and 9 were solved simultaneously, and the results are given in Fig. 7. The calculations demonstrate that the rate is truly mixed control, as indicated by the surface concentration being between zero and the bulk concentration. The results indicate that it would take approximately 40 min to remove 20 ppm; if the argon flow were doubled, it would still take over 20 min to remove 20 ppm. Although these calculations are rather crude, they indicate that it would be difficult to remove nitrogen by argon bubbling. For hydrogen, the chemical reaction and mass transfer are considerably faster than for nitrogen. If equilibrium between the metal and the gas bubbles leaving the system is assumed, it is possible to calculate the amount of hydrogen that can be removed by gas bubbling from thermodynamic considerations alone. The
100 0.012 80 0.010
p N + pH = 1 2
2
60
Nitrogen, %
Nitrogen, ppm
initial rate of nitrogen formation is primarily controlled by chemical kinetics at the surface, and surface-active elements, such as sulfur and tellurium, can reduce the rate. This information can be helpful in controlling the rate of the reaction. For example, if it is desired to remove nitrogen by argon gas flushing, the concentrations of these elements should be as low as possible. On the other hand, it may be possible to retard nitrogen evolution during solidification by the deliberate addition of these elements. Reactions During Solidification. During the solidification of most simple iron alloys, there is enrichment of the alloying element because there is greater solubility in the liquid than in the solid. The difference in the solubilities is expressed by the partition ratio kS/L, which is determined from the slopes of the solidus to liquidus lines on the phase diagram. For most simple alloys, the partition ratio is less than 1, resulting in liquid enrichment during solidification. For example, for an Fe-0.02wt% N alloy, the last portion of liquid to solidify will be enriched and will have a concentration of 0.045 wt% N, which is the solubility of nitrogen in liquid iron at 1 atm (Ref 12). Therefore, even though the liquid alloy is far from being saturated with nitrogen, during solidification, the concentration will increase to the point where the solubility is exceeded and nitrogen gas will be evolved. Researchers have measured the partition ratio for nitrogen in high-carbon iron alloys (Ref 4). The partition ratio was found to be 1.9 and 2.2 for stable and metastable eutectic solidification, indicating that there is greater solubility in the solid than in the liquid. Therefore, in this case, as the alloy solidifies, nitrogen is actually enriched in the solid, and the nitrogen content decreases in the liquid. It may appear that, because there is no enrichment during solidification, the thermodynamic pressure of nitrogen cannot increase. The thermodynamic pressure is defined by an expression similar to Eq 6. Carbon greatly increases the activity coefficient of nitrogen in the liquid and decreases its solubility; therefore, the thermodynamic pressure may increase because the carbon content of the liquid is increasing during solidification. For example, for an alloy containing 3.8 wt% C that is saturated at 1500 C (2730 F) with 110 ppm N just prior to eutectic solidification, the concentration in the liquid will decrease to 97.5 ppm because of enrichment in the austenite. However, the solubility of nitrogen in the liquid has decreased to 90 ppm primarily because of the increase in the carbon content. Consequently, the thermodynamic pressure of nitrogen exceeds 1 atm, and the evolution of nitrogen may cause a pinhole. When determining the possibility of pinhole formation, one must consider the total thermodynamic pressure of all the gases, and if the total pressure, pT, exceeds 1 atm, a gas pinhole may result. For example, for hydrogen and nitrogen, the total pressure is given by:
40
0.008
0.006
20
0 0
Fig. 6
2
6 4 Hydrogen, ppm
8
10
Thermodynamic pressure of nitrogen and hydrogen at the end of solidification equal to 1 atm for an Fe-3.8wt%C alloy
0.004 0
Fig. 7
20
60 40 Time, min
80
100
Rate of removal of nitrogen from an Fe-3.8wt% C alloy with argon bubbling at 0.005 m3/s (10 scfm) in a 10 Mg (11 ton) reactor
Gases in Metals / 67 equations are developed similarly to those for steel (Ref 12), taking the form: 1 1 ¼ kH V ½H ½Ho
(Eq 12)
where [H] is the hydrogen content after bubbling, [Ho] is the initial hydrogen content, and V is the total volume of gas used per ton of metal. The constant kH is related to the solubility of hydrogen and therefore depends on the alloy composition. For example, 10 Mg (11 tons) of an Fe-4.5wt%C melt at 1773 K containing 4 ppm H bubbling argon for 20 min at 0.005 m3/s (10 scfm) will reduce the hydrogen content to 1.6 ppm. The previous calculation assumes equilibrium and therefore the fastest rate possible and may be an overestimation of the amount of hydrogen removed. Generally, the efficiency of the flushing gas would be roughly 50% for the conditions encountered in practice. However, the calculation does indicate that it is possible to remove significant amounts of hydrogen by gas bubbling.
Aluminum and Its Alloys Compared to other nonferrous metals, aluminum is the most susceptible to hydrogeninduced defects. This has been attributed to two primary reasons. The first reason is that molten aluminum and its alloys readily absorb hydrogen by the reaction with water vapor: 2AlðlÞ þ 3H2 OðvÞ ! Al2 O3 ðsÞ þ 3H2 ðin AlÞ (Eq 13)
The standard free energy change, DGo, for this reaction is extremely high: Go ¼ 979; 100 719T log10 T þ 413T
(Eq 14)
where T is the temperature in K. Inserting this value of free energy change into the van’t Hoff reaction isotherm yields an activity coefficient that is exceedingly high and, for practical purposes, implies complete conversion to hydrogen of all traces of moisture contacting the metal. That is: Go ¼ RT ln Kp
(Eq 15)
where DGo is the standard free energy change, T is the temperature in K, R is the gas constant, and Kp is the activity coefficient. The activity coefficient is given by the equation: Kp ¼
pH2 aH2 O
(Eq 16)
where pH2 is the partial pressure of hydrogen, and aH2 O is the activity of the water vapor. The second reason is that, unlike other nonferrous metals, the partition coefficient, k, that is, the ratio of hydrogen solubility in the solid state to that in the liquid state, is relatively small. That is, the solubility of hydrogen in liquid aluminum is more than an order of magnitude greater than
that in the solid. The first reason is not unique since, for a similar reaction for magnesium and lithium and their alloys, the free energy change is even higher. This explains the greater propensity for hydrogen absorption and the high solubility of hydrogen in magnesium, lithium, and their alloys. However, the low partition coefficient is unique to aluminum and its alloys and partly explains why, despite the comparatively lower solubility of hydrogen in aluminum, it is still the most susceptible of the nonferrous metals to hydrogen-induced defects. Hydrogen Solubility and Reactions During Solidification. Hydrogen solubility is often a direct indication of the propensity for porosity formation and its evolution during solidification of a casting. It indicates the nature of portioning of hydrogen in both cast and wrought products and the effects of alloying elements on the susceptibility of aluminum alloys to hydrogen-induced defects. It is generally agreed that during solidification, hydrogen in the melt is rejected at the solid-liquid interface (at the liquid side) in the same manner in which alloying elements are rejected from the solid in alloys having an equilibrium partition coefficient k less than unity: k¼
Ss <1 Sl
(Eq 17)
where k is the equilibrium partition coefficient for hydrogen in aluminum, and Ss and Sl are the solubility of hydrogen in the solid and liquid, respectively (in the case of alloys, the solubility of hydrogen at the solidus/eutectic and liquidus temperatures, respectively). The solubility of hydrogen in both the liquid and solid states of aluminum has been measured by several researchers. Of the more than 26 sets of results (Ref 14) reported on the liquid state of the aluminum-hydrogen system, only 9 are assessed to be reliable (Ref 15–25). Combined regression analysis of these reliable sets of results yields the following recommended empirical equation for hydrogen solubility as a function of temperature, T (K), at 1 atm hydrogen partial pressure (Ref 26): Liquid aluminum-hydrogen : 2692 log10 S; cm3=100 g ¼ þ 2:7292 T
Of the 14 sets of results reported on the solubility of hydrogen in solid pure aluminum, only 7 are assessed to be reliable (Ref 14, 15, 26–33). Combined linear regression analysis of these results yielded the following empirical equation for the solubility of hydrogen in pure sold aluminum as a function of temperature at 1 atm hydrogen partial pressure: Solid aluminum-hydrogen: 2965:8 þ 1:8032 log10 S; cm3 =100 g ¼ T
where S is the hydrogen solubility in cm3/100 g, and T is the temperature in K. Hydrogen solubility in both liquid and solid pure aluminum follows Sievert’s law. That is, the solubility of hydrogen is directly proportional to the square root of the hydrogen pressure. Equations 18 and 19 yield values for the solubility of hydrogen in liquid and solid pure aluminum of 0.70 and 0.042 cm3/100 g, respectively, at the melting point, 933 K. This confirms the fact that the solubility of hydrogen in liquid pure aluminum is more than 16 times the solubility of hydrogen in solid aluminum, and the equilibrium partition coefficient of hydrogen in aluminum is significantly less than 1. Both recommended sets of results of hydrogen solubility in liquid and solid pure aluminum are presented graphically in Fig. 8. Alloying elements affect the solubility of hydrogen in aluminum alloys. Unfortunately, there are a limited number of published data on the solubility of hydrogen in aluminum alloys. This is especially the case for multicomponent and commercial aluminum alloys. However, alloying elements have a strong and varying influence on the solubility of hydrogen in aluminum. For example, in liquid aluminum, silicon, copper, zinc, and iron decrease hydrogen solubility, whereas lithium, magnesium, and titanium increase it (Fig. 9). The effects of temperature on the solubility of hydrogen in liquid binary aluminum alloys can be represented by the general equation of the same form as Eq 1 and 2: log10 S; cm3=100 g ¼
(Eq 18) 3
where S is the hydrogen solubility in cm /100 g, and T is the temperature in K. Unlike in the liquid state, published values of hydrogen solubility in solid pure aluminum are limited (Ref 14). This is because the solubility of hydrogen in the solid state of pure aluminum is significantly lower than that in the liquid state. Consequently, determination of the solubility of hydrogen in solid aluminum is a more formidable experimental task than the determination of the solubility of hydrogen in the liquid state. This problem is compounded by the fact that diffusion of hydrogen into aluminum in the solid state is several orders of magnitude slower than that into the liquid.
(Eq 19)
A þB T
(Eq 20)
where S is the hydrogen solubility in cm3/100 g, T is the temperature in K, and A and B are the solubility constants or coefficients that depend on alloy composition. The coefficients for the effects of temperature on the solubility of hydrogen in some liquid binary aluminum alloys, obtained from published data assessed to be reliable, are given in Table 1. In solid aluminum, published results show that silicon, magnesium, copper, lithium, and zinc increase the solubility of hydrogen (Fig. 10). The addition of lithium increases the solubility of hydrogen in solid aluminum by 1 to 2 orders of magnitude and that in liquid aluminum by two- to threefolds. Consequently, the partition coefficient of hydrogen in aluminum-lithium
68 / Principles of Liquid Metal Processing alloys is significantly greater than that in pure aluminum. This explains why aluminum-lithium alloys are less susceptible to gas porosity than other aluminum alloys. This is despite the comparatively higher hydrogen concentration in
lithium-containing aluminum alloys. Figure 11 shows the comparatively higher hydrogen solubility in the lithium-containing aluminum alloy 2090 than in other conventional aluminum alloys.
Reactions During Solidification of Aluminum and Its Alloys. Because the solubility of hydrogen is significantly higher in the liquid aluminum as compared with the solid, there will be enrichment of the liquid during solidification. Assuming no solid diffusion and complete liquid diffusion, at the end of solidification there will be a large increase in the hydrogen concentration. Therefore, even when the hydrogen content is far below the solubility for the bulk liquid during solidification, the concentration in the enriched liquid will increase, and the solubility limit may be exceeded. For example, consider pure aluminum containing a hydrogen concentration of 0.45 cm3/100 g at 750 C (1380 F). By the time the alloy is 90% solidified, the concentration in the remaining liquid will be approximately 4.03 cm3/100 g H, which will exceed the hydrogen solubility for 1 atm, resulting in the formation of hydrogen bubbles. Inert Gas Flushing. It is possible to remove hydrogen from liquid aluminum by inert gas (argon or nitrogen) flushing. As discussed pre viously for cast iron, it is possible to calculate the fastest possible rate, or the minimum amount of purge gas needed to remove hydrogen. The general form of the equation is: 1 1 kH Vt ¼ ½H ½Ho W
Fig. 8
where [H] is the hydrogen content in ppm at time t, [Ho] is the initial hydrogen content, W is the weight (in kilograms), t is the time (in minutes), kH is a constant related to the solubility, and V is the flow rate (in m s1). In practical units, at 750 C (1380 F):
Hydrogen solubility in liquid and solid pure aluminum at 1 atm hydrogen partial pressure
Temperature, ⬚F 1290 3.5
1380
1470
1560
1650
1740
Table 1 Empirical constants for the calculation of hydrogen solubility (in cm3/ 100 g) in the liquid state of some binary aluminum alloys at 1 atm hydrogen partial pressure
3.0
Hydrogen solubility, cm3/100 g
3.2 wt% Mg
4 wt% Cu Pure AI
2.5
(Eq 21)
Solubility constants, log10 S; cm3 =100 g ¼ A T þB A
B
Al-Mg alloys Al-3.2wt%Mg Al-6wt%Mg
2735.6 2695.5
2.9752 3.1452
Al-Zn alloys Al-1.8wt%Zn Al-3.5wt%Zn Al-6wt%Zn
3193.9 3058.1 2957.7
3.1392 3.0382 2.8602
Al-Si alloys Al-2wt%Si Al-4wt%Si Al-7wt%Si
2777.8 2941.3 2731.0
2.7702 2.9032 2.6562
Al-Cu alloys Al-2wt%Cu Al-4wt%Cu
2955.3 3047.7
2.9060 2.9392
Al-Li alloys Al-1wt%Li Al-2wt%Li Al-3wt%Li
2141.4 2850.7 2869.7
2.5952 3.3832 3.4862
Alloy
2.0
1.5 4.4 wt% Zn
1.0 7 wt% Si 0.5 700
Fig. 9
750
800 850 Temperature, ⬚C
900
950
Hydrogen solubility in liquid pure aluminum and binary aluminum alloys at 1 atm hydrogen partial pressure
Source: Ref 14
Temperature, ⬚F 570 0.10
660
750
840
930
1020
1110
Hydrogen solubility, cm3/100 g
0.08
0.06 7 wt% Si 5.3 wt% Mg 0.04
0.00 300
300
200
100
0
0
350
400
450 Temperature, ⬚C
500
550
1380
1470
Temperature, ⬚F 1560
1650
1740
600
1830
14.0 AA2090 12.0
10.0
ln 6.0
Pure Al AA7050
2.0 Duralumin B 0.0 700
750
800
850 Temperature, ⬚C
900
950
Hydrogen solubility in liquid multicomponent aluminum alloys at 1 atm hydrogen partial pressure
0.2
0.3
0.4
Gas removal ratio for equilibrium in aluminum at 760 C (1400 F)
1000
kg m3ðstpÞ
ðppmÞ s
(Eq 22)
For example, for 10 Mg (2.2 104 lb) of aluminum, a gas purge rate of 0.01 m3/s (0.35 ft3/s) for 100 s could reduce the hydrogen from 0.2 to 0.09 ppm. As the hydrogen content decreases, it becomes increasingly difficult to remove hydrogen (Ref 34), as indicated in Fig. 12. As the hydrogen content decreases, the ratio R of the purge gas to the hydrogen gas removed increases from approximately 15 at 0.4 ppm to over 500 at 0.1 ppm (1 cm3/100 g = 0.9 ppm). It should be emphasized that this calculation indicates the theoretical minimum amount of purge gas required. Due to kinetic factors, the amount of purge gas can be significantly higher. There are two limiting cases with respect to hydrogen removal: the thermodynamic limit discussed previously and the diffusion in the liquid film boundary layer surrounding the gas bubble. For the limiting case of diffusion control, the rate equations are similar to those discussed previously for nitrogen in iron, and the rate can be expressed by:
8.0
4.0
0.1
kH ¼ 62; 000
Hydrogen solubility in solid pure aluminum and binary aluminum alloys at 1 atm hydrogen partial pressure
1290
Fig. 11
400
Fig. 12 4 wt% Cu
Fig. 10
500
Hydrogen dissolved in aluminum (H), cm3/100 g
Pure Al 0.02
Hydrogen solubility, cm3/100 g
Gas removal ratio (R ), (standard ft3 N2)/(standard ft3 H2)
Gases in Metals / 69
%H km At ¼ H %Hi W
(Eq 23)
where %Hi is the initial hydrogen content, r is the liquid density, kmH is the mass transfer coefficient for hydrogen, A is the bubble surface area, and W is the weight of the melt. Both A and kmH increase with decreasing bubble size; consequently, the rate is sensitive to bubble size, with hydrogen removal favored by small bubbles. In general, the rate will be controlled by both diffusion and the thermodynamic limit in series. In an analysis of the case of mixed control, it was found that the purging efficiency depends on the bubble size and hydrogen content (Ref 35), as shown in Fig. 13. Purging efficiency is the ratio of the actual amount of gas required to the theoretical minimums given by Eq 23. Both
70 / Principles of Liquid Metal Processing Temperature, ⬚F
100
1110 10.0
1290
1470
1650
1830
2010
2190
2370
2550
0.05 cm3/100 g
Pure Cu
50
8.0
25 0.1 cm3/100 g 0
2 1 Bubble diameter (d B), cm
0
Fig. 13
3
Purging gas efficiency as a function of bubble size for 0.05 and 0.1 cm3/100 g of hydrogen
the interfacial area and the mass transfer coefficient increase with decreasing bubble size; consequently, the rate of hydrogen removal is enhanced by having small bubbles. The use of porous plugs helps improve the rate by providing small bubbles, but the bubbles generally coalesce in the bath, reducing the beneficial effect. The use of an impeller with a porous plug improves the rate further by dispersing the bubbles. It has also been found that the addition of chlorine (Ref 36) or Freon (E.I. Du Pont de Nemours, Inc.) (Ref 37) improves the rate in some cases. These halogen gases increase the mass transfer coefficient km and the rate. However, if the rate is being primarily controlled by the thermodynamic limit, the addition of these gases will have little effect. It is also possible that the halogen is actually taking part in a hydrogen removal reaction: 2H þ Cl2 ¼ HCl
(Eq 24)
Equation 24, under some circumstances, is thermodynamically favorable. This will increase the amount of hydrogen that can be present in the inert gas bubble; that is, hydrogen can be present as H2 or hydrogen chloride gas, thus increasing the rate of removal.
Copper and Its Alloys As in aluminum and its alloys, gas porosity is a problem in the casting of copper and copper alloys. Hydrogen is the most important dissolved gas in copper and its alloys. It forms dilute solutions in oxygen-free copper. However, in impure copper there are complex interactions among the gas-forming elements hydrogen, oxygen, and sulfur. Thus, in addition to hydrogen, gaseous compounds such as water vapor, carbon monoxide, and sulfur dioxide can also evolve during solidification. Hydrogen Solubility and Reactions. The solubility of hydrogen in pure copper has been experimentally determined by various investigators (Ref 38–47). Combined linear regression
Hydrogen solubility, cm3/100 g
Efficiency, %
75
6.0 5.9 wt% Sn
4.0 11.5 wt% Sn 21.7 wt% Sn 54.8 wt% Sn
2.0
40.2 wt% Sn
0.0 600
Fig. 14
700
800
900
1000 1100 Temperature, ⬚C
1200
1300
1400
Effect of temperature on the solubility of hydrogen in pure copper and copper-tin alloys at 1 atm hydrogen partial pressure
analyses of the sets of reported data for hydrogen solubility in liquid and solid pure copper are given by the following equations: Liquid copper-hydrogen: 2327:59 þ 2:4649 log10 S; cm3=100 g ¼ T Solid copper-hydrogen: 2057:76 þ 1:6829 log10 S; cm3=100 g ¼ T
(Eq 25)
(Eq 26)
Figure 14 is a graphical representation of the effect of temperature on the solubility of hydrogen in copper and copper-tin alloys. The difference in solubility between the liquid and solid pure copper is not as marked as for pure aluminum. At the melting point, 1083 C (1981 F), the concentrations of hydrogen in equilibrium with the gas at atmospheric pressure are 5.6 and 1.46 cm3/100 g for the liquid and solid, respectively. Thus, the partition coefficient at 0.26 for hydrogen in pure copper is comparatively higher than that in pure aluminum, where the partition coefficient is extremely low at 0.04. Hydrogen solubility in both liquid and solid pure copper is directly proportional to the square root of the hydrogen partial pressure, in accordance with Sievert’s law (Ref 38, 45). Published results on the effect of alloying elements on the solubility of hydrogen in copper are sparse. Nickel increases the solubility of
hydrogen in liquid pure copper. For example, at 1200 C (2190 F), the solubility of hydrogen in Cu-10wt%Ni is 11.2 cm3/100 g compared to 7.7 cm3/100 g in pure copper. However, large additions of aluminum, zinc, and tin decrease the solubility of hydrogen in pure copper. For example, at 1200 C (2190 F), the solubility of hydrogen in Cu-5.9wt%Sn is 6.2 cm3/100 g compared to 7.7 cm3/100 g in pure copper. The coefficients for the effects of temperature on the solubility of hydrogen in some liquid binary copper-tin alloys, obtained from data reported by Bever and Floe (Ref 38), are given in Table 2. Hydrogen can enter the copper directly from the hydrogen in the atmosphere, but most likely it enters according to: 2Cu þ H2 O ¼ Cu2 O þ 2H
(Eq 27)
H2 O ¼ 2H þ O
(Eq 28)
Reactions given in Eq 27 and 28 are favored by copper with low oxygen contents. If there are high concentrations of alloying elements with more stable oxides than cuprous oxide, such as tin, the alloying element may react to put hydrogen into solution, for example: Sn þ H2 O ¼ SnO þ 2H
(Eq 29)
The main cause of porosity in the casting results from the evolution of hydrogen or water
Gases in Metals / 71 Table 2 Empirical constants for calculating hydrogen solubility (cm3/100 g) in the liquid state of copper-tin alloys at 1 atm partial pressure Solubility constants, log10 S; cm3 =100 g ¼ A T þB Alloy
Cu-5.9wt%Sn Cu-11.5wt%Sn Cu-21.7wt%Sn Cu-40.2wt%Sn Cu-54.8wt%Sn
A
B
2284.8 2304.0 2551.0 4239.0 3410.8
2.3465 2.2962 2.3289 3.0562 2.3737
Temperature range
C
1100–1300 1000–1300 1000–1300 1000–1300 1000–1300
F
2010–2370 1830–2370 1830–2370 1830–2370 1830–2370
Source: Ref 38
vapor during solidification. The hydrogen and oxygen contents increase because of enrichment during solidification, as described previously for cast iron. As the concentrations increase toward the end of the solidification, hydrogen or water vapor is evolved by the reverse reactions of Eq 27 and 28, which become thermodynamically favorable because the hydrogen and oxygen contents in the last liquid to solidify are high. This mechanism is supported by the observation that the voids in the casting are normally distributed along the grain boundaries, which are the last parts to solidify. Other gases that can cause porosity are sulfur dioxide and carbon monoxide, which also form during solidification according to Eq 30 and 31, respectively: S þ 2O ¼ SO2
(Eq 30)
C þ O ¼ CO
(Eq 31)
The thermodynamic pressure of carbon monoxide gas can be very high, even at moderate oxygen and carbon contents, because the solubility of the carbon in copper is limited, and therefore, its chemical activity is high. Porosity in Copper Alloys. In copper ingots, gas porosity usually results from H2 and water evolution. The evolution of as little as 1.12 cm3/100 g of hydrogen results in a gas evolution in 44% of the volume of the metal. If the sulfur content exceeds 0.05 wt%, the evolution of sulfur dioxide may also contribute to gas porosity. Only if the oxygen content exceeds 0.01% is carbon monoxide considered to be the cause of porosity (Ref 48). Because phosphorus deoxidizes copper significantly, the oxygen content of phosphorized copper is very low, and water, carbon monoxide, and sulfur dioxide do not contribute to porosity; only H2 need be of concern. Copper-zinc alloys rarely have problems associated with gas porosity, primarily because zinc deoxidizes the metal. Consequently, brasses do not dissolve significant quantities of oxygen. Zinc may also remove dissolved gases because of its high vapor pressure, which results in the zinc vapor flushing hydrogen out of the melt. During casting of copper-zinc alloys, degassing is not usually required. This is because the quantities of hydrogen absorbed by industrially produced copper-zinc melts are usually insufficient to cause problems in castings. The lowest hydrogen contents are usually
found in brasses with the highest zinc contents. However, the copper-tin alloys are the copper alloys that are most susceptible to gas porosity. This is because tin significantly increases the solubility of hydrogen in copper, tin mildly deoxidizes copper, and copper-tin alloys have long freezing ranges. Degassing of Copper Alloys. The two primary methods of removing dissolved gases from copper alloys are oxidation-reduction and inert gas flushing. Although vacuum degassing is theoretically possible, it is rarely used because it is not cost-effective. In oxidation-reduction, the first step is to remove the hydrogen by oxidation using an oxidizing slag or oxygen-rich copper. This involves increasing the activity of oxygen in the melt by melting it under an oxidizing flame or covering it when molten with an oxidizing flux such as a mixture of cupric oxide-boric acid-sodium chloride-sodium fluoride (Ref 49). This drives off the hydrogen in the form of steam. Then, just prior to pouring, the melt is deoxidized by adding phosphorus or other deoxidizers. The phosphorus deoxidizer produces cuprous phosphate, a molten deoxidation product that separates fairly easily from the molten copper metal. Generally, an addition of 0.02 to 0.03 wt% P is sufficient to deoxidize bronzes and gun metals. Calcium boride, boron carbide, and lithium have also been used for deoxidation (Ref 50). A charcoal cover is often used for deoxidation and protection of the melt from reoxidation. For copper alloys that contain strong deoxidizers, such as phosphorus, zinc, and tin, it is not possible to remove hydrogen by oxidation, because these elements will form stable oxides. For these alloys, inert gas flushing with nitrogen is commonly used. The theoretical minimum amount of inert gas is determined by using relationships similar to those derived for iron and aluminum. In practice, 150 to 250 L (5.3 to 8.8 ft3) of nitrogen per 1000 kg (2200 lb) of copper alloy is generally recommended (Ref 51).
Magnesium and Its Alloys Magnesium and its alloys are susceptible to gas-induced porosity. Of the nonferrous metals considered in this article, magnesium dissolves more hydrogen and has the potential for forming hydrides. Magnesium hydride, MgH2, is stable only for very high hydrogen activities
that are not normally encountered in metallurgical operations. At comparable temperature and hydrogen partial pressure, the solubility of hydrogen in magnesium and its alloys is an order of magnitude of that in aluminum and copper and their alloys. However, the propensity for gas porosity formation in magnesium is less than that in aluminum and copper. The principal source of hydrogen in magnesium and its alloys is moisture in the surrounding atmosphere or introduced into the melt when corroded materials are charged or damp fluxes are used. Magnesium and its alloys readily absorb hydrogen by the reaction with water vapor: MgðlÞ þ H2 OðvÞ ! MgOðsÞ þ H2 ðin MgÞ
(Eq 32)
The standard free energy change, DGo, for this reaction is extremely high: Go ¼ 368; 600 19:8T log10 T þ 122T J=mol
(Eq 33)
where T is the temperature in K. Application of the van’t Hoff reaction isotherm yields a value for aH2 =aH2 O that is even higher than the corresponding quotient for aluminum. In the absence of hydride formation, hydrogen dissolution in magnesium and its alloys is in the atomic form. Hydrogen Solubility and Reactions During Solidification. The solubility of hydrogen in pure magnesium has been experimentally determined by several investigators (Ref 52–59). However, the measurement of the solubility of hydrogen in magnesium and its alloys is an experimentally difficult task. This is because magnesium is too volatile and reactive toward traces of water vapor for direct measurement with the Sievert’s method. Combined linear regression analyses of the sets of reported data considered to be reliable yielded the following empirical equations for hydrogen solubility in liquid and solid pure magnesium as a function of temperature: Liquid magnesium-hydrogen: log10 S; cm3=100 g 1290:5 þ 3:042 ¼ T =K
(Eq 34)
Solid magnesium-hydrogen: log10 S; cm3=100 g 1062:47 þ 2:6254 ¼ T =K
(Eq 35)
At the melting point, 651 C (1204 F), the concentrations of hydrogen in equilibrium with the gas at atmospheric pressure are 44.2 and 29.9 cm3/100 g for the liquid and solid, respectively. Thus, the partition coefficient at 0.68 compared to 0.27 and 0.04 for hydrogen in pure magnesium is much higher than that in copper and aluminum, respectively. Hydrogen solubility in both the liquid and solid pure magnesium
72 / Principles of Liquid Metal Processing Temperature, ⬚F 750
570 120.0
930
1110
1290
1470
1650
1830
Pure Mg
Table 3 Coefficients for calculating hydrogen solubility (cm3/100 g) in the liquid state of magnesium alloys at 1 atm partial pressure Solubility constants, log10 S; cm3 =100 g ¼ A T þB
5.3 wt% Al
2 wt% Sr 1.3 wt% Zn Hydrogen solubility, cm3/100 g
A
B
Mg-Al alloys Mg-5.3wt%Al Mg-8.3wt%Al
1127 1201.5
2.8219 2.8532
Mg-Zn alloys Mg-1.3wt%Zn Mg-3.2wt%Zn Mg-6.8wt%Zn
1253.6 1074.7 1262
2.9733 2.7958 2.9577
Mg-Sr alloys Mg-0.5wt%Sr Mg-2wt%Sr Mg-3.4wt%Sr
1198.9 1028.3 886.9
2.9473 2.8323 2.7487
Mg-Si alloy Mg-1wt%Si
1126.1
2.875
Alloy
1 wt% Si
100.0
80.0
60.0
Derived from results given in Ref 38
40.0
that in aluminum and copper alloys. This can be attributed to the following reasons: 20.0
The significantly higher partition coefficient
0.0 300
400
500
600
700
800
900
1000
Temperature, ⬚C
Fig. 15
Solubility of hydrogen in magnesium and magnesium alloys as a function of temperature and at 1 atm hydrogen partial pressure
Hydrogen solubility, cm3/100 g
60
55
50
Ag Sn Cu Zn Sb
45
40 0
2
6 8 4 Alloying element, wt%
10
12
Fig. 16
Effect of alloying elements on the solubility of hydrogen in binary magnesium alloys at 700 C (1290 F) and 1 atm hydrogen partial pressure
is directly proportional to the square root of the hydrogen partial pressure, in accordance with Sievert’s law (Ref 52, 53). Alloying elements affect the solubility of hydrogen in magnesium (Fig. 15, 16). With the exception of nickel, zirconium, and strontium, the addition of alloying elements to magnesium generally results in a decrease in the solubility of hydrogen in magnesium. For example, the solubility of hydrogen in both liquid and solid
magnesium decreases with an increase in the addition of zinc, silicon, aluminum, lead, silver, tin, manganese, antimony, and copper (Ref 57, 60). However, because it apparently forms hydrides in the molten aluminum, strontium increases the solubility of hydrogen in liquid magnesium (Ref 57). Although zirconium increases the solubility of hydrogen significantly in solid magnesium, at 700 C (1290 F) it does not have any discernible effect on the solubility of hydrogen in liquid magnesium (Ref 60). In liquid magnesium-zirconium alloys, zirconium apparently acts as a scavenger for hydrogen, precipitating as zirconium hydride, ZrH2, and consequently obviating formation of gasinduced porosity in the castings. In the solid state, the absorbed hydrogen is also precipitated within the structure as zirconium hydride. At above 1 wt%, the addition of nickel significantly increases the solubility of hydrogen in liquid magnesium. The coefficients for calculating the solubility of hydrogen as a function of temperature in some binary magnesium alloys are given in Table 3. Porosity in Magnesium and Its Alloys. Castings of magnesium alloys are susceptible to gas porosity. The prevalence of gas porosity apparently increases with an increase in the hydrogen content of the molten metal. However, the susceptibility of magnesium alloys to gas porosity formation is significantly less than
of hydrogen in magnesium and its alloys reduces hydrogen evolution during solidification. The nominally higher solubility of hydrogen in liquid magnesium alloys makes the nucleation of gas porosity during solidification rather difficult. Nucleation of the gas bubble requires that the hydrogen concentration in the solidifying melt should exceed the equilibrium solubility of hydrogen in the melt. The comparatively higher partition coefficient of hydrogen also makes the characteristic of gas porosity in magnesium alloys different from that in aluminum and copper alloys. The gas porosity in magnesium alloys is more interdendritic in nature, more like a shrinkageinduced porosity. This is due to the proportionately much lower decrease in hydrogen solubility during solidification, so that a hydrogen content in the residual liquid sufficient to nucleate gas bubbles is not reached until at a later stage in solidification, if at all. Degassing of Magnesium Alloys. Degassing of magnesium and its alloys is usually accomplished by bubbling nitrogen, helium, nitrogen-chlorine mixture, or chlorine through the molten metal (Ref 61). The most effective of the gas-flushing methods is the bubbling of chlorine through the molten metal at 720 to 780 C (1330 to 1440 F). It can be used to reduce hydrogen content to 2 to 3 cm3/100 g. Bubbling of nitrogen reduces hydrogen to approximately 4 cm3/100 g. Other degassing methods that have been shown to be effective in reducing the hydrogen content of magnesium and its alloys include vacuum degassing and treatment with a flux. Treatment with a flux consisting of 55 wt% KCl, 34 wt% MgCl2, 9 wt% BaCl2, and 2 wt% CaCl2 apparently decreases the
Gases in Metals / 73 hydrogen content to 3 to 4 cm3/100 g. Although more environmentally friendly, vacuum degassing is not as effective as the gas-bubbling methods.
Overcoming Gas Porosity Gas in cast iron, aluminum, copper, and magnesium can be a major problem. The porosity results from reactions with the environment or the evolution of gases during solidification. In determining if gas evolution can cause porosity, the thermodynamics for the reactions must be favorable. In this article, the solubilities of the common gases and the reactions during solidification were discussed. Because of enrichment during solidification, the concentration of the elements increases to the extent that the thermodynamic pressure of the dissolved gases may exceed 1 atm. In addition to the thermodynamics, kinetic factors and interfacial energies determine if the gases will evolve. The most common method for removing hydrogen from aluminum, copper, and magnesium, which can also be used for removing hydrogen and nitrogen from cast iron, is inert gas flushing. The theoretical minimum amount of flushing gas required for removing hydrogen and nitrogen can be computed. In practice, the amount of inert gas required is somewhat higher because of kinetic factors. The kinetics of inert gas flushing can be improved by the dispersion of small bubbles throughout the melt. ACKNOWLEDGMENT Updated from Richard J. Fruehan, Gases in Metals, Casting, Vol 15, ASM Handbook, ASM International, 1988, p 82–87. REFERENCES 1. R.J. Fruehan, in Proceedings of the Physical Chemistry of Foundry Processes Symposium (Warren, MI), General Motors Corporation, 1986 2. S. Katz and C. Landefeld, General Motors Research, private communication, 1986 3. Making, Shaping and Treating of Steel, 10th ed., United States Steel Corporation, 1985 4. A. Kagawa and T. Okamoto, Trans. Jpn. Inst. Met., Vol 22 (No. 2), 1981, p 137 5. R.D. Pehlke and J. Elliot, Trans. TMSAIME, Vol 227, 1963, p 894 6. M. Inouye and T. Choh, Trans. JISI, Vol 8, 1968, p 134 7. R.J. Fruehan and L.J. Martonik, Metall. Trans. B, Vol 11, 1980, p 615 8. P.C. Glaws and R.J. Fruehan, Metall. Trans. B, Vol 16, 1985, p 551 9. P.C. Glaws and R.J. Fruehan, Metall. Trans. B, Vol 17, 1986, p 317
10. M. Byrne and G.R. Belton, Metall. Trans. B, Vol 14, 1983, p 441 11. F. Tsukihashi and R.J. Fruehan, Trans. JISI, Vol 27, 1987, p 858 12. R.J. Fruehan, Ladle Metallurgy: Principles and Practices, Iron and Steel Society of AIME, 1985 13. R.J. Fruehan, B. Lally, and P.C. Glaws, in Proceedings of the Fifth International Iron and Steel Congress (Washington, D.C.), Iron and Steel Society of AIME, 1986 14. P.N. Anyalebechi, “A Comprehensive and Critical Review of Reported Hydrogen Limits in Aluminum and Its Alloys,” unpublished research, 2000 15. C.E. Ransley and H. Neufeld, J. Inst. Met., Vol 74, 1947–1948, p 599–620 16. W. Baukloh and F. Oesterlen, Z. Metallkd., Vol 30 (No. 11), 1938, p 386–389 17. W.R. Opie and N.J. Grant, J. Met. (Trans. AIME), Vol 188, 1950, p 1237–1241 18. C.E. Ransley and D.E.J. Talbot, Z. Metallkd., Vol 46 (No. 5), 1955, p 328–337 19. D.E.J. Talbot and P.N. Anyalebechi, Mater. Sci. Technol., Vol 4 (No.1), 1988, p 1–4 20. P.N. Anyalebechi, Ph.D. thesis, Brunel University, Middlesex, England, 1985 21. D. Stephenson, Ph.D. thesis, Brunel University, Middlesex, England, 1978 22. H. Shahani, Ph.D. thesis, The Royal Institute of Technology, Stockholm, Sweden, 1984 23. J. Kocur, K. Tomasek, and L. Rabatin, Hutn. Listy, Vol 44 (No. 4), 1989, p 269– 275 24. H. Liu, L. Zhiang, and B. Bouchard, Recent Developments in Light Metals, M. Gilbert, P. Tremblay, and E. Ozberk, Ed., The Metallurgical Society of Canadian Institute of Metallurgists, 1994, p 257–268 25. H. Liu, B. Bouchard, and L. Zhiang, Light Metals 1995, J. Evans, Ed., The Minerals, Metals and Materials Society, 1995, p 1285–1291 26. P.N. Anyalebechi, Light Metals 1998, B. Welch, Ed., TMS, 1998, p 827–842 27. W. Eichenauer and A. Pebler, Z. Metallkd., Vol 48, (No. 7), 1957, p 373–378 28. W. Eichenauer, K. Hattenbach, and A. Pebler, Z. Metallkd., Vol 52, (No. 10), 1961, p 682–684 29. W. Eichenauer, Z. Metallkd., Vol 59, 1968, p 613–616 30. M. Ichimura, M. Imabayashi, and M. Hayakawa, Nippon Kinzoku Gakkaishi, Vol 43, (No. 9), 1979, p 876–883 31. E. Hashimoto and T. Kino, J. Phys. F., Met. Phys., Vol 13, 1983, p 1157–1165 32. M. Ichimura, Y. Sasajima, and M. Imabayashi, Mater. Trans., JIM, Vol 32 (No. 12), 1991, p 1109–1114 33. M. Ichimura, Y. Sasajima, and M. Imabayashi, Mater. Trans., JIM, Vol 33 (No. 5), 1992, p 449–453 34. G.K. Sigworth, Trans. AFS, Vol 95, 1987, p 73–78
35. G.K. Sigworth and T.A. Engh, Metall. Trans. B, Vol 13, 1982, p 447–460 36. J. Botor, Metal. Odlev, Vol 6, 1980, p 21–27 37. J. Botor, Aluminum, Vol 56 (No. 8), 1980, p 519–522 38. M.B. Bever and C.F. Floe, Trans. AIME, Vol 156, 1944, p 149–159 39. A. Sieverts, Z. Phys. Chem., Vol 77, 1911, p 11 40. P. Rotgen and F. Moller, Metallwirtschaft, Vol 13, 1934, p 81, 97 41. E. Kato, H. Ueno, and T. Orimo, Trans. Jpn. Inst. Met., Vol 11, 1970, p 351 42. R.B. McLellan, J. Phys. Chem. Solids, Vol 34, 1973, p 1137–1141 43. C. Thomas, Trans. Met. Soc. AIME, Vol 239, 1967, p 485–490 44. F.G. Jones and R.D. Pehkle, Met. Trans., Vol 2, 1971, p 2655–2663 45. M. Mokaram, Ph.D. thesis, Brunel University, London, 1974 46. W. Eichenauer and A. Pebler, Z. Metallkd., Vol 48, 1957, p 373–378 47. W. Eichenauer et al., Z. Metallkd., Vol 56, 1965, p 287 48. R.H. Waddington, J. Inst. Met., 1948– 1949, p 311 49. W.T. Pell-Walpole, J. Inst. Met., Vol 70, 1944, p 127 50. E.R. Thews, Metall., June 1956, p 431 51. W.A. Baker and F.C. Child, J. Inst. Met., Vol 70, 1944, p 349 52. J. Koeneman and A.G. Metcalfe, Trans. ASM, Vol 51, 1959, p 1072–1082 53. T. Watanabe, Y.C. Huang, and R. Komatsu, J. Jpn. Inst. Light Met., Vol 46, (No. 2), 1975, p 76–81 54. Y.C. Huang, T. Watanabe, and R. Komatsu, Proc. Fourth Int. Conf. Vacuum Metallurgy, 1973, p 173–176 55. D. F. Chernega, Gotvjanskij and Prisjasznjokm Bestn. Kiev. Politexn. In-Ta, Kiev, Vol 17, 1959, p 55–58 56. V. Shapavalov, N.P. Serdyuk, and O.P. Semik, Dop. Akad. Nauk Ukr. RSR. Ser. A. Fiz.-Mat. Tekh. Nauki, Vol 6, 1981, p 99–101 57. E. Overlid, P. Bakke, and T.A. Engh, Light Metals 1997, Metaux Legers, Canadian Inst. of Metallurgists, 1997, p 141–154 58. Z.D. Popovic and G.R. Piercy, Met. Trans. A, Vol 6, 1975, p 1915–1917 59. M.V. Sharov and V.V. Serebryakov, Nauchnyye Doklady Vvyshe Shkoly, Metallurgiya, (No. 2), 1958, p 37 60. T. Watanabe, Y. Tachihara, Y.C. Huang, and R. Komatsu, Kei Kinzoku, Vol 26 (No. 4), 1976, p 167–174 61. T. Watanabe, Y.C. Huang, R. Komatsu, and T. Ito, J. Jpn. Inst. Light Met., Vol 26 (No. 6), 1976, p 266–272
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 74-83 DOI: 10.1361/asmhba0005193
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Inclusion-Forming Reactions Revised by Paul K. Trojan, University of Michigan—Dearborn; and Richard Fruehan, Carnegie Mellon University
INCLUSIONS can be defined as nonmetallic and sometimes intermetallic phases embedded in a metallic matrix. They are usually simple oxides, sulfides, nitrides, or their complexes in ferrous alloys and can include intermetallic phases in nonferrous alloys. In almost all instances of metal casting, they are considered to be detrimental to the performance of the cast component. For example, mechanical properties can be adversely influenced by inclusions, which act as stress raisers. There is no set pattern for the effect, but some properties are more sensitive to the presence of inclusions than others. Elongation or reduction in area is usually modified more significantly than ultimate tensile strength. As a result, ductility specifications are common quality-control indices in cast products. Similar observations are found with porosity due to gas or occluded shrinkage. The loss in casting properties measured by a tensile test may then reflect a combination of defects; of these defects, inclusions are only a single but very important source of poor performance. Because imperfections become areas of higher stress concentration, the percentage of property loss becomes greater when the strength requirement is higher. An important example is the decrease in fracture toughness when inclusions are present in higher-strength lower-ductility alloys. Similar pronounced property degradation is observed in tests that reflect slow, rapid, or cyclic strain rates, such as creep, impact, and fatigue testing. In all cases, the degree of property loss will depend on the parameters that control the characteristics of all microstructures. These metallurgical phase controls are the nature, amount, size, shape, distribution, and orientation of the phases. If inclusions are to be controlled, these characteristics must be suitably modified to minimize their impact on casting performance. Although inclusions are generally considered to be detrimental, their intentional introduction in larger quantities can lead to unique dispersion-strengthened materials. An example would be the dispersion of insoluble silicon carbide whiskers or alumina (Al2O3) fibers in a molten aluminum alloy that is subsequently cast into a final shape by squeeze casting methods (Ref 1).
Inclusion Types There are essentially two classifications for all inclusions: Exogenous: those derived from external
causes
Indigenous: those that are native, innate, or
inherent in the molten metal treatment process Exogenous Inclusions. Slag, dross, entrapped mold materials, and refractories are examples of inclusions that would be classified as exogenous. In most cases, these inclusions would be macroscopic or visible to the naked eye at the casting surface (Fig. 1). When the casting is sectioned, they may also appear beneath the external casting surface if they have had insufficient time to float out or settle due to
1
Fig. 1
density differences with respect to the molten metal. Indigenous inclusions include sulfides, nitrides, and oxides derived from the chemical reaction of the molten metal with the local environment. Such inclusions are usually small and require microscopic magnification for identification. They are often uniformly distributed within the microstructure. However, a grainboundary distribution of indigenous inclusions (Fig. 2) can be particularly damaging to mechanical properties. There are cases in which indigenous inclusions such as oxides may react with mold materials to promote an exogenous inclusion variety that otherwise may not appear in the absence of the indigenous oxide. To circumvent the strict definition as indigenous or exogenous, a number of investigators prefer to catalog all inclusions as either macroinclusions of microinclusions.
2
Ductile iron crankshaft segment essentially free of exogenous inclusions (1, left) and with numerous exogenous inclusions (2, right). Low pouring temperature and poor mold-filling practice were the cause of the inclusions in part 2.
Inclusion-Forming Reactions / 75 Oxygen, at.% 49 50
51
52
53
54
55
56
L2
L1 + L 2
(δ-Fe) 1394⬚ 1373⬚
59
60
61
62
63
64
L2 + O2 (g; 1atm) 28.30
3000
1585⬚
28.07 Fe O + O (g; 1atm) 3 4 2 1459⬚
22.60
1500 1400
58
1599⬚
1600 1529⬚
57
25.31
22.84
1426⬚
28.36
25.60 27.64
2800
30.04
2600
22.91 23.16
2400
1300 FeO (Wustite)
Temperature, ⬚C
1200
2200 Fe3O4 (Magnetite)
1100
2000 (γ-Fe)
Fe2O3 + O2 (g; 1atm)
1000 912⬚
23.10
900
1800
Fe2O3 (Hematite)
Temperature, ⬚F
1700
47 48
1600 Curie temperature
800
Fig. 2
Type II indigenous sulfide inclusions in a 0.25% C steel. Etched using 2% nital. Original Magnification: 500
600
Although such definitions describe the anticipated effect on mechanical properties, there can be a loss in source identification of the inclusion. Whatever definition is used, control of inclusions requires an understanding of their origin and the associated physical chemistry.
770⬚
1400
(α-Fe)
1200
700
560⬚ 23.26
1000
500 400 20
800
30.06 21
22
23
24
25
26
27
28
29
30
31
32
33
34
Oxygen, wt%
Physical Chemistry
Fig. 3
Phase diagrams, thermochemical relationships, and reaction rates must be considered for proper understanding of inclusion formation. The general concepts of each, along with their practical significance, are established in the following sections in this article. Phase Equilibria. Phase diagrams map the occurrence of phases to be expected under equilibrium conditions. Equilibrium also requires long periods of time to achieve. This is certainly not the case in castings, in which even moderate cooling rates may inhibit equilibrium phases. For example, in the iron-oxygen system shown in Fig. 3, wustite (FeO) produced at high temperatures should break down to a-iron and magnetite (Fe3O4) at 560 C (1040 F). However, iron scales may show considerable amounts of retained FeO at room temperature. A further limitation is that pure elements seldom exist in the casting process, and more than two elemental species are usually present. In Fig. 3, pure oxygen is one component, while air (20% O2) is more likely to be the oxidizing source during casting. Furthermore, a cast steel will contain carbon and other elements, such as silicon or aluminum, that will have a greater affinity for oxygen than iron. Therefore, phase diagrams become a tool for understanding, but the predicted equilibrium for inclusion occurrence may not always be observed
for the casting cooled under nonequilibrium conditions. Thermochemistry. Any chemical reaction will occur if the products are at a lower energy level than the reactants. The energy measurement is the chemical free energy, which can be treated simply as the energy of chemical position. Free energy values can be obtained directly from numerous sources or can be calculated from tabulated enthalpy, entropy, and heat capacity data. An equilibrium is then established between the products and reactants such that very little remaining reactant indicates a more or less complete reaction. A more thorough discussion of chemical free energy is provided in standard textbooks on chemical thermodynamics. By way of example, assume pure molten iron contains dissolved oxygen at 1600 C (2910 F). A small amount of aluminum is added to the melt, and the chemical reaction is: 2Al þ 3O ! Al2 O3 ðsÞ where the underlined indicates the elements to be dilute and dissolved in iron while the (s) indicates that Al2O3 is solid at 1600 C (2910 F). (The melting point of Al2O3 is 2072 C, or 3762 F.) For the reaction to occur, the standard free energy exchange at 1600 C (2910 F) must be negative. However, the quantity of each component in the reaction is also important, and the chemical activity of each species must
Iron-oxygen phase diagram
be known. An overabundance of aluminum will react stoichiometrically with oxygen, and the remainder may go into liquid solution or react with the surroundings. In the previous example, the Al2O3 product is insoluble in the molten iron and can result as inclusions trapped in the solidified casting. It is then possible to calculate the free energy exchange for the occurrence of various inclusion chemistries. Their relative stability is dictated by the most negative value of free energy. In the previous example, iron oxide would not occur, because Al2O3 has a lower free energy of formation. In fact, iron oxide would be predictably reduced by aluminum to form Al2O3. As with phase equilibria, thermochemical calculations are again based on equilibrium data, and castings may not be at equilibrium throughout the solidification process. Kinetics. There are several reasons why inclusion-forming reactions are usually incomplete and why reaction rates therefore become the most important practical consideration. Much of the understanding of reaction rates is associated with the following observations: Not all species diffuse at the same rate; when
coupled with the dependence of diffusion rates on temperature, nonuniform inclusion formation and distribution may be anticipated.
76 / Principles of Liquid Metal Processing Chemical free energy is also temperature-
Experimental kinetic data can be presented in several ways, including mathematical relationships associated with activation energy and first-order kinetic reaction rate theory. Perhaps more familiar are isothermal and continuous cooling transformation graphical interpretations consistent with nucleation and growth-controlled phenomena.
3990
2200
Liquid 3630
2000
1840⬚ 3270
1800 Cristobalite + liquid
Corundum + mullite solid solutiion
Mullite + liquid 1595⬚
1600
2910 Mullite solid solution
Cristobalite + mullite
1400 0
20
Temperature, ⬚F
1850⬚
40
60
SiO2
80 Mullite 3Al2O3.2SiO2
2550 100 Al2O3
Al2O3, wt%
Control of Inclusions The previous discussion suggests several practical means of controlling the occurrence of inclusions in any cast metal. All of the control methods can be cataloged as chemical, mechanical, or a combination of these. Several of the more common techniques are discussed separately as follows. Chemistry Control. Although control of alloy chemistry may be obvious, composition cannot always be modified to minimize inclusions. The metal compositions may be inherently reactive with the local environment. For example, the presence of magnesium is required in nodular cast iron; the propensity for inclusions is greater than for the same nominal chemistry of a gray cast iron. Similarly, the existence of inclusions is accepted because the alternative is worse. Aluminum added to steels may result in Al2O3 inclusions, but the lack of deoxidation if aluminum were not used would result in gas holes with a greater negative effect on casting performance. However, this does not imply that if a small addition is good then more is better. Phosphorus added to a copper-base alloy can be an effective deoxidizer, but excess phosphorus can be very aggressive and can react with mold materials commonly used with this alloy family. Therefore, the importance of chemistry control cannot be overemphasized if inclusions are to be controlled. Mold/Metal Interface and Refractory Control. Another chemical variable that can lead to inclusions is that of the mold materials and refractories. The SiO2-Al2O3 phase diagram is shown in Fig. 4. It is obvious that the solidus temperature increases with Al2O3
Corundum + liquid
Mullite + liquid Temperature, ⬚C
dependent and may not vary in the same fashion for all of the species under consideration. Available phase equilibrium data may predict the relative solubility of inclusion-forming species in the liquid state but may not treat decreasing solubility in the cooling liquid or the effects of the solidification process on the equilibrium. The surface energy of species will change with temperature and will influence the ability of an inclusion to float, sink, agglomerate, or react secondarily with its surroundings. Reaction products at a surface can become a barrier to further reaction and can block completion of the chemical reaction.
Fig. 4
SiO2-Al2O3 phase diagram. Source: Ref 2
content. Even if a refractory is to be used well below the solidus temperature, a high-Al2O3 refractory may still be preferred because Al2O3 plus mullite would be present rather than the silica (SiO2) plus mullite that would be present at compositions less than 72% Al2O3. The SiO2 does not possess the same thermochemical stability as Al2O3; therefore, aggressive liquid metals or oxides may react more readily with SiO2. Higher temperatures and longer holding times at temperature increase the possibility for reaction and the production of inclusions associated with refractory wear and erosion. Therefore, melt overheating, especially for long periods of time, is to be avoided. The scale on scrap surfaces can also lead to refractory erosion and exogenous inclusions. Rust on ferrous scrap can be considered a hydrated iron oxide scale that, when heated and dehydrated, can form low-melting complex oxides with the refractories. Inclusion difficulties have also been known to increase when adhering mold materials remain on charged revert scrap. The severity depends on the nature of the mold bonding agent; bentonite and organically bonded mold materials would display different tendencies toward refractory attack. The aforementioned observations can be applied to reaction between the mold materials and the molten metal. This is not to be confused with metal penetration, in which molten metal merely fills the void spaces between grains of
the mold material. However, the solution to the problem—for example, a mold wash— may be the same. In this case, the wash fills the voids between sand grains and isolates the mold material from contact with a potentially aggressive molten metal or oxide. Longer solidification times merely increase the possibility for mold-metal reaction and the accompanying inclusion-forming tendency. Separation Techniques. The choice for inclusion separation is between removal before the molten metal enters the mold or use of a removal system as a component part of the mold. Gating systems are discussed in the next section in this article. In both cases, inclusion control is basically mechanical. External methods include simple skimming of ladle slag or dross to minimize exogenous inclusion sources before pouring of the castings. A teapot ladle can be very effective for decanting clean metal from beneath a layer of floating slag or dross. Screens or sieves have been used in gating systems to filter inclusions. In general, metallic filters suffer from the tendency to alloy with the poured metal, and they are difficult to separate from the charge returns. Ceramic foam filters with consistent pore size and structural integrity are being increasingly used for those alloys from which definite cost benefits can be derived (Ref 3–5). Gating Methods. Gating and pouring practice has been known to create and to solve inclusion problems in castings. If exogenous
Inclusion-Forming Reactions / 77
500
475 68
64 425 60 400 56
Alternating stress, ksi
Alternating stress, MPa
450
375 Conventional 52 Calcium treated
350
48
325 104
105
106
107
Cycles to failure
Fig. 5
Comparison of axial fatigue data for untreated and calcium-treated rolled ASTM A516 steel. 51 mm (2 in.) thick plates tested with alternating stress ratio of 0.1. Source: Ref 8
Yield strength, ksi 30 320
50
70
90
110
130 240
280
180
200 160
120
120 Calcium treated 80
60
40 Conventional
0 −40 200
Charpy V-notch USE, ft · lb
240 Charpy V-notch USE, J
inclusions leave the ladle because of inadequate upstream control, the gating system should be designed to remove such inclusions in the pouring basin or in the runner system. Too often, a pouring basin is not designed to fill easily or to remain filled, and inclusions that may float are forced into the casting cavity. Whirl gates can be used to collect the inclusions and to ensure a better opportunity to have clean metal enter the casting proper (Ref 6). In addition, runners can have dead-end extensions that collect the first metal to enter the mold, which is the most likely to contain slag or dross (Ref 7). Pressurized gating systems are often recommended for preventing the aspiration of air that can cause the formation of inclusions during the pouring process. This technique is more important when the metal or alloy has a great tendency toward inclusion formation because of so-called secondary oxidation. More detailed information on correct gating is available in the article “Gating Design” in this Volume; many of the gating practices discussed in that article are used to minimize inclusions. Inclusion Shape Control. As pointed out previously, mechanical properties suffer when inclusions are present, with a more pronounced degradation in strain-rate-sensitive applications. For example, an inclusion at the surface of a casting subjected to cyclic strain can become a source for fracture initiation in fatigue. The most conservative approach in design is to determine the smallest inclusion at the surface that may be present without casting rejection. With no variation in inclusion continuity or size, a spherical shape (lower microstress concentration) would be less damaging than more angular shapes. Inclusions beneath the surface would exhibit the same dependence on shape for fracture propagation rather than initiation. As shown in Fig. 5, calcium treatment of a steel provides longer life under cyclic loading due to inclusion shape modification. Although the data in Fig. 5 are for a rolled constructional steel, similar results would be anticipated for a cast steel, where anisotropy also exists due to directional solidification. The advantage of a spherical inclusion shape is possibly more evident in Charpy V-notch impact data for constructional steels (Fig. 6). SEM fractographs show a pronounced difference in deformation mode with an inclusion shape modification (Fig. 7). Local ductility increases adjacent to the inclusions when the shape is spherical. This would then suggest that if inclusions are to occur and cannot be completely removed, a spherical shape is the least detrimental. In these examples, the MnS inclusions assume a planar shape that allows easier fracture propagation between inclusions or lower observed alloy ductility. The addition of calcium converts the inclusion to essentially CaS with a much higher melting point, spherical shape, and hence, higher alloy ductility. Inclusion shape control, therefore, is a useful technique that is often applied to more demanding casting applications in which added
300
400
500
600
700
800
900
0 1000
Yield strength, MPa
Fig. 6
Comparison of Charpy V-notch upper-shelf energies (USEs) for several grades and thicknesses of untreated and calcium-treated steel. Source: Ref 9
processing costs can be justified. The principle is a control of surface energy between the inclusion and the metal—normally through suitable chemistry modification of the metal. By analogy, the technique may not be unlike the addition of magnesium to cast iron to change the graphite shape from flakes to nodules.
available for modifying the viscosity of slag or dross, thus facilitating their ease of removal. On the other hand, treatment of steels and cast irons is likely to be different, and a few examples, along with the application of appropriate physical chemistry, are given as follows.
Steels
Inclusions in Ferrous Alloys Some applications for inclusion minimization or shape control are immediately obvious. For example, proprietary addition agents are
Oxides. The amount of carbon in iron controls the amount of dissolved oxygen (Fig. 8). The curve has been calculated from thermochemical data with variable CO/CO2 ratios. Experimental points would generally lie above
78 / Principles of Liquid Metal Processing
(a)
(b)
Fig. 7
SEM fractographs showing the effect of calcium treatment on the fracture morphology of ASTM A633C steel impact specimens. (a) Untreated steel with type II manganese sulfide inclusions showing evidence of brittle fracture. (b) Calcium-treated steel with spherical inclusions; the fracture is ductile. Courtesy of A.D. Wilson, Lukens Steel
Company
1.0
0.10
0.5 1705 ⬚C 1600 ⬚C 1515 ⬚C
0.2
0.08
2Fe
O+
0.1
Si
SiO
0.06
2 +2 Fe
0.02 FeO, %
Oxygen, wt%
0.05
0.04
0.01
3Fe
O+
0.005
2A
l Al
0.002 0.02
2O 3 +3 F
e
0.001
Equilibrium with CO + CO2 at 1 atm pressure and 1600 ⬚C
0.0005 0.0002
0
0
0.2
0.4
0.6
0.8
0.0001 0.001 0.002
0.005 0.01 0.02
0.05
0.1
0.2
0.5
1.0
2.0
5.0
10
20
50
Carbon, wt% Alloy remaining in steel, %
Fig. 8
Carbon-oxygen equilibrium at 1600 C (2910 F). Source: Ref 10
the theoretical curve, but the relationship between carbon and oxygen would be the same. Low carbon in solution, as in many cast steels, results in higher dissolved oxygen. The solid solubility of oxygen in iron is low, and during solidification, oxygen is rejected to the remaining liquid, where it can react with carbon to form CO and result in gas porosity. Although the porosity could occur in rimmed steel ingots (induced porosity that is “healed” during subsequent processing), the amount of dissolved oxygen in steel castings is reduced by additions of aluminum or silicon. The resultant oxides are insoluble and form inclusions.
Fig. 9
Equilibrium between FeO and dissolved aluminum and silicon. Source: Ref 11
Because Al2O3 has a lower free energy of formation than SiO2, aluminum is a more effective deoxidizer than silicon (Fig. 9). The FeO content is an index of the amount of dissolved oxygen in the steel where the oxygen can also be reduced by carbon, as shown in Fig. 8. At a given carbon content, the oxygen can then be reduced by aluminum or silicon to still lower levels, because the free energies for SiO2 and Al2O3 are lower than those for CO or CO2 at the reaction temperature. Once the oxide forms, the inclusion removal rate depends on the nucleation of particles, their growth, ultimate agglomeration, and rise to the surface. Although there is a significant
difference in density between the solid particles and the liquid metal, a simple application of Stoke’s law is difficult. The problem is dynamic and is influenced by the presence of other indigenous or exogenous inclusion species. Furthermore, gravitational effects are different while holding the treated metal in the ladle, pouring through air with an opportunity for the metal to reoxidize, and during solidification with the associated thermal convective currents. Therefore, it should not be surprising that oxide inclusions can become trapped in the final casting. Sulfides. Sulfur is also often present in steels and is controlled by the nature of the
Inclusion-Forming Reactions / 79 50 Reduction in area 40
54.2 Izod
40.7
30 Elongation 20
27.1 0.06% S
10 60 Reduction in area
50
67.8
Izod
40
54.2
Elongation 40.7
30 0.04% S
20 60 Reduction in area 50
67.8 Izod 54.2
40 Elongation
40.7
30
(b)
Izod impact strength, J
Izod impact strength, ft · lb Reduction of area and elongation, %
(a)
0.03% S
20 70 Izod
81.3
60 Reduction in area
67.8
50 40
54.2 Elongation
30 20
40.7 0.02% S 0
0.04
0.08
0.12
0.16
0.20
0.24
0.28
Aluminum added to heat, %
(c)
Fig. 10
Illustrations of common inclusion shapes and distributions found in steel. (a) Type I: globular silicates and oxides (no aluminum deoxidizer). (b) Type II: continuous eutectic sulfides and Al2O3 at grain boundaries (low aluminum addition). (c) Type III: duplex irregularly shaped sulfides and Al2O3 (large aluminum additions). Source: Ref 12, 13
slag-refining processes in the furnace. In general, as the process becomes more reducing (that is, the oxygen content of the metal is lower) and the slag more basic, lower sulfur will be present. Because the sulfur content is not reduced to zero, manganese is always present in steels to produce the more desirable manganese sulfide (MnS) inclusions rather than FeS inclusions with lower melting point. The resulting inclusions, although usually identified as manganese sulfides, may have several distributions and shapes, depending on the presence of other elements and the deoxidation procedure. Complex Inclusions. The various elements compete to form compounds with oxygen and sulfur. Furthermore, alloying elements in excess of those necessary to form compounds will modify the surface energy of the resultant inclusions.
Fig. 11
Combined effect of aluminum and sulfur on the ductility and impact strength of cast and normalized medium-carbon steel. Source: Ref 11–13
The shape and distribution of the inclusions will then also vary, depending on the variables discussed at the beginning of this article. Historically, inclusions have been classified into three general types, as shown in Fig. 10. Of these, type II has been demonstrated as particularly damaging to mechanical properties. The argument has been that sulfur, oxygen, and deoxidizer play a combined role in the inclusion type. Figure 11 shows this effect and the resulting decrease in structure-sensitive properties such as ductility and impact strength. The minimum in the curves is attributed to the formation of type II inclusions. The role of aluminum in the modification of inclusion shape and distribution also depends on the oxygen-sulfur equilibrium shift when the aluminum is added. Therefore, complex inclusions such as oxy-sulfides may be present in a cast steel but can be controlled by the quantity of oxide-sulfide-forming elements. A more complete discussion of oxide inclusion formation mechanisms and the relationship to mechanical properties is given in Ref 14.
Inclusion shape control becomes a consideration especially if fracture toughness, impact strength, or fatigue resistance must be optimized. Magnesium and calcium additions can be successfully used for shape control, that is, to produce more spherical shapes (Ref 15). It is noteworthy that both magnesium and calcium form stable sulfides and oxides as indexed by the free energy of formation of the compounds. Several production methods have been employed for calcium or magnesium additions, either as the metallic elements or as compounds. Because of the high vapor pressure of the pure metals and, to a lesser extent, the compounds at the addition temperatures, special procedures have evolved, including injection at the bottom of a filled ladle. In this case, the ferrostatic head provides a pressure that suppresses the boiling action and gives a longer reaction dwell time. The most common practice is to deoxidize the steel with aluminum followed by calcium for inclusion shape control. Solid Al2O3 particles are formed upon deoxidation, occasionally
80 / Principles of Liquid Metal Processing with complex aluminates such as spinels (Al2MgO4). These solids prefer their own kind and can therefore agglomerate into larger, more harmful inclusions. The addition of calcium lowers the melting range to form liquid calcium aluminates (40 to 60% CaO) that do not agglomerate but appear as small and more desirable spherical oxide inclusions. Rare earth elements added in conjunction with calcium, magnesium, and aluminum have also been reported to enhance the spheroidization and fineness of the inclusions by breaking up the Al2O3 galaxies (Ref 16, 17). However, difficulty in removal of the rare earth oxide inclusions, especially in scrap remelt, has limited the practice in present-day steel casting processes. Nitrides. Just as there are elements that preferentially form oxides or sulfides, there are differences in the free energy of formation of the nitrides. These elements control the amount of nitrogen that can be dissolved in both the solid 0.020 1260 ⬚C
Nitrogen, %
0.015 980 ⬚C
1150 ⬚C
0.010 1065 ⬚C 900 ⬚C
0.005
0 0
0.02
0.04
815 ⬚C
0.08 0.06 Aluminum, %
0.10
Fig. 12
0.12
Isothermal precipitation of aluminum nitride. Compositions to the right or above each isotherm cause precipitation. Source: Ref 12
and liquid states. For example, vanadium and manganese increase nitrogen solubility, while carbon and phosphorus decrease nitrogen solubility (Ref 12). The effect of aluminum on nitrogen solubility has been extensively studied, and criteria that result in the precipitation of aluminum nitrides have been developed (Fig. 12). Other nitride-forming elements, such as titanium, zirconium, and vanadium, would show similar but numerically different results. It is not uncommon to consider these nitride inclusions to be beneficial; they peg austenite grain boundaries, and this results in grain refinement. On the other hand, brittleness results when the nitrides become more or less continuous. Exogenous Inclusions. Numerous examples of inclusions occur because of the reaction between the molten steel and the adjacent environment. A few of these are considered as follows. When a steel is killed with aluminum, residual aluminum remains in solution. If the metal contacts a refractory that contains free SiO2, the aluminum can reduce the SiO2 because of the lower free energy of formation of the Al2O3. The eroded or partially dissolved refractory becomes a source of inclusions, as does the Al2O3 that is produced. Careful control of refractory chemistry and good pouring practice minimize this problem. Similar reactions can occur at the mold/metal interface, where the first equilibrium of importance is iron-oxygen-silicon. When conditions are oxidizing, FeO is present, and the corresponding form for silicon is SiO2. The FeOSiO2 phase diagram is shown in Fig. 13. The most significant observation here is the formation of
1800
3270 2 liquids ~1700⬚
1600
2910 SiO2 (Cristobalite) + L
1470⬚ 1400
2550
1371⬚ SiO2 (Tridymite) + L
2FeO·SiO2 + L FeO + L
1208⬚
1200
2190 1180⬚
1175⬚
FeO + 2FeO•SiO2 1000 FeO
20
2FeO·SiO2 (Fayalite) + Tridymite 40
60 SiO2, wt%
Fig. 13
FeO-SiO2 phase diagram. Source: Ref 10
80
1830 SiO2
Temperature, ⬚F
Temperature, ⬚C
Liquid
lower-melting liquid phases as the amount of FeO is increased. With minimal oxidation, burned-on sand and metal penetration will occur. However, when the oxidation level increases, local macroinclusions may form, especially if poor pouring practice, segregated bentonite, nonuniform mold ramming, and excessive mold moisture exist. A number of the problems and solutions associated with exogenous inclusions in steel have been summarized in Ref 18.
Cast Irons Gray Cast Iron. The earlier discussion of the carbon-oxygen equilibrium (Fig. 8) suggests that dissolved oxygen should not be a problem in cast irons because of their high carbon contents. Therefore, oxide inclusions should not occur in gray cast iron. However, because the silicon content of this material is also significant, the carbon-silicon-oxygen equilibrium is of interest. The temperature dependencies of the free energies of formation for SiO2 and CO are different for the two species. At temperatures above approximately 1535 C (2795 F), CO is the more stable compound, while below this temperature, SiO2 is more stable. The exact temperature depends on the carbon and silicon content of the metal. The phenomenon can be observed on the surface of a cast iron melt that appears clean at high temperatures and yet develops a slag covering at lower temperatures. The significance of the equilibrium shift is that skimming the slag from a melt surface may be unnecessary at high temperatures, because little may exist. Inclusions are then found in the casting because of pouring at lower temperatures and exposure of the metal to reoxidation during pouring. These inclusions are often silicates derived from the normal equilibrium change with temperature. Once the molten iron is in the mold cavity, there is less opportunity to form the silicates because conditions are usually less oxidizing. Nitride and sulfide inclusions also exist in gray cast iron but are not as severe a problem as in steel because of the inherently lower dissolved oxygen content of the cast iron. Special gray cast iron chemistries containing alloying elements with an affinity for nitrogen, oxygen, or sulfur may result in inclusions by further reaction with the refractories, mold, or atmosphere. The remedies are usually similar to those given for steel. Ductile Cast Iron. A number of foundries erroneously assumed that ductile (nodular) cast iron would be as easy to process as gray cast iron because the major difference was merely a lower sulfur plus a modest magnesium addition. Although many characteristics were found to be different, one of the more perplexing was the formation of dross usually not found in gray cast iron. The result was the so-called cope-side defect that was very prevalent in early ductile iron castings.
Inclusion-Forming Reactions / 81 Magnesium forms a very stable oxide and can therefore reduce oxides of lower stability, as determined by their free energy of formation. Furthermore, the alloys used to produce the nodular graphite found in ductile iron contain considerable amounts of silicon and are low in density, which allows them to float on ladle surfaces rather than reacting completely. These observations help to explain how dross particles form and can become entrapped in castings. In general, there are three types of dross inclusions in ductile cast iron (Ref 19). Type I consists of large dross particles generally derived from oxidation of the silicon and magnesium used in the nodularizing alloys. The dross particles are present on the treated liquid-metal surface and can be removed by skimming. Silica does not exist in the pure form; rather, magnesia (MgO) and forsterite (Mg2SiO4) are the predominant species. This suggests a driving force from the high-MgO end of the phase diagram shown in Fig. 14. Under some circumstances, especially after a late silicon inoculation, the dross also contains fayalite (2FeOSiO2 or Fe2SiO4) (Fig. 13). Fayalite and forsterite form a complete series of solid solutions known as the olivines and are used as molding materials. Type II inclusions are in a more or less stringlike configuration in conjunction with decayed graphite shapes. Chemically, they are predominantly MgO with some iron and silicon atom replacement for the magnesium in the MgO lattice. Reaction with the oxidizing components in the mold material is one source for this type of inclusion, as is an excess of nodularizing agent. Their position is usually just beneath the casting surface. Control of the mold
atmosphere and the amount of nodularizer usually minimizes the defect. Type III inclusions are fine dispersions of MgO and MgS, occasionally mixed with Mg2Si. These components would normally be expected only if the oxygen content in the immediate vicinity is very low. Their normal uniform distribution does not usually detract from the overall performance of the casting. However, undissolved nodularizer may be a source of type III inclusions, because Mg2Si is the active magnesium source in the nodularizing agent. Finally, the examples provided in this section are not intended to represent all of the inclusions to be anticipated in every ferrous alloy. On the other hand, they are sufficiently representative so that analogies for occurrence and remedy can be derived for the alloys not discussed.
Inclusions in Nonferrous Alloys Several important cases are given subsequently for aluminum, copper, and magnesium alloys. The principles discussed in this section can be extended to include other alloy systems. In the alloys discussed, both nonmetallic and intermetallic compounds are considered deleterious.
Aluminum Alloys Oxides. It is nearly impossible to prevent dross formation in the melting of aluminum alloys. The most common form of dross contains large quantities of Al2O3 and, when agglomerated, can be removed from the melt surface by skimming. Pouring operations should employ 3450
1900 1890⬚
MgO + L 1850⬚
Liquid 3270
1800
skimming ladles to prevent larger dross particles from entering the mold. Furthermore, a gating practice must be used that minimizes Al2O3 formation during mold filling. The fact that Al2O3 forms is not surprising, because the free energy of formation of Al2O3 is very low. Furthermore, the specific gravity of Al2O3 is not greatly different from that of the molten aluminum alloys. Although more dense, Al2O3 forms at an outside surface first if the melt is exposed to the atmosphere. The tenacious skin will float unless it is broken up into small particles, and it can become quite thick when melts are held for extended periods in the furnace. When small oxide particles are formed, they may not sink but may become suspended in the melt. In high-quality castings, these particles must be removed; ceramic filters have been used to remove the particles, as mentioned earlier in this article (Ref 3). The degradation of mechanical properties also depends on the size and distribution of oxide inclusions. Therefore, solidification rate is very important for minimizing property loss. When the casting solidifies rapidly, as in thin sections or in permanent molds, the oxide inclusions are smaller and more uniformly dispersed. This observation is consistent with the general constraints of nucleation and growth, where less time during solidification not only creates smaller particles but also prevents their agglomeration and growth. Complex Oxides. Other elements in the aluminum alloys may react with oxygen sources and produce complex oxides, with Al2O3 as one of the components. This is the case in magnesium-containing alloys, in which the spinel MgAl2O4 may form (Fig. 15). Here, MgO has a lower free energy of formation than Al2O3; the spinel results when they form together. Again, precautions in pouring, gating, and minimizing the reoxidation of the metal are important in inclusion control. The use of filtering systems is also advantageous, especially in high-strength castings in which fracture toughness characteristics are important.
1695⬚
3090
SiO2 + L
1600
2910 1557⬚
20
40
Fig. 14
MgO-SiO2 phase diagram. Source: Ref 10
Liquid + MgAl2O4
2600
2200
MgO + liquid
MgAl2O4 1800
1400
MgO·SiO2 + SiO2
4710
Liquid + spinel
Liquid + Al2O3
Spinel
MgO + MgAl2O4
3990
3270
Al2O3 + spinel
2550
(Enstatite)
60 SiO2, wt%
Liquid
1543⬚
2MgO·SiO2 + MgO·SiO2 1500 MgO
5430
3000
80
2730 SiO2
1000 0 MgO
Fig. 15
20
60 40 Al2O3, mol%
80
1830 100 Al2O3
MgO-Al2O3 phase diagram. Source: Ref 2
Temperature, ⬚F
1700
Temperature, ⬚C
2 liquids
Temperature, ⬚F
(Forsterite)
2MgO·SiO2 + L
Temperature, ⬚C
MgO + 2MgO·SiO2
2030
1100
2012
Temperature, ⬚C
1110
1090
L
1976
1080 1070 1060
1050 0.001
Fig. 16
1994
1958 α
α + Cu2O 0.01 0.1 Oxygen, wt%
Temperature, ⬚F
82 / Principles of Liquid Metal Processing
1940 1922 (0.7)
A portion of the copper-oxygen equilibrium diagram. Source: Ref 20
Other Inclusion Types. Inclusions in aluminum alloys can occur from grain-refining operations in which the grain-refining additions may not be adequately controlled. Titanium, vanadium, and boron are used in various combinations to produce grain refinement, and the diborides can become undesirable inclusions. Again, the fact that a small addition is beneficial does not mean that larger additions will provide greater rewards. Many of the commercial aluminum alloy compositions have been developed to exploit their age-hardening characteristics in conjunction with an alloyed matrix. They then become susceptible to the presence of seemingly minor elements such that undesirable intermetallic particles can be formed during cooling. In those cases of precipitation during solidification, the particles are placed in the category of inclusions. The intermetallic inclusions are particularly troublesome in aluminum alloys containing large quantities of zinc or magnesium. Complexes can be formed with the transition elements. These observations suggest that the remedy is close control of the metal chemistry. There are many examples in which control is inadequate and the problem is not identified until mechanical properties are marginally out of specification. Even aluminum carbide and iron-aluminum compounds have degraded properties by the formation of inclusions. Although the solubilities are low, the problem is additive in that scrap/revert procedures can cause a gradual increase in the impurity levels, and inclusion problems become more severe. The hazards associated with recycling then become apparent, especially when the final product is sensitive to the elements present and when strict mechanical property specifications are imposed. The development of new alloys, such as the aluminum-lithium alloys, brings its own set of metal-processing difficulties that begin with making ingots for forming. The question of casting more complex shapes then becomes a natural evolution. Here, the reaction of lithium with air requires melting under liquid salts that themselves are aggressive with the surroundings, such as refractories or mold materials. The opportunity for trapped inclusions is apparent.
Copper Alloys
Magnesium Alloys
A portion of the copper-oxygen system is shown in Fig. 16. The eutectic shows that the last liquid to solidify may contain 0.39% O, even though the initial melt may contain 0.01% O. Therefore, Cu2O inclusions may be more predominant internal to the casting (liquid enrichment in oxygen due to slow diffusion) and may also be present at grain boundaries. The key to control is to add elements that combine with dissolved oxygen (that is, to form oxides with free energies lower than that of Cu2O). Examples are lithium, boron, magnesium, or phosphorus, where control of the excess addition becomes important to modification of the electrical conductivity in the high-copper alloys. The excess deoxidizer goes into solid solution. A more thorough discussion of gas solubility is available in Ref 20 (see also the article “Gases in Metals” in this Volume). A number of the copper alloys are susceptible to dross formation, and the normal precautions in pouring and gating are absolutely necessary to minimize exogenous inclusions. The drosses are usually complex oxides of copper, zinc, tin, lead, or aluminum in the aluminum bronze family. Iron is added as both a dispersion strengthener and as a grain refiner in aluminum and manganese bronzes. However, iron can become an undesirable inclusion in the other copper alloys. When a foundry pours several different alloy chemistries, scrap segregation becomes important to inclusion control. Consideration can also be given to alternative grain-refining agents such as zirconium. In some cases, the intermetallic inclusions are rich only in iron, and complexes are formed. This is especially true when alloys are high in zinc, such as the high-strength yellow brasses. The problem becomes more severe when a melt is held close to its liquidus for extended periods. In this case, the solubility of the complex intermetallics is lower, and their formation is enhanced. High-zinc alloys offer fewer problems with dissolved oxygen because of the vaporization of zinc and the purging action of the melt. Oxides as inclusions are then a lesser problem. However, zinc oxide (ZnO) can react with refractories and mold materials to produce inclusions during pouring. In addition to careful pouring and gating practices, exogenous inclusions can be removed by filtering processes. However, the practice has not been as well received for copper alloys as for aluminum alloys. Although not strictly an inclusion from the previous definition, lead in many of the copper alloys can form oxides that will react with mold bonding agents such as bentonite. These lead complexes can become leachable in standard tests, and waste sand streams that contain lead compounds may be classified as toxic, with the attendant difficulty in disposal (Ref 21).
Of the nonferrous alloys considered in this section, magnesium has the lowest free energy of formation for its oxide. Therefore, the complete prevention of oxide inclusions is impossible, and a number of methods have been developed to minimize oxide inclusion formation. Agents that evolve gas to partially insulate the melt from the atmosphere and fluxes that agglomerate the oxide (plus other impurities that are to be removed) are two examples of control of the melting practice. Again, great care must be exercised in gating to minimize reoxidation. The magnesium alloys are light in weight, and many of the potential exogenous inclusions are collected as sludge in the bottom of the furnace and ladle. Desludging operations are then necessary. The use of filtering systems, whether metallic or ceramic, is common to all of the higher-quality magnesium castings. As with aluminum alloys, excess grain-refining agents can result in undesirable inclusions that can have complex chemistries because of secondary reactions. The fluxes themselves must be reactive in order to perform their intended function. Excess flux can then react with the atmosphere, refractories, or the mold materials to produce inclusions. On the other hand, insufficient flux causes metal reaction with the atmosphere to produce oxides and nitrides that may be difficult to remove if not in an agglomerated form. Temperature, time at temperature, alloy composition, pouring practice, and gating procedures dictate the optimal fluxing process. So-called fluxless melting has been suggested for magnesium alloys in order to overcome the environmental and metallurgical problems associated with fluxes (Ref 22). Fluxless melting uses nonreactive gaseous atmospheres above the melt. The most effective protective atmosphere is CO2 plus the addition of trace inhibitors such as SO2 and SF6. On the other hand, the potential addition of fluorine-containing gases to a gas waste stream must also be of concern. Finally, the general observations and processing techniques in this article have assumed the presence of air to generate the bulk of the inclusions. Alternative melt processing can exclude air for the primary purpose of inclusion prevention. Both vacuum and inert gases are used to process alloys for more demanding applications such as metal bearings, superalloys, and metal alloy prosthetics. Further discussion is included under specific alloy systems and melt methods elsewhere in this Volume.
ACKNOWLEDGMENT From: P.K. Trojan, Inclusion-Forming Reactions, Casting, Vol 15, ASM Handbook, ASM International, 1988, p 88–97
Inclusion-Forming Reactions / 83 REFERENCES 1. L. Ackermann, J. Charbonnier, G. Desplanches, and H. Kaslowski, Properties of Reinforced Foundry Alloys, Trans. AFS, Vol 94, 1986, p 285–290 2. R.A. Flinn and P.K. Trojan, Engineering Materials and Their Applications, 3rd ed., Houghton-Mifflin, 1986 3. F.R. Mollard and N. Davidson, Experience with Ceramic Foam Filtration of Aluminum Castings, Trans. AFS, Vol 88, 1980, p 595–600 4. L.A. Aubrey, J.W. Brockmeyer, and P.F. Wieser, Dross Removal from Ductile Iron with Ceramic Foam Filters, Trans. AFS, Vol 93, 1985, p 171–176 5. L.A. Aubrey, J.W. Brockmeyer, P.F. Wieser, I. Dutta, and A. Ilhan, Cast Steel Improvement by Filtration with Ceramic Foam Filters, Trans. AFS, Vol 93, 1985, p 177–182 6. P.K. Trojan, P.J. Guichelaar, and R.A. Flinn, An Investigation of Entrapment of Dross and Inclusions Using Transparent Whirl Gate Models, Trans. AFS, Vol 74, 1966, p 462–469
7. D.G. Schmidt, Gating of Copper Base Alloys, Trans. AFS, Vol 88, 1980, p 805–816 8. A.D. Wilson, Calcium Treatment of Plate Steels and Its Effect on Fatigue and Toughness Properties, Proc. 11th Annual Offshore Technology Conference, OTC 3465, May 1979, p 939–948 9. A.D. Wilson, Effect of Calcium Treatment on Inclusions in Constructional Steels, Met. Prog., Vol 121 (No. 5), April 1982, p 41–46 10. The Making, Shaping and Treating of Steel, 9th ed., United States Steel Corporation, 1971 11. C.W. Briggs, The Metallurgy of Steel Castings, McGraw-Hill, 1946 12. R.A. Flinn, Fundamentals of Metal Casting, Addison-Wesley, 1963 13. C.E. Sims and F.B. Dahle, The Effect of Aluminum on the Properties of Medium Carbon Cast Steel, Trans. AFS, Vol 46, 1938, p 65–132 14. W.O. Philbrook, M.C. Flemings, L.H. Van Vlack, A. Gittins, and J. Lankford, Oxide Inclusions in Steel, A series of five papers, Int. Met. Rev., Sept 1977, p 187–228 15. Low Sulfur Steel, Symposium proceedings, AMAX Inc., 1986
16. S.K. Paul, A.K. Chakrabarty, and S. Basu, Effect of Rare Earth Additions on the Inclusions and Properties of Ca-Al Deoxidized Steel, Metall. Trans. B, Vol 13 (No. 2), June 1982, p 185–192 17. M. Olette and G. Gatellier, Effect of Calcium, Magnesium or Rare Earth Additions on Cleanness of Steel, Rev. Metall. Cah. Inf. Tech., Vol 78 (No. 12), Dec 1981, p 961–973 18. R.A. Flinn, L.H. Van Vlack, and G.A. Colligan, Mold-Metal Reactions in Ferrous and Nonferrous Alloys, Trans. AFS, Vol 94, 1986, p 29–46 19. D.R. Askeland and P.K. Trojan, The Approach to Equilibrium and Dross Formation in Nodular Cast Iron, Trans. AFS, Vol 77, 1969, p 344–352 20. Casting Copper Base Alloys, American Foundrymen’s Society, 1984 21. T.R. Ostrom, B.P. Winter, and P.K. Trojan, Lead Transfer from Copper Base Alloys into Molding Sand, Trans. AFS, Vol 93, 1985, p 757–762 22. J.W. Fruehling and J.D. Hanawalt, Protective Atmospheres for Melting Magnesium Alloys, Trans. AFS, Vol 77, 1969, p 159–164
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 87-98 DOI: 10.1361/asmhba0005195
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Electric Arc Furnace Melting*
THE MELTING OF STEEL is performed in both arc furnaces and induction furnaces. The arc furnace is the primary melting method and is used by over 80% of steel foundries (Ref 1). However, arc furnaces are not good holders or efficient at superheating metal, so it is not uncommon to use induction furnaces to superheat steel after the arc melters. Induction furnaces used to melt steel require cleaner charge materials to meet metallurgical requirements and therefore require low-carbon charge to produce lower slag levels. The basic concept of an electric arc melting furnace is a refractory-lined steel vessel having a bowl-shaped hearth and a domed-shaped refractory roof. The hearth and sidewalls are typically lined with bricks, monolithic materials, or a combination of both. Acid refractories (silica-based), basic refractories (magnesitebased), and, in some cases, neutral refractories (alumina-based) are used to line the furnace sidewalls, hearth, and roof. All major additions are made to the furnace through the roof by swinging the roof out of the way and charging scrap, fluxes, and carbon from loaded buckets. The furnace has two or three electrodes mounted vertically through the roof of the furnace that float just above the surface of the cold ferrous scrap or molten bath. Over 80% of arc furnaces used for steel castings are alternating current furnaces with three electrodes. These electrodes strike an arc on the cold metal scrap or molten metal to impart energy to the bath and melt or superheat the furnace. Control of the electrode level above the bath is critical to the melting process efficiency. Arc melting furnaces are a widely used method for melting steel scrap. They can melt a wide variety of ferrous scrap materials and can accept certain levels of dirty scrap. They also use chemical energy to assist in the melting process and to lower the carbon levels of the molten metal. Since steel castings require low-carbon-level metallurgy, carbon from scrap metal feedstock and carbon electrodes must be removed from the molten metal. Injecting oxygen into the molten bath oxidizes carbon and uses the exothermic reaction to assist in melting and superheating the metal. This is referred to
as supplying chemical energy to the furnace. In some instances, different forms of oxyfuels are used to provide additional energy to speed up the melting process. Although refractories line the furnace shell and roof, significant energy losses occur through the surfaces of the furnace and from exposed surfaces during charging and slagging operations. Energy losses from the surface of the bath are due to the large surface-area-to-volume ratio of the furnace, resulting in significant energy losses during delays in the operation. In addition, the large surface area of the furnace results in large losses through the refractory walls, hearth, and roof. Nearly all steel foundries blow oxygen into the bath after most of the scrap is melted, which generates chemical energy from exothermic decarburization, assisting in the melting process. Some operations use additional sources of chemical energy to decrease tap-to-tap time. Electric arc furnaces are used as melters and holders in duplex operations and as melting and refining units. Since its first appearance as a production tool at the beginning of the 20th century, the electric arc furnace(EAF) has become one of the primary melting tools used by foundries and steel mills. The early furnaces had capacities of 910 to 14,000 kg (1 to 15 tons), but new technology has vastly increased EAF productivity. There are currently approximately 100 minimills in the United States capable of producing 50 million or more tons per year, with numerous furnaces capable of producing in excess of 100 tons per hour (Ref 2). This article focuses on the construction and operation of these furnaces and their auxiliary equipment in the steel foundry (shape casting) industry. Primary steelmaking practices differ from that of smaller foundry furnaces, and most primary steelmaking EAF operations have adopted separate practices since the 1980s. For example, methods of improving the efficiency of arc melting furnaces in the primary steel industry are not directly transferable to the smaller furnaces used by steel shape casting producers. Oxygenenriched burners and foaming slag methods are not considered effective on the smaller furnaces typical of steel foundry operations. For primary
melting, the reader is referred to The Making, Shaping, and Treating of Steel, 11th ed., published by The AIST Foundation. More information on ladle metallurgy is also contained in the article “Steel Melt Processing” in this Volume.
Power Supply The current applied to the EAF is supplied by a local electric utility company and passes through an electrical substation that is designed specifically for the furnace(s) to be supplied. A simplified diagram is shown in Fig. 1. The power supplied to the furnace during melting is provided by an electrical arc
Fig. 1
Schematic of electrical network for the electric arc furnace
*Revised from Melting Furnaces: Electric Arc Furnaces, Casting, by Nick Wukovich, in Vol 15, ASM Handbook, ASM International, 1988, p 356–368 from Reviews by Peter Glaws and Craig Darragh, The Timken Co.
88 / Melting and Remelting established from three carbon or graphite electrodes. During the meltdown portion of the heat, the three arcs or phases act as a single phase. Two arcs can strike the furnace charge and draw current without a current going to the third arc or electrode. Because of the type of scrap used and the arc lengthening and shortening that take place, there is a great fluctuation in the current during meltdown, which causes a significant variation in the electrical supply system. If the same power source is the supply for other plant power, a flickering of lights and voltage fluctuation will be noticed on machinery and electrical equipment. During the refining cycle or after meltdown, the arcs tend to stabilize, partially because of a slag cover on the liquid metal and the flatness of the bath. In addition, the arcs have been shortened to direct their energy in a smaller area. A great amount of energy is produced during this meltdown and refining of ferrous metals. Controls for harnessing and directing the energy of the arcs are required in order to produce molten iron and steel in the electric furnace without destroying the furnace refractories. The energy requirements for melting various carbon levels in iron or steel are shown in Fig. 2. Power Factor. The efficiency with which power is transferred to the melt is called the power factor. The power factor, PF, is the ratio of watts, W, divided by volt-amperes, VA: PF ¼
W VA
have been chosen. The electrical system for the furnace includes the transformer and the control panel. Very small furnaces (<910 kg, or 1 ton) are not used in current production schedules; heats of these sizes are normally made in induction furnaces. Furnaces are described as supplying so many tons or pounds per heat under normal conditions, but larger heats can be made by modifying
the furnace. This is done by stopping off the tap hole and raising the breast on the working and charge door (if the furnace has one). The furnace is lined with refractories that determine the type of melting practice (acid or basic) that will be used (Fig. 3). Table 1 lists various refractory furnace materials used in acid and basic melting practices. Depending
100
One method of measuring the average power factor over a period of time requires the use of two separate meters: a watt-hour meter to accumulate the useful power and a var-hour meter to accumulate the reactive power. The var-hours (in a given time) divided by the watt-hours (in the same time) will give the tangent of the power factor angle. The cosine of this angle is the power factor. The local utility company can be contacted for this information and for recommendations on the effect on energy usage.
Fig. 2
Power consumed in melting iron and steel in the electric arc furnace. Values will vary depending on scrap, transformer, lining, and so on. The melting point of pure iron (0.0% C) is 1535 C (2795 F); of iron containing 4.3% C, 1130 C (2066 F).
Arc Furnace Components The list of equipment associated with an EAF can be extensive. This list is assembled as if a proposal for the furnace has been made. The locations for the furnace foundation and pits or elevated platforms are selected. Factors that also must be considered include the location of the furnaces in the plant, flow and access to raw materials, storage of melt materials, ladle construction, ladle preheating, crane runways, water, air, electrical lines, transformers, laboratory, and temperature equipment. This list is not all-inclusive. Fume and dust collection equipment can be added, as well as used-refractory and slag removal equipment. For the melting portion, a supply of oxygen, hoses, pipe for oxygen blowing, gages, valves, thermocouples, and so on will be needed. The furnace itself is the primary concern after the location, cement work, and electrical supply
Fig. 3
Cutaway view of electric arc furnace showing the typical refractories used
Electric Arc Furnace Melting / 89 on whether the process is basic or acid, slag color may be sufficient to determine the iron oxide content of the slag, as shown: Color
FeO, %
CaO/SiO2
Acid slag Gray-black Dark green Streaked dark green Blue-green (jade green)
30–40 20–30 15–20 <15
... ... ... ...
Basic slag Black Dark brown-gray Light brown Gray White White
14.85 11.50 1.05 1.05 0.56 0.51
3.33 2.19 1.91 1.88 2.01 2.53
The silica content of basic slag can profoundly affect the melt. If present in sufficient (quantities, silica can acid constituent) can significantly reduce the slag’s ability to effectively desulfurize the steel. The size of the furnace can vary from 450 to 91,000 kg (0.5 to 100 tons) or larger. Typical furnace data are given in Table 2.
Table 1
delta section of the roof and can cause excessive refractory wear in this area. The use of water-cooling rings around the electrode ports can help to reduce this wear. The electrodes should be vertical from all sightings around the furnace. When the electrodes are not vertical, the distance and arc flare on the sidewalls are not uniform, and a hot spot may cause the refractory wear to be greater in the areas closer to the electrodes. In addition, because the arc works between the electrodes and the bath, the current per phase can fluctuate (Fig. 5). The roof structure is arched to be selfsupporting and to keep the greatest distance between the arc action in the delta section. Refractory roofs start with a roof ring that is designed as a channel or an angle beam that is rolled or cast into a ring. The ring generally has a water-cooling chamber on all or a portion of its height (Fig. 6). Newer designs of watercooled roofs are cast in gray iron except for the delta section and the exhaust and oxygen lance holes on large roofs, and they are cooled by a network of steel pipes running throughout the structure. The delta section and electrode ports are
The electrode can be graphite or carbon. The graphite electrode has greater current capacity than the carbon unit for the same size electrode. Table 3 lists some data on electrode sizes, weights, and current capacities. The electrodes are supported on three separate arms and held to the electrode arm by a spring-clamped, pneumatically released electrode holder. On older furnaces, the electrode holder had a mechanical wedge-type clamp; on newer, larger furnaces, power-operated clamping devices are used. The electrode arms support the electrodes, holders, bus tubes, and transformer power cables. Smaller bus tubes and power cables can be used if they are air or water cooled. The electrodes are raised and lowered by manual or automatic controls. These controls can be operated manually when the bottoms of the electrodes are raised to the roof or when electrodes are shifted. The electrodes are raised and lowered automatically during the melting and refining process (Fig. 4). The furnace designer will make the electrode circle as small as possible to increase the speed of melting. This will intensify the heat in the
Refractories as a source of slag based on typical composition Composition, %
Type
Fireclay Super duty Medium duty Semi silica Alumina type (high) 50% 60% 70% 80% 90% Mullite Corundum Silica type Silica super duty Conventional Basic type Chrome Magnesite Magnesite high periclase Chrome-magnesite
Table 2
Softening point
SiO2
Al2O3
Cr2O3
TiO2
CaO
MgO
Fe2O3
Other
C
F
49–56 57–70 72–80
40–44 25–38 18–26
... ... ...
1.5–2.5 1.3–2.1 1.0–1.5
... ... ...
... ... ...
... ... ...
2.5–4.0 4.0–7.0 1.0–3.0
1745–1765 1660–1685 1640–1685
3175–3210 3020–3065 2985–3065
41–47 31–37 20–26 11–15 7½–9 18–34 0.2–1.0
47½–52½ 57½–62½ 67½–72½ 77½–82½ 89–91 60–78 98–99.5
... ... ... ... ... ... ...
2.0–2.8 2.0–3.3 3.0–4.0 3.0–4.0 0.4–0.8 0.5–3.1 Trace
... ... ... ... ... ... ...
... ... ... ... ... ... ...
... ... ... ... ... ... ...
3.0–4.0 3.0–4.0 3.0–4.0 3.0–4.0 1.0–2.0 1.0–3.0 0.3–1.0
1785 1805–1820 1820–1850 1865 1930 1850 2000
3245 3280–3310 3310–3360 3390 3505 3360 3630
95–97 94–97
0.15–0.35 0.45–1.20
... ...
... ...
2.5–3.5 1.8–3.5
... ...
0.3–2.2 0.3–0.9
0.02–0.10 0.10–0.30
1680–1700 1635–1665
3060–3090 2975–3025
3.0–6.0 0.7–1.0 0.5–5.0 4–8
15–34 0.3–1.5 0.2–1.0 16–27
28–33 ... ... 18–28
... ... ... ...
... 1.0–3.5 0.5–1.5 0.7–1.5
14–19 85–93 92–98 27–53
11–17 0.3–7.0 0.2–1.0 ...
1.0–2.0 0.5–1.0 0.0–0.6 ...
1290–1425 1480–1675 1595–1705 1595–1675
2350–2600 2700–3050 2900–3100 2900–3050
Typical dimensions and capacities of foundry-sized electric arc furnace components Refractory
Shell dimensions (inside) Diameter m
1.5 1.8 2.20 2.4 2.67 2.74 3.05 3.35 3.81
Dimensions
Depth
Melting rate
ft
m
ft
5.0 6.0 7.25 8.0 8.75 9.0 10.0 11.0 12.5
1.14 1.4 1.7 1.8 2.01 2.06 2.3 2.51 2.84
3.75 4.5 5.5 6.0 6.58 6.75 7.5 8.25 9.33
kg/h
450–640 910–1,270 1,360–1,910 1,810–2,540 2,270–3,180 2,720–3,810 3,630–5,440 5,440–8,160 7,260–10,900
Sidewall
Liquid metal capacity
Bottom
Normal
Maximum
Electrode diameter
lb/h
mm
in.
mm
in.
kg
lb
kg
lb
mm
in.
Transformer capacity, kVA
1,000–1,400 2,000–2,800 3,000–4,200 4,000–5,600 5,000–7,000 6,000–8,400 8,000–12,000 12,000–18,000 16,000–24,000
254 254 254 254 368 368 368 368 368
10 10 10 10 14.5 14.5 14.5 14.5 14.5
305 381 432 457 483 483 508 508 533
12 15 17 18 19 19 20 20 21
1,360 2,810 4,260 6,080 9,070 9,840 14,500 22,700 29,900
3,000 6,200 9,400 13,400 20,000 21,700 32,000 50,000 66,000
1,810 3,750 5,670 8,120 12,100 13,200 19,100 29,900 39,900
4,000 8,260 12,500 17,900 26,700 29,000 42,000 66,000 88,000
75 152 178 203 229 229 254 305 356
3 6 7 8 9 9 10 12 14
1,000–1,500 1,500–2,000 2,000–3,000 3,000–5,000 4,000–6,000 4,000–6,000 5,000–9,000 7,500–12,500 10,000–15,000
90 / Melting and Remelting
Table 3 Electrode sizes, weights, and current capacities Diameter length mm
Approximate weight of electrode and nipple in.
Graphite electrodes 102 1015 4 40 152 1220 6 48 152 1520 6 60 178 1220 7 48 178 1520 7 60 203 1220 8 48 203 1520 8 60 229 1520 9 60 254 1220 10 48 254 1520 10 60 305 1520 12 60 305 1830 12 72 356 1520 14 60 356 1830 14 72 Carbon electrodes 203 1520 254 1520 305 1520 356 1520 356 1830
8 60 10 60 12 60 14 60 14 72
Approximate current capacity, A
kg
lb
1,800–3,300 3,500–5,800 3,500–5,800 4,400–7,500 4,400–7,500 5,500–9,300 5,500–9,300 6,700–11,300 8,000–13,300 8,000–13,300 11,300–17,000 11,300–17,000 18,000–25,000 18,000–25,000
13.6 37.6 47.1 46.3 59.9 62.1 77.1 96.6 97.1 118 171 201 228 273
30 83 104 102 132 137 170 213 214 261 378 444 502 602
2,500–4,500 3,100–6,300 4,500–7,900 5,400–10,000 5,400–10,000
79.8 125 178 240 293
176 275 392 528 647
usually constructed of a high-alumina refractory brick, rammed material, or precast shapes. The roofs of all modern electric furnaces are designed for top charging. The roof and electrodes are raised several inches above the top of the furnace shell and swung to one side. The charge is dropped into the furnace, and the roof is swung back into place. The raising of the roof is done by electrohydraulic systems. Schematics of roof-lifting devices are shown in Fig. 7. Tilt Control. The electric furnace is tapped out by tilting the furnace into a position in which the metal and slag will run out of a hole above the slag line in the furnace wall down into a spout and into a ladle. Tapping of the furnace requires that the roof and electrodes remain attached to the furnace shell; therefore, safety locks or stops are found welded to the shell and roof ring. The tilting mechanism can be mechanical for very small furnaces or can be motor driven (either by a screw pitman or a rack-and-pinion mechanism) or hydraulic, and it has rockers attached to the furnace and rails attached to the foundation. Emergency equipment is usually provided to pour the heat from the furnace in case of a power failure. The furnace shell is typically made of carbon steel sections that are rolled and welded to the diameters listed in Table 2. The diameter and height define the furnace size. The furnace bottom is formed and welded to complete the pot. The shell will have door openings cut into it and a cooling frame welded around the door. The door-lifting supports will also be welded onto the shell. The furnace rockers and other lifting equipment attachments are welded onto the shell. With the rockers and door aligned, the spout opening is cut out of the shell and the spout frame attachment welded onto
the shell. To ensure that the furnace is properly grounded, straps, rods, or mesh are attached to the bottom on the inside and ground on the outside prior to lining of the furnace. The choice of refractories will depend on the melting practice selected (Fig. 3 and Table 1). Furnaces built in the 1980s and beyond have used water-cooled panels in the furnace sidewalls above the slag line (Fig. 8). Spout and Tap Hole. The length and shape of the furnace spout have historically created design issues. The melter and metallurgist want the shortest possible spout, while the design engineer is confronted with a network of electrodes, crane cables, pit dimensions, and ladle and tap hole dimensions needed for positioning the ladle in order to obtain a fast and smooth tap from the furnace, using an acceptable spout. New designs and changing the locations of tap holes may reduce the physical and mechanical problems encountered with the spout. One of these alternatives is the side slide gate tap hole (Fig. 9), and another is eccentric bottom tapping with the use of a slide gate (Fig. 10). These alternative devices are primarily used to eliminate furnace slag carryover. In order to do so, they must use a hot heel practice. Eccentric bottom tapping does a very good job of preventing slag carryover at the beginning of pouring. Both do a good job when an effective hot heel practice is employed. Tap-Out and Back-Slagging Pits. The design of the furnace pit should include sufficient depth for the slag boxes, and along with this depth, provision must be made for preventing ground water from entering the pit. Sump pumps are needed to minimize the presence of water and to ensure an environment that is safe from steam explosions. The length and width of the pit should be sufficient to allow movement
of the ladle; this is necessary if the furnace slag is to be kept out of the ladle so that ladle metallurgy or secondary refining can be done following tap out. Water-Cooling System. Water-cooled sidewall panels have become common practice. Even before this innovation, water cooling was used on: Electrode holders, arms, and clamps Roof rings and electrode rings in the delta
section
Around the door Cooling panels positioned on the outside of
the furnace to cool the furnace wall refractories Preheat and Furnace Scrap Burners. Preheating scrap with natural gas or flue gases helps reduce the electrical cost of melting a charge. On the electric furnace, this preheating may also help stagger furnace operation to keep electrical demand peaks to a minimum. Furnace oxygen/natural gas burners can help melt the charge and reduce electrical charges and electrode usage while shortening the melt time from tap to tap. Improper use of these burners can cause excessive oxidation of the heat and the electrodes and can reduce the life of the refractory lining. Fume and dust collection at the electric furnace has been a required practice in melting during the last 25 to 30 years. The complexity of the fume and dust collection equipment varies from a simple fourth-hole connector to an entire separate room for the melting furnace in which the fumes and dust are collected and where the furnace is tapped into a ladle that is brought to the furnace and then removed through a dust protection door. Figure 11 shows typical dust and fume collectors. The charge bucket is sized to fit within the refractory walls of the furnace and is open at the top, with a lifting bale that folds back or is removable from two trunions. The bottom of the charge bucket can be designed to drop the charge in one of several ways: The clam shell bottom is actuated with a trip
lever.
The orange-peel bottom, which is set in a
stand to fold the leaves back and is tied with a rope, burns off when the charge bucket is placed into a hot furnace. The preheat charge bucket uses a trip lever and, after being charged with scrap, is preheated to some temperature above 260 C (500 F) to up to 815 C (1500 F). Figure 12 shows a typical charge loaded in two different types of charge buckets. Ladles may be part of the equipment assigned to the electric furnace, or they may be assigned to the pouring crew or the mold department. Nevertheless, the heat is not finished until the metal is removed from the furnace.
Electric Arc Furnace Melting / 91 slag will attack the refractory, as in Hadfield manganese steel. In foundry operations, the type of ladle used varies. For holding ladles, bottom-pour, teapot, and lip-pour ladles are used. For pouring ladles, bottom-pour, teapot shanks, lip-pour shanks, and drum ladles are used. The ladle linings have changed from refractory lining to the one-heat type with insulating linings made of insulating board or of one-piece designs on shank ladles to eliminate preheating. The repeated use of the refractory reduces cast steel cleanliness and quality, but the board ladles have overcome this problem (Fig. 13).
Weighing and Chemical Analysis Scales. Charge and alloy scales are required for weighing the charge and alloy additions. The charge scale can be the standard beam scale type or the newer load cell type. These scales should be checked daily or weekly to ensure accurate weighing of the furnace melt. The alloy scale can be a small spring or a large balance beam type capable of measuring in pounds and ounces. These should be periodically checked for accuracy. Quality-Control Testing Equipment. The demands for faster and more accurate chemical analysis methods have spawned advancements in the spectrographic, x-ray defraction, and scanning electron microscope analysis of metals. Carbon, sulfur, alloy, and residual analysis is quick and accurate. Gas analysis, oxygen probes, and nitrogen and hydrogen analysis equipment are within the reach of the average melt shop to assist in quality control. Temperature-control immersion thermocouples that are easy to use and recorders that are accurate to within 3 or 6 C (5 or 10 F) are available.
Scrap and Alloy Storage Purchase scrap, shop returns, and alloys should be stored in a covered area. Wet scrap will produce small errors in weight and increase the gas content of the melt. Purchase scrap should be stored by size and the density of its bulk. Scrap should be stored according to specification or alloy content. Alloys and melt additions, that is, slag formers, fluxes, and deoxidizers, should be organized by usage.
Acid Melting Practice Fig. 4
Schematics of apparatuses for raising and lowering electrodes. (a) Arrangement of electrode and mast. (b) Electrode arm arrangement with power-operated electrode holder in which clamp shoe is held in place by a push rod mechanism. (c) Typical arrangement of electrode arm using power-operated electrode holders; pressure on clamp shoe is exerted by spring-operated pull rod.
In foundries and steel mills, the ladles are refractory lined—usually with an alumina lining. More than 72% alumina is used if a clean
steel low in inclusions due to ladle refractory is desired. In one application, magnesite or olivine is used in situations where the metal and
The melting operation from charge to tap out will vary, depending on the type of charge materials and alloy additions used to produce steel in the acid melting practice. The furnace lining will be made from silica or alumina refractories. These refractories determine the type of slag generated during melting, along with the dirt or silica sand included in the charge. The acid practice is a
92 / Melting and Remelting
Fig. 5
Effect of arcs on furnace refractories due to electrodes not being vertical (a) on section left of center and (b) running too high a tap after meltdown
Lower refractory costs due to the initial cost
and the refractory life
Lower power consumption as compared to a
two-slag basic practice
Greater tonnage per unit of time
Acid Steelmaking Steelmaking practices can be classified as partial-oxidation or complete-oxidation operations. These practices are differentiated by the amount of carbon removal during oxidation compared to the final desired carbon content of the melt. The advantages of acid melting practices using partial oxidation are: Rapid steel production Good refractory life due to lower iron oxide
content in the slag
Lower power input Lower electrode consumption No recarburization required
The disadvantages include: Insufficient removal of gases Occasional poor fluidity due to oxides Occasional poor mechanical properties
The advantages of acid melting practices using complete oxidation are:
Fig. 6
Typical roof ring with cooling face. (a) Designed for use with special-shaped or cut starter brick. (b) Designed for use with skewback starter brick. (c) Designed for use with special-shaped or cut starter brick
fast oxidizing operation that does not remove phosphorus or sulfur from the melt. Therefore, charge materials with low phosphorus and sulfur levels should be selected in order to meet the final chemical analysis of the steels produced. The acid melting practice is selected for the following reasons: Flexibility in producing small quantities of
steel to various chemical compositions
Low gas content Good fluidity Good mechanical properties Elimination of nearly all silicon and most of the manganese to a uniform base metal to which additions can be calculated more closely
The disadvantages are: Slower steel production Shorter refractory life due to higher iron
oxide content in the slag
Higher production cost
Fig. 7
Roof-lifting devices and swing mechanisms used on top-charging electric furnaces. (a) Roof-lifting device and swing arm mounted separately from furnace. (b) Roof-lifting device and swing mechanism attached to furnace. (c) Roof lifting by cylinder and cable/swing system by rotating roof with a kingpin arrangement
In partial oxidation, the charge carbon content is 0.05 to 0.15% above the finishing carbon content. If the carbon is above this level, a short blow with the oxygen lance is required. No slag-forming materials are added to the bath or the charge, and a fast melting practice is used. The rust and scale in the charge, along with the oxidation of the charge during melting, are sufficient to produce a light boil. When melting is complete, the carbon content will be approximately 0.25 to 0.35%, and the silicon content will be 0.15 to 0.20%; the actual value depends on the amount of oxidation from the charge and the oxidation during meltdown. By the time the laboratory report on the melteddown carbon and manganese is prepared, the carbon content will have dropped another two
Electric Arc Furnace Melting / 93 or three points, and by the time the tap temperature is reached, the carbon will have dropped so that any alloy additions will bring the carbon content to the desired finishing range. Slag additions are seldom required. The slag will be 3 to 5% of the charge weight because of silica dirt and refractory erosion. Foundries using partial oxidation have minimal mechanical test properties to meet. They
use a good grade of scrap, and their primary concern is the high production of molten steel from their furnace capacity. This practice is losing favor because of quality standards and product liability laws affecting the finished product. Complete Oxidation. Most steel foundries using an acid melting method employ the complete-oxidation practice. This is because nearly all steel castings must meet minimum
mechanical test requirements, along with magnetic-particle, visual, and radiographic inspection. In this process, rapid melting is necessary to start a boil during meltdown, followed by oxidation with the use of the oxygen lance. Some melters add iron ore or mill scale to the charge to obtain a partial boil during meltdown. This boiling action or oxidation is accomplished with the following reaction: FeO þ C ! Fe þ CO
Fig. 8
Fig. 9
Furnace design incorporating a water-cooled roof and upper sidewalls. The location of the water-cooling passages depends on manufacturer’s suggestions.
Furnace slide valve used to hold back slag
Fig. 10
Slag-forming materials are not added to the charge or the bath. Slag is formed from silica and dirt in the charge and erosion from the furnace refractories. A fluid black slag should be obtained by prudent additions of mill scale to ensure consistent silicon and manganese contents at meltdown. Meltdown Period. The meltdown carbon content should be 0.20 to 0.40% above the aimed carbon level. The meltdown carbon level will drop 20 to 40 points after oxygen lancing to approximately 0.20% C. The silicon and manganese contents at meltdown will be approximately 0.13 to 0.20% and 0.25 to 0.35%, respectively. After oxygen lancing, the silicon content will be 0.02 to 0.10%, and the manganese content will be 0.07 to 0.15%. A vigorous boil (oxygen lancing) will eliminate carbon, most of the hydrogen, and small amounts of nitrogen from the bath; the elimination of oxygen depends on the carbon-iron oxide reaction during this period. The slag should be green when the boil is completed. After the boil, silica sand can be added to increase the slag volume and to reduce its iron oxide content. Lime is sometimes added to the slag to reduce some of the iron oxide in the slag. Most slag formers are added to the bath during the refining period. Lime is sometimes added to thin the slag. To minimize the loss of oxidizable elements—silicon and manganese—during
Eccentric bottom tap hole furnace (a) with slide gate attached to furnace bottom (b)
94 / Melting and Remelting melting, the percentages of these elements in the charge should be low. Overoxidizing of the heat should be avoided, and alloying elements should be added immediately before tap out. The Refining Period. After decarburization, the boil will slow to a gentle roll, and recarburization and slag additions can be made if required. The recarburizers may be low- silicon, low-phosphorus pig iron, sorel pig iron, highcarbon steel, broken electrodes, petroleum coke, and so on, and occasional dipping of the carbon electrodes. The addition of carbon gives a carbon boil: O ¼ COðgÞ
This helps keep the gas content down. Small additions of ferrosilicon and ferromanganese or silicomanganese are made to the bath at the end of this boil to hold the bath at the desired carbon level. This is known as a block. The heat should be tapped as soon as possible after this procedure and after the addition of any required alloying elements. Deoxidation is required for preventing the reaction of dissolved oxygen and carbon in producing gaseous products such as CO or CO2, into which hydrogen can find a nuclei and into which it can accumulate and form pinholes in the cast product. The first of these additions is silicon in the form of ferrosilicon or silicomanganese and then aluminum, a powerful deoxidizer. Special deoxidizers such as titanium and
Fig. 11
zirconium help tie up nitrogen, while complex deoxidizers containing silicon, aluminum, calcium, and other elements help modify inclusions.
Melting Heats and Raw Materials The preceding discussion is a general description of the acid melting process. This section discusses the heat and the materials used. The charge will consist of 30 to 50% shop returns (heads, gates, and scrap castings) and 50 to 70% purchase scrap (plate, clean punchings, railroad wheels, and so on). The weight of each group of materials is recorded on the furnace log. The charge distribution into the furnace will normally be heavy scrap and casting on the bottom, followed by light scrap and purchased plate. Nonoxidizable materials such as nickel and molybdenum can also be added with the scrap charge. Carbon raiser additions to the furnace, if desired, are normally done before the charge is placed into the furnace. The roof is swung into place after an inspection of the furnace top to ensure that the scrap does not contact the roof ring, and the roof is then locked down. Melting should proceed as soon as the charge is in the furnace. The No. 1 tap will normally be used for meltdown. To protect the roof from excessive erosion, the No. 2 tap can be used to start the electrodes into the charge, followed by a change back to the No. 1 tap. This procedure will reduce roof wear in the delta section.
Four methods of dust and fume control in electric furnaces. (a) Prepollution control ventilation for dust and fume removal. (b) Direct furnace dust and fume collection (both front view and top view are shown). (c) Total furnace hood for fume and dust collection. (d) Canopy hood for fume and dust collection
A pool of metal will form on the bottom of the furnace as the electrodes burrow through the charge. Once the electrodes have reached this pool of metal, the remainder of the charge will be melted by radiation from the pool and the arc splash off of the electrodes. Rust and any iron ore added to the charge will help in forming a watery slag with the sand, eroded refractory, and slag remaining from the previous heat to form an oxidizing slag on the pool of metal produced from the melting portion of the heat. When the charge is completely melted, the tap is changed to No. 2 or lower to reduce arc splash on the furnace walls and roof. A laboratory sample is taken to determine the carbon level of the bath in preparation for the blowdown, or oxygen lancing. The oxidation following meltdown should not begin until a temperature of at least 1540 C (2800 F) is reached. When heats containing chromium are to be made, temperatures of 1595 to 1620 C (2900 to 2950 F) should be reached before lancing to minimize the loss of chromium to the slag, from which it cannot be recovered. Depending on the amount of carbon, the bath temperature, and the amount of oxygen present in the bath, a decrease of 0.08 to 0.10% C will occur for each minute of oxygen injection. Generally, a steel pipe is used as the lance (ceramic-coated steel pipes are also available). The pipe must be clean and free of grease and oil. The blow time is recorded on the furnace log. After the oxygen blow, a lab sample is taken, and the heat is prepared for finishing and tap out. When the lab test is returned to the melter, adjustments to the carbon level will be made by recarburizing or by additional lancing to the required carbon level. A final check on carbon is made, and the heat is blocked with ferrosilicon. This silicon addition will bring the bath silicon content up to approximately 0.10% or slightly higher to stop any remaining carbon boil. Final additions are made to the furnace to adjust the bath to the final chemistry. These additions may include ferrosilicon, ferromanganese, ferrochromium, ferromolybdenum, silicomanganese, nickel, or other alloys. Following these additions, the bath is heated for a few minutes to ensure that the alloys have melted and that the desired tap temperature has been attained. The bath is quickly tapped out into a ladle and the refractory is preheated to minimize temperature loss and skulling from the refractories. Any alloy additions required to meet the final chemistry can be added to the ladle before tap out. The furnace should be tapped so as to hold back the furnace slag and to give a clean full metal stream. The furnace spout should be smooth, clean, and dry in order to obtain a full metal stream without a spray into the ladle. Deoxidation of the steel in the ladle can be accomplished in several ways. One method is to add the aluminum as ingots to the bottom of the ladle just before tap out with other ladle additions. In another method, a steel bar with stars or doughnut ingots attached to it is plunged into the ladle when the ladle is one-half to
Electric Arc Furnace Melting / 95 of rammable material. Silica materials or materials high in silica will cause melting and refining problems and will attack the basic lining materials. The slag additions for basic melting are lime and high-lime-containing minerals. The basic practice can remove phosphorus from the steel. Sulfur removal with the basic practice can be very successful if a very high lime-to-silica ratio (2.5 to 3.5) is maintained in the refining slag or if a synthetic desulfurizer is used at tap out in the ladle. Because of the ability of slag to remove phosphorus (by oxidation) and sulfur (by reduction), lower-quality scrap can be used in the charge. A great portion of the oxidizable elements can also be recovered from the lime slag during refining. For these reasons, the basic melting practice is the preferred method for the production of highly alloyed steels, stainless steels, high-manganese steels, and specialty alloys. However, the basic steelmaking practice will be slower than the acid melting practice when its abilities to reduce the phosphorus and sulfur from the steel and to reduce alloys from the slag are employed to capacity.
Basic Steelmaking
Fig. 12
Schematics of foundry charge buckets. (a) Orange-peel bottom. (b) Clam shell bottom
three-quarters full. Small additions of less than the weight of an ingot are made with aluminum shot or wire clippings weighed into a bag and tossed into the metal stream at tap out. The amount of aluminum used for deoxidation ranges from 1.0 to 1.5 kg/metric ton (2 to 3 lb/ ton); the higher-carbon heats require less aluminum, while the lower-carbon heats require more aluminum because of their higher oxygen levels in the steel. All alloy additions should be recorded on the furnace log with the wattmeter reading, along with any comments on the heat. Other deoxidizers are often used to provide a complex deoxidation of the steel. These include titanium and zirconium to tie up nitrogen plus calcium and barium to modify inclusions. The type and amounts of these supplemental deoxidizers are determined according to the need and choice of each foundry. After tap out, the furnace should be inspected for slag line and bottom erosion and should be patched with silica sand or a mixture of lowgrade alumina and silica. This prepares the furnace for the next heat. Loss of Alloys. In the acid melting process, alloying elements that are oxidizable are not
completely recoverable, and care must be taken to estimate their loss when the additions are made to the bath and the ladle, as well as the loss while the heat is being melted. This loss of alloying element can be compensated for based on the data given in Table 4. Alloy Additions. The typical analysis of alloy additions made to the electric furnace is shown in Table 5. Because this is an incomplete list, alloy manufacturers should be consulted for an expanded list and additional detail with regard to residual elements, material size, and packaging.
Basic Melting Practice The melting operation from charge to tap out will vary, depending on the type of charge materials and the alloy additions used to produce steel in the basic melting practice. The lining of the furnace bottom and sidewalls is made of basic materials, such as dolomite, magnesite, chrome magnesite, bricks, and/or rammable materials. The roof is generally made of highalumina brick, and the delta section is made
The basic steelmaking practice is divided into two parts: the oxidizing period and the reducing (refining) period. Both periods are important because of the reactions that occur in each portion. The charge for the foundry practice will consist of 30 to 50% shop returns, such as heads, gates, and casting scrap; the remainder will be 50 to 70% purchase scrap. Along with these materials, a carbon raiser and lime or limestone are added to the charge. The amount of carbon raiser added should be sufficient to increase the meltdown to 0.20 to 0.45% C over the finish carbon level. In addition to the charge to the furnace, a small amount of limestone or lime is normally added to the furnace bottom along with the recarburizer. (It is sometimes considered preferable to add lime to the middle or top of the melt.) Heavy scrap is first placed on the bottom, followed by light scrap and finally medium scrap on the top of the charge. Every effort is made to get the entire charge into the furnace at one time to reduce the time loss caused by delays in backcharging the required weight into the furnace. When stainless steels are being produced, a greater portion of the chromium addition is held out of the charge to conserve chromium, to help keep carbon levels down, and to cool the bath after the carbon boil when temperatures may reach 1870 to 1925 C (3400 to 3500 F) to achieve concentrations of 0.02 to 0.08% C. Because of the negative effect of silica in the basic practice, a conscious effort should be made to keep sand out of the charge. Oily and wet scrap should also be avoided in the charge materials to minimize gas pickup. The oxidizing period will require the addition of lime to form a slag on the bath during melting until approximately 1.5 to 2.0% CaO is
96 / Melting and Remelting A high lime ratio to negate the effect of the sil-
ica in the slag; because this slag must also be fluid, fluorspar additions may be required. High concentrations of manganese oxides also hinder the satisfactory removal of phosphorus. The desired reaction for the phosphorus is: 3CaO þ 5FeO þ 2P ! ðCaOÞ3 P2 O5 þ 5Fe
Fig. 13
Typical types of insulating ladles with cold disposable liners used in foundry operations. (a) Teapot. (b) Bottom pour. (c) Teapot Shank. (d) Lip-pour shank
Table 4
Alloy loss in acid melting practice Loss from charge after blow, %
Loss when making addition, %
Manganese Silicon Nickel Chromium
80–85 90–95 0.0–2 35–50
20–30 5 0.0–2 0.0–3
Molybdenum Niobium Copper
0.0–2 0.0–2 0.0–2
0.0–2 0.0–10 0.0–2
Element
Required element weight to be added as furnace or ladle addition to make up desired weight, %
125 95 100 For less than 1% Cr, and 150% to charge; over 1% Cr, add blow for a total of 120% for >4% Cr alloy. 100 100 100
present in the charge. Once the charge has been melted and temperatures of 1540 to 1595 C (2800 to 2900 F) have been reached, the oxygen lancing can begin. If phosphorus is to be removed, another approach requiring a doubleslag practice should be used. The following conditions must be met to remove phosphorus from the melt during the oxidation period:
after
C (2800 F), because the process is reversible with high temperature. High iron oxide content in the bath and the slag; the use of iron oxide in the early stages of meltdown is required as a source of oxygen to oxidize the phosphorus. Low silicon and silica in the bath and the slag Temperature below 1540
1=3
Tricalcium phosphate is formed in the slag. This slag must then be removed from the furnace to minimize the chance for phosphorus reversion which occurs upon de-oxidation of the melt. This is accomplished by raking off the slag. Limited success is achieved in reducing the phosphorus content, as manifested by the film visible around the sidewalls of the bath following slag removal. This film may contain a major portion of the tricalcium phosphate remaining in the bath, which, if left in the furnace, allows for only minor phosphorus reductions and the reversion of phosphorus to the bath with higher temperatures. The use of slag rakes or a slag flush have become the common methods for removing the slag. After the oxidizing slag has been removed, a new lime-based reducing slag must be added to the melt. Any oxidizable element lost to the first slag must also be added back to the bath during the reducing and final tap out. In the single-slag process, oxygen lancing will remove silicon, carbon, manganese, and chromium from the bath, with manganese and chromium partially reduced from the slag in the reducing period. To minimize chromium loss, a high temperature is required before lancing; therefore, a starting oxygen lancing temperature of 1595 C (2900 F) is suggested. The silicon and manganese contents after oxygen lancing will be between 0.01 to 0.05% and 0.03 to 0.20%, respectively, with carbons in the 0.10 to 0.20% range. An addition of ferrosilicon to bring the melt chemistry to 0.10% Si will stop the boil. Reducing or refining of the heat follows the oxygen lancing of the bath. At this time, a deoxidizer is added to the bath to lower the soluble contert of the melt, after which a reduction in the sulfur can be made. With the temperature of the bath increased after lancing, additions of lime or calcium oxide and fluorspar are made to the furnace to prepare for sulfur reduction and alloy recovery. Sulfur reduction depends on: A high lime-to-silica ratio in the slag, called
the V-ratio
High temperature Low oxygen content in the steel and slag
The reaction may be expressed as follows: ðCaOÞ þ S $ ðCaSÞ þ O
This reaction is reversible. An increase in silica will reduce the activity of CaO, thus retarding the reaction, while highly oxidizing conditions in the slag (high FeO in the slag) will limit the amount of sulfur that can be removed.
Electric Arc Furnace Melting / 97 Table 5
Analysis of alloy additions for the electric furnace Composition, %
Alloy(a)
50% ferrosilicon 75% ferrosilicon Standard ferromanganese 67% silicomanganese Pig iron Low-carbon ferrochromium Molybdenum oxide Low-carbon ferromolybdenum High-carbon charge chromium Standard ferromolybdenum Electrolytic nickel Electrolytic manganese Ferroniobium Copper Ferrovanadium Armco iron Manganese calcium-silicon Calcium silicon Calsibar alloy Hypercal alloy
C
Mn
Si
Cr
Ni
Mo
Cu
Nb
Al
Ca
P
S
V
Fe
Ba
... ... 47–51 ... ... ... ... ... 1.25 max 0.20 max 0.04 ... ... ... ... ... ... 74–79 ... ... ... ... ... 1.25 max 0.60 max ... ... ... ... ... 7.0 78 min 1.0 max ... ... ... ... ... ... ... 0.35 max ... ... ... ... 1.50–2.00 65–68 16–18.5 ... ... ... ... ... ... ... 0.20 max ... ... ... ... 4.0–4.5 0.30–0.50 1.00 max ... ... ... ... ... ... ... ... ... ... ... ... 0.025 max ... 2.00 max 67–72 . . . ... ... ... ... ... ... ... ... ... ... ... ... 4–7 ... . . . 55–60 0.70 max ... Fe + Al ... ... ... ... ... ... 0.12 max 0.25 max 1.5 max ... . . . 65 min 0.70 max ... ... ... ... ... ... ... ... 5–7 ... 2–5 65–70 . . . ... ... ... ... ... 0.03 max 0.07 max ... ... ... 2.5 max 0.25 max 1.5 max ... . . . 60 min 0.80 max ... ... ... ... ... ... ... ... ... ... ... ... 99.9 ... ... ... ... ... ... ... ... ... ... 0.005 max 99.9 0.001 ... ... ... ... ... 0.002 max ... ... ... ... ... ... 0.10 ... 1.3 ... ... ... ... 55.0 + 3.0Ta ... ... ... ... ... ... ... ... ... ... ... ... ... 99.7 ... ... ... ... ... ... ... ... 0.20 max ... 1.50 max ... ... ... ... ... ... ... ... ... 55–60 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 98.0 ... ... 14–20 49–55 ... ... ... ... ... 1.75 max 14–20 ... ... ... ... ... ... ... 60–65 ... ... ... ... ... ... 28–32 ... ... ... 5.0 max ... 1.00 max ... 57–62 ... ... ... ... ... 1.5 max 14–17 0.035 ... ... 7.0 max 14–18 ... ... 38–40 ... ... ... ... ... 19–21 10–13 ... ... ... ... 9–12
(a) Aluminum is not included in this list, but it is the primary deoxidizer used in steelmaking. The melter needs to understand that there are various grades of aluminum. For example, primary aluminum in the form of ingots and shot has a purity of 99%, while grades I, II, III, and IV can contain 86–96% Al plus other nonferrous alloy elements. Care should be taken to purchase and add this deoxidizer on the weight contained and not to buy on lowestcost basis.
The slag in the basic melting practice is the key to the production of quality steel. The reduction of iron oxide by the use of coal, aluminum, graphite, or low-sulfur coke is often done to deoxidize the slag and to reduce the content of the less oxides. With this deoxidation, much of the manganese and chromium in the slag are reverted (returned to the bath). After a check on the melt or bath chemistry, alloy additions are made to adjust the chemistry with ferrosilicon, silicomanganese, and ferromanganese, along with chromium, nickel, molybdenum, and other alloys if required. During tapping, final alloy additions are made to the ladle, and the heat is deoxidized with aluminum as the primary deoxidizer plus special deoxidizers as desired.
Melting Heats and Raw Materials The following discussion is a brief review of the basic melting practice. A heat and the raw materials used are also described. Historically, the charge was placed into the furnace with the heavy scrap carbon raiser and limestone on the bottom, the light scrap in the middle, and the medium scrap on the top. Today (2008), it is often common practice to use a cushion of busheling scrap on the bottom. Nonoxidizable materials such as nickel and molybdenum can also be charged into the furnace at this time. The roof is swung into place with an inspection of the top of the charge to ensure that the scrap does not contact the roof ring, and it is then locked down. Melting should proceed as soon as the charge is in the furnace on the No. 2 tap to burrow a hole into the scrap, followed by a change to the No. 1 tap until the charge is melted. The initial use of the No. 2 tap will reduce wear on the furnace delta section.
A pool of metal will form on the bottom of the furnace as the electrodes burrow through the charge. Once the electrodes have reached this pool, the charge will be melted by radiation from the pool and the arc splash off the electrodes. Rust on the scrap, along with oxidizing scrap from the melting charge, will react with the lime and eroded refractories to make a fluid slag. When the charge is completely melted, additional limestone or lime is added to the furnace to increase the volume of slag for an estimated 1½ to 3% of the charge weight (approximately 25 kg/metric ton, or 50 lb/per ton, of lime). At this point, the melt can be processed as a single-slag heat without phosphorus removal, or it can be processed as a double- slag heat where phosphorus removal is attempted. This procedure is described earlier in this section. If a double-slag practice is chosen, the oxygen content of the slag and the bath must be increased while holding down the bath temperature. The old slag must then be removed and a new slag built on the bath. Oxygen lancing can be done before or after this stage to lower the carbon content to the required level, or it could have been done to increase the oxygen content of the bath. Reducing the Heat. The slag must be reduced in iron oxide, and because only a small amount of slag has been added to the furnace up to this point, another 1½ to 3% of charge weight of lime is added to this slag, along with a measured amount of fluorspar. This additional slag volume will dilute the iron oxide content in the slag to reduce the oxygen content and to improve the ability of the slag to retain sulfur. A deoxidizer such as coke, coal, or aluminum can be added to the slag if desired to speed the reaction. Sulfur removal and alloy recovery are the primary concerns while the bath is being heated in preparation for finishing. If a specific sulfur level is needed, further laboratory testing and lime additions may be required
for achieving a higher lime-to-silica ratio, along with fluorspar to thin the slag. If the sulfur level is acceptable, the heat, blocked with ferrosilicon and alloy additions, is finished, and a temperature check is made prior to tap out. If desulfurization is required, it should be done after the heat is blocked. The melt is tapped into a ladle that has been preheated if it is refractory lined, or it is tapped into a cold ladle liner of insulating basic material or boards. Any further alloy additions are made to the ladle prior to tap out. The furnace should be tapped to hold back the slag and to provide a clean full stream. The furnace spout should be clean, smooth, and dry to obtain a full metal stream without a spray into the ladle. Additional sulfur removal may be achieved by adding a desulfurizing agent during tapping. Reductions in sulfur content by 20 to 50% are possible. Final deoxidation of the steel should be accomplished in the ladle; this can be done in several ways. One method is to add the aluminum to the bottom of the ladle and tap onto it. Wire feeding is a popular method that has become a common practice in many steelmaking shops. Another method (poling of aluminum) is to attach stars or doughnuts of aluminum to a steel rod and plunge the rod into the ladle when the ladle is onethird to three-quarters full. (Today, 2008, poling is usually done in the ladle. Historically, it was done in the furnace with even greater effectiveness.) Small additions of aluminum should be in the form of shot or wire clippings weighed and placed in a bag and thrown into the stream while tapping. Finally, ferroaluminum additions can also be used during tapping. The amount of aluminum used for deoxidation is approximately 1.0 to 1.5 kg/metric ton (2 to 3 lb/ton), although higher carbon heats require less, and lower carbon heats require more because of the oxygen levels in the steel.
98 / Melting and Remelting After tap out, the furnace should be inspected for slag line and bottom erosion and then patched with dolomite or other materials. Patching is done by shoveling in or gunning on the patch material. This prepares the furnace for the next heat. Loss of Alloys. A wide range of melting techniques are used in the basic melting practice by foundries—for example, single slag, double slag, deoxidation of slags, desulfurization, and alloy production. Therefore, the individual foundry and melter will need to adjust the recovery and additions to meet particular needs and final chemical analysis based on experience.
Newer Technologies Arc Melting of Iron. In addition to steel melting, electric arc melting of iron is practiced. The first continuous submerged direct current electric arc (Contiarc, SMS Demag AG) furnace for melting iron was installed by American Cast Iron Pipe Company in July 2001. The key to the EAF molten iron production is that it can tap, charge, and arc continuously and simultaneously. As a result, the power does not need to be turned off to tap or charge the furnace. The submerged arc design and the continuous nature of the furnace operation make it more energy efficient than the cupola. In addition, the furnace can take either low-grade steel scrap or direct reduced iron and/or hot briquette iron and combine it with coal and silica to melt steel or smelt iron in the same furnace system, producing 73 Mg/h (80 ton/h). EAF Steelmaking. Several technologies have become common practice in most high-volume
EAF steelmaking shops as well as some foundries. For more expansive treatment of these technologies, the reader is referred to either the article “Steel Melt Processing” in this Volume or to The Making, Shaping, and Treating of Steel, 11th ed. published by The AIST Foundation. The direct current arc furnace has been introduced into steel mills and should be changing the arc melting practices of the future from the three-phase arc furnace to the single-arc direct arc practice (Fig. 14). Use of the direct current arc furnace can depend on the stability of the electricity supply in a given local area. Steel mills using the direct current arc furnace have lower electrode consumption and slightly lower operating costs. The direct current arc furnace was developed for two main purposes: An alternative to the conventional three-
phase alternating current electric arc
For the final stage in practices for melting
prereduced material for the production of liquid iron Since its development, this furnace has been used in the production of ferroalloys as well as carbon, low-alloy, and stainless steels. An existing three-phase furnace can be modified into a direct current arc furnace by the addition of a thyristor converter to its controls and the incorporation of a water-cooling system. Modification of the existing furnace requires the removal of the two outboard electrodes and the positioning of the middle electrode to the center of the roof. It has also been shown that a larger electrode will require additional
hardware to conduct current and to provide support. The electrode is the cathode in this system. In the three-phase system, current passes from one electrode to another through the charge; in the direct current system, the current passes from the electrode through the charge or bath to the anode located at the bottom of the furnace. The current from the bottom of the furnace passes through graphite- or tar-impregnated refractories to a copper-base plate to the outside cables, or through pins of steel connected to outside current cables (Fig. 14). The heat distribution in the direct arc is 360 , while in the three-phase furnace, the heat pattern is off the edge of the electrode, causing localized refractory heating and erosion. The electrode usage per ton of steel is reported to be 50 to 60% of a conventional three-phase furnace. The shell bottom of the furnace needs to be modified to accept the current-carrying loads that will result and should be insulated to protect it from shorting out. Thermocouples and cooling air will be required for monitoring overheating or loss of current-carrying ability in the bottom. The furnace foundation may also require modification to accept power cables and air cooling ducts. Patching of the refractory is limited to the slag line, sidewalls, and roof. Noncurrent refractories on the bottom will cause poor conduction. Another consideration to be made by the melter is the heat sequence, because direct arc melting requires long start-ups or a heel of metal over the bottom to make electrical contact. Ladle metallurgy has become the processing method of choice in all large EAF steelmaking shops and several foundries. As their name implies, multipurpose ladle refiners have several important capabilities: precise temperature control, vacuum degassing, inert gas bubbling/ stirring for inclusion agglomeration/flotation and compositional uniformity, and alloy addition systems for precise compositional control. Argon stirring of heats of steel in ladles as a method to promote the removal of nonmetallic inclusions via agglomeration and flotation has become a primary tool for making clean steel. The need for dry gas is very important, because argon containing more than 5 ppm of moisture will induce gas pickup. Systems for the accurate, real-time assessment of hydrogen in the liquid metal have become commonplace and are an invaluable aid to the steelmaker. Calcium wire and other wire and powder injection systems have become important tools for specific deoxidation practices, compositional modification, and inclusion engineering. REFERENCES
Fig. 14
Cutaway view showing the basic components of a direct current arc furnace
1. J.F. Schifo and J.T. Radia, Theoretical/Best Practice Energy Use in Metalcasting Operations, U.S. Department of Energy, May 2004 2. Chap. 2, Steel Industry Technology Roadmap, AISI, Dec 2001
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 99-107 DOI: 10.1361/asmhba0005197
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Cupola Furnaces William L. Powell, ThyssenKrupp-Waupaca (Retired) Alan P. Druschitz, University of Alabama at Birmingham Jim Frost, ACIPCO Melting Department
THE CUPOLA is basically a simple cylindrical shaft furnace that burns coke for fuel. The heat generated is intensified by the blowing of air through the heated bed of coke. Alternate layers of scrap metal and coke are charged into the top of the cupola. In its slow descent, the scrap metal is heated to the melting temperature by direct contact with the upward flow of the hot gases from the coke combustion. The moltenmetal droplets collect in the inner portion of the cupola known as the well, where it is discharged into the foundry from the taphole by intermittent tapping or by continuous flow of molten metal and slag. Melting in cupola furnaces dates back several centuries to old cask-type units that were blown with hand-pumped bellows and hand charged with bits of metallic iron and scraps of cast iron. The cupola as we know it today (2008) was patented in 1794 by John Wilkinson. Prior to the 1950s, the cupola was the predominant furnace for iron melting because of its simplicity and the low cost of molten iron produced. Progressive improvements were made in electric induction furnace equipment since the 1950s. At their inception, induction furnaces required little or no pollution-control equipment because mostly clean, new scrap was charged. This mix of new scrap iron and pig iron yielded a low-sulfur iron that was satisfactory for the fast-growing ductile iron industry. Electric induction furnaces also began to replace many cupolas in small-tonnage operations where both gray and ductile iron or several other types of irons were being produced. The ability to make new chemical compositions in each melted batch provided more flexibility to the foundry operator. Electric induction melting was found to have a lower melting cost and better chemical control under many of these conditions. The number of cupolas in use in the United States has decreased from 4000 in the early 1950s to less than 1400 today (2008). However, the cupola remains the predominant tonnagemelting method, with over 60% of today’s iron tonnage still melted in cupolas. Through the years, the cupola has been progressively refined
and moved toward more efficient combustion, better chemical composition control, and more flexibility in the use of widely varying raw materials. However, during the melting process, the cupola emits considerable amounts of smoke and particulate matter. Known as emissions, this waste gas stream is now cleaned to meet the highest environmental standards, but this has required much engineering effort and considerable expense. With environmental laws requiring more elaborate pollution-control equipment, the cupola lost much of its advantage of low cost and simplicity, especially on small cupolas. In high-iron-tonnage operations, the cupola remains the most efficient source of continuous high volumes of iron needed to satisfy highproduction foundries or the multiple casting machines of centrifugal pipe producers. The continuous stream of cupola-melted iron is not interrupted by charging, slag removal, or chemical composition alteration. On a per ton basis, cupolas usually require less labor to operate and can use lower-cost forms of scrap. In such high-tonnage, continuous-demand situations, the cupola has remained the lowest-cost method of melting. Although the cupola is more difficult to control, many consider it to have favorable iron qualities, such as a low tendency toward chilling, good fluidity, better machinability, and improved shrinkage characteristics compared to electrically melted cast irons.
Progressive Refinements in Cupola Equipment Since the early days of the simple cupola, continuing technological refinements have been progressively applied to the cupola melting process to improve efficiency, flexibility, and control. Many of the refinements have stood the test of time and greatly added to the cupola melting technology. Some other ideas have not proved to be technically sound or have lost appeal because of ever-changing economic conditions, industry needs, state and federal law changes, and a host of other reasons to abandon
them by the melting community. Some of the more successful improvement technologies include the following. Preheated air blast was found to improve efficiency of the combustion reaction by allowing the input of British thermal units (Btus) from an external source, thus increasing melting rate, temperature, and carbon pickup with the same amount of coke, or alternatively, the amount of coke used to melt a ton of iron can be reduced for the same melt rate, temperature of the iron exiting the cupola, or greater carbon pickup. Early blast air heating systems could preheat the incoming air only 150 to 260 C (300 to 500 F), but even this limited additional temperature demonstrated the advantages of the preheated blast. The most successful hot blast systems have been externally fired with gas and are capable of up to approximately 650 C (1200 F) preheat. Higher blast temperatures would continue to improve performance, but the availability of heat-exchange tube materials has a practical economic limit of approximately 650 C (1200 F). Recuperative hot blast systems have begun to replace the externally fired hot blast system. The recuperative hot blast system is an integrated system, which combines both the blast air heating and gas cleaning system. Emission gases are usually first combusted to raise the heat content of the offgas, then the heat is extracted via a gasto-gas heat exchanger (the hot blast portion of the system), and finally, the cooler gases are cleaned of dirt particles. In some designs, the gases are cleaned and then burned. Both have distinct advantages and disadvantages but are effectively used to preheat and clean the offgas stream. Oxygen enrichment and injection, which adds 1 to 15% O2 to the blast air, also intensifies the combustion reaction by increasing temperature, which in turn increases the melting rate. The efficiency is derived from blowing in a higher number of oxygen molecules per volume of air and lower amounts of inert nitrogen molecules. Enrichment is limited to oxygen increases of approximately 2% and is accomplished by adding pure oxygen to the blast ductwork. Injection allows much higher additions of
100 / Melting and Remelting oxygen. Increases of 4 to 15% can be achieved. A pipe is inserted into the center of each tuyere for adding the oxygen. Often, a nozzle is added to the tip of the pipe to achieve supersonic flow rates for the oxygen. This increased velocity allows blast air and oxygen to penetrate deeper into the coke bed. Injection of fine coke and other materials, such as silicon carbide, through the tuyeres has proved to be effective in both increasing carbon and silicon content and shortening response time for chemical changes in the melt composition. The effectiveness of this technology is somewhat dependent on the operating parameters of the particular melting operation. Water cooling of the cupola shell has extended the length of the melting period beyond 24 h and has permitted a single cupola to be operated for weeks before requiring a refractory repair shutdown. Water cooling of the shell allowed the removal of the refractory lining and its effect on the slag chemistry. The same cupola then could operate with a choice of basic, neutral, or acid slag chemistry simply by regulating the amount of basic flux added. The other benefit is the reduction of sulfur in the molten iron, allowing those ductile iron treatment methods requiring low-sulfur-base irons to be used with no external desulfurizing step in the molten-metal processing. Duplex electric holders are large holding furnaces that receive the iron after leaving the cupola. They have ensured that a constant supply of iron is available at all times while also reducing variations of composition and temperature of the iron. Large electric channel induction furnaces hold between 30 and 90 min of iron. These electric channel furnaces can provide a modest boost in iron temperature or maintain iron temperature over cupola melting down-periods. Computer applications have been increasingly used to record, monitor, and control more cupola conditions. Today’s (2008) melting operations are as highly computer controlled as any manufacturing operation. Virtually all control of the motors, valves, flow devices, dampers, and so on are computer controlled. All the various inputs of temperature, flow, pressures, status of valving, and so on are fed back to a computer for data storage and analysis. Continuous analyses of top gases have made it possible to monitor combustion efficiency and to signal any water leaks to avoid explosive gas combinations. Basic slags have made possible the production of low-sulfur iron directly from the cupola and have increased the carbon pickup. By regulating slag basicity, a wide range of iron compositions can be obtained from a wide range of charge materials. Technology Failures. Some technologies have not passed the test of time in melting operations. These have failed the test of providing a true increase in efficiency, being technologically sound, or may have simply been a passing fad. Many are still reflected in
the technical literature of the time, and care must be taken by individuals to determine if they are truly an improvement to the cupola melting operation. Some of these technology failures include the following. Divided-blast cupolas divide the air between two rows of tuyeres in controlled proportions. This was reported to increase combustion efficiency and accomplish either coke reduction or increased melt rate and temperature at the same coke usage. Reversed taper shell was a cupola design meant to concentrate the air across the entire cupola cross section for more uniform melting area and lower coke usage. Ground rubber tires were tried as a coke substitute. Rubber tires are largely composed of carbon and have a high Btu content. Considered a waste product, their use in the cupola would be a way to reduce both a national waste item and coke cost. However, the volatile gases produced did not provide a carbon source and also proved difficult to scrub from the emission gases. Natural gas enrichment to either the blast or the offgases was proposed as a method to raise the Btu content of the blast air. However, the water vapor released during combustion proved to have a negative effect on the heat content within the cupola. Humidity control of the blast air has been employed on some cupolas to eliminate the effect of varying humidity when very close control was important. However, the cost of drying the absorption media proved to exceed the benefits of water removal.
melting needs and inputs: type and size of scrap available in the market place, desired melting rate, melting rate flexibility needed, carbon pickup desired, use or nonuse of additional oxygen, length of desired campaign, length of time for relining, chemistry control limits, and so on. Optimum cupola height has been well studied, and 8.5 to 10 m (28 to 34 ft) of effective melting height is universally accepted as the optimum height for a cupola. Cupola shell design (both diameter and height) as well as tuyere placement, offgas take-off location, and so on require detailed forethought and planning during the design and installation of a new cupola. The foundation of the cupola structure must be adequately engineered for the size of the cupola and the loads it will carry. Footing loadings will be for both the cupola itself and the burden contained within the cupola. A large cupola will often contain 136 to 181 Mg (150 to 200 tons) of scrap iron, coke, stone, and molten iron. An often overlooked component is the
Cupola Construction and Operation This section discusses the shell, intermittent or continuous tapping, tuyere and blower systems, refractory lining, water-cooled cupolas, emission-control systems, and storage and handling of the charge materials.
Shell Design The cupola shell is a vertical, cylindrical steel shell and is the container of the combustion and melting operation. Figure 1 shows a typical modern-day cupola. The size, or specifically, the diameter, of the cupola shell basically defines the limits of operation of the cupola. It limits the minimum and maximum melting rate per hour of operation, the carbon pickup ability, the length of the melting campaign, the maximum size of the scrap that can be charged, the size of the coke purchased for melting, and a host of other parameters. In the past, cupolas were purchased as somewhat standard pieces of equipment. For the past 30 years or more, cupolas have been very much a custom-designed piece of melting equipment. Cupola manufacturers can suggest an optimum diameter based on the user’s specific anticipated
Fig. 1
Typical modern-day cupola showing the construction and arrangements of the principal cupola components
Cupola Furnaces / 101 dynamic loading place on the cupola during charging. Charge drops of 4 to 7 Mg (4 to 8 tons) of scrap falling 1.5 to 6 m (5 to 20 ft) are not uncommon. Charge doors today (2008) are almost always not a door but are actually the opening at the top of the cupola. Traditionally, cupolas were side charged near the top of the cupola. Today’s emission systems demand controlled volumes of offgases. The side charge door allows uncontrolled volumes of air to enter the gas streams, thus causing erratic cleaning of the waste gases. The top-charged cupola employs a “plug” of scrap at the top of the cupola to effectively limit the amount of air being pulled into the cupola stack. A second benefit is the ability to soft-drop a charge onto the burden, thereby reducing the crushing of the coke charge during the charge drop. Bottom doors at the bottom of the cupola stack facilitate emptying of the stack of coke ash and residual metal at the end of an operational melting campaign in preparation for refractory repairs and the resumption of melting. The doors are usually constructed of heavy steel or ductile cast iron and are massive enough to support the entire burden of the charge material and coke as well as the dynamic forces encountered during charging. These doors cover the entire bottom cross section so as not to impose any restrictions during the dropping of the bottom. During the melting operation, the doors are held rigidly in place by heavy steel posts or similar mechanisms. There can be no flexing of the doors, either mechanically or thermally, during the melting operation, or cracks in the bottom refractory will occur that may result in an iron runout. A sand bottom is rammed over the closed bottom doors to contain the melted iron in the well. This sand bottom can either be all sand, all refractory, or a combination of both. In smaller cupolas with short-duration melting campaigns, the bottom can be as simple as foundry molding sand. For larger cupolas with longer melting campaigns, special sands and clays are preferred, with the addition of a refractory layer. There is no general agreement among melters on bottom sand materials and arrangements. What can be agreed upon is that once a system is determined, strict controls must be put in place to assure quality of installation every time a new bottom is put in service. A sand bottom is sloped toward the taphole to ensure that iron and slag continuously flow to the taphole and out of the cupola.
Intermittent or Continuous Tapping Cupolas Iron flow from a cupola can be made to be intermittent or continuous. The first cupolas were intermittently tapped by opening and closing the taphole as in a blast furnace operation. For intermittent tapping, a refractory plug is pushed into the open taphole after iron has
stopped flowing, and the plug is left in the taphole until a volume of iron has accumulated in the cupola well, then the plug is pulled out and iron begins to flow. This process is repeated for the duration of the melting campaign. Slag is accumulated within the cupola well area and flows out a back slag hole. Most modern-day cupolas are tapped continuously, with iron flowing under a refractory dam and over a lip in the cupola spout. The lighter slag floats on top of the stream of molten iron. A notch in the side of the cupola spout (also called the well) allows the slag to flow out of the front spout area. This slag flows continuously to a slag disposal area. Cupola slag has traditionally been quenched in a water basin or a trough of flowing water. However, due to wastewater rules prohibiting the discharge of process water, a variety of dry slagging methods have been developed. The various dams and notches in the front spout area are constructed of refractory materials. Due to refractory wear, the levels of metal dam and slag notch must be carefully calculated and maintained throughout the melt campaign. The dry-bottom cupola is a recent European development in which iron and slag are not collected in the well of the cupola but run into a special vessel outside and beside the cupola. In this outside vessel, the accumulated metal and slag are separated by gravity as they flow through the vessel. The advantages claimed for this method include more consistent iron analysis and easier repair of refractories. Cupolas often have two vessels attached to the front, allowing the repair of one while melting iron through the second vessel. This arrangement allows for longer melting campaigns.
Tuyere and Blower Systems Air for the combustion of the coke is introduced through holes in the shell that have water-cooled pipes, called tuyeres, protruding into the cupola. Tuyeres are constructed of either cast copper or fabricated copper coils twisted around a large piece of pipe. They are equally spaced around the cupola, with the number and size proportioned to cupola diameter. Tuyeres are sized to deliver a specific relationship of air volume, pressure, and velocity. On water-cooled hot blast cupolas, the desirable blast air velocity range is 4600 to 7600 m/min (15,000 to 25,000 ft/min). This velocity has been determined to give the blast air sufficient velocity to maximize the penetration of the coke bed. Blast air blowers generate the necessary blast air volume and pressure to melt iron. The air required to melt 900 kg (1 ton) of iron is 620 to 680 m3 (22,000 to 24,000 ft3) at standard temperature and pressure or 770 to 835 kg (1700 to 1840 lb). Air from the blast air blower is transmitted to the tuyeres through a large-diameter duct system to the cupola wind box. Blast pressure, usually called backpressure, normally ranges from 3.3 to 8.5 N (20 to 65 oz/in.2) or more, depending
on charge density. Blast blowers can be of several designs, such as positive displacement blowers, fans, or centrifugal-type blowers. The latest engineering design preference on large cupolas has been to use two or three blowers in parallel with each other and use automatic air weight controls to control the air volume produced by them. Blast blowers typically consume the second-largest amount of electrical energy in a melting department, second only to the emission system fans. The biggest waste of efficiency in the melting department is due to leakage in the blast air system. This is due both to missing insulation allowing heat loss before the heated blast air reaches the cupola, and leaks in the blast air ductwork system caused by thermal and mechanical movement.
Refractory Lining Conventional cupolas are usually lined with refractories to protect the steel shell against abrasion, heat, and oxidation. Figure 2 shows a cross section of a cupola with the various refractory types indicated. Lining thicknesses range from 114 to 305 mm (4.5 to 12 in.). Most linings are either fireclay brick or a sprayed or gunned alumina material. As the melting campaign progresses, the refractory lining in the melt zone is progressively fluxed away by the high temperature and oxidizing atmosphere, reducing the lining thickness. Early cupola melting practice was to stop melting, drop the bottom, and repair the lining at the end of the day’s melt. After cooling the cupola interior, repairs were made to the refractory lining, restoring it to its original diameter. Often, a pair of cupolas were alternated each day to facilitate the repair and preparation of the cupola lining after each day of use. Today’s (2008) improved refractories and shell water cooling have, for the most part, eliminated the need to have two operational cupolas. Many methods and materials are employed for cupola refractory repair. The most widely used method of lining repair uses a pneumatic gun to blow a mixture of clay and ganister through a nozzle containing water that becomes entrained in the stream as it exits the mixing nozzle. New methods such as shot creting have been satisfactorily used for cupola refractory patching and for extending the life of the cupola refractory. Shot creting creates a denser refractory layer and allows for a greater variety of refractory materials to be used. For repair of the taphole area, refractories containing silicon carbide particles are rammed in with pneumatic hammers. For the upper stack portion of the cupola, brick or blocks are layered with a thin mortar joint. Specialty refractories are being developed for every segment of the cupola to meet the exacting needs of the specific cupola operation. The advent and growth of ductile iron in the 1950s made it desirable to find a way to remove sulfur inside of the cupola. This was
102 / Melting and Remelting accomplished by using a sufficiently basic slag, as is done in basic steelmaking furnaces. However, a highly basic slag cannot be maintained in contact with acid refractories. After considerable experimentation, suitable basic refractories were successfully developed for use in cupolas. Originally, linings were either magnetite or dolomite brick. Patch materials composed of dolomite or magnetite aggregate were then
Fig. 2
developed that could be rammed or blown in using the pneumatic gun.
Water-Cooled Cupolas When the melting operations of foundries were extended to a continuous 16 to 20 h of operation, there was little or no refractory lining
remaining in the melting zone at the end of the melting campaign. When three-shift melting operations became a necessity, the water-cooled cupola was developed. The early water-cooled shells were simple tanks or jackets added to the outside of the cupola shell. Figure 3 shows an early water-cooled cupola design. Water circulated through the tank with little velocity, and cooling was basically just through heat conduction into the water. A limiting factor to the tank method of water cooling was that if a hot spot developed on the shell, it could not be observed nor could additional water cooling be added to the area during the melting campaign. Later water-cooling developments led to the cascading water shell, where a thin layer of water flows rapidly down a cupola shell, wiping heat from the shell in the process. At the top of the water-cooled area, a water curtain is applied to the shell by spray rings encircling the cupola. The water curtain cools the shell as it cascades down the shell, and the water is collected in a trough or moat at the well. One advantage of the open external cooling is that the outside of the cupola shell can be sand blasted at intervals to remove excess deposits of lime and other built-up materials. This sand blasting also restores the thermal conductivity of the steel shell. With increasing dependence on water cooling, it was found that the refractory lining could be eliminated from the melting zone and that the cupola could be operated continuously without a refractory lining. However, the advantages of water cooling are achieved at a loss in thermal efficiency. With demands on lower cost for coke and greater operating efficiencies,
Fig. 3 Cross section of a modern-day cupola showing typical refractories used in lining a cupola
Early water-jacketed cupola used to extend operating hours of a melting campaign
Cupola Furnaces / 103 today’s (2008) cupolas are a hybrid of water cooling and refractory linings. When iron is needed for only two shifts, the water-cooled cupola can be banked during the downshift and brought back on stream when needed. These water-cooled cupolas can even be banked over a weekend as well as operated without the need for refractory repairs for 2 to 4 weeks (or even longer under some operating conditions) before requiring the dropping of the cupola bottom for repairs. Some of the largest cupolas are equipped with double tapholes and double rear siphon slag spouts to enable the cupola to be operated continuously for even longer melting campaigns.
Emission-Control Systems Equipment has been successfully engineered to enable the present-day cupolas to meet the most rigid of environmental regulations. The two most common types of collectors currently in use today (2008) are the high-energy wet scrubber and the dry bag house. However, current environmental regulations have placed even higher demands on the cupola operator for cleaner air emissions, and only dry bag house systems are able to meet these new regulations. High-energy wet scrubbers use high-velocity water jets to impinge atomized water droplets on effluent dust particles passing through a venturi scrubber. The result is a dirty water sludge. This dirt and water sludge requires thickening, separation, and dewatering. Due to the various hazardous chemical elements (typically cadmium and lead) present in the cupola gas stream, the sludge and water are deemed toxic and require further chemical treatment to neutralize the hazardous elements. This is a complex science and a materials handling
challenge. Figure 4 illustrates a typical wet scrubber emission-control system connected to a recuperative hot blast system. The dry bag house environmental system is the best cleaning system for meeting today’s (2008) air emission standards. In this unit, the dust particles are collected on the surface of fabric bags, and then the bags are shaken to free them of the dust. Gases leave the cupola at temperatures ranging from 95 to 1100 C (200 to 2000 F), depending on what is happening in the cupola. The cloth bags can only withstand temperatures of approximately 230 C (450 F); thus, a system for cooling the gas stream before reaching the bag house is necessary to protect the bags from excessive heat. The two most popular heat-exchange systems are the air-to-air and air-to-liquid heat-exchange systems. The engineering for either is complex and a challenge for mass flow and heat-removal requirements. To better understand the operation of the recuperative, dry cleaning system, it can be broken into smaller subsystems. Figure 5 shows a modern dry bag house cupola emission system. The system design is basically simple in terms of what is being accomplished in each stage. Cupola offgas exits the cupola at a temperature between 95 and 540 C (200 and 1000 F) under normal operating conditions. This gas is rich in carbon monoxide (CO) but depleted in oxygen and has the potential for combustion. Air is added to the mixture in a controlled amount, and the mixture is ignited via pilot burners in a large combustion chamber. This combustion results in a stable temperature of approximately 820 C (1500 F). The entire gas stream is now passed through either an air-to-air or air-to-liquid heat exchanger to heat the blast air going to the cupola, or a secondary heat-removal step can be added to heat air for
building heat and/or for other processes in the foundry operation. The gas stream exits the cooling tower at a temperature of 135 to 165 C (275 to 325 F) and is now ready to enter the bag house at a temperature that is low enough not to harm the filter bags. While the pores of the cloth are very small, the particles in the gas stream are much smaller, so, in theory, the dust particles would pass through the bags and into the atmosphere and not be trapped. However, the dust layer that builds up on the bags traps the smaller dirt particles in a porous cake of dust. The bags are mechanically shaken or popped with a shot of air to remove some of the dust cake. As in the wet system, the dust may be a health hazard, depending on the concentration of hazardous elements present. Process chemicals may be added to the dust/gas stream before the bag house to neutralize these hazardous elements. Each lot of bag house dust is tested and then properly disposed of in a suitable landfill.
Charge Materials Storage and Handling of Charge Components Essential to efficient cupola operation is a good materials storage and handling system. Metal charge components are received via dump trailers, rail cars, or bulk containers. Unloading into storage bins can be by selfunloading dump trucks or by using an overhead charge crane with a magnet. The size of scrap storage bins varies widely with how much is to be melted each day, distance from the scrap-processing source, reliability of the transportation system, severe weather issues, and so on. Each charge material needs to be transferred to a weigh hopper, where a designated weight is measured out and transferred to a charge bucket for travel to the top of the cupola. The design possibilities for the charging area are endless. Figure 6 shows a typical charge yard design. Coke and limestone are usually stored in covered overhead storage bins and are discharged into a weigh hopper and on into the charge buckets. One factor in charge yard design that is often overlooked is planning for flexibility. Charge yards are expensive to construct and are longterm structures. The needs of the foundry and the availability of charge materials are likely to change over the lifetime of the facility, so it is necessary to build in some flexibility in design to accommodate for an unknown future need.
Coke
Fig. 4
Typical wet scrubber with heat recuperation hot blast used to clean emission gases
As the major fuel, coke is the most important charge material. High-quality, foundry-grade coke is essential for optimum cupola performance. The amount of coke in the charge mix usually ranges within 8 to 16% of the metal charge, depending on the iron temperature
104 / Melting and Remelting
Fig. 5
Fig. 6
Typical dry bag house emission and recuperative blast air heating system used to clean emission gases
Charge yard arrangement showing scrap delivery, storage bins, weigh hoppers, and charging bucket
required, the percentage of steel scrap used, and many other variables. Foundry coke must be strong to support the charge materials in the cupola stack and have sufficient impact strength to maintain its size during the impact of charging. These mechanical properties are important since they affect hot blast penetration and carbon pickup. Experience has shown that performance is optimized when coke size is consistent. Fine particles of coke should be screened out of the coke portion of the charge, because they fill the interstitial spaces between the coke and scrap iron pieces; this limits air distribution and overloads the emission system with excess carbon monoxide gas. On large cupolas, usually the largest available coke size is preferred, but on small cupolas, coke screened to any consistent size range provides good performance. Coke is manufactured to coke industry standards. Table 1 is a typical foundry specification for coke. Typical foundry coke should have a low ash content (<10%) and a chemical structure that provides for optimum carbon pickup and combustion efficiency. Low sulfur content (<0.70% S) is required for preventing excessive sulfur pickup by the molten iron. Volatiles should be less than 1.0%, and fixed carbon should be above 90% for maximum cupola performance. Physical properties of coke also must be controlled. Coke strength must be high to prevent crushing in the cupola stack. A dropshatter test best measures coke strength. Abrasion resistance, as measured by stability or hardness, must be as high as possible. Efforts have been directed toward the use of limited amounts of less-expensive substitutes
Cupola Furnaces / 105 Table 1 Typical specifications for foundry coke Fixed carbon Volatile Ash Sulfur Moisture Shatter strength 50 mm (2 in.) screen 75 mm (3 in.) screen 100 mm (4 in.) screen Stability index Coke size
91.5–94.5% 0.1–1.0% 5.0–7.5% 0.4–0.7% 0.1–5% 96.0% min 92% min 78% min 38.0% min As required by cupola size
for quality coke, such as briquetted coke fines, anthracite coal, injected fine coke, and various substitute fuels. However, all of these substitutes reduce cupola efficiency, and the effectiveness of each must be weighed against total fuel cost-savings.
Metallic Charge Materials The modern cupola can melt a wide range of metallic charge materials, most of which can be classified as scrap iron and steel, various cast irons, and other ferrous materials. Through its use of scrap materials, the casting industry contributes greatly to environmental cleanup and resource reutilization. Steel scrap was first melted in low percentages to lower the carbon content of gray iron in order to produce higher-strength irons. Today’s (2008) cupolas use steel scrap at a rate of 25 to 90% of the charge mix. Obsolete scrap iron is processed in a variety of ways. It can be cut with a cutting torch, sheared with a large blade shear, shredded in a hammer mill, or bailed into cubes. Each method produces a useable material for cupola melting, but each has its own melting characteristics. Cupolas are generally flexible enough in design to melt any of these prepared materials, but an optimized cupola design can make any scrap preparation method more efficient. Each preparation method has a unique cost associated with it and is local market dependent, so there is not one method or material that is preferred. Cast iron scrap is a common cupola charge material. It is broadly divided into four or five classifications of automotive, machinery, heating, and pipe. The grades useable in any particular foundry are largely dependent on the foundry products being produced. Scrap size and chemical composition are the basic source of the different grades of cast iron. Cast iron has a high carbon content and low melting point. It generally gives a consistent chemistry to the melt and lowers the chill tendency of the melt. Due to its good melting properties and limited availability, it is generally higher in price than scrap steels. Foundry returns, such as gates, runners, and internally generated scrap castings, usually constitute 30 to 50% of the pouring weight of a mold, and thus, they are returned to the charge
yard for melting in approximately this ratio. A melting operation will need to melt all of the foundry returns sent to it. As in cast iron scrap, carbon content is high and melting temperature low. Since this is internally generated scrap, it should have the most consistent chemical content of all metallic charge materials. Pig iron is considered a premium charge material because of its uniformity of composition and size, and it is free of detrimental tramp elements. In the early years, pig iron was the predominant charge component for ensuring quality iron with sufficient carbon content and low quantities of detrimental elements. Despite economics that encourage the use of more scrap and less pig iron, pig iron has remained a basic charge material because of its consistent size and quality. A wide range of scrap metals can be melted in the cupola. However, some forms of scrap metal are not practical to use in a cupola charge. Cast iron borings charged directly into the cupola are excessively oxidized if not blown out of the stack. With briquetting, limited proportions can be used. Steel turnings are more impractical because they are so thin that they readily oxidize high up in the cupola stack. Thin sheet steel has a very low density as received from the source and cannot be used without being compressed into bundles. Compact punchings tend to clog interstitial coke passages and hinder combustion efficiency. Very thick steel parts and large-diameter shafts are not completely melted within the melting zone as the charge descends into the cupola, and this results in low iron temperatures. Recently, a manufactured product consisting of cast iron borings, steel turnings, grinding dust, coke breeze (optional), and a binder has been successfully pressed into a brick. This allows the use of materials previously not usable in the cupola. Brick use must be limited, but use rates of up to 30% of the charge are possible.
Ferrous Alloys Any number of alloys can be added to the cupola for melting into the base iron. Most commonly, these alloys are purchased preblended with an iron-base material. Silicon, which is added to yield the desirable final silicon content, is usually added in any of the following forms: Lump ferrosilicon: 50 or 75 wt% Si content Silvery pig: 16 to 20 wt% Si content Silicon briquettes: composed of silicon fines
briquetted with a suitable binder, often portland cement, with the briquette containing approximately 40 wt% Si Silicon carbide briquettes: made of silicon carbide plus carbon briquetted together with a suitable binder Manganese briquettes of various compositions can be charged into the cupola in several forms to obtain the desired final manganese
content. The briquettes are generally composed of ferromanganese granular material and a cement binder. Contained manganese percentages are usually 50 to 75 wt%. Alloys of chromium, nickel, molybdenum, and other required alloys can be purchased and added to the cupola. Care must be taken in the selection of alloys to be sure that the size of the additive is sufficiently large to minimize oxidation and losses to the emission system. Briquettes must have sufficient strength to survive charging operations and the descent into the cupola melting zone. Binders must be evaluated to have sufficient hot strength to prevent premature breakdown of the binder system before entering the melting zone.
Fluxes Charge materials contain large amounts of nonmetallic materials, such as dirt on scrap, sand adhering to returns, slag attached to pig irons, and so on. Additionally, coke ash, oxidation products from melting, cement binders, and other noniron materials must be fluxed into low-melting-point compounds. These slags must be sufficiently fluid to be able to exit the cupola and float to the top of the iron so they can be skimmed off and not sent to the holding furnace. Limestone or dolomite additions of 2 to 8% of the metallic charge (or 30 to 45% of the coke portion of the charge) are generally needed in an acid cupola to contribute enough basic constituents of calcium oxide and magnesium oxide to impart sufficient fluidity for the slag to flow freely through the coke spaces, into the well, and down the front slagging trough. In the basic slag cupola, more basic fluxes are added to net an excess of basic slag constituents. Limestone is a natural calcium carbonate or dolomite stone. A lump size of 25 to 75 mm (1 by 3 in.) is preferred. Limestone composition varies widely across the country, so selection of stone for a melting operation is critical. Screening of limestone fines is mandatory for a good flux stone. For limestone (calcium carbonate limestone), CaO contents should be above 45%. For dolomite stones (those containing high levels of calcium-magnesium carbonate), CaO levels should be approximately 30% and MgO levels approximately 20%. Silica is an impurity and should be less than 2%. Several special or supplementary fluxes are occasionally used. When used, they are usually added very sparingly in addition rates of 0.2 to 2.0% of the metallic charge weight. Soda ash (Na2CO3) is a strong slag liquefier and is added as briquettes to the bed to supplement the limestone. Fluorspar (CaF2) is a powerful fluidizer that is used to clean up a cupola clogged with excessively viscous acid slag. In highly basic slags, fluorspar is needed to maintain fluidity. However, excessive fluorspar can cause environmental problems.
106 / Melting and Remelting Calcium carbide (CaC2) serves the multiple functions of deoxidizing, raising carbon, raising temperature, and providing a strong calcium oxide (CaO) flux contribution.
Cupola Control Principles Coke Bed. Preparation of the coke bed is vital to the initial start of the melt and to the maintenance of the melt at the specified level in the cupola combustion zones. Before metal is charged, a bed of good-quality, properly sized coke is ignited and blown to a red heat either internally in the cupola or externally and charging it hot. The height of the coke bed is adjusted to a specific height above the tuyeres that is determined by experience and based on blast volume, cupola diameter, and desired melting conditions. Bed flux and metal charges are then added, the blowers are turned on, and the melting begins within a few minutes. With a high coke bed, melting occurs higher in the cupola stack, producing slower melting rates and hotter iron with less oxidation and more carbon pickup. A low bed melts charge materials lower in the cupola, with faster melting rates and colder iron temperatures with more oxidation. Charge coke is the coke intended to replace the coke burned in the melting of each layer of charge materials. It is weighed and added to each charge to maintain the bed level within the cupola during the melting cycle. While the exact height of the bed during melting cannot be observed, a skilled operator knows its relative position by observing melting rates, slag temperatures, iron temperatures, and final iron chemistry. Melt Rate. A cupola is expected to maintain a melting rate sufficient to meet the needs of mold production. The cupola size must be chosen for the midrange of production demand. Several general rules can predict melting rates. One rule is that 4.5 kg (10 lb) of iron can be melted per hour for each square inch of crosssectional area. A second rule is that 3.2 to 3.3 Mg (3.5 to 3.7 tons) of gray iron can be melted with 0.5 m3/s (1000 ft3/min) blast air, and 2.9 to 3.0 Mg (3.2 to 3.3 tons) of ductile iron can be melted with 0.5 m3/s (1000 ft3/min) blast air. By increasing or decreasing the blast air rate, melt rate can be varied by approximately þ 20% with reasonably good efficiency and control. Beyond these limits, overblowing a cupola increasingly oxidizes the iron, causing lower carbon and silicon levels and decreasing the iron temperature. Underblowing will lower the iron temperature because the blast will not penetrate the cupola with sufficient air distribution for optimum combustion efficiency. Changes in melting rate can also be accomplished by changes in the charge mix ratio of cast iron components to steel and by varying the hot blast temperature. If oxygen injection is used, changes in oxygen volume can widen the melt rate control range.
Iron Temperature. The iron temperature at the cupola spout must be sufficient to overcome temperature losses due to metal handling from the cupola to the electric furnace or holding ladle, the holder to the pouring ladles, alloying additions, the treatment process (for ductile iron), and radiation losses from the pouring ladles. Iron temperature can be measured by optical pyrometers, immersion thermocouples, and/or continuously recording radiamatic instruments. Carbon and silicon concentrations are most important in the control of iron properties. Sufficient concentrations of carbon and silicon must be maintained to avoid chill and to encourage a microstructure suitable for the section size of the casting and the strength level required. In ductile iron, sufficient carbon and silicon concentrations must be maintained within controlled ranges for proper production of the nodular form of graphite and control of the matrix structure. Excessive carbon content can result in carbon flotation in ductile iron castings. Carbon content of the iron is influenced by charge mix, carbon content of the metallic scrap, the coke-to-metal ratio, and slag basicity. Carbon pickup is increased by increasing the blast temperature and by increasing the oxygen injection percentage. The final carbon content can be altered by changing any one or a combination of blast control parameters, which affect the combustion process. Silicon is controlled in the cupola with the addition of ferrosilicon alloy and/or SiC briquettes. Maintaining the desired silicon content in the final iron requires accounting for the oxidation losses of silicon in the charge. Sulfur in gray iron is a naturally occurring element of melting in a cupola. It primarily comes from the sulfur contained in the coke. Sulfur is not detrimental below approximately 0.150 wt% if the manganese content is maintained at a sufficient ratio in relation to the sulfur content. Sulfur inhibits the conversion of carbon to ductile iron. Ductile irons are treated with a magnesium-containing alloy, which preferentially reacts with sulfur, forming inert MgS compounds. The small amount of magnesium remaining in solution in the iron promotes the nodular graphite shape. Thus, sulfur in ductile base iron should be below 0.030 wt% and preferably below 0.015 wt% for efficient conversion costs. Manganese is needed in gray iron at an amount of 1.7 times sulfur plus an additional 0.20 to 0.30 wt% to ensure that the sulfur is precipitated as manganese sulfide rather than iron sulfide. In pearlitic ductile iron, manganese is added as an alloy to produce a pearlitic matrix. In ferritic ductile iron and austempered ductile iron, manganese should be kept as low as possible to minimize pearlite in the as-cast structure and to minimize segregation, respectively. Phosphorus is controlled by using a variety of charge materials that are low in phosphorus. In gray iron, phosphorus contents should be below 0.10 wt% and even below 0.05 wt%, because high phosphorus levels are detrimental
to machinability and casting soundness in complex shapes. In ductile iron, phosphorus content below 0.05 wt% and preferably 0.03 wt% is desirable for maximum impact resistance and ductility. For low phosphorus levels, careful selection of charge material is required. Steel scrap is generally the lowest-phosphorus-content charge material. Residual Alloys and Tramp Elements. The effects of various inadvertent alloys and tramp elements should be understood, and control ranges and maximum levels should be established. Rapid analysis by modern spectrometers makes it possible to monitor any number of trace element levels. Any scrap source contributing excessive concentrations of detrimental elements should be removed from the cupola charge. Certain tramp elements, such as lead and bismuth, are detrimental to graphite shape (particularly the nodular graphite shape) and must be carefully monitored and maintained below the maximum level allowed. Titanium in trace amounts is beneficial to gray iron but is detrimental to the nodular iron, since it inhibits the formation of the nodular graphite shape. Inadvertent alloys, such as chromium, copper, tin, and molybdenum, if maintained within controlled ranges and combinations, can contribute to the pearlitic matrix structure of high-strength gray irons and high-strength pearlitic ductile irons. However, for ferritic ductile iron requiring the highest ductility, acceptable limits for these residual alloys are much lower.
Control Tests and Analyses The chill test was one of the earliest control tests developed and is still a good indicator of the graphitizing balance of an iron. A chill specimen is cast either as an elongated wedge in a sand mold or as a bar cast against a metal chiller on one side. After a few minutes, the test specimen is broken, and the depth of carbidic white iron chill indicates the balance between carbon, silicon, and other graphitizing elements against carbide-forming elements. This chill depth signals the need for more or less silicon, more or less carbon, and/or an adjustment in the postinoculation. Eutectic arrest methods were developed to provide carbon equivalent and carbon content as well as calculated silicon content. These thermal analysis techniques are rapid and make possible immediate corrective action, that is, adjustment of carbon and silicon. Optical spectrometer equipment greatly facilitates the control of iron chemistry, since all of the important elements can be determined within a few minutes. In large plants, samples are conveyed to the laboratory in pneumatic tubes, and analysis reports are conveyed to the melting center by phone or computer within a few minutes of taking the sample. Mechanical tests of several types are conducted to ensure the uniformity of iron properties and compliance with specifications.
Cupola Furnaces / 107 Tension tests of gray iron ensure that the iron meets the strength level required for the specified class of iron. In ductile iron, ultimate tensile strength, yield strength, and percent elongation are usually measured. Brinell hardness testing of castings monitors part-to-part and batch-tobatch uniformity and is an indicator of the general matrix structure. For ductile irons, microscopic evaluation of each treatment batch is essential to ensure graphite nodularity before castings reach the shakeout operation and lose their identity. Nodular graphite structure in a casting can also be verified with ultrasonic velocity test equipment. Ultrasonic velocity testing of all castings is performed by some companies to ensure the nodularity of safety-critical parts.
Specialized Cupolas Cokeless Cupola. In the 1970s, a cokeless cupola fired with gas or oil was developed in Europe and put into operation in several countries. It is a refractory-lined shaft furnace with a water-cooled grate supporting a bed of refractory spheres, as illustrated in Fig. 7. Heat for melting was generated by either gas or oil burners located below the grate and refractory bed. The iron melted in a cokeless cupola has low sulfur content and low carbon content, both of which are due to the absence of coke. The iron is carburized to a specified carbon content by injecting a fine carburizer (either petroleum coke or graphite) into the iron in the well. Pollution from the cokeless cupola is minimal and reportedly would meet environmental standards in some countries without expensive emission-control equipment. In countries where coke must be imported at a high price and oil or gas is readily available, the cokeless cupola is economically attractive. Production experience has been limited to small cupolas in the 4.5 to 9.1 Mg/h (5 to 10 ton/h) range. No practical experience has been reported on large cokeless cupolas. Attractive fuel (oil or gas) cost is available in some countries, but the high cost of refractory sphere replacements and the restrictions on the percentage of steel scrap used limit the attractiveness of this equipment in many markets. Plasma-Fired Cupola. In the 1980s, a pilot plant plasma-fired cupola was built, and a number of experimental heats were run, with some encouraging results. In the 762 mm (30 in.) inside diameter cupola, a single 2 MW plasma torch was projected into the cupola at tuyere level, as shown in Fig. 8. In the water-cooled torch, an electric arc from a high-amperage direct current is maintained across a gap between two copper sleeves, and the arc is rotated by a magnetic field. Air (or gas) is passed through the electric arc, where it is ionized and heated to very high temperatures unattainable with current hot blast systems. The experimental heats conducted indicated several advantages of using a plasma-fired cupola:
Coke (fuel) could be considerably reduced
due to the supplementing electric power from the torch. Lower-quality coke could be used; blast furnace coke, fine waste coke, and anthracite coal were melted successfully. Direct melting of borings and turnings proved practical because of lower air velocity and lower oxidation level. Silicon could be produced by injecting silica sand plus coke breeze. Iron could be produced by injecting as much as 20% iron oxide through the tuyeres.
The many possibilities of the plasma-fired cupola merited further exploration. Economically, the most attractive immediate application was the direct melting of iron borings or steel turnings. In March of 1989, the first commercial plasma cupola started production. This cupola was 3963 mm (156 in.) in diameter and had six conventional tuyeres. A 1.5 MW plasma torch was mounted to the end of each tuyere. Compressed air at 0.7 MPa (100 psig) was used as a plasma source. A booster compressor increased the air supply to 0.9 MPa (130 psig). The compressed air was passed through driers and filters and then to a distribution manifold. After being collected in the manifold, the air was dispersed into the six torches. The operating consumption of air was approximately 0.08 standard m3/s (160 standard ft3/min) per torch. Deionized water was necessary for the cooling of the torches. A conventional 700 C (1300 F) gas heat exchanger was used to preheat the blast air. The air at the tip of the plasma torch was approximately 1650 C (3000 F). The iron temperature at the top of the coke bed was approximately 1540 C (2800 F). The air blast volume was 6.0 to 7.0 m3/s (13,000 to 15,000 ft3/min) at the tuyeres. The development of the plasma cupola for melt operations was to offer melters the opportunity to use low-cost charge materials such as borings and steel turnings. This production plasma cupola was able to use an average of 40% loose borings/turnings in the charge makeup. The highest level achieved was 70%. This offered a significant cost-savings. The increased electric energy cost was partially offset by reduced coke usage.
Fig. 7
Cokeless cupola showing scrap charging, watercooled grates, burner location, and so on
ACKNOWLEDGMENT S.F. Carter, Melting Furnaces: Cupolas, Casting, Vol 15, ASM Handbook, ASM International, 1988, p 383–392. SELECTED REFERENCES Conference on Cupola Operation, American
Foundrymen’s Society, June 1980
Cupola Design, Operation, and Control,
BCIRA, 1979
Cupola Handbook, American Foundrymen’s
Society, 1975 and 1984
Fig. 8
Early version of an experimental plasma cupola used to confirm the concept of plasma heating as a melting technology
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 108-115 DOI: 10.1361/asmhba0005196
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Induction Furnaces Revised by Ralph Y. Perkul, Foundry Solutions & Design LLC
INDUCTION FURNACES have become the most widely used means for melting iron and, increasingly, nonferrous alloys as well. The key to the ready acceptance of induction furnaces has been its excellent metallurgical control coupled with its relatively low pollution operation compared to cupola furnaces. Induction furnaces are very efficient and are made in many sizes. Induction furnaces are relatively simple, and small quantities can be melted quickly. The melting time is relatively short, allowing metal to be delivered at small, regular intervals. A wide range of metals can be melted, although little refining of the metal is possible. Currently, induction furnaces are available in a wide variety of sizes. Coreless units range in capacity from a few pounds, favored by the precision cast metal producers, to 68 metric tons (MT) (75 tons) powered at 23,000 kW. Channel-type units have been built with a capacity of over 2200 MT (2400 tons) powered by 6 inductors rated at 6000 kW each for holding applications. Induction furnaces require much cleaner scrap than cupola furnaces and somewhat cleaner scrap than electric arc furnaces. The capital costs are higher than those of electric arc furnaces, but the operating costs are lower due to reduced refractory wear. In the past, large, high-powered induction furnaces operated at 60 Hz, the frequency of the incoming power. Thanks to major breakthroughs in electronic, solid-state frequency conversion techniques, it is now possible to build a highly efficient furnace operating at medium and high frequencies (70 to 10,000 Hz). Coreless units of 20,000 kW in power for ferrous metals and 7000 kW for nonferrous metals have been built in medium frequencies, and even channel-type units have recently been equipped with solid-state, medium-frequency power supplies. The advantage of medium frequency is that the power density of the furnace system can be increased substantially without increasing its size. Much more melting power can be applied, while maintaining the stirring action at desirable levels. Heat losses, which are a function of furnace size, are reduced, and
overall efficiency of the system is improved. The furnace is also easier to operate because a single potentiometer will typically provide efficient and stepless control of the power.
Types of Furnaces The basic principle of induction furnaces is that a high voltage in the primary coil induces a low-voltage, high current across the metal charge, which acts as a secondary coil. Because of electrical resistance in the metal, this electrical energy is converted into heat, which melts the charge. Once the metal is in its molten state, the magnetic field produces a stirring motion. The power and frequency applied determine the stirring rate. This is controlled to ensure complete melting of the charge and adequate mixing of alloy and fluxing materials, and to minimize temperature gradients in the charge. Excessive stirring, on the other hand, can increase lining damage, increase oxidation of the alloys, generate excess slag, and increase inclusions and gas pickup. There are two classifications of induction furnaces: coreless and channel. Cross sections of each are shown in Fig. 1 and 2. In a coreless furnace, the refractory-lined crucible is completely surrounded by a water-cooled copper coil, while in the channel furnace the coil surrounds only a small appendage of the unit, called an inductor. The term channel refers to the channel that the molten metal forms as a loop within the inductor. It is this metal loop that forms the secondary of the electrical circuit, with the surrounding copper coil being the primary. In a coreless furnace, the entire metal content of the crucible is the secondary. Although both channel and coreless induction furnaces employ similar electrical principles, the two types of furnaces are quite different in their capabilities and operation. The coreless furnace is more widely used for melting and superheating, whereas a channel furnace is better suited for superheating, duplexing, and holding. The channel furnace has been used by nonferrous foundries for many years. Recently,
however, this type of furnace has been increasingly replaced by the coreless and resistance furnace in aluminum-melting applications. It is still commonly used for melting copper alloys. Technological innovations in the channel furnace have been slower in developing than advances in the coreless furnace. Today (2008), medium-frequency power supplies are being installed on some new channel furnaces. Many older furnaces still rely on 60 Hz power supplies and use power-regulating autotransformers for control. Inherent in the channel concept is the fact that only the small amount of metal in the inductor channel receives energy through the surrounding coil. This necessitates superheating the metal in the loop and pumping it out of the channel into the main vessel, where it mixes with the cold metal and heats it by conduction. The ratio of heat transfer is limited, as is the stirring action in the main vessel. This is the primary reason this type of furnace has seen only limited service as a melter. The power that can be connected to a vessel is much smaller than for a similarly sized coreless vessel. The mild stirring action also makes it difficult to get alloy and carbon additions dissolved and mixed rapidly, particularly important in the production of synthetic irons. In addition, the channel furnace must be started with a supply of molten metal, or a melt-out form and a metal heel must be continuously maintained in the furnace. The furnace must not be shut down even for holidays or for other extended periods, because the metal will freeze, and considerable refractory damage will be incurred. The channel furnace is more economical to manufacture in large sizes and uses electric power more efficiently than a coreless furnace. If more power is desired, additional inductors can be connected to the main vessel. Several geometric shapes of vessels have been used. These include configurations that are described as vertical (upright cylindrical), drum (horizontal cylindrical), and low profile (spherical). Typically, all of these units are hydraulically tilted to permit the emptying of the contents.
Induction Furnaces / 109
Fig. 1
Components of a coreless-type induction furnace. (a) Operational elements. (b) Cross section showing water-cooled copper induction coil and key structural components. The entire molten metal bath (which serves as the secondary) is surrounded by the coil (the primary) that encircles the working lining.
Because the magnetic field is transmitted in all directions, special vertical laminations of transformer iron, which form the magnetic yokes (shunts), are evenly spaced around the circumference of the coil to provide additional strength and to pick up the stray flux that would otherwise heat up the furnace frame. Coreless furnaces are typically tilted about the spout through a 95 to 100 angle to empty their contents. The medium-frequency units can be completely emptied for alloy changes or factory shutdowns and quickly started up again with a cold charge. It is not necessary to maintain a molten heel or to use precast starter blocks, as is the case with the 60 Hz units.
Electromagnetic Stirring Fig. 2
A cross section of a channel-type induction furnace showing the water-cooled copper induction coil, which is located inside of a 360 loop formed by the throat and channel portion of the molten metal vessel. It is the channel portion of the loop that serves as the secondary of the electrical circuit, in which the copper coil is the primary.
The coreless furnace is completely surrounded by a spiral of copper tubing of a special cross section that features a water-cooling channel in its center. It must also provide good electrical coupling and strength to withstand substantial electromagnetic forces, yet with provision for thermal expansion.
When alternating current is applied to an induction coil, it produces a magnetic field, which in turn generates a current flow through the charge material, heating and finally melting it. The amount of energy absorbed by the charge depends on the magnetic field intensity, electrical resistivity of the charge, and the operating frequency. A second magnetic field is created by the induced current in the charge. Because these two fields are always in opposite directions, they create a mechanical force that is perpendicular to the lines of flux and cause metal movement, or stirring, when the charge is liquified. The mechanical force stays perpendicular to the
field only in the center of the coil; on both ends of the coil it changes direction. The metal is pushed away from the coil, moves upward and downward, and flows back. Figure 3 shows the four-quadrant stirring. It is this stirring that allows excellent alloy and charge absorption and aids in producing a melt that is both chemically and thermally homogeneous. The stirring is directly determined by the amount of power induced and is inversely proportional to the square root of the frequency. Therefore, the higher the power and the lower the frequency, the more intense the stirring for a given size furnace. The reasons for an increased interest in induction melting for aluminum applications are many, but perhaps most significant is the question of metal loss. Because of the inherent electromagnetic stirring action in the coreless furnace, charge materials quickly become immersed in the bath, minimizing oxidation losses. This effect is particularly pronounced when melting light-gage scrap. Thus, the growing emphasis on recycling favors induction melting as a cost-effective alternative to other metal melting processes.
Power Supplies 60 Hz Line Frequency. In order to meet the range of melting needs, the proper power supply must be selected. Because the vast majority of existing furnaces use line frequency, this arrangement is discussed first (Fig. 4).
110 / Melting and Remelting
Fig. 3
A cross-sectional view of a coreless-type induction furnace illustrating four-quadrant stirring action, which aids in producing homogeneous melt
Fig. 4
Modifying 60 Hz line frequency supplied by utility company to serve as power supply for induction furnaces
Protection of the furnace and the power system, phase balancing, and power correction are required in order to operate an effective installation. The typical 60 Hz (line frequency) power supply basically consists of:
Primary switchgear Furnace transformer Mainline contactor Starting resistor and bypass contactor (for installations with off-load tap changers) Phase balancing system Power factor correction capacitor bank Instrumentation Controls and supervision
A primary transformation from the highvoltage grid of the local utility is usually
required because the voltage applied to the coil is between 500 and 3000 V. Because of its high inductance, an induction furnace coil has a low power factor. In order to bring the power factor to near unity, capacitors are installed and connected in parallel to the furnace coil. When starting a furnace, in order to reduce the inrush, a soft start is provided. When the transformer with an on-load tap changer is installed, the tap changer travels down to a lower tap whenever the furnace is switched off. This provides a start with a low voltage, resulting in low inrush current. When an off-load tap changer is applied, a starting resistor bank with bypass contactors is provided. Because the induction furnace is a singlephase furnace but its feed is commonly a three-phase line, a phase balancing system consisting of capacitors and reactors must be provided. In ferrous applications, refractory lining wear decreases the coupling distance between the coil and charge, resulting in a power increase. Because power input is limited, the voltage must be decreased to keep power from exceeding its maximum. This is typically done by using steps on a transformer tap changer. As the lining wears and the power increases, the voltage must be reduced by going to a lower tap. Additionally, in order to compensate for the amount of metal in the furnace and the condition of the refractory lining, a portion of the capacitor bank that is installed in parallel with the coil needs to be switched by adjusting the coarse and fine settings on the electrical panel. Medium Frequency. In the modern medium-frequency power supplies, (Fig. 5), many of the functions described are now performed automatically or have been eliminated altogether. The typical conversion is in an ac-dc-ac sequence. While frequency conversion entails electrical losses, modern power supplies incorporate solid-state frequency converters, and the increase in thermal efficiency usually outweighs the conversion losses. Modern converters exceed an efficiency of 97%. In the power conversion commonly used, the incoming 60 Hz ac power is rectified to dc and then chopped by an inverter to a higher frequency. Series and Parallel Inverters. Two types of inverters are typically used, namely, series and parallel. In each case, the furnace and compensating capacitor bank (fixed, in this case) form part of a resonance circuit. In the voltage series converter, the rectifier produces dc voltage, and the energy is stored in a filter capacitor. In a current parallel converter, the rectifier produces dc current, and the energy is stored in a smoothing reactor. The melting power for each of the two types of converters is also controlled differently. In the voltage series converter, the output is controlled entirely by variation of the inverter commutation frequency. The voltage is not controlled and is defined only by the line voltage.
In the current converter, the melting power is controlled by the variations of this inverter frequency as well as the magnitude of the dc voltage. The latter is regulated by variation of the conduction angle or phase control in the rectifier section. In both series and parallel circuits, it is possible to get power control that allows full power to be drawn regardless of the level of the charge in the furnace or the condition of the refractory lining. The midfrequency converters operate on a variable frequency. The load-commutated converter increases frequency as the charge contents of a furnace increase, while it simultaneously compensates for lining wear. No capacitor switching is required as in the line frequency units. In addition, a single potentiometer offers stepless and infinite control, eliminating the need for tap changer systems for power control. Modern induction furnaces take advantage of the latest technology, and many new installations are equipped with programmable controllers in lieu of relay logic. An even more recent advance has been the use of industrial computers for controlling the furnace and auxiliaries. With the variety of software that is currently available, the operator can choose screens such as systems data, automatic lining sintering, automatic cold starting of the furnace after a prolonged shutdown, automatic melting programs for various alloys, continuous calculated-temperature monitoring, fault detection, demand limiting, tabulation of energy consumption, melting cost data, metal inventory control, and chemistry adjustment. Also, by tying the computer to support systems such as charge make-up and hot-metal delivery, the utilization of the furnace can be increased. Packaging of Systems. Until the early 1970s, the electrical components of coreless induction furnaces were typically assembled on site in a concrete vault. Field construction costs made this installation method quite expensive. Modern component assembly techniques have allowed for the power supply to be packaged into steel modules at the equipment manufacturing plants. For furnaces up to 3.6 MT (4 tons) in capacity, current practice allows the manufacturers to completely assemble, wire, and pipe all components on a steel frame, to test, and subsequently to ship the unit as a single module, thus minimizing installation and start-up costs. For furnaces with power above approximately 1250 kW, a step-down isolation transformer from the main distribution voltage is normally used. Below 1250 kW, a 480 V distribution line backed by a disconnecting device can be used. The step-down transformers, when required, can be of the oil-filled outdoor type, which is the least expensive, or the indoor dry or silicon-filled type. Some manufacturers use the 480 V directly in their power supply; others choose to provide a step-up transformer and operate at the 575 V level.
Induction Furnaces / 111 In addition, with channel furnaces, standby generators are often employed to maintain cooling to the inductor.
Lining Material
Fig. 5
Typical medium-frequency induction power supply incorporating (a) a parallel inverter and (b) a series inverter
Water Cooling Systems Because the furnace coils and power supplies of the induction furnaces need to be cooled, a closed-loop water system is generally called for. In a medium-frequency power supply, the electronic components of the system contain a dc leg; it is therefore important to have clean, demineralized water circulating, typically with conductivity set at 50 mS (50 mmho) in order to prevent electrolysis from taking place. The heat exchanger to be used depends entirely on local circumstances such as climate, availability of raw water, and so on. The most widely used are the evaporative, the water-towater, and the air-to-water types. The latter has been frequently adapted to heat recovery systems. In those applications, the hot water from the furnace power supply and coil is piped to a location in the plant where heat is required. The water is circulated through a tube bundle, and a fan dispenses the warm air to the
surroundings. Considering that up to 10% of the power from a channel furnace and 20% of the power from a coreless furnace is lost to the water system, heat recovery systems for induction furnaces are often practical. In all coreless- and channel-type furnaces, the water is typically circulated through the power supply and then routed through the coil and on to the heat exchanger. Sometimes the water system is separated into two — one for the power supply and another for the coil. Usually, two pumps are provided, one as a standby with an automatic switchover, because it is extremely important to circulate water through the coil continuously to prevent it from being damaged. To protect induction furnace equipment during momentary power interruptions or prolonged power failures, manufacturers provide a special pressure-operated valve that, when activated, allows emergency deployment of city water to circulate through the coil. Such a system is shown in Fig. 6.
The selection of lining material is determined by the metallurgical requirements, the operating temperatures, and the type of operation. Refractories used in induction furnaces are oxides of minerals. Typically used are silica (SiO2), alumina (Al2O3), or magnesia (MgO) linings. From a chemical point of view, the silica is classified as an acid, alumina as a neutral, and magnesia as a basic material. Silica is the clear choice in iron melting, because it does not readily react with the acid slag typically produced in high-silica iron, has an extremely forgiving expansion curve, and is easy and economical to use. Alumina is the usual choice for aluminummelting furnaces, and steel-melting furnaces usually employ alumina or magnesia lining. Each of the refractory groups mentioned previously has a different thermal expansion characteristic, as shown in Fig. 7. As can be seen, alumina and magnesia have a nearly linear expansion. Silica, however, completes its expansion at approximately 815 C (1500 F) and remains constant at higher temperatures. With the selection of silica lining, the furnace can be shut down, cooled, and restarted without running the risk of metal penetrating thermal cracks. Magnesia is very sensitive to thermal shock and has the greatest expansion. Once cracked, this type of lining is not likely to seal itself, thus resulting in metal penetration. Alumina has approximately one-half the expansion of magnesia and is therefore less crack sensitive. It is important, however, to heat the furnace lining slowly when undergoing a cold start. This programmed start-up allows the cracks to seal before the metal adhering to the side walls melts and penetrates. This applies particularly to aluminum-melting furnaces because the metal melting temperature is relatively low. In the case of aluminum and some copper alloys, the oxide and metal buildup on the crucible not only poses a potential refractory penetration problem but also effectively increases the lining thickness and reduces the power that goes into the furnace. If not removed, this layer, which contains a small amount of metal, absorbs energy, and becomes superheated. The superheated metal can penetrate the lining and cause premature failure. For these reasons, the furnace crucible walls must be cleaned periodically. Crucible Wall Scrapers. The task of scraping the crucible wall has been made easier with the introduction of a mechanical scraper (Fig. 8), which cleans the sides in a full, hot furnace. The device essentially consists of a number of oscillating blades assembled on a steel frame. To start the operation, the scraper is placed on top of the furnace, and the
112 / Melting and Remelting
Fig. 6
A layout of a closed-loop cooling system showing routing of furnace-cooling water under both normal (black arrows) and emergency conditions (white arrows). Standby emergency pump is activated if the primary pump fails.
Temperature, ⬚F 32 2.0
392
752
1112
1472
1832
2192
2552 A
1.8 1.6
Linear thermal expansion, %
B 1.4
C D E F
1.2 1.0
Fig. 8
A cross section of a patented mechanical scraper used to clean the refractory lining of induction furnaces that melt nonferrous metals
G 0.8
H I K
0.6 0.4 0.2 0 0
100
200
300
400
500
600
700
800
900
1000 1100
1200
1300
1400
Temperature, ⬚C
Fig. 7
Thermal expansion curves of various refractory brick oxide materials used for linings in induction furnaces: A, magnesia; B, chrome magnesia; C, chromite; D, silica; E, zirconia; F, corundum 99; G, corundum 90; H, fireclay; I, sillimanite; J, zircon; K, silicon carbide
oscillating tools travel down the depth of the crucible. Radial pressure on the furnace is adjustable. The loose materials float to the top of the bath and are skimmed off. Ramming Mixes. Most induction furnaces operating in iron foundries are, as stated
previously, lined with a silica ramming mix. Alumina, magnesia, or zircon are typically used for melting steel and alloys. Alumina is typically used for aluminum. The current practice is to use granular materials. For example, the silica ramming mix is granular material combined with the sintering agent boron oxide (B2O3). The amount of boric
acid depends on the operating temperature and the wall thickness of the crucible. The flux or sintering agent fuses the silica particles to each other, forming a hard, glazed, ceramic surface. Ideally, the lining, when sintered, consists of one-third hard, sintered ceramic; one-third fritted material, in which fine particles are sintered and larger particles can still be identified; and one-third loose particles. This loose material has its purpose. When starting up a furnace from a cold start, the lining naturally expands. The loose material provides a cushion to handle the pressure. All granular materials are rammed on the furnace bottom using an electric vibrator equipped with a tamping plate. For the side walls, the lining material is spread behind a lining form that is vibrated with a pneumatic vibrator attached to it. For ferrous applications, the lining form is consumed in the first heat. For nonferrous applications, the lining form can be removed and reused. It may be tapered or collapsible.
Induction Furnaces / 113
Fig. 9
The principal components of a patented push-out device used to extract deteriorated refractory lining from coreless induction furnaces
The sintering process itself takes some time because it is necessary to heat the lining at the rate of approximately 110 C/h (200 F/h) until a temperature of about 30 C (50 F) above operating temperature is achieved. This temperature is held for 1 h before the furnace is ready for production. Lining Push-Out Device. When the linings are to be replaced, they can be chipped out with pneumatic hammers, and the lining procedure is repeated. A push-out device is available that eliminates the need to chip out the spent linings manually, saving both labor and time. The key components of this device, a schematic of which is provided in Fig. 9, are a locking plate that is bolted to the bottom of the furnace and a hydraulic ram assembly that is brought into position and latched to the plate. Once activated, the hydraulic ram moves forward, exerting pressure on a specially designed refractory cone installed in the bottom of the furnace. This distributes the force over the cross section of the lining. The push-out is aided by a slip joint between the coil grout and the working lining. The whole operation takes but a few minutes, with additional time being saved because the cool-down time is cut in half.
Melting Operations There are two distinct ways of operating a coreless induction furnace. One method is a batch operation, in which the entire contents of the furnace are emptied and the unit is
recharged with solids, typically using a conveyor or bucket charger. The other method involves a tap-and-charge operation, in which a portion of the furnace contents, typically one-third to one-half, is tapped and the identical weight in solids is recharged using a charge bucket or charge conveyor. Batch operation has now become feasible because medium-frequency power supplies allow the furnace to be started with ordinary scrap, rather than with starting blocks, which are required of 60 Hz (line frequency) furnaces. In terms of electrical consumption, the batchmelting method provides a 7% higher efficiency in iron-melting furnaces and a 4% higher efficiency in nonferrous-melting units over the tap-and-charge method. This is because the cold charge material couples with the magnetic flux more effectively than does molten metal. In addition to having greater electrical efficiency, a medium-frequency batch operation allows the drying of wet or oily charge material in the furnace, minimizing the need for a scrapdrying system. The tap-and-charge operation is still the most popular approach, however, since the chemistry is much easier to control. Less than one-half of the furnace contents are replenished at any one time, making alloying simpler because most of the furnace contents are already at the current specification. Finally, in order to use a batch operation effectively, the contents of the unit must be quickly and completely discharged, usually into a holding furnace or large ladles. This option may not be possible if small amounts of molten metal need
to be taken out at defined times and the capital and operational costs of a holding furnace cannot be economically justified. In either method of operation, it is particularly important to use the proper power supply efficiently and keep the cover closed to minimize potentially high radiation losses. Utilization is calculated in terms of the time that the power is on. It is obvious that power cannot be on constantly, because certain operations, such as charging, slagging, sampling, and tapping, are best and most safely done while the power is off. The objective is to keep the power on at least 75% of the time. Sometimes, a second furnace added to the power supply improves utilization, because the power can be switched from one furnace to the other while the above-mentioned operations take place. Utilization can also be improved by quickly charging and tapping to minimize waiting time and heat losses as well as by increasing the charge size. Today (2008), induction power supplies can operate two or more furnaces at the same time without the use of mechanical switches. These designs allow the system utilization to go to over 90%. Although medium frequency has made the need for heavy charge materials obsolete, it is still good practice to choose materials that have a reasonably high bulk density, are relatively clean and dry, and are not too long. In order to avoid bridging, the length of the pieces should be no more than two-thirds the diameter of the crucible. If longer pieces such as casting gating systems and structural steel sections are available, they should be broken or cut first. A further aid to utilization efficiency is the recent method of conveyor charging. Made possible by furnace covers acting as close-capture hoods, the furnace lid is kept up for much of the charging cycle. The charging is done continuously, adding more solid charge material as the contents of the furnace liquefy and collapse, allowing room for additional material. This adds safety to the often hazardous charging operation by removing the operator from the proximity of the furnace and allows control to take place from the melt shop control room. There are a variety of conveyor charging concepts. They are usually vibrating steel decks that can swivel between furnaces or extend-retract and move laterally between furnaces. Dry charge material is best suited for melting operations. Oils on the charge surface generate fumes from the furnace, which may need to be exhausted. Moisture on the charge also can be a problem, because a steam explosion can result. If dry materials are not readily available, a charge dryer is often employed to drive off any moisture remaining on the charge. The most common dryer consists of a refractorylined hood equipped with gas burners, which heat the solids on a vibrating conveyor by means of forced convection of the hot gases through the material. Figure 10 shows the various major components of the system. It is
114 / Melting and Remelting
Fig. 10
The major components of a dryer used to preheat induction furnace charge material. Integrated scrap preheat process combines (1) weigh hoppers, (2) preheat hood, (3) material transfer mechanism, and (4 and 5) furnace-charging apparatus into a single automated process.
Fig. 11
A mechanical clamshell-type power skimmer used to remove slag before each tap of an induction furnace
important to get the preheated charge into the furnace quickly to take advantage of the thermal energy in the material and to reduce the electrical energy output required for melting. The average temperature used for drying is typically 315 C (600 F); for preheating, it is 540 C (1000 F). When applied properly, the preheating of the charge material can increase the melt rate by 10 to 15%. Direct impingement
Fig. 12
Robotic attending of furnace. This robot can slag the furnace, take the bath temperature, as well as sample the metal.
Induction Furnaces / 115 of the flames on the charge material, particularly if it is thin, should be avoided because oxidation losses can occur. When dirty material is charged into the furnace, slag forms from the dirt and rust in the scrap as well as from the foundry return sand. Slag is undesirable in induction furnaces because it absorbs energy and erodes the lining. Therefore, it is removed each cycle with a skimmer to prevent slag accumulation. Mechanical Skimmers. Furnace operators can increase the power supply utilization by the use of mechanical skimmers (Fig. 11). This is particularly important in large furnaces because mechanical slagging units can minimize the time that power is either turned off or reduced for skimming during each charge-to-tap cycle.
A power skimmer can easily boost the output and productivity of the melting furnace. A skimmer is typically mounted on a jib. This tool has a long handle, which keeps the operator farther away from the hot furnace. The best and most effective type of slagger is the clamshell type. The skimmer is positioned over the furnace and lowered into the slag. With an automatic raking motion, the jaws of the unit close around the slag, which is then lifted clear of the bath and deposited into a container. The blades are typically dipped into a slurry to reduce slag adhesion during the next charge-to-tap cycle. Today (2008), most larger coreless units use back-slagging, where the furnace is back-tilted to allow the slag to be more easily removed.
A further development has been introduced recently that makes use of an industrial six-axis robot for the first time. In this innovation, the operator, from the safety of the melt deck control room, opens the cover and deploys the robot to skim the furnace, using preset programs (Fig. 12). As an added benefit, this system allows the operator to remotely obtain bath temperature, take samples, and add alloys. This further adds safety to the melting operation. ACKNOWLEDGMENT Updated from: R.Y. Perkul, Melting Furnaces: Induction Furnaces, Casting, Vol 15, ASM Handbook, ASM International, 1988, p 368–374
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 116-123 DOI: 10.1361/asmhba0005200
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Vacuum Induction Melting MELTING UNDER VACUUM in an induction-heated crucible is a tried and tested process in the production of liquid metal. It has its origins in the middle of the 19th century, but the actual technical breakthrough occurred in the second half of the 20th century. Commercial vacuum induction melting (VIM) was developed in the early 1950s, having been stimulated by the need to produce superalloys containing reactive elements within an evacuated atmosphere. The process is relatively flexible, featuring the independent control of time, temperature, pressure, and mass transport through melt stirring. As such, VIM offers more control over alloy composition and homogeneity than other vacuum melting processes. Vacuum induction melting can be used to advantage in many applications, particularly in the case of the complex alloys employed in aerospace engineering. The following advantages have a decisive influence on the rapid increase of metal production by VIM: Flexibility due to small batch sizes Fast change of program for different types of
the melting process. Accordingly, the vacuummelted superalloys (compared to EAF/AODmelted alloys) are improved in fatigue and stress-rupture properties. Control of alloying elements also may be achieved to much tighter levels than in EAF/ AOD products. However, problems can arise in the case of alloying elements with high vapor pressures, such as manganese. Vacuum melting also is more costly than EAF/AOD melting. The EAF/AOD process allows compositional modification (reduction of carbon, titanium, sulfur, silicon, aluminum, etc.). In vacuum melting, the charge remains very close in composition to the nominal chemistry of the initial charge made to the furnace. Minor reductions in carbon content may occur, and most VIM operations now include a deliberate desulfurization step. However, the composition is substantially fixed by choice of the initial charge materials, and these materials are inevitably higher-priced than those that are used in arc-AOD.
The charge generally consists of three portions: A virgin portion, which consists of material
that has never been vacuum melted
A refractory portion, which consists of those
virgin elements that are strong oxide formers and have the tendency to increase the Crucible Power supply
Induction coil
To vacuum pumps
steels and alloys
Easy operation Low losses of alloying elements by oxidation Achievement of very close compositional
tolerances
Precise temperature control Low level of environmental pollution from
dust output
Removal of undesired trace elements with
high vapor pressures
Removal of dissolved gases, for example,
hydrogen and nitrogen Vacuum induction melting is indispensable in the manufacture of superalloys. Compared to air-melting processes such electric arc furnaces (EAF) with argon oxygen decarburization (AOD) converters, VIM of superalloys provides a considerable reduction in oxygen and nitrogen contents. Accordingly, with fewer oxides and nitrides formed, the microcleanliness of vacuum-melted superalloys is greatly improved compared to air (EAF/AOD)-melted superalloys. Additionally, high-vapor-pressure elements (specifically lead and bismuth) that may enter the scrap circuit during the manufacture of superalloy components are reduced during
Process Description A VIM furnace is simply a melting crucible inside a steel shell that is connected to a highspeed vacuum system (Fig. 1). The heart of the furnace is the crucible (Fig. 2) with heating and cooling coils and refractory lining. Heating is done by electric current that passes through a set of induction coils. The coils are made from copper tubing that is cooled by water flowing through the tubing. The passage of current through the coils creates a magnetic field that induces a current in the charge inside the refractory. When the heating of the charge material is sufficient that the charge has become all molten, these magnetic fields cause stirring of the liquid charge. The optimal induction coil frequency for heating the charge varies with the charge shape, size, and melt status (liquid or solid). Older equipment used a single frequency, but newer power supplies are able to be operated at variable frequencies and are adjusted throughout the melt to obtain the most rapid heating/melting conditions.
Fig. 1
Basic elements of a vacuum induction melting furnace
Shunt Heating coil
Brick crucible
Cooling coil Ground detection
Fig. 2
Schematic of vacuum induction melting crucible (shell, coil stack, backup lining, and working lining)
Vacuum Induction Melting / 117 Bulk charger Crucible (pouring)
Melting chamber
Cover (removed) Launder Tundish car
Operating platform
Mold chamber
Electric room Power supply
Molds
Shop floor
Mold car
Fig. 3
Vacuum system
Schematic of a top-opening, double-chamber vacuum induction melting furnace
Table 1
Typical refractories used to line vacuum induction melting crucibles Maximum melt temperature
Refractory
MgO Al2O3 MgO-spinel Al2O3-spinel ZrO2 Graphite
C
1600 1900 1900 1900 2300 2300
F
2910 3450 3450 3450 4170 4170
Refractory density g/cm3
lb/in.3
Resistance to thermal shock
2.8 3.7 3.8 3.7 5.4 1.5
0.101 0.134 0.138 0.134 0.195 0.054
Good Good Poor Relatively good Poor Excellent
solubility of oxides and nitrides in the virgin charge A revert (or scrap) portion, which consists of both internal and external scrap that previously has been vacuum melted Vacuum-melted scrap has already had its gas content reduced to levels consistent with vacuum production. Scrap, however, has the possibility of having become contaminated during the production process, and care (expense) must be taken in the segregation and preparation of scrap materials for vacuum melting. In most VIM furnaces there is a vacuum lock bulk charger located directly over the crucible (Fig. 3). Charge material may be added to the heat through the bulk charger while melting is in process in the crucible. The material to be added is placed in bottom-opening buckets, placed in the bulk charger, and the charger is evacuated. The valve isolating the charger from the melt chamber is opened, and the bucket is lowered to a point close to the crucible top and the bottom opened so as to drop the charge material into the crucible. In constant operation, if a furnace is not to be opened to the atmosphere, all charge material for a heat will be added by this process.
Applications
Superalloys, high-quality Superalloys, high-quality Superalloys, high-quality Superalloys, high-quality Superalloys, high-quality Copper, copper alloys
steels steels steels steels steels
Older VIM furnaces may have been designed as single-chamber systems with the mold put inside the furnace before the beginning of the melt. The molten charge is then poured into the mold inside the furnace. Single-chamber furnaces thus must be opened after each heat to extract the molds and put in the new molds. Most furnaces have some system of large vacuum locks for transferring the prepared molds into the melt chamber. In double-chamber furnaces (Fig. 3), there is a separate chamber for the molds. The molten metal is transferred via launders (refractory-lined steel troughs) to a refractory tub (tundish). Some systems are designed to pour directly from the crucible into the tundish. The tundish contains a considerable volume of metal and allows residence time for entrained slag to float to the top of the tundish and be removed from the pour stream. The pour stream exits the bottom of the tundish. The pour time is regulated by pour temperature and the nozzle diameter of the tundish. The typical tundish is designed to provide a low-velocity flow path from the point of entry of the metal to the bottom nozzle. Refractories. Crucibles with a capacity of approximately 4500 to 22,500 kg (10,000 to 50,000 lb) are generally built up from refractory brick. Smaller furnaces, used for production of
master melt, may use single-piece crucibles. Refractory brick linings are usually two layers. The backup lining protects the induction coil in the event of a failure of the outer or working lining. The working lining is the primary interface with the metal and is replaced when erosion of the lining becomes excessive. Refractory life is also affected by the expansion of the refractory during the repeated melting cycles. Refractory brick is chosen with regard to resistance to erosion and expansion. Commercially available refractory brick may be incompletely sintered and expands during use, causing loss of crucible integrity. The refractory material used for the crucible lining is based on oxides such as Al2O3, MgO, CaO, or ZrO2 (Table 1). The lining is almost always rammed and sintered; prefabricated brick is used in larger furnaces. Dried silicate, combined with small oxide additions, appears to be very suitable for crucible lining because of its thermal characteristics. Because of an irreversible thermal expansion of 8% above 1000 C (1830 F), a high densification of the lining takes place during sintering. For this reason, this active lining is suitable for foundries. The behavior of the lining refractory with regard to stability at high temperature under vacuum must also be considered. The melting crucible material is not inert and is actually another source of oxygen and other impurities, depending on refractory type and condition. Therefore, both melt refining temperature and refining duration are carefully scrutinized. Proper melt stirring is integral to the deoxidation process and must be optimized through proper furnace power frequency and application procedure to prevent refractory lining erosion, a potential problem particularly during the controlled but more vigorous CO boiling portion of the process. Process Sequence. Figure 4 shows a typical process profile for the VIM of nickel- and cobalt-base superalloys. Before operation or at the completion the preceding heat, the melt chamber is isolated from the mold chamber and the vacuum integrity of the furnace is verified. Specific practices with regard to vacuum measurements will differ in detail. The deterioration rate of the vacuum is a measurement of the inherent vacuum integrity of the furnace. This integrity cannot be measured by vacuum alone, because the large pumping capacities of the pumps used in VIM are able to achieve low vacuum pressures even with significant leakage into the furnace. It is considered bad practice to continuously draw air into the vacuum furnace and across the melt. Immediately after ensuring that the furnace is vacuum-tight, the virgin portion of the charge is placed into the VIM furnace first. This may be done by opening the furnace or, more commonly, by charging the furnace through hoppers lowered through a large vacuum lock (bulk charger) located over the crucible. The furnace is capable of quickly pumping down to or maintaining vacuum levels below 100 mm (and often into the
118 / Melting and Remelting 105
1000
Installed power 1200
104
100 Pyrometer temperature
900
1200 1100
10
1.0
1000 900 800
2732 10 Melting power Pressure rise test
800
1.0
700 Casting end Casting start
600 500
All liquid
400
Additions/Sampling Pressure rise test
300 0.1
200 100
0.01
0
0.1
2552
2192 2012 1832 1652
0.01
1472
10−3
392
Additions Sampling
200
212
100 0
2372
Pyrometer temperature, ⬚F
1300 Melting power, kW
100
1400
1000
Pyrometer temperature, ⬚C
Melt chamber pressure, Pa
103
1100
Melt chamber pressure, mbar
1500
1. Charge 0
3. Charge
2. Charge 1
Melting
2
3
4 Refining
5
6 Superheating
10−4
7
32
Casting
Time, h
Fig. 4
Typical vacuum induction melting protocol for nickel- and cobalt-base superalloys
<10 mm) range. The virgin material is melted by application of current to the induction coils surrounding the refractory crucible. When the virgin material has been completely melted (all molten), it outgasses. The outgassing is monitored until it is complete. The outgassing is a response of the gas in solution in the melt to the low partial pressures of gases in the vacuum chamber. Some degassing is accomplished because carbon in the charge will form CO with the oxygen and also be evolved from the melt. The progress of degassing is followed by measuring the leak-up rate at set time intervals. When a constant leak-up rate is obtained, this indicates that degassing (refining) of the melt is complete. After degassing is complete, the reactive and revert charges are added to bring the charge to its planned weight. A ladle sample is taken after all additions are complete. Based on the analysis of this sample, trim additions are made to bring the melt into a very precise compositional range. Because there are no ongoing chemical changes in the melting, as there are in EAF/AOD, the compositional requirements of a melt may be met as closely as allowed by the reproducibility of chemical analysis. After the trim additions have been made, the temperature of the heat is brought precisely to the desired point, and the heat is poured. The heat, although produced in vacuum, will still have generated significant amounts of slag from the products of deoxidation, desulfurization, and the deterioration of the refractory crucible lining.
Process Controls. Process computerization and automation continue to enhance better reproducibility of the melts and computer modeling. Using charge and alloy calculations, it is easy to achieve the required chemical analysis with minimal cost. For this application, in the ideal case, the computer is linked to the analysis system computer so that the additives can be calculated and, if necessary, weighed and added directly after the analysis. If this alloy calculation is laid out as a charge calculation, the charges and the amounts of scrap necessary can be optimized. The complete calculation system enables a calculation of the alloying elements, starting with the amount of scrap needed and proceeding to the necessary alloy addition at the end of the refining period. A computer-controlled mass spectrometer system, specially developed for the VIM process, can make a significant contribution to the optimization of the melting process and its economics. Using a computer evaluation of the gas composition in the furnace chamber, information about the state of the degassing process and of the chemical reactions in the melt can be continuously obtained. The melting process can thus be controlled more exactly. Process parameters that can be closely controlled include: Determination of the correct time for the
addition of chemically active elements and the sequence of addition of such elements
Determination of the process steps, such as
completion of the refining period and the suitable time for tapping Advance detection of leaks in the furnace chamber, cooling water system, or hydraulic system Monitoring of the state of degassing of the crucible refractory lining Remelting by VIM for Shape Casting. Vacuum furnaces for precision casting are used to remelt alloys that have already been treated in vacuum. The remelted alloy is then precision cast into preheated ceramic molds. Heating and melting in such furnaces almost always takes place by induction, although some electron beam melters and vacuum arc furnaces are used. Electron beam furnaces offer the advantage of ceramic-free melting, but they do not permit a sufficiently accurate and reproducible temperature control for charges of more than approximately 5 to 7 kg (11 to 15 lb). Vacuum arc furnaces are used for titanium precision casting. Remelting with VIM furnaces for precision casting is done with two-chamber furnaces (Fig. 5a), where the melting crucible and mold are in separate chambers. The mold chamber is preferably placed under the melting chamber so that, in connection with another chamber for the charge or melting material, semicontinuous operation with minimal cycle times is possible. For cobalt- or nickel-base superalloys, the melting chamber pressure is of the order of 0.01 Pa
Vacuum Induction Melting / 119 (104 mbar, or 7.5 105 torr), while the mold lock would be evacuated to approximately 5 Pa (0.05 mbar, or 0.038 torr) within 1 min. In addition to the casting of equiaxed, directional solidification, and single-crystal blades for gas turbine engines, vacuum precision casting can be used for large parts (melt weights up to 1000 kg, or 2200 lb) or for the large-scale production of small parts. Examples of large parts produced using vacuum precision casting include compressor sections of jet engines and structural parts for stationary turbines.
A typical small casting produced by this method is a turbocharger wheel for automotive engines (Fig. 5b). Automotive turbocharger wheels weighing 400 to 500 g (0.9 to 1.1 lb) are produced in automatic precision casting machines. Total cycle time, including 40 s for melting, is typically 90 to 100 s. In the production of such small parts, the crucible is often integrated into the mold. Casting yields of 80 to 90% are obtained. Filters. One of the most critical stages with respect to cleanliness is the pouring of the melt.
Ceramic foam filters are used in some master metal operations to remove relatively large melt inclusions by means of entrapment. Foam filters are most effective where extremely high pour rate conditions and gross cleanliness problems prevail. Filter performance often varies because of the occasional use of filters with poor mechanical strength and/or thermal shock resistance. Foam cell particle breakage often results from handling during shipment or tundish installation and, if undetected, results in filter particulate in the alloy bar stock and subsequently cast components. Optimized VIM technology and practice without filters provide a clean alloy, without the inherent risks associated with filter use when applied to master metal production.
VIM Metallurgy
(a)
Fig. 5
(b)
Shape casting with vacuum induction melting, (a) Computer-controlled vacuum furnace with mold chamber. (b) Precision-cast turbocharger wheels for automotive engines. From left: mold with integrated crucible, bar stick, cast part, machined turbocharger wheel
Fig. 6
Vacuum induction melting is often done as the primary melting operation followed by secondary melting (remelting) operations such as electroslag remelting (ESR) and/or vacuum arc remelting (VAR).Various processing routes following VIM are illustrated in Fig. 6. Some superalloys are produced by a triple-melt sequence (VIM/ESR/VAR), where the VIM ingot is typically referred to as an electrode for subsequent ESR and/or VAR operation. The casting weight can vary from a few kilograms to 30 Mg (33 tons), depending on whether the VIM furnace is being used for
Potential processing routes for products cast from vacuum induction melting (VIM) ingots or electrodes. VAR, vacuum are remelting; ESR, electroslag remelting; EB, electron beam; HIP, hot isostatic pressing. Source: Ref 1
120 / Melting and Remelting precision casting or for the production of ingots or electrodes for further remelting. There are many different metallurgical factors that arise with induction furnace melting. The crucible material has an extraordinary effect on the metal/slag reaction because the ceramic outer wall reacts with the liquid metal and with the slag. In a VIM furnace, slag would be transported to the crucible wall by the characteristic bath movement. The result is that the slag solidifies at the wall and therefore has an insufficient reaction with the melt. It is therefore beneficial to use a slag that is more or less saturated with the oxides of the crucible lining in order to minimize heavy slag attack of the lining. In contrast to the ladle metallurgy processes with EAF, the crucible wall lining is susceptible to significantly higher erosion than the brick wall of a ladle. For this reason, metallurgical operations such as dephosphorization and desulfurization are limited. The metallurgy of VIM is primarily limited to the pressure-dependent reactions, such as carbon, oxygen, nitrogen, and hydrogen, and the evaporation of undesired elements with high vapor pressures, such as copper, lead, bismuth, tellurium, and antimony. Cleanliness can be significantly improved if a reactive liquid slag capable of absorbing oxides and sulfides is in contact with the melt. Vacuum induction furnaces are generally not operated with active slags. Therefore, reaction products can precipitate only on the crucible walls, and the melt may not be as clean as with other processing methods. Because the various alloys produced in vacuum induction furnaces must meet the highest quality requirements and because the vacuum induction furnace is primarily a consolidation unit and only secondarily a refining unit, the following methods are used to produce clean melts: Selection of a more stable refractory mate-
rial for the crucible lining
Rinsing of the melt with inert gas Minimizing the contact time of the melt in
the crucible
Exact temperature control to minimize cru-
cible reactions with the melt
Fig. 7
Vacuum induction refining process
Suitable deslagging and filtering techniques
during pouring Conception of a suitable tundish and launder system for good oxide removal For particular applications, however, the quality of the material produced by VIM will not be sufficient to satisfy the highest quality requirements with respect to cleanliness and primary structure. In this case, the VIM-produced material must undergo secondary melting processes (Fig. 6). VIM Refinement Process. The primary process of melt refinement is the removal of meltcontained oxygen by means of a reaction with carbon to form carbon monoxide (CO). The reaction occurs most readily at or near the melt surface, with the reaction kinetics being affected by crucible geometry and melt stirring. The CO bubbles form along the walls and, sometimes, bottom of the melt/lining-refractory interface (Fig. 7). This occurs preferentially at small crevices existing in the lining, with the bubbles growing during movement toward the molten metal/vacuum interface (Ref 2). Actual bubble formation is dependent on the number of gas molecules present, the pressure in the liquid at the level of the bubble, the temperature of the gas, and, for very small bubbles, the interfacial tension between the gas and the liquid metal. Following formation, bubble growth and mass transport within the liquid toward the liquid/vacuum interface is dependent on: The quantity of the dissolved gas The decreased pressure exerted on the bub-
ble as it rises in the melt
The bath temperature The time it takes for the bubble to rise
through the melt to the surface, which, in turn, is a function of melt stirring The pressure above the melt The interfacial tension between the bubble and the liquid metal The relatively vigorous, but controlled, portion of the boiling process results in the greatest CO removal. Concurrently, a slight nitrogen
loss is realized because of scavenging associated with the CO bubbles, and a slight sulfur reduction may occur during the CO supersaturation stage via sulfur dioxide (SO2) evolution. Minor tramp elements such as lead, silver, bismuth, selenium, and tellurium are partially evaporated during this period as well as throughout the entire refining process (Ref 3). Some undesirable elements, however, such as arsenic and tin, must be controlled through raw materials selection because they are not removed by vacuum refining. Bath refining from a somewhat vigorous CO boil is undertaken at a temperature and duration long enough to reach the so-called system equilibrium conditions, the assurance of which is provided by the attainment of consistent furnace leak-up rates. At this point, there is the addition of the refractory charge, which includes elements with reactivity toward oxygen (for example, aluminum, titanium, zirconium, and hafnium) that were withheld from the virgin charge because of their reactivity. Revert charge is also added. Once the boiling subsides, surface desorption of additional CO occurs, and it is during this nonboiling period that nitrogen removal (desorption) is most effective (Ref 4). Trace Element Removal. The removal of undesired volatile trace elements, such as arsenic, antimony, tellurium, selenium, bismuth, and copper, from the vacuum induction furnace is of considerable practical importance. These elements must be held to very low concentrations to avoid the risk of premature part failure, particularly for the production of superalloys for critical applications, such as jet engine parts. Figure 8 shows the influence of some trace elements on the stress-rupture properties of alloy 718. Because of the high vapor pressures of most of the undesirable trace elements, they can be kept to very low levels by distillation during melting under vacuum. Figure 9 shows how the different trace elements behave under vacuum. In a nickel-chromium melt, arsenic, antimony, and tin cannot be reduced over the gas phase, while copper, lead, selenium, tellurium, and bismuth can be reduced to a level far below 10 ppm. Nitrogen Degassing. Nitrogen gas can be decreased because its solubility at constant temperature is directly proportional to its partial pressures. As long as no strong nitride-forming elements are present in the melt, there is no difficulty in reducing the nitrogen content to 20 ppm (Ref 7–9). When nitride-forming elements such as chromium, vanadium, aluminum, and titanium are present, the activity of the nitrogen is very low; therefore, removal of nitrogen under high vacuum is difficult. If low nitrogen contents are desired in alloys containing nitride-forming elements, raw materials low in nitrogen must be used. Figure 10 shows the reduction in nitrogen levels obtained in the VIM of a die steel. From an average level of approximately 400 ppm at the beginning of treatment, the nitrogen has
Vacuum Induction Melting / 121 200
99.9 99.5 99
Silver Concentration, %
Rupture life, h
Selenium 120 Lead 80
Frequency, %
160
97.5 95 90
10
40
Fig. 0
10
20
30
50
40
60
70
80
Effect of trace elements on the stress-rupture properties of alloy 718. Test conditions: 650 C (1200 F), 690 MPa (100 ksi). Source: Ref 5
(0.5)
99.9 Arsenic Tin Frequency, %
0.1
Concentration, %
Antimony
Selenium 0.01
99.5 99 97.5 95
After treatment With argon bubbling
90
Without argon bubbling
80 70 60 50 40 30
Before treatment
20
Copper
10 0
100
200
300
400
500
Nitrogen content, ppm Bismuth
Lead
10−3 Tellurium (3⫻10−3)
0
20
40
60
80
100
120
Time, min
Fig. 9
Evaporation of trace elements from Ni-20Cr melts under vacuum. Source: Ref 6
been reduced to a level of 50 ppm. Additional argon purging during melting of the steel did not improve nitrogen removal. In the case of iron-nickel melts, however, the additional argon purging reduces nitrogen to much lower values. Hydrogen Degassing. Like nitrogen, hydrogen can be decreased via the gas phase because its solubility is directly proportional to its partial pressure. The solubility of hydrogen at atmospheric pressure and at 1600 C (2910 F) in iron or nickel melts is in the range of 30 ppm. During VIM, the hydrogen can be removed to very low concentrations. Figure 11
Before treatment
0
1.0
2.0 3.0 Hydrogen, ppm
4.0
5.0
11
Reduction in hydrogen content of X38CrMoV51 die steel after vacuum induction degassing
90
Impurity content, ppm
Fig. 8
80 70 60 50 40 30 20
Bismuth
0
After treatment
Fig. 10
Reduction in nitrogen content of X38CrMoV51 (Fe-0.38C-5.2Cr-1.3Mo-0.4V-1Si-0.4Mn) die steel after vacuum induction melting processing
shows the hydrogen contents of a die steel before and after vacuum treatment. It is evident from Fig. 11 that final hydrogen contents below 1 ppm can be routinely achieved. The reduction in hydrogen during the degassing treatment amounts to approximately 80%. Argon purging had no perceptible influence. The bath agitation in the case of this furnace, which is operating at normal frequency, is sufficient for hydrogen removal. Similar results were also obtained for the VIM processing of superalloys. Deoxidation of the melt in the vacuum induction furnace can be done via the generation of CO gas (C + O ! CO). The removal of oxygen from the melt as CO is favored by decreased melt chamber pressure, elevated bath temperature, and increased carbon activity (Ref 10). Proper melt stirring is integral to the deoxidation process and must be optimized through proper furnace power frequency and application procedure to
prevent refractory lining erosion, a potential problem particularly during the controlled but more vigorous CO boiling portion of the process. The final attainable oxygen content for a given carbon content is directly proportional to the CO partial pressure. The theoretical equilibrium values for oxygen are below 1 ppm in carbon-containing iron or nickel melts at 1600 C (2910 F) and at a pressure of 0.1 Pa (103 mbar, or 7.5 104 torr). However, the actual values are up to 1 order of magnitude higher. These higher oxygen contents result from impurities in the crucible lining, crucible outgassing and leakage, and, above all, the fact that the reaction does not reach equilibrium with decreasing oxygen content because of the difficulty of CO nucleation. In addition, the hydrostatic pressure of the liquid metal in the melt must be taken into account because only a relatively small bath surface area is in direct contact with the prevailing vacuum pressure. However, the influence of the liquid-metal hydrostatic pressure can be minimized through the use of additional agitation effects. The CO reaction occurs in two stages (Ref 1). The first stage is boiling, that is, the formation of CO bubbles within the melt along with a strong bath agitation as a result of this gas formation. The second stage is desorption, in which no more CO bubbles form inside the melt, and CO formation takes place only at the bath surface. Deoxidation with carbon is also a decarburization reaction, which is used for the production of low-carbon high-chromium steels. By using the CO reaction under vacuum, very low carbon contents can be achieved without a noticeable loss of chromium. The oxygen content in such alloys is higher because of the decrease in oxygen activity due to chromium. In this case, precipitation deoxidation using aluminum, silicon, or titanium must be applied. The CO partial pressures at which the lining oxides are reduced by carbon in the melt are:
122 / Melting and Remelting Carbon monoxide partial pressure Lining material
CaO ZrO2 MgO Al2O3 SiO2
kPa
0.04 0.13 0.53 0.53 81.1
torr
0.3 1.0 4.0 4.0 610.0
Source: Ref 1
It is evident that CaO represents the stable crucible material, while SiO2 is reduced at relatively high pressure. Apart from that, SiO2 can be reduced by alloying elements such as manganese, chromium, aluminum, titanium, or zirconium. This can cause a heavy chemical erosion of the lining, accompanied by an undesired silicon pickup in the melt. Because of these undesired reactions with a silica lining, a spinel-forming basic refractory material is used for the melting of high-grade steels and superalloys. The spinel formation during sintering leads to volume growth; therefore, as with the oxidic material, a densification of the rammed lining takes place. The vacuum induction degassing and pouring (VIDP) furnace (Fig. 12) is a variation of the conventional VIM furnace. The VIDP furnace design employs a modular concept that allows for connecting it to casting units for ingot casting, horizontal and vertical continuous casting, or powder production. Because of the smaller volume of the VIDP furnace compared to the VIM furnace and significantly lower desorption and leakage rates, it is possible to obtain very low pressures with lower pumping capacity. The lower part of the furnace can be decoupled and replaced rapidly; therefore, the VIDP furnace enables faster replacement of different furnace vessels with changes of alloy.
Melting Bath Agitation. As noted, melt agitation is an important process factor. Agitation is caused by the induction coil itself, depending on the power input and the installed frequency. Stirring is directly proportional to the amount of induced power and inversely proportional to the square root of the frequency. Therefore, a more intense stirring for a given size furnace occurs at a higher power and a lower frequency. Bath agitation from the induction coil is basically limited to the meltdown period. It can be done with the induction coil alone only if there is a high-power input and therefore strong bath heating. This high temperature increase is not always metallurgically desirable. For this reason, two agitation systems are offered in stateof-the-art vacuum induction furnaces:
vacuum. This ensures that the admitted purging gas flows through the melt and does not take the path of least resistance through the lining of the crucible. Both agitation processes offer advantages with regard to higher reaction rates, correct temperature adjustment, homogenization with simultaneously lower erosion of the crucible lining, and better cleaning of the melt. Figure 14 shows the three different agitation methods operating in a typical VIM process sequence. The agitation effect at 200 to 500 Hz in the melting phase assists in shortening the melting time. In the refining and superheating periods, it is better to use electromagnetic agitation at 50 to 60 Hz or argon purging through a porous plug.
Electromagnetic agitation with an additional
coil
Production of Nonferrous Materials
tom of the crucible
Apart from melting high-grade steels and superalloys, VIM is being increasingly used for the production of nonferrous metals and alloys. Table 2 shows some examples for possible use in nonferrous metallurgy. Aluminum alloys with additives such as zirconium, titanium, beryllium, cerium, tellurium, and cadmium must be melted under vacuum or under inert gas atmosphere because of their high reactivity with air and, in some cases, their toxicity. Aluminum-lithium alloys are also candidates for VIM processing. Copper Alloys. The production of highpurity copper having less than 2 ppm O can be accomplished only in a vacuum induction furnace. Oxygen content influences the electrical conductivity of copper alloys; the lower the oxygen content, the higher the electrical conductivity (Ref 11). For the production of
Agitation by argon purging through the bot-
The difference between the electromagnetic agitation by a separate coil and the agitation caused by the electromagnetic forces of the main coil lies in the fact that the bath can be more vigorously stirred without a temperature increase. A second transformer would be needed. Another method of bath agitation is derived from the principle of argon purging usually applied in ladle metallurgy. Argon purging has a long history of use in atmospheric melting furnaces. A characteristic of this purging is that the porous plug is not in direct contact with the liquid melt. Instead, the plug is covered with refractory material identical to that used for the crucible lining (Fig. 13). The basic difference is that the porous plug is in direct contact with the melt and is suitable for operating under
Crucible lining
Exchangeable porous plug
Basic set
Fig. 12
Schematic of the vacuum induction degassing and pouring furnace
Fig. 13
Argon purging system for vacuum induction melting furnaces
Vacuum Induction Melting / 123 oxygen-free copper, melting and casting must be carried out under vacuum. Selective Evaporation of Alloying Elements. The use of vacuum metallurgy is primarily linked with degassing and decarburization. A side effect of these treatments is the evaporation of
elements with high vapor pressures. In nonferrous metallurgy, this effect is used for the distillation of metals—for example, for the separation of lead and zinc in lead refining, in zinc production, and for the reduction of magnesium and nonalkali metals. Similarly, copper can be refined
from copper scrap by using the vacuum to evaporate volatile elements such as lead and zinc (Ref 12). ACKNOWLEDGMENTS Portions of this article were adapted from: A. Choudhury and H. Kemmer, Vacuum
Melting and Remelting Processes, Casting, Vol 15, ASM Handbook, ASM International, 1988, p 393–425 M. Donachie and S. Donachie, Superalloys: A Technical Guide, ASM International, 2002 G.L. Erickson, Polycrystalline Cast Superalloys, Properties and Selection: Irons, Steels, and High-Performance Alloys, Vol 1, ASM Handbook, ASM International, 1990 REFERENCES
Fig. 14
(a) Melting and (b) stirring modes of the vacuum induction melting process
Table 2 Possible applications for vacuum induction melting in the processing of nonferrous metals and alloys Metal or alloy
Metallurgical results
Aluminum, aluminum alloys High-purity metals
Low oxygen content, effective removal of oxides Oxygen-free copper (2 ppm O)
“New” alloys
Al-Mg, Al-Ce, Al-Zr, Al-Hf alloys; copper alloys with beryllium, cobalt, titanium, zirconium, and lithium; avoidance of lithium losses in production of Al-Li alloys Production of calcium, barium, lithium, strontium, and magnesium; recycling of hard metals; purification of alloys or melts; removal of zinc and lead
Metal production and recycling
Achieved by
Vacuum tank degassing without additions of polluting gases or materials Melting and pouring under steady, controlled atmosphere or vacuum; gas stirring; carbon as reducing agent Furnace and casting system under inert gas atmosphere; vacuum-tight, highly automated furnace
Thermic/alumino-thermic vacuum distillation; leaching with zinc and subsequent vacuum distillation melting; vacuum distillation
1. J.W. Pridgeon et al., in Superalloys Source Book, American Society for Metals, 1984, p 201–217 2. D. Winkler, Thermodynamics and Kinetics in Vacuum Metallurgy, Vacuum Metallurgy, O. Winkler and R. Bakish, Ed., Elsevier, 1971, p 42–54 3. Evaporation of Elements from 80/20 Nickel-Chromium During Vacuum Induction Melting, Transactions of the Vacuum Metallurgy Conference, American Vacuum Society, 1963 4. V.M. Antipov, Refining of High-Temperature Nickel Alloy in Vacuum Induction Furnaces, Stal0 , Feb 1968, p 117–120 5. W.B. Kent, Int. Voc. Sci. Technol., Vol 11 (No. 6), 1974, p 1038–1046 6. P.P. Turillon, in Transactions of the Sixth International Vacuum Metallurgy Conference (Boston), American Vacuum Society, 1983, p 88 7. G.A. Simkovich, Int. Met., Vol 253 (No. 4), 1966, p 504–512 8. H. Katayam et al., in Proceedings of the Seventh International Conference on Vacuum Metallurgy (Tokyo), The Iron and Steel Institute of Japan, 1982, p 933–940 9. A. Choudhury et al., World Steel and Metalworking Manual, Vol 9, 1987–1988, p 1–6 10. D.R. Gaskell, Introduction to Metallurgical Thermodynamics, Scripta Publishing, 1973, p 268–273 11. O. Kamado et al., Method of Producing Electrical Conductor, European Patent 0121152, 1986 12. J.G. Kru¨ger, Proceedings of the Fifth International Vacuum Metallurgy Conference (Munich), 1976, p 75–80
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 124-131 DOI: 10.1361/asmhba0005201
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Electroslag Remelting ELECTROSLAG REMELTING (ESR) is a secondary refining process (Fig. 1), where a consumable electrode is melted by immersion in a molten flux (slag) consisting primarily of fluorspar (CaF2). The heating is generated by an electrical current (usually ac) that flows from the base plate through the liquid slag and the electrode. Oxides may be added to the slag to raise the resistivity (so as to make melting more efficient electrically) or to modify the slag chemistry with regard to its reactivity with the metal being melted. Electroslag remelting is similar to vacuum arc remelting (VAR), except that ESR is carried out at normal atmospheric pressure and has a greater melting rate than VAR. When the tip of the electrode is melted by the molten slag, metal droplets fall through the liquid slag and are collected in the water-cooled mold (Fig. 1). The slag not only shields the metal pool from contaminating gases but can also be so chosen that it acts to segregate and absorb impurities from the molten metal in its passage through the slag from the electrode to the molten pool. During the formation of the liquid film, the metal is refined and cleaned of contaminants, such as oxide particles. The high degree of superheat of the slag and of the metal favors the metal/slag reaction. Melting in the form of metal droplets greatly increases the metal/slag interface surface area. The intensive reactions between metal and slag refine the melt. The remaining inclusions are very small and are evenly distributed in the remelted ingot. The absence of macrosegregation or heterogeneity in the distribution of nonmetallic inclusions in ESR ingot improves yield to justify the additional cost of remelting. In addition, the very clean and smooth surface of the ESR ingot helps to reduce production costs, because surface conditioning before hot working is not necessary. Another special feature of the ESR process, as in VAR, is the directional solidification of the ingot from bottom to top. The macrostructure is marked by an extraordinarily high density and homogeneity as well as by the absence of segregations and shrinkage cavities. Electroslag remelting is commonly used to produce the highest levels of quality in plate steels, particularly in thick plates. The homogeneity of the ingot results in uniform mechanical properties in the longitudinal and transverse directions after hot working. For steels, the slag
composition can be adjusted to serve as a desulfurizer or dephosphorizer as well as a reservoir for floating inclusions. Sulfur can be lowered to below 0.002%. The cleanliness and homogeneity of ESR ingots results in steels with excellent mechanical properties and superior toughness (Fig. 2) compared to conventionally manufactured steel. In superalloys, ESR is
Fig. 1
often used as an intermediate remelt in a socalled triple-melt process (VIM-ESR-VAR), where the ESR electrode is produced by vacuum induction melting (VIM) and then remelted a third time by VAR after ESR (Ref 4). The theory of ESR was developed in the 1930s and was the subject of a U.S. patent in
Schematic of an ESR furnace with multiple electrodes for large ingot production
Electroslag Remelting / 125 1945 (Ref 5). However, it took more than 30 years for a general breakthrough on a production scale (Ref 6–12) in primary (mill) products. Currently, ESR metallurgy is a well-established production technology (Ref 13). Global production capacity of ESR steel is over a million (106) tons, which includes low-alloy highstrength steel, bearing steel, tool and die steel, and stainless and heat-resistant steels (Ref 14). Ingots from high-quality heavy ESR are over 200 tons, and production cost of heavy ESR manufacturing is significantly reduced by electroslag hot topping. Related technologies include electroslag casting, electroslag turning casting, Boher electroslag tapping, and electroslag centrifugal casting. Process variations of ESR technology include vacuum ESR and pressurized ESR. Electroslag remelting under vacuum (VACESR) is a process variation that has received some attention, because the traditional ESR process is essentially atmospheric and thus ineffective in reducing dissolved gases such as hydrogen and nitrogen. Vacuum ESR is not a superalloy production technology, given the predominance of triple-melt (VIM-ESR-VAR) for superalloys. Pressure electroslag remelting Temperature, ⬚F −4
Impact strength, J
80 60
32
68 104 140 176 212 73.8 ESR (tangential rim) 59.0
ESR (axial core)
44.3
40
29.5
20 Conventional (tangential rim)
0 −40 −20
20
0
40
Conventional (axial core)
60
Impact strength, ft-Ib
−40 100
is a process variation that increases dissolved nitrogen in the melt of austenitic steels for strengthening. Another ESR process variation is electroslag rapid remelt (ESRR). This technique allows production of smaller cross-sectional shapes (rounds and squares), which can be directly processed into smaller shapes on conventional open-die presses and rolling mills, saving conversion time and costs as compared to starting with typical larger round ingots. The ESRR technology is also being developed for nucleated casting (Ref 15), which is a type of spray casting.
ESR Furnaces Major components of modern singleelectrode ESR plants are shown in Fig. 3 (with fixed mold) and Fig. 4 (with a retractable bottom plate). In the case of a retractable bottom plate furnace, current feedback does not take place through a closed copper pipe; instead, it is accomplished with four symmetrically arranged current feedback tubes. The ESR process, like VAR, is a fully automated operation. Unlike a VAR crucible, the ESR crucible is normally self-contained in that the inside and outside shells through which the cooling water runs are assembled as a single piece. A watercooled stool is assembled to the shell to form the bottom of the crucible. A starter or striker plate is usually enclosed in the junction between the crucible shell and the crucible stool.
For crucibles to be used with hot slag starts, the crucible may contain an opening at the bottom (a “mousehole”) through which the molten flux would be introduced. The top of the crucible does not mate to the rest of the furnace, so it operates exposed to air. The electrode is attached to the ram through a stinger (Fig. 5). The ESR stinger is welded to the electrode outside the furnace. The stinger is often tubular, because only the periphery of the electrode is involved in the welding process. The top of the ESR furnace contains the ram drive and load cells. It is connected electrically to the crucible stool through, typically, four vertical bus bars located at 90 intervals around the crucible. The ESR head commonly is able to be translated in the horizontal direction by an XY drive, which allows centering of the electrode in the crucible prior to the start of the melt. (Note that XY capability is inherently different in ESR compared to VAR.) The power supply is most often ac. Most ESR furnaces have two melting stations (crucible setups). While an electrode is melting in one station, the next station is prepared for melting. When one melt is finished, the ESR head is rotated in the horizontal plane to be located over the
14.8
0 80 100
Temperature, ⬚C (a)
167
212
257
3822
in.
122
3500 ESR (0.002% S)
3185
2800
2548
Fracture toughness, ksi
Fracture toughness, MPa
m
Temperature, ⬚F 77
4200
1911
2100 Conventional (0.001% S)
1400
1274 637
700 25
50
75
100
125
Temperature, ⬚C (b)
Fig. 2
Comparison of properties of steel rotor forgings made from ESR and conventionally melted ingots. (a) Impact strength of grade X22CrMoV121. (b) Fracture toughness of grade 30CrMoNiV511. Specimen orientation and location are indicated next to curves. Source: Ref 1–3
Fig. 4 Fig. 3
Basic design of a modern ESR furnace with a fixed mold. 1, ram drive system; 2, electrode ram; 3, XY adjustment; 4, load cell system; 5, sliding contact; 6, four bus tubes; 7, pivoting drive; 8, electrode; 9, mold assembly; 10, coaxial bus tube; 11, base plate; 12, multicontacts
Schematic of an ESR furnace with retractable bottom plate. 1, electrode drive system; 2, load cell system; 3, stinger crosshead; 4, two ball screws; 5, three bus tubes; 6, mold assembly; 7, ingot; 8, ingot withdrawal table; 9, three bus tubes; 10, three ball screws; 11, three sliding contacts; 12, base plate; 13, furnace head drive; 14, electrode; 15, three sliding contacts; 16, XY adjustment
126 / Melting and Remelting second crucible/electrode stinger setup. The ESR process is also capable of using multiple electrodes, which can be melted into a single mold. Electrode preparation is not as critical a part of quality ESR operations as for VAR. Oxides on an electrode surface are incorporated into the slag. Thus, electrodes for ESR are generally not ground. An exception to this is for production of premium-quality material where the end-user specifications require grinding to ensure the removal of potential iron-mold pullout on the electrode surface. Also, the shrinkage cavity at the top of the electrode is not a problem, if the electrode is melted top down. Thus, superalloy ESR electrodes are not commonly hot topped in VIM or cropped prior to melting. There are two possible start-up scenarios: cold start and hot start. In cold start, the slag and small particles of the alloy to be melted (usually machining chips) are placed on the starter plate. The electrode is touched into the slag-alloy mix and backed out to establish an arc. High melt power is used in this start-up phase. The arc melts down both the slag and the metal particles, at which point the electrode becomes immersed in the slag, and the melt process shifts to melting of the electrode surface by the slag. It is necessary that the start-up power be sufficiently high that actual welding (melt-in) occurs between the embryonic ingot and the starter plate. This is required so that the electrical conduction path will be predominantly through the electrode, then the slag to the ingot, from the ingot through the starter plate, and then to the stool. In hot slag starts, the slag is melted externally by electric arc in graphite crucibles. The molten slag is introduced, generally through a bottom mousehole, into the crucible. The electrode is lowered into the slag, and melting
Fig. 5
ESR electrode with stinger
commences. Although high initial melt power is not required to melt slag and metal starter material, a high-power profile start-up is generally used to ensure the melt-in of the ingot to the starter plate (to ensure good electrical conduction paths). ESR Control Parameters. Control of the ESR process is considered from two different aspects: control of pool depth and shape (control of solidification structure) and development of a satisfactory ingot skin. Pool depth and shape are controlled primarily by melt rate (power input). Electroslag remelting generally is run using melt rate control, with continuous amperage adjustments to maintain a uniform melt rate. Due to the high thermal inertia of the ESR process, the melt rate of ESR is more variable than in VAR. Secondary control of pool depth and shape is achieved by maintaining a uniform, shallow immersion of the electrode in the slag. The passage of molten metal droplets through the slag affects slag resistivity and thus causes variation in the operating voltage of the system. The operating voltage appears as a band of values due to this variation. The width of that band is called the volt swing. The electrode immersion is proportional to the volt swing, and the ram drive of the ESR advances the electrode to maintain the desired swing value. Production of satisfactory ingot skin is not actively controlled during the melt but is due to the choice of variables. The most important of these are melt rate, depth of the slag pool, electrode immersion depth in the slag (analogous to arc gap in VAR), ingot and electrode diameter (as in VAR), and choice of slag composition. Higher melt rates produce better ingot surfaces. Because higher melt rates also produce deeper molten pools and move the process closer to the conditions for freckle formation, selection of the correct balance of parameters is critical. The choice of ingot and electrode diameter defines the clearance (annulus) between the crucible and the electrode. Insufficient clearance is generally regarded as a problem in processes using cold slag starts, in that the slag-chip mix may get hung up in the annulus. However, a tight annulus may result in more uniform heat distribution across the molten pool. There is no published information to support either viewpoint. Thus, annulus becomes, much more than in VAR, a parameter dictated by the individual producer’s philosophy. Electrode immersion is generally thought of as the depth of penetration of the electrode into the slag. There is no definitive method for measuring the immersion. In superalloy melts, observation of interrupted melts suggests that the meniscus of slag reaching up the side of the electrode will generally not exceed 6.4 mm (0.25 in.) (Ref 4). Thus, schematics of the ESR process are misleading when they imply significant immersion into the molten slag. However, the degree of immersion, as affected by the ram drive, can produce drastically
different electrical signals. The quality of the melt has been found to be responsive to these signals. Thus, while there is no realistic understanding of what depth of immersion means physically, the parameter that controls it can be demonstrated to have a significant effect on melt quality. The electrode immersion plus the thickness of the slag cap may be considered as the resistance (ac impedance) in the circuit. Changing the weight of slag in the system or the slag composition changes the impedance of the circuit and thus changes the amperage and voltage required to maintain a given power input (for constant melt rate). The slag cap thickness is the primary driver, while electrode immersion may account for 10 to 20% of the total impedance. Measurement of the changes in this resistance or of the corresponding changes in voltage may be used to control the immersion of the electrode in the slag. This suggests that the voltage and impedance changes associated with electrode immersion are actually the impedance of the circuit associated with the interfacial resistance between the electrode melt surface and the slag surface. It thus should be realized that in ESR of superalloys, the electrode immersion is always extremely shallow, and, in fact, the most common electrode positioning is that of “skittering” on the top of the molten slag cap. The degree of this skittering is commonly controlled by the repetitive change in voltage in the process (volt swing). The electrode melt rate is the other important parameter to be controlled. The melt rate is dependent on the melt current. Although some VAR processes are run in constant current rather than in melt rate control, ESR processes are exclusively melt rate control. Melt rates are generally calculated using rolling averages. Load cells are subject to the generation of false signals and to errors in the absolute value. Thus, load cell design and maintenance are important factors in maintaining the quality of ESR. In ESR, the melt current responds to changes in the melt rate and attempts to hold the melt rate constant at the set point. Because of the presence of the slag, which must be heated or cooled, in the system, the response time of ESR to the melt current change is much slower than that of VAR. A fourth factor controlling melt quality and unique to ESR is the choice of slag and the volume of slag used. The resistance of the slag cap depends on both the resistivity of the slag and the thickness of the slag cap through which the melt current must travel. An additional consideration is the degree of chemical reactivity of the slag with elemental components of the electrode. Figure 6 is a nominal plot of process parameters for the remelting of a nickel-base alloy. Figure 7 shows voltage swings and the major control parameters for a hypothetical ESR melt. The melt begins with a high-amp start-up in amperage control and, at the completion of the start-up, is placed into melt rate control.
Electroslag Remelting / 127 6
Melt rate, kg/min
Ingot diam: 460 mm Ingot weight: 2070 kg
Set rate 3.5 kg/min
4
6.6
3 2
4.4
Start
1 Power control 0 5 Power, kW ⫻ 10−2 Voltage, V ⫻ 10−1 Current, kA ⫻ 10−1
8.8
2.2
Hot topping
Melting rate control
Melt rate, Ib/min
11.0
5
0 Power control
4
Voltage
3 Power 2 Current
1 0
Ensuring consistent remelting over a period 0
1
2
3
4
5
6
7
8
9
10
Time, h
Fig. 6
Electroslag Remelting of Heavy Ingots. At the end of the 1960s, the concept of using ESR plants to manufacture large ingots weighing up to 350 Mg (385 tons) gained acceptance. Increasing demands for energy and the trend toward larger electrical power-generating units required cast ingots weighing 100 Mg (110 tons) and more for the manufacture of generator and turbine shafts. With ESR, it is possible to achieve higher yields and to avoid such internal defects as porosity, macrosegregation, and the accumulation of nonmetallic inclusions. An early example of a large ESR furnace in the 1970s had a retractable bottom plate and manufactured ingots 2300 mm (90.5 in.) in diameter, 5000 mm (197 in.) in length, and weighing up to 160 Mg (176 tons) (Ref 2) (Fig. 8). Remelting of large ingots of this size in retractable bottom plate furnaces demanded the solution of specific technical and metallurgical problems, such as: of several days without interruption
Prevention of steel and slag breakout when
remelting with a retractable bottom plate
Achieving a good ingot surface Ensuring directional solidification to avoid
ESR process parameters for the melting of a nickel-base alloy
macrosegregation and shrinkage cavities
Controlling the steel and slag compositions
Volt swing - Ram travel (in.) - Resistance Voltage swing and resistance
over the entire ingot height
Ram
Adjusting for low hydrogen content Adjusting for a low aluminum content
iion Posit
Volt Swing Resistance
Voltage
Melt rate - Voltage - Amps 40 35 30 25 20 15 10 5 0
Voltage
Melt Rate Amperage
Elapsed time during melt
Fig. 7
Melt trace for ESR showing variation of melt current, melt rate, ram travel, slag resistance, voltage, and volt swing
Note that the change in voltage is the result of the deliberate change in amperage. Formation of a shrinkage cavity several hundred pounds deep into the ingot can occur from solidification of the large molten pool after the end of the melt, if the electrode is completely consumed or if the power abruptly shuts off. This cavity must be removed during subsequent processing. To minimize the amount of material that must be cropped because of the end-of-
melt shrinkage, it is common to step down the melt current as the end of melt approaches and leave a small amount of electrode. Unlike VAR, this biscuit is generally not used to put electrical power into the melt. Rather, it simply prevents radiation loss from the slag. Because of the presence of the slag as a heat source at the top of the ingot, ESR hot topping practices are not as extended as are those for VAR.
(<0.010%) for rotor steels; this has a decisive influence on the creep-rupture properties of the rotors After these problems had been solved, heavy rotors for electrical generators could be manufactured from ESR ingots. Interior defects such as macrosegregation, shrinkage cavities, and nonuniform distribution of inclusions can be avoided in large ingots only if directional solidification is ensured over the entire ingot section. By maintaining the correct melting rate and temperature of the slag, directional solidification can be achieved for large ingot diameters. Accordingly, the ESR ingot is free of macrosegregation in spite of the large diameter. Several electrode changes are necessary to produce such large ingots. In the case of a multi-electrode operation in a single mold, the formation of the ingot can be concluded by a hot topping procedure performed by withdrawing all but one of the electrodes and continuing the process in single-phase mode with the remaining electrode centralized in the mold, for example, by appropriate sideways displacement of the latter on a wheeled support. In such a case, loss of bottom contact to the ingot during the multiphase operation could prevent the subsequent single-phase hot topping step from proceeding. Moreover, loss of contact is more likely to occur with such a multiphase operation than with a single-phase operation, because in the latter the flow of current between the ingot bottom and the mold base has some tendency to keep a thin slag film
128 / Melting and Remelting
Fig. 8
160 Mg (176 ton) ESR ingot before forging into a generator rotor
Fig. 9
Schematic of solidification relationships in VAR (left) and ESR (right) melting processes. Source: Ref 16
between the ingot and mold base in an electrically conductive state. The refining process for producing steel ingots of large sizes (e.g., 200 to 400 ton class) is possible by controlling the rate of solidification in electroslag remelting or with electroslag hot topping techniques.
Ingot Solidification The solidification structure of an ingot of a given composition is a function of the solidification rate and temperature gradient at the
solid/liquid interface. When remelting a consumable electrode in a water-cooled copper mold, the solidification rate is constant because of the constant heat transfer through the cooling water; therefore, a relatively high temperature gradient at the solidification front must be maintained during the entire remelting period to achieve a directional dendritic structure. The structures in the solidified ingot are governed by the same principles as those governing a solidifying VAR ingot. However, the nature of the heat transfer (through an oxide skin) and the presence of a heat reservoir (the molten
slag) at the top of the ingot/molten pool make ESR structure inherently different from that of VAR. Figure 9 schematically illustrates the relationship between the ESR electrode, the solidifying ingot, and the shape and depth of the molten pool. A major feature of ESR solidification is that a “skin” of oxide is incorporated between the ingot surface and the crucible wall. Similar to VAR, the distance between the side of the electrode and the crucible wall is the annulus. The depth of the slag pool is an important operating parameter that controls the electrical resistance of the system. The depth of immersion of the electrode into the slag is an important parameter in controlling the molten pool shape. As shown in Fig. 9, the heat extraction in ESR is of a much greater magnitude than for VAR. Due to the insulating characteristic of the ingot slag skin, heat extraction in ESR is inherently less efficient than in VAR. Thus, ESR is a process with high thermal inertia and slow responses (change in pool shape) that occur in response to changes in power input. The presence of an ingot slag skin not only improves the ingot surface quality but also, more importantly, eliminates the presence of a solute-lean shelf layer on ESR ingots. (As oxide particles from the electrode are dissolved into the slag, they do not agglomerate on the ingot surface.) Because ESR ingots do not have an ingot shelf, they are inherently free of the formation of drop-in (discrete) white spots from this source, but drop-in electrode may still be a source of white spots in ESR. Solidification white spots have not been reported in ESR product. However, one consequence of the ingot slag skin is that heat extraction in ESR is less efficient than for VAR. This is shown schematically in Fig. 9 by the difference in pool shape between the two processes. ESR pools tend to be both deeper and more V-shaped (compared to a U-shaped VAR pool). Deep pools favor freckle formation, which can only be counteracted by reducing ingot size. Consequently, the maximum size of ESR ingot produced (for any given alloy system) is smaller than one that can be produced by VAR. Figure 10 schematically illustrates the structure from an ESR molten pool (the relative sizes of annulus and immersion are not to scale). Again, the most important feature shown is the depth of the molten pool (where the depth and angle of the liquid-solid zone also controls freckle formation in superalloys). Compared to VAR, few studies have been done on the nature of the molten pool in ESR. The pool differs fundamentally from a VAR pool in that the molten slag cap provides a source of heat to the system, and the slag skin on the solidifying ingot provides insulation, reducing heat extraction. Consequently, for comparable melt rates and crucible diameters, an ESR ingot will have a deeper, steeper-sided pool. Compared to VAR, ESR ingots have traditionally been regarded as being more prone to the formation of positive segregation effects. However, even though the ESR pool may inherently be more
Electroslag Remelting / 129 prone to positive segregation, that shape may be altered by judicious choice of melting parameters. Even in the case of directional dendritic solidification, the microsegregations increase with increasing dendrite arm spacing. For optimal results, the objective should be a solidification structure with dendrites oriented as close to parallel to the ingot axis as possible. However, this is not always possible. To achieve a good ingot surface, there is a minimum energy requirement, and therefore a minimum melting rate, which is a function of ingot diameter. This means that the melting rate for large-diameter ingots cannot be maintained for axis-parallel dendrite growth. Like VAR, the melting rates for different steels and alloys are a function of ingot diameter. Ingot Defects. In spite of directional dendritic solidification, remelted ingot may include various defects, such as the formation of: Tree ring patterns Freckles White spots
These types of defects also occur in VAR ingots and are discussed in more detail in the article “Vacuum Arc Remelting” in this volume. In Fig. 10, the schematic for the edge of a solidifying ingot indicates a processing advantage for ESR that has made it desirable to develop this process for use in critical components. The insulating nature of the slag
creates conditions at the edge of the ingot that do not produce a shelf. Thus, with no arc instabilities of concern and no shelf to be undercut, the formation of discrete white spots in ESR product could only occur by some infrequent occurrence, such as drop-in of pieces from an unsound electrode. Positive Segregation. Alloys that have highdensity interdendritic liquids may form unique structures in ESR. The high-density liquid tends to flow downhill toward the center of the ingot. Although channel defects may not be formed, large pools of high-solute alloy may be formed. These regions may form in widely separated parts of an ingot and be separated by “normal” material. Because of the greater number of variables inherent in ESR, the underlying causes for such segregation have not been determined. The nature of these areas and their effect on properties have not been reported. Electroslag Remelting Ingot Surface. The interaction of the molten metal with the slag cap should be understood with regard to obtaining good surface on ESR ingots. As shown in Fig. 10, the molten slag solidifies against the water-cooled crucible side just as the molten metal would. The molten metal pool, in addition to being V-shaped (containing the mushy zone), also contains a fully molten straightsided component or metal head. The metal head remelts some of the solidified slag on the crucible wall, so a thinner slag layer is then contained between the ingot and the crucible wall. Should this slag skin be locally penetrated by the metal head, a thin stream of metal will be driven through the hole. Such a metal fin or bleedout is illustrated in Fig. 11. Note that the penetration of the slag is through a circular region that is essentially the size of a dime.
Metal head
Liquid metal penetration of slag
Ni balls Metal fin
Fig. 11
Fig. 10
Longitudinal schematic of the structure developed in an ESR ingot during melting
Longitudinal section through a metal fin that was formed by a bleedout on the surface of an ESR nickel superalloy ingot. Note: The circular, darketching features are nickel balls introduced in an experiment to delineate the shape of the molten pool. They are not related to the bleedout mechanism.
The molten metal not only moves down the ingot surface in response to gravity but is forced upward from the penetration due to the combined hydrostatic pressure of the molten metal head and the molten slag cap. Bleedout formation is sensitive to melting conditions and slag composition. While there is no published association of bleedouts with detrimental structure, there is an economic factor involved, in that an ingot with bleedouts will require grinding prior to forging or other hot working.
Process Variations Several variations of ESR have been developed: remelting under high pressure, known as pressure electroslag remelting; remelting under reduced pressure; and electroslag rapid remelting. Pressure Electroslag Remelting. The strength of austenitic steels can be increased by alloying with nitrogen as long as the nitrogen is interstitially dissolved in a stable condition. Nitrogen solubility, however, is limited in the case of melting under atmospheric pressure for a given steel composition. Because the solubility of nitrogen in iron is proportional to the nitrogen partial pressure in the furnace atmosphere, there is the possibility of alloying nitrogen into the melt under higher pressure and allowing it to solidify. The ESR process is suitable for this technique. The first trials in this area were carried out in 1971 (Ref 17). For some years, a high-manganese steel has been successfully remelted under a pressure of 4.2 MPa (610 psi). Up to 0.5% interstitial solubility of nitrogen in steel has been achieved. The nitrogen-alloyed austenitic steels are used for the manufacture of retaining rings in generators. Electroslag remelting under vacuum (VAC-ESR) is another interesting process. Alloy 718 has been remelted in a VAC-ESR pilot plant with a fluoride-free slag based on lime and alumina (Ref 18). Because the remelting is carried out under vacuum, problems of oxidation of the melt do not arise. In addition, dissolved gases, such as hydrogen and nitrogen, can be removed. Thus, the advantages of both ESR and VAR are combined in one process (Ref 18). Nonetheless, VAC-ESR is not a production practice for superalloys. Electroslag rapid remelt is a process technology where electrodes are ESRed into smaller cross-sectional shapes (rounds and squares) that can be directly processed into smaller shapes on conventional open-die presses and rolling mills, saving conversion time and costs as compared to starting with typical larger round ingots.
Steel ESR There are a number of publications that deal primarily with improvements in the physical properties of remelted steel (Ref. 6–12).
130 / Melting and Remelting However, they discuss not only the directional solidification of the metal but also the influence of liquid, superheated, and therefore reactive slags. The extent and direction of the metallurgical reactions in the ESR process are determined by the steel composition, the slag used, and the atmosphere used (inert gas, air, etc.). The principal feature of the ESR process is the slag bath. A continuous transport of liquid metal through the slag takes place. During this transport, the slag and the metal compositions change, according to the kinetic and thermodynamic conditions. To perform its intended function, the slag must have some well-defined properties, for example: Its melting point must be lower than that of
the metal.
It must be electrically efficient. Its composition should be such that the
desired reactions, such as removal of sulfur and oxides, are ensured. It must have suitable viscosity at remelting temperature.
of easily oxidizable elements, such as aluminum and silicon, during remelting. To counteract this oxidation, the slag should be continuously deoxidized, preferably with aluminum. With proper slag composition and
Table 1 Compositions of slags commonly used in ESR Composition, % Slag designation(a)
CaF2
CaO
MgO
Al2O3
SIO2
Comments
100F 70F/30
100 70
... 30
... ...
... ...
... ...
70 70 50 70
20 15 20 ...
... ... ... ...
10 15 30 30
... ... ... ...
Electrically inefficient; use when oxides are not permissible. Difficult starting; high conductivity; use when aluminum is not allowed; risk of hydrogen pickup
40 60 80 60 ...
30 20 10 10 50
... ... ... 10 ...
30 20 10 10 50
... ... ... 10 ...
70F/20/0/10 70F/15/0/15 50F/20/0/30 70F/0/0/30 40F/30/0/30 60F/20/0/20 80F/0/10/10 60F/10/10/10/10 0F/50/0/50
Good general-purpose slags; medium resistivity Some risk of aluminum pickup; good for avoiding hydrogen pickup; higher resistivity
g
Good general-purpose slags Moderate resistivity; relatively inert Low-melting, “long” slag Difficult starting; electrically efficient
60 ⬉10% Al2O3
50 40 30 20
(Eq 1)
(Eq 2)
It is evident that a saturation of the slag with sulfur does not take place; therefore, the desulfurization capacity of the slag remains intact throughout the entire remelting process. With a highly basic slag (CaO/SiO2 > 3), more than 80% of the sulfur can be removed. Behavior of Oxygen. As mentioned previously, the ESR process is usually carried out under a normal air atmosphere. Oxidation of the metal is unavoidable. Oxygen can be transferred into the metal in several ways: Oxidation of the electrode surface above the
slag bath Oxidation on the slag surface of elements with variable valences, such as iron and manganese Oxides attached to the electrode surface Transfer of oxygen into the slag also takes place, and oxygen is transferred into the metal according to Eq 1. This results in losses
for C16 10 Average oxygen content, ppm
The second reaction is the slag/gas phase reaction. In this case, the sulfur absorbed by the slag is removed by the oxygen of the gas phase in the form of gaseous sulfur dioxide: 3 CaS þ O2 ðgasÞ ! CaO þ SO2 ðgasÞ 2
g
(a) The universal designation system for slags is the percentage of constituents listed in the order CaF2/MgF2/CaO/MgO/Al2O3. Thus, a common base slag such as 60%CaF2/20%CaO/20%Al2O3 would be referred to as 60/0/20/0/20. However, industrial designations often omit the 0s. Thus, the previously mentioned slag is generally referred to as 60/20/20. Source: Ref 7
Slags for ESR are usually based on calcium fluoride (CaF2), lime (CaO), and alumina (Al2O3). Magnesia (MgO), titania (TiO2), and silica (SiO2) are also added, depending on the alloy to be remelted. Table 1 shows the compositions of some common ESR slags. Behavior of Sulfur. One of the primary advantages of the ESR process for steel is the good desulfurization of the metal. The final desulfurization is determined by two reactions. The first is the metal/slag reaction, in which sulfur is transferred from the metal to the slag: FeS þ CaO ! CaS þ FeO
remelting techniques, oxygen contents of less than 20 ppm in unalloyed steel are possible. Figure 12 shows the influence of slag composition on the final oxygen content of remelted ingots.
0 40 15–25% Al2O3 30 20 10 0 40 30–45% Al2O3
30 20 10 0
1
2
3
4
5
6
7
8
9
Slag basicity
Fig. 12
Effect of slag composition on the oxygen content of the remelted ingot. Slag basicity is the ratio of CaO to SiO2.
Electroslag Remelting / 131
Superalloy ESR Generally, superalloy ESR uses a cast VIM electrode as a consumable electrode. Unlike VAR, high-vapor-pressure tramp elements are not removed by ESR. Thus, magnesium may be retained at higher levels than is possible in VAR (with beneficial effects on hot workability). A beneficial chemical reaction related to workability is the removal of sulfur. Detrimental compositional changes may occur in reactive elements (notably titanium and aluminum), because these elements change to attain chemical equilibrium with components of the slag. Oxides incorporated in the electrode being melted are effectively removed by the ESR process, producing an ingot that is generally cleaner than a VAR ingot. All ESR slags contain CaF2 as the primary constituent. Oxides may be added to the slag to raise the resistivity (so as to make melting more efficient electrically) or to modify the slag chemistry with regard to its reactivity with the metal being melted. Additions of CaO, MgO, and Al2O3 are made to modify both the resistance of the slag and its melting point. All commercial-purity CaF2 contains SiO2. During melting of titanium-bearing alloys, an exchange will take place between the titanium in the alloy and the silicon in the slag. The slag becomes enriched with TiO2, and the silicon content of the alloy increases until an equilibrium is established, usually within the first few hundred pounds of melting. To minimize this exchange, some commercial slags are buffered with intentional additions of TiO2. Titanium and zirconium are superalloy elements that often interact with the slag during ESR. Many proprietary slags contain additions of TiO2 and/or ZrO2 to prevent excessive loss of these elements into the slag during melting. The loss is due to exchange with any oxide in the slag that is less stable than the element in the electrode. A common problem with titanium-bearing superalloys is that commercially available CaF2 contains SiO2. During ESR, the SiO2 is converted to silicon, which is picked up by the metal. Thus, the slag is enriched in TiO2, which reduces the titanium content of the metal. This reaction generally results in a titanium gradient in the ingot, because the SiO2 content is consumed early in the melt and no longer plays a part in the reaction after the first few hours of melting. When
ESR slag is recycled, it is already depleted in SiO2, and this reaction no longer occurs. The slag always tries to move to chemical equilibrium with the molten metal at the operating temperature of the slag. Thus, repetitive recycling of slag with a given alloy composition produces a slag that is naturally buffered against causing elemental loss from the electrode during ESR. Unfortunately, slag loss (left as a skin on the electrode during the process) generally keeps recovery levels of the slag below 100%, so some fresh slag is always needed in the ESR process.
ACKNOWLEDGMENTS Reviewed by Scott A. Balliett, Latrobe Specialty Steel Company, with portions adapted from Ref 4 and from: A. Choudhury and F. Knell, Electroslag Remelting (ESR), Metals Handbook, 9th Edition, Volume 15, Casting, 1988, pp 401–405.
REFERENCES 1. R. Jauch, A. Choudhury, F. Tince, and H. Steil, Herstellung und Verarbeitung von ESU-Blo¨cken bis zu 160 t, Stahl Eisen, Vol 101, 1981, p 41–44 2. A. Choudhury, R. Jauch, and F. Tince, Low Frequency Electroslag Remelting of Heavy Ingots with Low Aluminium Content, Proceedings of the Sixth International Conference on Vacuum Metallurgy (San Diego), 1979, p 785–794 3. A. Choudhury, R. Jauch, H. Lo¨wenkamp, and F. Tince, Application of the Electroslag Remelting Process for the Production of Heavy Turbine Rotors from 12 %Cr-Steel, Stahl Eisen, Vol 97, 1977, p 857–868 4. M.J. Donachie and S.J. Donachie, Superalloys: A Technical Guide, ASM International, 2002 5. R.K. Hopkins, Method and Apparatus for Producing Cast Metal Bodies, U.S. Patent 2,380,238, 1945 6. B.E. Paton, B.I. Medovar, and W.E. Raton, Ein neues Verfahren—Das Elektroabgiessen von Blo¨cken, Bull. Tech. Informazil NTO, No. 1, Maschprom., Kiew, 1956
7. G. Hoyle, Electroslag Processes, Applied Science, 1983 8. W. Holzgruber and E. Plo¨ckinger, Metallurgische und verfahrenstechnische Grundlagen des Elektroschlacke-Umschmelzens von Stahl, Stahl Eisen, Vol 88, 1968, p 638–648 9. M. Wahlster, A. Choudhury, and K. Forch, Einfluss des Umschmelzens nach Sonderverfahren auf Gefu¨ge und einige Eigenschaften von Sta¨hlen, Stahl Eisen, Vol 88, 1968, p 1193–1202 10. M. Wahlster and A. Choudhury, Beitrag zum Elektroschlacke-Umschmelzen von Sta¨hlen, Rheinstahl Technik, Vol 5, 1967, p 31–37 11. M. Wahlster, G.H. Klingelho¨fer, and A. Choudhury, Neue metallurgusche und technologische Ergebnisse einer 10 t ESUAnlage, Radex-Rundsch., Vol 2, 1970, p 99–111 12. G.H. Klingelho¨fer, A. Choudhury, and E. Ko¨niger, Etude compare´e des caracte´ristique des aciers elabore´s a` l’air et des aciers refondus selon le proce´de´ sous laiter electroconducteur (ESR) pour les cylindres de laminage a` froid et les cylindres calandreurs, Rev. Metall., Vol 67, 1970, p 512–522 13. B.I. Medovar, Arc-Slag Remelting Steel and Alloys, Cambridge International Science Publishing, England, 1995 14. T. Li, Spec. Steel (China), Vol 25 (No. 5), Sept-Oct 2004, p 1–5 15. W.T. Carter, J.J. Jackson, R.M. ForbesJones, and R. Minisandram, Clean Metal Nucleated Casting of Superalloys, Sixth International Symposium on Superalloys 718, 625, 706, and Derivatives, Oct 2–5 2005 (Pittsburgh, PA) 16. G.E. Maurer, Primary and Secondary Melt Processing—Superalloys, Superalloys, Supercomposites and Superceramics, Academic Press, 1989, p 49–97 17. Ch. Kubisch, Druckstickstoffsta¨hle—eine neue Gruppe von Edelsta¨hlen Berg.—und Hu¨ttenma¨nnische Monatshefte, 1971, p 84–88 18. A. Choudhury, “Vacuum Electroslag Remelting—A New Process for Better Products,” Paper presented at the Vacuum Metallurgy Conference, June 1986 (Pittsburgh, PA)
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 132-138 DOI: 10.1361/asmhba0005202
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Vacuum Arc Remelting THERE ARE TWO COMMONLY USED remelting processes for metal refinement: electroslag remelting (ESR) and vacuum arc remelting (VAR). In both processes, an electrode is melted as it advances into the melting region of the furnace. As the working face of the electrode is heated to the melting point, drops of liquid metal that fall from the electrode face are collected in a lower crucible and rapidly solidified. However, ESR and VAR have very different melting methods, which have implications regarding the magnitude of cooling rates obtained and the nature of defects created by the remelting process itself. The VAR process also is carried out in a vacuum, while the ESR process typically occurs under atmospheric conditions—although electroslag remelting under vacuum (VACESR) or under pressure is a process variation of ESR (see the article “Electroslag Remelting” in this Volume for a brief description of VAC-ESR). The VAR process is widely used to improve the cleanliness and refine the structure of standard airmelted or vacuum induction melted (VIM) ingots (typically referred to as consumable electrodes
Fig. 1
Key components of a vacuum arc remelting furnace. Source: Ref 1
for remelting). The VAR process is also commonly used in the triplex production (VIMESR-VAR) of superalloys. The VAR process was the first commercial remelting process for superalloys. It was used in the late 1950s to manufacture materials for the aircraft industry. Current applications of VAR processing include the production of steels and superalloys as well as the melting of reactive metals such as titanium and zirconium alloys. The melt cleanliness and homogeneity from VAR provides benefits in both primary ingot and foundry shape casting for high-integrity applications where improved fatigue and fracture toughness of the final product are essential. This includes aerospace applications, high-strength steels, ball-bearing steels, die steels, tool steels (cold and hot work steels), power generation, defense, and medical and nuclear industries.
Process Description The basic process of VAR (Fig. 1) is the continuous melting of a consumable electrode by means of a direct current arc (electrode negative, melt pool positive) in a vacuum on the order of 0.1 to 1 Pa (7.5 104 to 0.0075 torr). In some cases, the melting is carried out under inert gas with a pressure up to 1000 Pa (7.5 torr). Evaporation losses of volatile alloying
Fig. 2
elements are minimized at this pressure. Remelting currents are up to 40 kA. Melting rates with VAR furnaces are as high as 1150 kg/h (2500 lb/h), and ingot diameters are as large as 1.5 m (5 ft). Figure 2 shows melting rates for various steels and alloys as a function of ingot diameter. These are empirical values that were obtained from experience in operation. These melting rates gave low microsegregation while achieving acceptable surface quality. The diameter of the electrode and its relationship to the crucible is critical and must be matched for the melting rate. The VAR crucible is immersed in a tank of cooling water. Often, a tube is used to further enclose the crucible and limit the thickness of the water stream moving past the crucible surface. This increases the water velocity and thus increases heat removal from the crucible surface. Such tubes are called water guides. The crucible base (stool) has provision for the inlet of a gas (usually helium). The top of the crucible mates with the VAR head. The head contains the vacuum ports and the ram, which drives the electrode into the crucible as it is consumed. The electrode position is controlled to maintain an arc gap between the melting electrode tip and the molten pool on the top of the solidified ingot in the water-cooled copper mold. When an ingot is solidifying in the crucible, it pulls away from the crucible wall as it
Typical vacuum arc remelting rates for various steels and nickel- and cobalt-base superalloys
Vacuum Arc Remelting / 133 cools. This reduces heat conduction to the crucible wall. To maintain the highest possible heat removal from the solidifying ingot, a high-heat-capacity gas such as helium is introduced into the gap that is formed by the ingot shrinkage. The ram in newer designs is capable of X-Y movement. In VAR, this is not movement in the horizontal plane, which would allow centering of the stinger/electrode assembly in the crucible. It is actually angulation of the ram from the vertical, so that if the stinger is not perfectly parallel with the electrode, then the ram may be angled to make the electrode parallel to the crucible sides. The VAR head mates to the crucible flange through an O-ring seal. Most VAR furnaces have two melting stations (crucible setups). While an electrode is melting in one station, the next station is prepared for melting. When one melt is finished, the VAR head is rotated in the horizontal plane to be located over the second crucible/electrode setup. The electrode is attached to the ram through a stinger, which is welded to the electrode. Although the basic design of a VAR furnace has remained largely unchanged over the years, significant advances have been made in the fields of control and regulation of the process with the object of achieving a fully automatic melting procedure. Modern VAR furnaces (Fig. 3) make use of computer-controlled process automation with continuous weighing of the consumable electrode and closed-loop process control of melting parameters such as remelting rate and arc gap based on arc voltage or drop-short pulse rate. This in turn has had a
decisive positive influence on the metallurgical properties of the products. Close control of all remelting parameters is required for reproducible production of homogeneous ingots with a controlled solidification structure free of macrosegregation. Process Capabilities. Because of the high arc temperature and the small pool of liquid metal, sound ingots with dense crystal structure, low hydrogen and oxygen contents, and minimal chemical and nonmetallic segregation are produced. The primary benefits of remelting a consumable electrode under vacuum are: Removal of dissolved gases, such as hydro-
gen, nitrogen, oxygen, and CO
Reduction of undesired trace elements with
high vapor pressure
Improvement of cleanliness by removal of
oxides
Achievement of directional solidification of the
ingot from bottom to top, thus avoiding macrosegregation and reducing microsegregation In VAR, center porosity and segregation in the ingot resulting from ingot casting are also eliminated because of progressive melting and solidification. Directional solidification of the ingot from bottom to top avoids macrosegregation and minimizes microsegregation. Oxide inclusion removal is optimized because of the relatively short reaction paths during melting of the hot electrode end and because of good drop dispersion in the plasma arc. Oxide removal is achieved by chemical and physical processes. Less stable oxides or
nitrides are thermally dissociated or are reduced by carbon present in the alloy and are removed via the gas phase. However, in special alloys and in high-alloyed steels, the nonmetallic inclusions (e.g., alumina and titanium-carbonitrides) are very stable. Some removal of these inclusions takes place by flotation during remelting. The remaining inclusions are broken up and evenly distributed in the cross section of the solidified ingot. Because the molten metal drops immediately by gravity into the molten pool, it is inherently impossible to superheat the metal in the VAR process. This, coupled with the very high heat-extractive capability of the process, makes VAR the choice for economic manufacture of the largest-diameter ingots of segregation-prone superalloys. Additionally, the exposure of small volumes of molten metal to high vacuum is capable of removing detrimental high-vaporpressure elements. Unfortunately, beneficial high-vapor-pressure elements such as magnesium are also reduced greatly in concentration. Melt Pool in the VAR Process. Figure 4 illustrates the features of a molten pool in a VAR ingot. The most important feature shown is the depth of the molten pool. It generally is assumed that the depth of the liquid zone is representative of the depth of the liquid-solid zone. For a given crucible diameter, the size (depth) and shape of the liquid-solid zone has a direct relationship with heat input (melt rate) and heat extraction (water cooling). The depth and angle of the liquid-solid zone also controls freckle formation for any given alloy. The melt rate is directly proportional to the applied melt current (Fig. 5). However, as seen in Fig. 6, the relationship between pool depth and melt current is direct but is divided into
Fig. 4
Fig. 3
Modern vacuum arc remelting (VAR) furnace. (a) 30 ton VAR. (b) Operational components: 1, electrode feed drive; 2, furnace chamber; 3, melting power supply; 4, busbars/cables; 5, electrode ram; 6, water jacket with crucible; 7, vacuum suction port; 8, X-Y adjustment; 9, load cell system. Courtesy of ALD Vacuum Technologies GmbH
Molten pool region in a vacuum arc remelted ingot. The relative sizes of annulus and arc gap are not to scale. Competing currents are illustrated for stirring by thermal buoyancy (rising of lower-density hot metal along the ingot centerline) versus the stirring by electromagnetic fields (Lorentz stirring) that drive liquid metal from the edge of the crucible to the center of the crucible and down the centerline. Source: Ref 1
134 / Melting and Remelting
Fig. 5
Melt rate versus vacuum arc remelting current (50 cm, or 20 in., ingot; a composite graph from several independent sources). Source: Ref 1
liquid metal from the edge of the crucible to the center of the crucible and down the centerline. Figure 6 thus has two regimes, where the inflection point (nominally 6600 A for a 50 cm, or 20 in., IN-718 ingot) marks the point when Lorentz stirring becomes stronger than thermal buoyancy stirring. The apparent depth of the molten pool is altered not because there is a change in the volume of molten metal, but because the centerline stirring at higher amperages causes the pool to depart from a nominal U-shape and thus become deeper in the center. Due to Lorentz stirring, high VAR amperages thus produce deeper pools and also alter the angle of the liquid-solid zone, with respect to the ingot axis. In addition to deepening the molten pool and increasing the angle of the liquid-solid zone, there is another ramification when Lorentz stirring is dominant. At low currents, oxides and nitrides in the electrode are melted out, drop onto the molten pool surface, and are swept to the sides of the electrode. When melting at high currents, these particles may not migrate to the edge of the ingot but may be incorporated to some extent in the general structure of the ingot, perhaps with a concentration along the centerline. It is thus considered inadvisable to melt at VAR currents in the Lorentz-dominated region. As the volume of metal on the stool (the ingot) builds up, an equilibrium state is reached in which there is a solid ingot, a mushy zone of both liquid and solid above that, and then a zone that is totally liquid. Because the heat extraction in the steady-state condition of melting is fastest through the sidewalls and slower down the ingot and through the stool, the mushy zones and liquid zones are shallower near the sidewalls and deeper in the ingot center. The thickness and growth angles of the mushy zone determine if the interdendritic liquid regions will coalesce to form positive segregation channel defects. The thickness of the mushy zone at the center of a VAR ingot is controlled by the efficiency of heat extraction, the size (diameter) of the crucible, and the melt rate of the electrode face. The melt rate is controlled by the magnitude of current passed through the electrode. Other factors that are controlled so as to positively affect the solidifying structure are the distance of the melting face above the molten pool (arc gap) and the clearance of the sides of the electrode from the crucible (annulus).
Solidification Fig. 6
Vacuum arc remelting pool depth versus melt current. Source: Ref 1
two separate regions. The reason for this is indicated by the two lines in Fig. 4, which indicate the direction of liquid metal flow in the VAR pool. Competing currents are generated
by the tendency of lower-density hot metal to rise along the ingot centerline (thermal buoyancy stirring) versus the tendency of electromagnetic fields (Lorentz stirring) to drive
The solidification structure of a VAR ingot is a function of the local solidification rate and the temperature gradient at the liquid-solid interface. To achieve a directed dendritic primary structure, a relatively high temperature gradient at the solidification front must be maintained during the entire remelting process. The growth
Vacuum Arc Remelting / 135 direction of the cellular dendrites conforms to the direction of the temperature gradient, that is, the direction of the heat flow at the moment of solidification at the solidification front. The direction of the heat flow (and thus the growth direction of the dendrites) is always perpendicular to the solidification front (Fig. 4). The edge of a VAR ingot also is the first metal to solidify. It is thus alloy-lean, as dictated by the phase rule. It is present as a skin around VAR ingots and is known as shelf. The shelf, in addition to being alloy-lean, contains both the oxides and nitrides collected from the melt pool surface as well as portions of the vapor deposits and splash BBs that have been deposited on the crucible wall above the advancing molten pool. If not adequately removed, this shelf material carries over from the ingot. As pool depth increases with the remelting rate, the growth angle of the dendrites, with respect to the ingot axis, also increases. In extreme cases, the growth of the directed dendrites can come to a stop. The ingot core then solidifies nondirectionally, for example, in equiaxed grains, leading to segregation and microshrinkage. Even in the case of directional solidification, microsegregation increases with dendrite arm spacing. A solidification structure with dendrites parallel to the ingot axis yields optimal results. However, a good ingot surface requires a certain level of energy input, resulting in respective remelting rates. Optimal melt rates and energy inputs depend on ingot diameter and material grade, which means that the necessary low remelting rates for large-diameter ingots cannot always be maintained to achieve axis-parallel crystallization. In spite of directional solidification, defects do occur. This can lead to rejection of the ingot, particularly in the case of superalloys. For the most part, these defects also occur in discrete regions, as opposed to continuous defects. Solute-lean (negative segregation) defects are not inherently as detrimental to the properties of a component as are defects resulting from positive segregation. Types of ingot defects include: Tree ring patterns Freckles White spots
Tree ring patterns can be identified in a macroetched transverse section as light-etching rings. They usually represent a negative crystal segregation. Tree ring patterns seem to have little effect on material properties. Freckles and white spots have a much greater effect on material properties as compared to tree ring patterns. Both defects can represent a significant cause for premature failure of turbine disks in aircraft engines. Tree ring patterns are the result of a wide fluctuation in the remelting rate. In modern remelting plants, however, the remelting rate is maintained at the desired value by means of an exact control of the melting rate during
operation, so that the melting rate exhibits no significant fluctuations unless caused by electrode defects. Freckles are dark-etching circular or nearly circular spots that are generally rich in carbides or carbide-forming elements. The formation of freckles is usually a result of a high pool depth or movement of the rotating liquid pool. The liquid pool can be set into motion (rotation) by stray magnetic fields during remelting. Freckles can be avoided by maintaining a low pool depth and by avoiding disturbing magnetic fields through the use of a coaxial current supply. White spots are typical defects in VAR ingots. They are recognizable as light-etched spots on a macroetched surface. There are several mechanisms that could account for the formation of white spots: Residues of unmelted dendrites of the con-
sumable electrode
Pieces of the ingot crown that fall into the
pool and are not dissolved or melted
Pieces of the shelf region transported into
the solidifying interface All three of the aforementioned mechanisms, individually or combined, can be considered as sources of white spots. This indicates that white spots cannot be completely avoided during VAR. To minimize the frequency of occurrence of these defects, the following conditions should be observed during remelting: Use the maximum acceptable metal rate
permitted by the ingot macrostructure.
Use a short arc gap to minimize crown for-
mation and to maximize arc stability.
Use a homogeneous electrode substantially
free of cavities and cracks.
Use a proper melting power supply.
Under ideal conditions, the arc is distributed uniformly across the surface of the molten pool. (This is called a diffuse arc.) However, the introduction of conductive ionic species into the melt or increases in the resistance (arc gap) in the system may cause the arc to become locally concentrated (constricted arc). Constricted arcs do not dwell in one place on the molten pool surface but move about the surface. When constricted arcs undercut the crown/ shelf, small pieces of the shelf may fall into the molten pool. These pieces move by gravity down the pool profile. They do not readily remelt into the pool, because they have a higher melting point than the pool in general. Ultimately, they are not dissolved and are incorporated into the ingot as regions of low solute content, containing stringers of oxide and nitride. When detected in the final product, they are seen as light-etching defects containing stringers of oxide and nitride and are known popularly as dirty white spots. More correctly, these regions are called discrete white spots, in keeping with their nature as a distinct but isolated structure within the ingot matrix.
Theoretically, a piece of shelf, free of oxides and nitrides, may be undercut and form a discrete white spot that is not dirty. Often, because of the lack of solute, the grain size in discrete white spots will be larger than that of the matrix. When present, dirt stringers act as stress raisers. Thus, discrete white spots, which have a high probability of being largegrained and dirty, are acknowledged to be detrimental to properties in superalloys. Unfortunately, although best practice for electrode preparation and for VAR melting may minimize the frequency of occurrence of discrete white spots, there is no way to guarantee their elimination in VAR products. The probability of their occurrence in the final component must be considered in the design of the component. Currently, there is no method of detecting a subsurface defect except to the extent that, combined with forging deformation, a crack may be generated that will be detectable by ultrasonic inspection. Macroetching of the surface of critical components is a common method for detection of discrete white spots and any other melt-related structure.
Process Variables The manufacture of homogeneous ingots with minimal segregation requires careful control of remelting parameters. The three major parameters defining the melt process are ingot and electrode diameter, arc gap, and melt rate. The choice of ingot and electrode diameter defines the clearance (annulus) between the crucible and the electrode. Insufficient clearance, either through a poor choice of annulus or through an off-center set-up, will lead to excessive current loss from the electrode to the crucible wall, rather than through the arc to the ingot. The choice of ingot diameter controls the amount of possible heat extraction and dictates the choice of melt rate so as to avoid positive segregation defects in the ingot. These choices, of course, are long-term choices dictated in the purchase of VIM molds and VAR crucibles. Control of the VAR process also includes: Pool depth and shape (control of solidifica-
tion structure)
Arc stability (control of drop-in defect
formation) As noted, the depth and angle of the solidification front (pool shape) needs to be maintained so that unacceptable solute-rich solidification structures do not occur. Pool shape and depth are controlled by regulating heat input and heat extraction. Of these, the melting current density has the largest influence on the melting bath geometry and conditions of solidification. Melt Rate. The melt rate is dependent on the melt current. Many producers choose to select a given melt amperage and hold it constant. Other producers use load cells built into the
136 / Melting and Remelting VAR furnace to measure electrode weight and thus the change in weight. The change in weight is used to calculate a melt rate, and the applied current is varied to attempt to maintain uniform melt rate. Individual melt rate measurements will vary widely depending on the time frame chosen for the measurement. Melt rates are generally calculated using rolling averages. A common averaging time is 20 min. Modern VAR furnaces are equipped with a load cell system to measure the weight of the electrode at a particular interval of time. Load cell design and maintenance are important factors in maintaining the quality of material melted in melt rate control, because load cells are subject to the generation of false signals. The actual values of the melt rate are compared by computer with the desired set values. Any difference between the measured melt rate and the desired value is eliminated by the proper accommodation of the power input. Figure 7 shows the melt rate and the melting current at startup, during steady-state melting, and during hot topping. Startup and hot topping are usually controlled based on time. The melting phase is controlled based on weight. Hot topping begins when a preselected residual electrode weight is reached. A computer controls the melting parameters, and melting rates can be controlled with a precision of better than þ2%. Heat input is determined by the melt rate of the electrode. Melt rate is often monitored and corrected over very short time periods by changes in melt current (melt rate control). Alternatively, control of the median melt rate over the duration of a melt may be maintained by operating at constant melt amperage (amperage control). Heat extraction is a function of the contact surface/ingot volume (ingot size) being melted. Larger ingots have poorer heat extraction (lower surface/volume) and must be melted at proportionately lower melt rates to prevent the formation of large liquid + solid zones and thus the formation of unacceptable positive segregation. While, for a given ingot size, high
Fig. 7
melt rate may promote positive segregation, a melt rate that is too low may cause the formation of visible solute-lean structures. The mechanism of formation of these solidification white spots is not completely understood. They differ from discrete white spots in that they are a solidification effect, not a shelf or electrode drop-in, and therefore do not contain oxide/ nitride stringers. Nevertheless, such solute-lean regions are potentially detrimental because they may make it difficult to control grain growth in subsequent processing. Arc Stability. Active control of arc stability is primarily by control of the arc gap throughout the melting of an electrode. The arc gap is the nominal distance of the melting electrode surface above the molten pool at the top of the solidifying ingot. Commercially useful gaps range from a minimum of approximately 2.5 mm (0.1 in.) to a maximum of 12 mm (0.5 in.). The arc gap may be considered as the resistance in this circuit. Because the melt current is held constant, any increase or decrease in the resistance of the circuit (change in arc gap) is seen as a change in voltage. Early VAR controls actually did measure these changes in voltage as a means for maintaining a uniform arc gap throughout the melt. Modern systems control the arc gap by measurement of drip short frequency (DSF). When a drip (drop) of molten metal is formed on the melting electrode surface, there is a point in time at which it is in contact with both the electrode and the molten pool. This causes a sudden decrease in voltage (an electrical short) across the arc gap. The duration of the short is measured in milliseconds. Control systems measure the number of drips of a given duration. It has been demonstrated that the frequency of drip shorts (usually in drips/min) is related to the width of the arc gap. The relationship is not linear over the whole range of melt conditions that may be experienced but does allow useful measurement and control of the arc gap in the ranges that are currently used
Process control parameters and set point functions in vacuum arc remelting
commercially. The higher the DSF, the smaller the arc gap. Making the gap smaller increases the duration of a short and thus increases the number of shorts counted. The relationship between arc gap and DSF also changes with changes in the melt current. At higher currents (higher melt rate), the drips are larger but less frequent. Thus, as in Fig. 8, raising the current requires a reduction in DSF to maintain equivalent arc gap. Drip short frequency is measured over a period of time, usually fractions of a minute. Individual DSF values vary widely. Thus, control signals for the ram drive (electrode feed speed) fluctuate. Normally, the ram is in continuous motion. The response to changes in DSF is to increase or to decrease the ram speed. Monitoring of the consistency of the arc gap may be accomplished by computing a rolling average or by simple observation of the width of the band generated in recording each individual DSF measurement. For melting of superalloys, arc gaps are generally controlled in the range of 6.5 to 13 mm (¼ to ½ in.). To maintain constant arc gap, the electrode is fed into the melt either faster or slower (ram drive) according to the fluctuation in voltage that occurs when the resistance of the system changes in response to changes in the arc gap. Modern control systems for superalloys generally do not use voltage for arc gap control but rather use a related voltage measurement: the DSF. A drip short is the transient downward voltage spike that occurs when a molten metal droplet is in contact with both the electrode and the molten pool. The number of shorts occurring in a given time period is proportional to the arc gap. The relationship for arc gap versus DSF is not melt rate independent. Passive control of arc stability is maintained by controlling the annulus and controlling the quality of the electrode being remelted. Electrodes for VIM generally contain porosity and low levels of residual refractory, both of which may act to destabilize the arc. Vacuum induction melting and pouring practice
Fig. 8
Drip short frequency as a function of arc gap and melt current. Source: Ref 1
Vacuum Arc Remelting / 137 is thus a significant factor in ensuring VAR arc stability and in minimizing the formation of drop-in-type defects. Vacuum Arc Remelting Control Anomalies. Variations in the traces of VAR parameters may indicate changes in melt conditions that will be detrimental to the quality of the solidification structure. Two of the more easily recognized problems are dead shorts and melt rate excursions (MREs). Dead shorts occur when the electrode is driven into the top of the molten pool, causing direct transfer of the melt current to the solidifying ingot. Such a problem may be caused by an unusually shaped primary shrinkage cavity in the electrode. If the normal ram drive programming cannot accommodate the unusual shape, the electrode may be driven into the pool. This will show on the recorded parameters as a sudden reduction of voltage to zero and an immediate ram back-out from the dead short situation. Drip short frequency, of course, would also go to zero while the electrode face was in contact with the pool. (This would not show as zero DSF, because DSF is measured over a time period.) Dead shorts will cause a major change in solidification characteristics in the molten pool in that region. Another cause of multiple dead shorts is the presence of high volumes of oxides and nitrides in the VAR electrode. If the oxide/nitride volume being melted onto the molten pool surface exceeds the capability of the process to sweep the “dirt” to the side and incorporate it into the ingot surface, then the ability to sustain an arc will be deteriorated. Melt rate will drop, and the ram travel, unable to react rapidly to sudden changes in arc gap, will drive the electrode into the pool, causing a short. This may happen repetitively as the ram backs out of the pool and then advances into it again. Transverse cracks in the electrode cause MREs, the most readily definable VAR chart anomaly. Electrodes may develop transverse internal cracks due to thermal stresses. This may happen in the cooling of the electrode, or, often, due to the thermal stresses generated as the VAR melt front moves up the electrode. When the melt front approaches a transverse crack, heat transfer across the crack is diminished, and the material below the crack becomes hotter than normal. Thus, the melt rate begins to increase. The ram speed, which responds to a decreasing DSF, does not respond fast enough to maintain arc gap, and thus, the arc gap increases.
Superalloy VAR The VAR process typically uses a cast (generally VIM or VIM-ESR) electrode. Elements such as bismuth and lead (which can be present in low concentrations even after VIM) are highly undesirable but are reduced to negligible levels by the VAR process. Magnesium, which is considered desirable for improved workability (and
is an addition element in VIM), is reduced in concentration but not completely removed. With the exception of high-vapor-pressure elements, the VIM chemistry is representative of the VAR chemistry. The flotation of oxides and nitrides to the melt surface in VAR improves the cleanliness and reduces the gas content of material processed through VAR. A common factor affecting consumable remelt quality for both ESR and VAR is the quality of the electrode. Electrode quality is affected by composition, cleanliness (low oxide and nitride content), and soundness (freedom from porosity and cracks). The most important characteristic of an electrode is composition. With the exception of reductions in volatile elements, VAR does not change the composition of the electrode. For high-nitrogen electrodes, some reduction of nitrogen is accomplished by flotation of nitrides. As the solubility levels of TiN in the alloy are approached, this mechanism is no longer effective. Electroslag remelting reduces sulfur content but may also cause minor changes in composition through reaction of titanium, aluminum, zirconium, and silicon with components of the slag. These changes are predictable in a mature process. Thus, there is no practical way to achieve composition modification once the master heat has been cast, but departure from normal practice in ESR may change titanium, aluminum, zirconium, or silicon content. Cleanliness. The cleanliness of the cast electrode with regard to entrained oxide and nitride is an important characteristic for VAR electrodes. The introduction of a high volume of oxide or nitride onto the molten pool in VAR may degrade the efficiency of melting of the arc as well as disturb the electrical characteristics measured to maintain control of the arc gap. This is not a problem in electrodes to be melted by ESR. Porosity. The porosity (secondary shrinkage) of an electrode may be of some concern in VAR. Vacuum arc remelting of electrodes with centerline porosity results in the preferred melting of the face in the porosity area. Thus, the melt face is no longer flat. The effect of this with regard to gap control during melting has not been quantified. A secondary problem connected with centerline porosity is the concern that projecting dendrites in this region may become detached from the electrode and fall, unmelted, into the molten pool. Then, they may be incorporated into the structure without being remelted. The composition of such a region thus will be that of a primary dendrite (alloy-lean) rather than representative of the bulk composition. Theoretically, these concerns may also apply to ESR processes, but the necessity for melted or unmelted material to pass through the slag makes ESR products inherently less sensitive to electrode porosity. Cracking. The final electrode characteristic affecting remelt quality is the freedom of the electrode from transverse cracking. In highly alloyed superalloys, the thermal stresses
generated upon cooling of electrodes or upon heating in the remelt process may be sufficient to form transverse cracks. When the melt front of either VAR or ESR approaches such a crack, the heat transfer up the electrode is diminished, because of the thermal barrier presented by the crack. Thus, the material in front of the crack becomes hotter than would occur under equilibrium conditions and will tend to melt at a faster rate. After the melt front has gone through the crack, it encounters cold material, and the melt rate will tend to drop. In both VAR and ESR, the result is that the controls respond with changes in applied power to try to keep the melt rate constant. This, however, is seldom accomplished without some discernable disruption to the continuous nature of the growth of the solidification front. The only defense against these MREs (also called events) is to produce an electrode free of cracks. In the future, it is hoped that more sophisticated melting controls will be able to detect the melt rate changes earlier and thus compensate for them more efficiently.
Titanium VAR Since the commercial introduction of titanium alloys in the 1950s, VAR has been the principal method for the production of titanium ingots. The process is capable of high melt rates with titanium, but steady-state solidification is not possible, and the molten metal pool is relatively deep. Vacuum arc remelting has many advantages, including high purity, good control, and reproducibility in the product. However, if high-density components of an electrode fall into the melt as solids, they can survive in the melt by rapidly dropping to the bottom of the molten pool and solidifying before they can be melted and homogenized in the solidifying ingot. An alternate to VAR is cold hearth melting. The electrodes for making titanium ingots are compacted aggregates (compact or briquettes) of sponge and alloy elements, including both master melt and elemental materials. For commercially pure titanium, the sponge is compacted, and the electrode is melted with appropriate oxygen control. When alloys are to be produced, granules of titanium sponge are mixed and then blended with appropriate alloy-containing material. The resultant alloy mixture is pressed into a compact or briquette. Most titanium and titanium alloy ingot is melted twice using the VAR process. The procedure is known as the double consumable-electrode VAR process. Double melting is considered necessary for all applications to ensure an acceptable degree of homogeneity in the resulting product. For certain critical applications, a third, or triple, melting step was specified at times, especially after the middle of the 1960s when defects were found in rotor-grade titanium alloys produced by double melting. Triple melting achieves even better
138 / Melting and Remelting uniformity of chemistry and structure. Triple melting also reduces oxygen-rich or nitrogen-rich inclusions in the microstructure to a very low level by providing an additional melting operation to dissolve them. In the two-stage process, titanium sponge, revert, and alloy additions are initially mechanically consolidated and then are melted together to form an ingot. Ingots from the first melt are used as the consumable electrodes for secondstage melting. Ingots from the second melt are used as the consumable electrodes for thirdstage melting when it is done. Processes other than consumable-electrode arc melting may be permitted in some instances for first-stage melting of ingot for noncritical applications.
Usually, all melting is done under vacuum, but in any event, the final stage of melting for noncritical applications must be done by the consumable-electrode vacuum-arc process.
“Vacuum Arc Remelting,” ALD Vacuum
Technologies, GmbH
REFERENCE ACKNOWLEDGMENT
1. M. Donachie and S. Donachie, Superalloys: A Technical Guide, ASM International, 2002
Portions of this article were adapted from: A. Choudhury and E. Weinga¨rtner, Vacuum
SELECTED REFERENCES
Arc Remelting (VAR), Casting, Vol 15, Metals Handbook, 9th ed., ASM International, 1988, p 406 M. Donachie and S. Donachie, Superalloys: A Technical Guide, ASM International, 2002
M. Blum, Vacuum Metallurgy, Vulkan, 2005 M. Donachie and S. Donachie, Superalloys:
A Technical Guide, ASM International, 2002
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 139-141 DOI: 10.1361/asmhba0005203
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Skull Melting Reviewed and revised by G. Keough, Consarc
SKULL MELTING refers to the use of furnaces with water-cooled crucibles that freeze a solid “skull” of material on the crucible wall. This technique is useful when melting highly reactive metals (e.g., titanium and zirconium) that would otherwise attack the linings of a crucible with refractories. The solidified metal creates a corrosion-resistant protective layer on the containment vessel surface. Two methods of skull melt casting have evolved. The first, vacuum arc skull melting, was originally developed in the early 1950s by the U.S. Bureau of Mines in Albany, Oregon (Ref 1, 2). Both methods are used to melt significant quantities of contamination free titanium for ingots or near net shape castings. Because the processes are used to melt reactive metals, they are normally carried out in a vacuum, but, in some cases, may also be done in an inert atmosphere. Vacuum arc skull melting was once the dominant method of melting titanium. Currently, other methods for melting titanium have become prevalent. Vacuum arc refining (VAR) furnaces with static crucibles are used in titanium primary melting. Recently, much of the melting of titanium sponge has been converted to plasma cold hearth (PCH) melting. Also, some of the largest rectangular ingots produced for rolling into sheet are now made on large multimegawatt electron beam (EB) furnaces. The total tonnage of material melted by VAR, PCH, and EB is significantly higher than that melted in vacuum arc skull caster furnaces. Vacuum arc skull melting works reasonably well for medium-to-large pour weights, typically in the range of 110 kg (250 lb) and up, but it suffers from a relatively large skull (typically 20% of metal melted). This lowers the thermal efficiency of the system and produces little superheat, so that precision and fine detail are relatively difficult to reproduce. The consumable reactive metal electrode also is expensive to produce, because the alloy shape and size must be preformed. An alternative technique is induction skull melting. Induction skull melting furnaces can melt a wide range of material forms including sponge (compacted), scrap, revert, or master alloy billet. The induction field also promotes vigorous stirring, which ensures uniform alloy composition throughout the melt
and increases superheat (compared to vacuum arc skull casting). Induction skull melting is also used in atomization of melts for powder production of reactive metals such as titanium.
Vacuum Arc Skull Melting Although the possibility of titanium skull melting was considered as early as 1948 and 1949, the first castings were made in 1953. In the late 1950s, this technology was applied by research institutes, which were looking for a practical means of liquefying and pouring uranium into graphite molds, for example, to produce uranium carbide. An early industrial vacuum arc skull melter was built in 1963 for the continuous production of uranium carbide. This furnace had a crucible volume of approximately 0.01 m3 (0.35 ft3) and used a nonconsumable graphite electrode to liquefy the uranium pellets fed into the crucible. The crucible tilting system was hydraulically driven. The molds were stationary. It was not until 1973 that one of the first skull melters for titanium went into operation; this furnace started production in 1974 in West Germany. State-of-the-art titanium vacuum arc skull melting furnaces are often equipped with turntable systems for centrifugal casting (up to 350 rpm). Casting weights of more than 1000 kg (2200 lb) are possible. Vacuum arc skull melting and casting is used for many titanium investment castings for aircraft, aerospace, medical, and chemical applications. Electron beam skull melting is also used for titanium alloys (see the article “Electron Beam Melting” in this Volume).
Furnaces Vacuum arc skull casting furnaces basically consist of a vacuumtight chamber in which an electrode is driven down into a water-cooled copper crucible. The direct current power supply provides the fusing current needed to strike an electric arc between the consumable electrode and the crucible. Because the crucible is water cooled, a solidified skull forms at the
crucible surface, thus avoiding direct contact between melt and crucible. The mechanism of skull formation is similar to the slag splashing on a refractory lining in the steel industry or the formation of a “frozen ledge” on refractories in the aluminum industry. Once the predetermined amount of liquid titanium is contained in the crucible, the electrode is retracted, and the crucible is tilted to pour the melt into the investment casting mold positioned below. For optimal mold filling, the mold can be preheated and/or rotated on a centrifugal turntable. Figure 1 shows the basic components of a vacuum arc skull melting furnace. Operating pressure is in the range of 1 Pa (0.0075 torr), and the specific working current ranges from approximately 1 kA/kg for small furnaces to approximately 0.2 kA/kg for large pouring weights. This batch-type skull melting furnace allows for cycle times of approximately 1 h for a full 50 kg (110 lb) pumping/melting/casting cycle. In principle, three consecutive pours can be obtained from one electrode. This furnace basically consists of a vacuum chamber, an arc voltage-controlled electrode drive system, a skull crucible, a centrifugal casting system with stepless adjustable turntable speed, an automatically sequenced vacuum pump system, a power supply, and an electrical control system with control desk. The cylindrical vacuum chamber is equipped with two large dished doors that support the crucible with the tilting mechanism, the mold platform with the centrifugal casting system, and the casting tundish with its cover. The crucible support system with an additional detachable device is also used for electrode loading. The chamber is jacketed for water cooling in regions that are subject to heat radiation. A top flange with a throat carries the electrode chamber and the electrode feeding system. Viewing ports allow for video monitoring of the melting and pouring. A vacuum pumping port is located in the cylindrical portion of the chamber. Figure 2 shows a semicontinuously operated vacuum arc skull melter for charge weights of up to 1000 kg (2200 lb). The principal difference between this furnace and the smaller model (Fig. 1)—apart from capacity-related
140 / Melting and Remelting
1
2 3 4
5
layout features—is the rectangular vacuum chamber. Again, the crucible and the tilting mechanism are carried by a dished door. In this furnace design, the centrifugal casting system is introduced from the bottom of the chamber to allow the horizontal connection of a separate cooling chamber, if desired. Molds are loaded from the back side into the chamber and can be discharged through a front door, which also allows the use of a cooling and charging chamber with lock valves for continuous mold transport flow. Modern vacuum arc skull melting furnaces are usually equipped with: Coaxial power feed directly to the skull
6 7
8 9 10
Fig. 1
Schematic of a vacuum arc skull melting and casting furnace. 1, fast retraction system; 2, power cables; 3, electrode feeder ram; 4, power supplies; 5, consumable electrode; 6, skull crucible (50 kg, or 110 lb); 7, tundish shield; 8, mold arrangement; 9, centrifugal casting system; 10, chamber lid carriage
crucible to avoid electromagnetic fields that can disturb the melt bath Programmable control systems for crucible tilting to allow repeatable pouring profiles for consistent parameters Highly accurate electrode weighing system for precise determination of pouring weights X-Y adjustment system for coaxial positioning of the electrode in the skull crucible Compact air-cooled power-supply modules of high capacity for achieving high melt rates, thinner skulls, and correspondingly increased yields above 80% Proven vacuum pumping and measuring systems Forced argon cooling systems for faster mold cooling
Induction Skull Melting 1 2
3 4
5
6
7
8
9 10
11
Fig. 2
Schematic of a semicontinuously operating vacuum arc skull melter for charge weights of up to 1000 kg (2200 lb). 1, fast retraction system; 2, power cables; 3, power supplies; 4, electrode feeder ram; 5, consumable electrode; 6, skull crucible; 7, crucible carriage; 8, tundish shield; 9, mold arrangement; 10, vacuum pumping system; 11, centrifugal casting system
In the 1970s, the Bureau of Mines developed an induction melting process for reactive metals known as the induction slag melting process. However, the drawback of that process was that a slag based on calcium fluoride was used to inhibit electrical shorting between segments of the crucible. Later, the Duriron Company further developed the process and, in particular, the crucible design to allow the slag-free melting of highly reactive metals and alloys in what is now referred to as the induction skull melting (ISM) furnace. Induction skull melting has evolved into a process with unique applications. As noted, ISM furnaces have some important advantages compared to vacuum arc skull furnaces. Induction skull furnaces can be charged with various material forms (scrap, revert, billet), and custom alloys can be alloyed in the ISM crucible because the induction field promotes vigorous stirring of the melt. A combination of stirring and levitation ensures uniform alloy composition throughout the melt and more uniform heating and increased superheat (compared to vacuum arc). As an indication of the flexibility of the ISM process, it is claimed that over 3000 alloys have been melted in this process. Part of this flexibility comes from the fact that the process is able to operate at pressure levels from vacuum through 101 kPa (1 atm). The process
Skull Melting / 141
Fig. 3
A 10 kg (22 lb) induction skull melting/casting unit in a vacuum/inert gas chamber
ac coil ISM crucible
Top dc coil
Water in Water out Bottom dc coil Pole piece (a)
ac coil
ISM crucible
Top dc coil
Bottom dc coil Water in Water out Pole piece (b)
Fig. 4
Cross section of induction skull melting (ISM) crucible with two different configurations of alternating current (ac) and direct current (dc) coils used in trials. Source: Ref 3
is also very rapid. A 10 kg (22 lb) charge, for example, can be loaded, melted, cast, and prepared for the next charge on a 15 min cycle. The key design elements of a reliable ISM crucible include effective water cooling and a segmented crucible that allows the field from the induction coil to couple to the load. Modern ISM crucibles use narrow gap technology to minimize the risk of liquid metal penetrating in between the segments and causing a short circuit. Each of the segments is intensively water cooled. A well-designed crucible will typically last 15,000 to 18,000 heats before extensive repair or replacement is required. Figure 3 shows a 10 kg (22 lb) induction skull furnace mounted in a vacuum/inert gas chamber. The chamber and vacuum or inert gas backfill is essential to avoid excessive oxidation of the reactive alloy elements, particularly titanium. Directly below the furnace is a centrifuge table that the molds are mounted on so that the molds can be centrifuged during filling. Skull weights in the range of 5 to 10% of the charge weight can be achieved in a welldesigned ISM furnace. Subsequent charges of the same alloy can be charged into the skull of a prior melt, thus minimizing the material requirements. At the end of a run for a particular alloy, the skull can be removed, identified, and saved for use on a future run. To optimize the efficiency of the process, a relatively high power together with a corresponding frequency are used to achieve both rapid melting and effective stirring of the liquid metal. The individual segments are intensively cooled to protect the copper from erosion and to minimize the risk of contamination of the melt. The combination of high power and high frequency also achieve higher superheat in the melt. This allows the process to be used for smaller pour weights (down to approximately 1 kg, or 2 lb) and for high-precision castings. Typically, the ISM process is recommended for pour weights in the range of 1 to 100 kg
(2 to 220 lb), although further extension of the range for higher pour weights is being developed. Of particular interest is a relatively new development using a combination of direct current electromagnetic fields (Fig. 4) with the traditional alternating current electromagnetic fields to achieve a significant further increase in available superheat of the molten charge (Ref 3). The ISM process also offers other advantages in casting. It can, of course, be used for conventional tilt-pour casting into both static molds and molds mounted on a centrifuge. However, unlike other methods, it can also be used for rollover casting as well as countergravity casting. Rollover casting, in particular, has been shown to be one of the least turbulent methods of skull casting, with promising results for casting complex and thin-walled near-net shape parts. ACKNOWLEDGMENTS Updated from: F. Mu¨ller and E. Weinga¨rtner, Vacuum Arc Skull Melting and Casting, Volume 15, Casting, Metals Handbook 9th Edition, 1988. REFERENCES 1. R.A. Beahl, F.W. Wood, J.O. Borg, and H.L. Gilbert, “Production of Titanium Castings,” Report 5265, U.S. Bureau of Mines, Aug 1956, p 42 2. W.J. Kroll, C.T. Anderson, and H.L. Gilbert, A New Graphite Resistor Vacuum Furnace and Its Application in Melting Zirconium, Trans. AIME, Vol 175, 1948, p 766–773 3. R.A. Harding, M.Wickins, G. Keough, R.J. Roberts, K. Pericleous, and V. Bojarevics, “The Use of Combined DC and AC Fields to Increase Superheat in an Induction Skull Melting Furnace,” LMPC Conference 2005 (Santa Fe, NM), ASM International
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 142-148 DOI: 10.1361/asmhba0005204
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Electron Beam Melting Revised by David W. Tripp, TIMET, USA
ELECTRON BEAM MELTING AND CASTING TECHNOLOGY is accepted worldwide for the production of niobium and tantalum ingots weighing up to 2500 kg (5500 lb) in furnaces with electron beams of 200 to 1500 kW. Other applications in Germany and former Soviet bloc countries are the production of steel ingots weighing 3.3 to 18 Mg (3.6 to 20 tons) using electron beams of up to 1200 kW. Furnaces of up to 4200 kW in electron beam power have been used since 1982 for recycling titanium scrap to produce 16 Mg (17.5 ton) slabs 1420 mm (56 in.) wide. Furnaces of 200 to 1200 kW are used to refine nickel-base superalloys. Other metals, such as vanadium and hafnium, are melted and refined in furnaces between 60 and 260 kW. Many furnaces with melting powers ranging from 20 to 300 kW are in operation in research facilities. These furnaces are used in the development of new grades and purities of conventional and exotic metals and alloys, for example, uranium, copper, precious metals, rare earth alloys, intermetallic materials, and ceramics. The total power
of industrial installed electron beam melting and casting furnaces worldwide was approximately 35,000 kW at the end of 2007. Electron beam melting and casting includes melting, refining, and conversion processes for metals and alloys. In electron beam melting, the feedstock is melted by impinging highenergy electrons. Electron beam refining takes place in vacuum in the pool of a water-cooled copper crucible, ladle, trough, or hearth. In electron beam refining, the material solidifies in a water-cooled continuous casting copper crucible or in an investment ceramic or graphite mold. This technology can be used for all materials that do not sublimate in vacuum. Competing processes include sintering (for example, for refractory metals), vacuum arc melting and remelting (for reactive metals and superalloys), and electroslag melting and vacuum induction melting (for superalloys, specialty steels, and nonferrous metals). Some advantages and limitations of the competing vacuum processes are given in Table 1. Additional information on some of these processes is available in
the articles “Vacuum Arc Remelting,” “Electroslag Remelting,” and “Vacuum Induction Melting” in this Volume.
Process Description The characteristics of electron beam melting and casting technology are: The flexibility and controllability of the pro-
cess temperature, speed, and reaction
The use of a wide variety of feedstock materi-
als in terms of material quality, size, and shape
The different methods of material processing
available
Product quality, size, and quantity
Contamination of the product is avoided by melting in a controlled vacuum and in watercooled copper crucibles (Fig. 1). The energy efficiency of electron beam processing exceeds that of competing processes because of the control of the beam spot dwell
Table 1 Comparison of characteristics of electron beam melting and competing processes Sintering Metal
Advantages
Vacuum arc melting Limitations
Tungsten, molybdenum
Small grain size; most often used
Refining limited; small batches; high energy consumption
Tantalum, niobium
Small grain size; good workability
Same as above; rarely applied
Hafnium, vanadium
...
Same as above
Zirconium, titanium
...
Not used
Advantages
Limitations
Refining limited; costly Moderate grain size; electrode preparation; acceptable melting dangerous workability; large ingots; low energy consumption Alloying; moderate grain Refining limited; expensive size; large ingots; low electrode; melting energy consumption dangerous Alloying during Almost no refining; costly remelting electrode preparation; melting dangerous
Electron beam melting Advantages
Limitations
Highest possible purity; economical feedstock preparation; large ingots; low energy consumption
Large grain size; brittle product; very rarely applied
Same as above; most frequently used
Alloying limited
Good refining; economical feedstock preparation and ingot production; most often used Economical feedstock Very low contamination; Limited refining; expensive feedstock preparation; only preparation; refining of wide range of alloying round ingots high-density inclusions; possible; large ingots melting of slabs, ingots, low energy and rods; high production consumption; rate; low energy economical melting consumption
High melting costs
Alloying limited; material losses from splatter; high furnace investment
Electron Beam Melting / 143 Large power range of 0 to 1200 kW Long free beam path of 250 to 1500 mm (10
to 60 in.) and the adjustable beam power distribution Beam deflection angle of þ45 and spot frequency up to 500 Hz Usable vacuum pressure range between 1 and 0.0001 Pa (102 and 106 mbar, or 7.5 103 and 7.5 107 torr) Reliability of the gun and cathode system The power control and distribution system allows for very accurate distribution of beam power and energy for achieving the required heating for material melting, superheating, refining, and electrothermal effects.
Drip Melting
Fig. 1
Schematic of the electron beam melting process
Fig. 2
Examples of electron beam melting and casting processes. (a) Button melting with controlled solidification for quantitative determination of low-density inclusions. (b) Consolidation of raw material, chips, and solid scrap to consumable electrodes for vacuum arc or electron beam remelting. (c) Drip melting of horizontally or vertically fed feedstocks. (d) Cold hearth refining/melting. (e) Floating zone melting. (f) Investment casting. (g) Pelletizing (manufacture of pellets from scrap and other materials for scrap recycling). (h) Atomization and granulation of refractory and reactive metals
time and distribution at the areas to be melted or maintained as liquid. In addition, unnecessary heating of the ingot pool, as occurs in vacuum arc remelting, for example, is avoided. Power losses of the electron beam inside the gun and between the gun nozzle and the target are very small, but losses up to 40% of the beam power can be incurred because of beam reflection, radiation of the liquid metal, and heat conductivity of the water-cooled trough and crucible walls. From the large variety of electron beam melting and casting processes shown in Fig. 2, only the processes illustrated in (c) and (d) are related to processes used in foundry technology:
Drip melting process for the preparation of
refractory and reactive metal feedstock material for electron beam and vacuum arc remelting skull melting and casting Cold hearth melting process for the feedstock refining of superalloys for vaccum induction melting and electron beam casting Electron Beam Heat Source Specifications. For all electron beam melting and casting processes, except for the crucible-free floating zone melting process, Pierce-type electron beam guns with separately evacuated beam generating and prefocusing chambers are used. The essential features of these guns are:
The drip melting processes (Fig. 3c and d) are primarily used for the production of clean, mostly ductile ingots of refractory and reactive metals. The feedstock for the first melt (Fig. 3c) can be compacted sponge, granular, powder, or scrap, which may be presintered in a vacuum-heating furnace. In some cases, loose raw materials can be consolidated in a watercooled copper trough (Fig. 3a). The consolidated ingot can then be fed horizontally for drip melting. Raw material that is continuously consolidated in a water-cooled copper crucible with a retractable bottom plate (Fig. 3b) can be fed horizontally or vertically for drip melting. In both consolidation processes, only 20 to 80% of the material is melted. Refining and losses of material by splattering are negligible. Drip melting of horizontally fed compacts is the most frequently used process for the production of ingots from refractory or reactive metals. The resulting ingot is of sufficient purity but has an area of inhomogeneity caused by the shadow of the horizontally fed bar. Two or more electron guns are used in drip melting to make use of reflected electron beams and to reduce evaporation and splattering. The end of a compact is welded to the front of the following one to avoid dropping semisolid material into the pool. Table 2 lists processing parameters that have been successfully used to electron beam melt various reactive and refractory metals. Vertical Feeding of Ingots. Refractory and reactive metal ingots of high purity, homogeneity, and smooth surface are remelted by vertical feeding (Fig. 3d). The molten metal droplets run down the conical, rotating electrode tip, are refined, and then drop into the pool center. The crucible pool is normally of the same diameter as the electrode, but it is not necessary. The crucible pool is kept in the liquid state to allow final refining and to guarantee ingot homogeneity. Because two or more electron guns are used, the entire pool can be equally bombarded; thus, shadow effects of the electrode can be eliminated. Simultaneous melting of horizontally and vertically fed electrodes (Fig. 3e) can be used
144 / Melting and Remelting
Fig. 3
Schematics of electron beam consolidation and drip melting processes. (a) Consolidation of coarse and solid scrap. (b) Continuous consolidation of raw material, chips, and solid scrap by direct feeding into a continuous casting crucible. (c) Drip melting of horizontally fed compacts, sintered bars, or consolidates for initial melting of reactive and refractory metals. (d) Drip melting of vertically fed vacuum induction melted or conventionally melted electrodes. (e) Drip melting of horizontally and vertically fed materials for the production of alloys from feedstocks with very different melting points
Table 2 Melting and refining data of refractory and reactive metals and alloy steels gained in laboratory and pilot production furnaces
Metal
Feedstock size, mm (in.)
Tungsten
40 (1.6) diam
Tantalum
55 (2.2) diam, 100 (4) long 60 (2.4) diam
60 (2.4) square, 160 (6.3) long Molybdenum 100 (4) diam 100 (4) diam 120 (4.7) square, 180 (7.1) long Niobium 80 (3.2) square, 150 (6) long 120 (4.7) square, 180 (7.1) long Hafnium 60 (2.4) square 80 (3.2) diam 100 (4) diam 100 (4) diam Zirconium 60 (2.4) square, 100 (4) long 60 (2.4) square, 150 (6) long Vanadium 60 (2.4) diam, 80 (3.2) long Titanium 100 (4) diam Ti-6Al-4V ... Ti-8Al100 (4) diam 1Mo-1V 4340 steel 80 (3.2) diam
Ingot diameter, mm (in.)
Ingot weight, kg (lb)
Integral melting rate, kg/h (lb/h)
Electron beam power of second melt, kW
... 60 (2.4) ... 115 (4.5) ... 80 (3.2) ... 160 (6.3) ... 100 (4) ... 180 (7.1) ... 150 (6) ... 180 (7.1) ... 80 (3.2) ... 130 (5.1) ... 100 (4) ... 150 (6) ... 80 (3.2) ... 100 (4) ... 100 (4) ... 150 (6) 150 (6)
... 37 (80) ... 200 (440) ... 65 (145) ... 523 (1150) ... 64 (140) ... 408 (900) ... 227 (500) ... 326 (720) ... 40 (90) ... 173 (380) ... 19.5 (43) ... 179 (395) ... 7.7 (17) ... 28.2 (62) ... 28.2 (62) ... 31 (70) 31 (70)
... 10.5 (23) ... 20 (44) ... 16.7 (37) ... 38.4 (85) ... 12.5 (27.5) ... 50.2 (111) ... 17.6 (39) ... 13.2 (29) ... 1.7 (3.7) ... 7.5 (16.5) ... 14.3 (31.5) ... 73 (161) ... 2.8 (6.2) ... 45.2 (100) ... 22.5 (50) ... 10 (22) 80 (176)
... 119 ... 300 ... 130 ... 371 ... 130 ... 290 ... 240 ... 218 ... 80 ... 110 ... 80 ... 250 ... 80 ... 87 ... 60 ... 52 80
Operating vacuum pressure of last melt, Pa (torr)
... 8 103 (6 105) ... 2 102 (1.5 10 4) ... 8 103 (6 10 5) ... 3 103 (2 10 5) ... 8 104 (6 10 6) ... 103 (7.5 10 6) ... 102 (7.5 10 5) ... 5 103 (3.8 10 5) ... 5 103 (3.8 10 5) ... 4 103 (3 10 5) ... 8 103 (6 10 5) ... 2 102 (1.5 10 4) ... 8 103 (6 10 5) ... 0.8 (6 10 3) ... 0.4 (3 10 3) ... 4 103 (3 10 5) 2.0 (0.015)
Interstitial elements in feedstock and final ingot, ppm
Total specific melting energy, kW 7 h/kg
Material yield, %
C
O
N
... 10.3 ... 9.9 ... 6.0 ... 8.30 ... 10.4 ... 5.2 ... 12.3 ... 15.9 ... 38 ... 14.7 ... 4.6 ... 2.6 ... 3.1 ... 1.91 ... 2.66 ... 1.67 1.0
... 93.1 ... 90 ... 92 ... 91.8 ... 95 ... 96.8 ... 87 ... 96.8 ... 93.1 ... 93.5 ... 88.2 ... ... ... 91 ... 99 ... 98 ... 93 99
70 10 200 10 100 10 ... 6 170 12 200 10 160 12 80 6 ... ... 500 100 ... ... ... ... ... ... ... ... ... ... 3900 4300 3622
4100 8 5000 8 1200 75 650 15 810 10 750 12 5220 106 4500 111 1870 170 900 200 950 545 950 735 1045(a) 235(a) 2180 1850 890 730 63 2.4 10
30 5 50 7 140 30 13 13 50 10 60 11 554 60 330 52 95 25 100 50 95 30 95 48 210(a) 95(a) 100 80 70 50 100 26 78
H
10 1 20 1 10 1 10 2 10 2 10 2 35 5 40 8 2 1 2 1 30 3 30 9 16(a) 13(a) 30 16 ... ... 0.3 0.08 0.10
Hardness, HB
... 200 ... 210 ... 65 ... 69 ... 140 ... 140 ... 77 ... 66 ... 160 ... 170 ... 120 ... 125 ... 30–100 ... ... ... ... ... ... ...
(a) The reproducibility of the refining data could not be confirmed.
for the production of critical alloys. In this case, the feedstock should be of the desired purity. Horizontally Fed Ingots. In this process, the feedstock size is smaller than the pool diameter to minimize the shadow effect of the horizontally fed bar. In production units, feeding can be carried out from two opposite sides. Other Process Considerations. To ensure the production of clean, homogeneous metals and alloys in electron beam drip melting furnaces, various aspects of material processing and handling must be controlled. Key considerations include:
Dimensions and quality of the feedstock,
and the feeding system used Ingot cooling and unloading during melting of another ingot Passivation and removal of condensates from the melt chamber Planning of melt sequences to minimize the number of furnace cleanings required Routine preventive furnace maintenance to ensure reliability Material yield and energy consumption
Equipment for Drip Melting. The essential equipment groups required for drip melting—
melting furnaces, control systems, and power supply units—are all important for achieving optimum productivity. The melting furnace (Fig. 4) includes the electron beam gun as the heat source, material feeding and ingot withdrawal systems, a crucible for material solidification, and a vacuum system to maintain the low pressure. Process observation, both visually and with video systems, is possible through viewports. The melt chamber flanges are equipped with x-ray-absorbing steel boards, and interlocking systems prevent operation failures and accidents.
Electron Beam Melting / 145
Fig. 4
Single-gun furnace for horizontal drip melting
Table 3 Principal applications for vacuum-arc-remelted (VAR), electron-beam-melted (EB), and powder metallurgy (PM) reactive and refractory metal ingots Metal
Reactive metals, VAR and EB melting Hafnium Vanadium
Zirconium Titanium
Refractory metals, EB melting and PM Tungsten
Tantalum
Molybdenum
Niobium
Applications
Glow discharge tubes for the electronics industry; control rods and breakoff elements in submarine nuclear reactors Targets for high-deposition-rate sputtering processes in the electronics industry; breakoff elements, fixtures, and fasteners in nuclear reactors; standards for basic research; alloying element for certain high-purity alloys Getter material in tubes in the electronics industry; fuel claddings, fasteners, and fixtures for nuclear reactors Components for bleaching equipment and desalination plants in the chemical industry; superconductive wires; turbine engine disks, blades and housings, rain erosion boards, landing legs, wing frames, missile cladding, and fuel containers in the aircraft and aerospace industries; shape memory alloys; biomedical fixtures and implants; corrosion-resistant claddings Heating elements, punches and dies, and nonconsummable electrodes for arc melting and gas tungsten arc welding for metal processing equipment; targets for x-ray equipment and high-sputtering-rate devices such as very large-scale integrated circuits, cathodes and anodes for electronic vacuum tubes in the electronics industry; radiation shields in the nuclear industry; cladding and fasteners for missile and reentry vehicles Condensers, autoclaves, heat exchangers, armatures, and fittings for the chemical industry; electrolytic capacitors for the electronics industry; surgical implants; fasteners for aerospace applications Dies for conventional and isothermal forging equipment; electrodes for glass melting; targets for x-ray equipment; cladding and fasteners for missile and reentry vehicles Superconductive wire for energy transmission and large magnets for the electrical and electronics industries; heavy ion accelerators and radio frequency cavities for nuclear applications; components for aircraft and aerospace applications
The control system allows the adjustment and control of such operating process parameters as electron beam power, operating vacuum level, material feed rate, and ingot withdraw speed. The control system also records and logs the process data.
Power Supply Units. One or more high-voltage power supply units are needed to supply the electron beam guns with the required continuous voltage (30 to 50 kV). The beam power of each gun can be adjusted between zero and maximum power with an accuracy of þ2%.
Other Equipment. Large-production furnaces are equipped with lock-valve systems to allow simultaneous melting and unloading of ingots without breaking the vacuum in the melt chamber. Production is thus limited only when the condensate remaining in the melt chamber requires cleaning or when a different alloy is to be melted. Characteristics of Electron Beam DripMelted Metals. Electron beam melted and refined material is of the highest quality. The amount of interstitials present is very low, and trace elements of specific high vapor pressure can be reduced to very low values (Ref 1, 2). Tantalum and niobium ingots have smooth surfaces and are of sufficient ductility that they can be cold worked, and sheets and wires can be produced. Tungsten and molybdenum ingots are also of the highest possible purity, but the ingots are brittle because of the very large grain size and the concentration of impurities at grain boundaries. Hafnium. Electron-beam-melted hafnium is of higher ductility than the vacuum-arcremelted metal (Ref 3). The main application of electron-beam-melted hafnium is as control elements for submarine nuclear reactors. Vanadium is refined by electron beam drip melting. The alumino-thermically produced feedstock is drip melted in several steps. During this procedure, the ingot diameter is reduced at each step by approximately 30 to 40 mm (1.2 to 1.6 in.) to obtain an ingot 30 to 40 mm (1.2 to 1.6 in.) in diameter, regardless of the initial ingot diameter. The clean vanadium ingots are primarily used in nuclear reactor applications (Ref 4). Applications for electron-beam-melted refractory and reactive metals are listed in Table 3.
Cold Hearth Melting The cold hearth melting process (Fig. 5) was developed approximately 10 years after drip melting (Ref 5). Cold hearth melting is mainly used for refining and recycling reactive metal scrap, especially titanium turnings machined with high-density tungsten carbide tool tips (Ref 6) and for refining specialty steels and superalloys. Principles of Cold Hearth Melting. Cold hearth melting (Fig. 6) is the most flexible vacuum metallurgical melting process. It is a two-stage process. The first step (material feeding, melting, and refining) takes place in a water-cooled copper trough or hearth. In the second step, solidification occurs in one of several round, rectangular, or specially shaped water-cooled continuous copper crucibles. Both process steps are nearly independent from each other; they are linked only by the flow of the liquid metal stream. The major refining actions are carried out in the hearth, but some
146 / Melting and Remelting
Fig. 5
Schematic of the cold hearth melting process
postrefining takes place in the pool of the continuous casting crucible, similar to the drip melting of horizontally fed billets. Refinement in cold hearth melting occurs by vacuum distillation in the hearth pool, superheating, and stirring of the molten metal pool. Removal of Impurities. Most impurities with densities lower than that of the melt (for example, oxides in superalloys) can be segregated by flotation and formed into a slag raft. The raft is then held in place by either mechanical or electrothermal means. Impurities denser than the melt, such as tungsten carbide tool tips in titanium, are removed by sedimentation. Inclusions with densities such that efficient flotation or sedimentation does not occur can be partially removed by adhesion to the slag raft or by dissolution in the liquid pool. Hearth dimensions are based on the type and amount of refining required. For example, hearths for vacuum distillation should be nearly square and relatively deep to allow sufficient melt stirring. For flotation refining, the hearth should be
long and narrow. Hearths for titanium alloy scrap recycling can be relatively short if all the materials can be transported to the pool of the hearth rather than to the ingot pool. Longer hearths provide for longer residence times, improving the removal of impurities by dissolution. Feeding. Material feeding criteria include 100% homogeneous material transportation to avoid uncontrolled evaporation of alloying elements and correct feeding into or above the hearth pool. Horizontal feeding of compacted, premelted, or cast material is most often used. Loose scrap and raw material can be used. Feeding of liquid metal was used in one of the first cold hearth melting furnaces to produce ferritic steel in a vacuum induction furnace (Ref 7). Postrefining was carried out in a cascade of five hearths 1.5 m (60 in.) long and 1 m (40 in.) wide. Casting and Solidification. The criteria for material casting and solidification include the shape of the final product and the solidification rate required to avoid ingot tears or other
defects and to ensure a homogeneous ingot structure. The multiple casting of small ingots is sometimes used, especially when forging is impossible because of the brittleness of the solidified material (for example, MCrAly wearresistant coating alloys). The casting of round and rectangular ingots and slabs is common practice. Cold Hearth versus Drip Melting. Table 4 compares the essential features of drip melting and cold hearth melting. Generally, cold hearth melting is used for titanium and superalloys especially when flotation or sedimentation of inclusions is required. Drip melting is used for other refractory metals because of the high melting points and the resulting high heat losses to the water-cooled copper crucible. Depending on production quantity, double or triple drip melting may require less energy than a single cold hearth melt of some materials, such as niobium. Refining and Production Data. Data on cold hearth electron beam melting and refining in laboratory and pilot production furnaces are given in Table 5. The data demonstrate the effectiveness of the process in reducing impurities and interstitial elements. It can also be seen that the selective evaporation of chromium from superalloys can be controlled by the distribution of beam power at the trough pool and by controlling trough pool area and melt rate. The selective evaporation of aluminum from Ti6Al-4V alloy is much more difficult to control; additional aluminum must be used to compensate for the aluminum evaporated. Equipment for Cold Hearth Electron Beam Melting. The equipment required for cold hearth melting is different from that used in drip melting, mainly because of the hearth and the larger melting chamber. In addition, because of the materials often melted in the cold hearth process (superalloys and titanium alloys), additional instrumentation is often provided. This may include an ingot pool level control system, metal vapor and partial pressure analyzers, a two-color temperature control system, and a data logging system. Accurate beam power distribution is achieved in multiple-gun furnaces by microprocessor control, which allows the splitting of a single beam to many locations and the adjustment of dwell time at each location. The beam spot at each of the locations can be scanned to create spot patterns as required by the application. With such systems, the required heating can be achieved without unnecessary power consumption and evaporation of alloying elements. Process observation is accomplished with a video monitoring system. Samples can be obtained from both the trough pool and the ingot pool for nearly continuous control of material quality. Feeding systems for cold hearth furnaces must maintain homogeneity along the length of the feed material. The trough and crucible should be easily accessible for convenient maintenance, especially when different alloys are to be melted in the same furnace.
Electron Beam Melting / 147 Characteristics of Cold-Hearth-Melted Materials. Titanium ingots and slabs can be produced from titanium scrap contaminated with tungsten carbide tool tips. The electronbeam-melted product will not contain any tungsten carbide particles. Oxygen content can be controlled by mixing titanium sponge with titanium revert material in the appropriate ratio. Cold-hearth-melted titanium ingots can be remelted in a vacuum arc remelting furnace. In some cases, depending on alloy and application, the product of the cold hearth furnace can be used without additional melt processing. Superalloys. Cold-hearth electron-beammelted superalloy ingots 150 to 200 mm (6 to 8 in.) in diameter are often used in vacuum induction melting investment casting furnaces. Such ingots are nearly free of nonmetallic inclusions and trace elements (Ref 9).
ACKNOWLEDGMENT Revised from article by W. Dietrich and H. Stephan, Electron Beam Melting and Casting, Metals Handbook, 9th Edition, Volume 15, 1988 pp 410–418
REFERENCES
Fig. 6
Four-gun 1200 kW combined electron beam drip melting and cold hearth melting furnace
Table 4
Comparison of the characteristics of drip melting and cold hearth melting
Characteristic
Power density Inclusions Ingot shape and structure Mass production Competitive economical processes Preferred method
Refractory metals
High Irrelevant Round; coarse grain Low Vacuum arc remelting Drip melting
Reactive metals, superalloys, and specialty steels
Soft; smoothly distributed Must be removed Round or flat; fine grain, segregation-free High Vacuum arc remelting; electroslag remelting Cold hearth melting
1. R.E. Lu¨ders, Tantalum Melting in a 800 kW EB Furnace, Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R. Bakish, Ed., 1983, p 230–244 2. J.A. Pierret and J.B. Lambert, Operation of Electron Beam Furnace for Melting Refractory Metals, Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R. Bakish, Ed., 1984, p 208–218 3. H. Sperner, Hafnium, Metallwis. Technik., Vol 7 (No. 16), 1962, p 679–682 4. R. Ha¨hn and J. Kru¨ger, Refining of Vanadium Aluminium Alloys to Vanadium 99.9% by EB Melting, Proceedings of the
148 / Melting and Remelting Table 5 Refining and production data for the melting of reactive and refractory metals and stainless steels in laboratory and pilot production furnaces
Metal
Feedstock size, mm (in.)
Hafnium
60 (2.4) square
Zirconium
100 (4) square
Zirconium
80 (3.2) square
Vanadium
50 (2) square
Ti-6Al-4V
Swarf
Ti-6Al-4V
Solid scrap
Ti-6Al-4V
125 (5) diam
Commercially 160 (6.3) pure titanium Commercially Sponge pure titanium Stainless steel 150 (6) diam Alloy 718
133 (5.2) diam
AISI type 316 stainless steel
150 (6) diam
Trough size, mm (in.)
Ingot size, mm (in.)
Ingot weight, kg (lb)
Melt rate, kg/h (lb/h)
Electron beam power, kW
120 250 (5 10) 120 300 (5 12) 120 300 (5 12) 120 300 (5 12) 120 300 (5 12) 120 300 (5 12) 150 400 (6 16) 150 250 (6 10)
100 (4) diam
83.0 (183)
40 (88)
180
150 (6)
90.5 (200)
42 (92.5)
185
100 (4)
40.2 (89)
80 (176)
140
Operating pressure, Pa (torr)
Specific melting energy, kW 7 h/kg
Composition of feedstock and product
C, ppm
4 10 2 (3 10 4) 3.5 10 2 (2.6 10 4) 3.5 10 2 (2.6 10 4) 1.5 10 2 (1.1 10 4) 2 10 2 (1.5 10 4) 7 10 2 (5.3 10 4) 6 10 2 (4.5 10 4) 6 10 2 (4.5 10 4)
4.5 4.4 1.75
O, ppm
N, ppm
H, ppm
Al, %
V, %
Cr, %
... ... ... ... ... ... ... ... 400 200 1520 ... ... ... ... ...
900 . . . 600 . . . 950 95 540 30 4000 800 1520 210 1045 210 277 50 2600 110 2700 110 1520 75 1320 76 ... ... ... ... ... ... ... ...
... ... 30 3 10 3 10 3 84 22 15 8 ... ... ... ...
... ... ... ... ... ... ... ... 6.0 4.4 6.0 4.8 6.0 3.6 ... ...
... ... ... ... ... ... 99 99 4.0 4.2 4.0 4.1 4.0 4.3 ... ...
... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ...
100 (4)
...
20 (44)
130
150 (6)
62.6 (138)
40 (88)
122
150 (6)
62.6 (138)
70 (154)
140
2 75 (3) diam 100 400 (4 16)
2 32 (70.5)
91 (200)
147
96.4 (213)
86.3 (190)
148
150 500 (6 20)
100 400 (4 16)
103.0 (227)
41.2 (91)
226
8 10 2 (6 104)
5.5
...
...
...
...
...
...
...
150 400 (6 16) 150 400 (6 16) 150 400 (6 16)
2 75 (3) diam 2 75 (3) diam 3 65 (2.6) diam
2 55 (121)
136 (300)
144
1.06
2 57 (126)
136 (300)
156
3 41.5 (91.5)
136 (300)
156
6 10 2 (4.5 10 4) 6 10 2 (4.5 10 4) 6 10 2 (4.5 104)
701 536 417 363 ...
97 33 14 17 ...
155 68 52 34 ...
... ... ... ... ...
... ... ... 0.72 ...
... ... ... ... ...
18.25 18.11 19.11 18.73 ...
6.5 3.0 2.0 1.61 1.71
1.15 1.15
Source: Ref 8
Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R. Bakish, Ed., 1986, p 53–67 5. C.d’A. Hunt and H.R. Smith, Electron Beam Processing of Molten Steel in Cold Hearth Furnace, J. Met., Vol 18, 1966, p 570–577 6. H.R. Harker, Electron Beam Melting of Titanium Scrap, Proceedings of the Bakish Conference on Electron Beam Melting and
Refining (Reno, NV), R. Bakish, Ed., 1983, p 187–190 7. C.d’A. Hunt, H.R. Smith, and B.C. Coad, The Combined Induction and Electron Beam Furnace for Steel Refining and Casting, Proceedings of the Vacuum Metallurgy Conference (Pittsburgh, PA), June 1969, p 1–22 8. H. Stephan, Production of Ingots and Cast Parts from Reactive Metals by Electron
Beam Melting and Casting, Proceedings of the Third Electron Beam Processing Seminar (Stratford, U.K.), 1974, p 1b1–1b69 9. M. Krehl and J.C. Lowe, Electron Beam Cold Hearth Refining for Superalloy Revert for Use in Foundry Production, Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R. Bakish, Ed., 1986, p 286–296
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 149-154 DOI: 10.1361/asmhba0005205
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Plasma Melting and Heating Reviewed and revised by Robin Lampson and Todd Telfer, Retech Systems
PLASMA MELTING is a material-processing technique in which the heat of a thermal plasma is used to melt a material. Plasma torches produce heat from an electrical current that passes through a partially ionized gas (or plasma). The plasma is generated by the phenomenon known as gaseous discharge, whereby a gas is ionized by electrical breakdown. The electrical breakdown of the gas can be accomplished by an applied voltage between two electrodes (as in the case of fluorescent lights). Plasmas also can be generated by various “electrodeless” methods that involve radio frequency (RF) discharges, microwaves, high-frequency induction coils, shock waves, lasers, or highenergy particle beams. Plasma arc torches have several advantages. Unlike heating by consumable electrodes (such as tungsten arc or graphite electrodes), plasma arc heating involves a gas medium instead of a consumable electrode. The gas environment and flow rates can be controlled, and only a small percentage of ionized gas (less than ½% of the total gas flow) is sufficient to create heating temperatures. Without a consumable electrode, plasma arc can generate higher temperatures than the electric arc methods. Products of combustion are eliminated or reduced. In addition, high inert gas pressure (5 kPa, or 50 mbar, or 37.5 torr, and higher) can be obtained with plasma torches. This is an important prerequisite for the avoidance of selective evaporation of alloying elements from the melt, an aspect that is particularly important in the case of titanium and superalloys. Plasma arc torches also supply electric heat to a ladle of molten metal to either maintain the temperature or raise the temperature of the melt. For numerous steel grades, especially those with extremely low carbon contents, plasma arc ladle heating is an alternative to heating with graphite electrodes and the risk of carbon pickup. Highpower plasma torches can be used with minimum modification to the existing ladle design. Prior to their integration into steel ladle metallurgy stations, plasma torches had been used primarily for scrap melting and heat-loss compensation before and during continuous casting. Only thermal (hot) plasmas are considered in this article. The term thermal differentiates
these plasmas from cold (nonequilibrium) plasmas, which are characterized by high electron temperatures and low sensible temperatures of the heavy particles. Cold plasmas are produced in various types of glow discharges, low-pressure RF discharges, and corona discharges. Cold plasmas are not suitable for melting processes. Thermal plasmas are suitable for melting in the production of ingots, slabs, castings, or powders.
Plasma Torches Breakdown (that is, the creation of ionic charge carriers) establishes a conducting path for resistance heating within a plasma torch. Most plasma generators (plasma torches) for melting processes use an electric arc to produce gaseous discharges. The characteristics of an electric arc include relatively high current densities, low cathode fall, and high luminosity of the column. A typical potential distribution along an arc is shown in Fig. 1. The column cross section at the cathode is smaller than at the anode, with contrary current and power densities. This fact influences the design principle of plasma torches. Most plasma heaters operate on direct current (dc) and are designed as either nontransferred or transferred arcs (Fig. 2). However, threephase alternating current plasma torches are employed in plasma heating and degassing applications. These units do not need a bottom electrode but do require more sophisticated power supplies than dc plasma torches do in order to limit the torch current and to ensure a quick, uninterrupted current pass through zero. In a nontransferred plasma torch, the gas is the only resistive element that generates heat. Nontransferred arcs are desirable for hot gas contact in shaft heaters, melters (cupolas), and reactors such as blast furnaces. This mode is suitable for the melting of conductive and nonconductive materials, but it has the disadvantage of lower efficiency compared to the transferred mode. Examples of nontransferred plasma torches are illustrated in Fig. 3. Magnetic coils are used to direct the flow of the plasma in the torch. For the electrodeless RF
torch, the plasma does not depend on the physical connection of electrodes. Transferred Arc Plasma Torches. In the transferred mode (Fig. 2b), one electrode is inside the torch, and the counterelectrode is the material to be melted. With the workpiece being one of electrodes, it provides the medium for resistive heating of both conductive and nonconductive materials. Most nonconductive materials become conductive when heated. Transferred arc heaters are useful for bulk heating of conducting solids and melts, where the gas phase is not a reactant but can be used as an inert shield. Figure 4 shows two typical design principles for torches in the transferred mode: the tungsten tip design and the hollow copper electrode design. In the tungsten tip design (Fig. 4a), the torch electrode is connected as the cathode because high current densities are required. The cathodes are usually made of tungsten, along with small additions of thoria to lower the thermionic work function of tungsten. Still, electron emission requires high electrode temperatures (3500 to 6000 K, or 5800 to 10,300 F) at the attachment of the arc. Therefore, the cathode material is liquid at the arc attachment point even with water cooling at the rear of the
Fig. 1
Typical potential distribution along a plasma arc. Va, anode voltage; Vc, cathode voltage; da, anode current density; dc, cathode current density
150 / Melting and Remelting
Fig. 2
Configurations of (a) Nontransferred arc plasma torch and (b) transferred arc plasma torch
Fig. 3
Examples of some typical nontransferred plasma torch configurations. (a) Electrode-based torches. (b) Radio frequency “electrodeless” plasma torch
cathode. For this reason, oxidizing gases generally cannot be used in direct contact with the tungsten tip. Argon shielding is one solution. Further, it should be kept in mind that portions of the small liquid pool at the cathode tip are ejected into the material being melted. The high melting point of tungsten can result in inclusions in the product; such defects are not tolerable in high-strength metals such as titanium alloys and superalloys for aircraft applications. With argon as the plasma gas, the service life of the cathode in the tungsten tip design is typically between 30 and 150 h. The inert gas consumption and plasma column stabilization of such torches are low. The hollow copper electrode design is shown in Fig. 4(b). The electrode material, usually copper alloyed with chromium or zirconium, is intensively water cooled. Copper is a field emitter of electrons (as opposed to tungsten, a thermionic emitter), because its boiling point is substantially below that required for thermionic electron emission. Current densities at the attachment points on cold cathodes are double those for thermionic emission on hot cathodes. Cathode spots are of high intensity, rapidly moving on the surface of the cold cathode and causing erosion by locally vaporizing copper. For this reason, the polarity is usually reversed in plasma torches with cold electrodes. The rear electrode acts as the anode, and the workpiece or melt is the cathode. The anodic arc attachment is smoother, and erosion is reduced. The plasma-forming gas is blown in tangentially between the rear electrode and the nozzle to create a swirl or vortex to stabilize the arc and to rotate the arc attachment in the rear electrode. With the movement of the arc attachment, a substantially lower power density can be achieved in the rear electrode. Electrode life of 1000 h has been demonstrated in torches using argon or helium as the plasma gas. All inert and reactive gases, or mixtures of these, can generally be used in such torches, but the influence of the gases on electrode erosion varies. The gas consumption of vortex-stabilized torches is higher than that of other designs. The efficiency of plasma torches is less affected by design principles than by process parameters and by the type of plasma gas used. Efficiency is defined as Pw/Pt, where Pt is the electrical power input of the torch, and Pw is the thermal power applied to the workpiece or melt. The efficiency of a 650 kW torch measured in a cold furnace ranges from approximately 25 to 50%, depending on the plasma gas used, furnace pressure, and the distance between torch and workpiece. The highest efficiency is obtained with helium as the plasma gas and a short distance between material and torch. The power loss in torches with hollow water-cooled electrodes is approximately 25%. Remaining additional 25% losses are due to radiation and convection of the plasma column between torch and workpiece. These losses are primarily a function of
Plasma Melting and Heating / 151
Melting Processes
Fig. 4
Design concepts for plasma arc torches in the transferred mode. (a) Torch with tungsten tip and concentric gas flow. (b) Torch with hollow copper electrode and vortex generator
the distance between torch and workpiece, but radiative losses of the plasma column also increase with pressure.
Furnace Equipment Plasma melting furnaces usually consist of double-wall, vacuum-tight, water-cooled constructions that can withstand the radiative and convective heat transfer from the plasma column. Furnaces also usually have material feeding systems. These can consist of vibratory or rotary feeders with bins (for bulk materials) or bar feeding systems. If the furnace is to operate continuously, the feed material must be “locked in” to prevent disturbance of the furnace atmosphere. The furnace must also have a vacuum pump so that it can be evacuated before backfilling. If the furnace is operated under reduced pressure, an offgas pump with an on-line gas control valve is required.
Atmosphere Control Plasma melting furnaces are usually operated under slightly positive pressure to prevent the potential atmospheric contamination by oxygen and nitrogen. However, state-of-the-art furnaces are vacuum-tight and can be operated at pressures between 5 and 200 kPa (50 and 2000
mbar, or 38 and 1500 torr). Vacuum-tightness is essential because the back diffusion of oxygen, nitrogen, and moisture through small leaks can be easily demonstrated even with positive pressure in the furnace. As noted, plasma torches are normally operated with argon or helium plasma gas. A typical gas purity of 99.999% indicates 10 ppm gaseous impurities and corresponds to an absolute pressure of 1 Pa (0.01 mbar, or 0.0075 torr) in a vacuum process. Additional sources of atmospheric contamination are the moisture and gases in the feed material, desorption of the furnace wall, and leaks in the furnace construction. In the case of metals with high affinities for oxygen or nitrogen, these gases are totally absorbed by the melt and will be found as oxide or nitride inclusions in the product. To obtain a clean initial furnace atmosphere, the vacuum-tight plasma furnace is evacuated with a mechanical pump to final pressure (typically approximately 1 Pa, or 0.01 mbar, or 0.0075 torr). The furnace walls can be degassed with the help of warm water running between the double walls of the furnace to enhance moisture desorption. The furnace is then backfilled with high-purity inert gas up to 1 kPa (10 mbar, or 7.5 torr), evacuated again to 1 Pa (0.01 mbar, or 0.0075 torr), and backfilled to operating pressure. This procedure dilutes residual gas impurities by a factor of 1000.
Plasma furnace equipment is of great interest when selecting a suitable heat source for melting, remelting, casting, and atomizing of reactive metals, titanium, and superalloys. Plasma torches, however, are the only nonconsumable heat sources for melting under high inert gas pressures in skull crucibles. High pressure is an essential requirement for preventing the selective evaporation of alloying elements that are characterized by high activity coefficients and/or vapor pressures, such as chromium and manganese in superalloys and aluminum in titanium alloys. Various plasma melting processes have been developed, including plasma consolidation, plasma arc remelting, plasma cold hearth melting, and plasma casting. Plasma cold hearth melting and plasma cold crucible casting are thought to have the most potential for industrial application. Table 1 lists some typical data on feed materials, operating parameters, and products of furnaces for some of these processes. With plasma furnaces, several companies are producing 3 to 5m (120 to 200 in) slabs with sections 50 to 89 cm (20 to 35 in.) thick and 100 to 150 cm (40 to 60 in.) wide. Plasma Consolidation. Consolidation implies that low-density feed material is converted into a high-density product. Plasma consolidation is used for titanium scrap and sponge to produce an electrode for further remelting in vacuum arc remelting (VAR) furnaces. The material is usually fed directly into a watercooled copper crucible and melted by plasma heat from one or several plasma torches (Fig. 5). The water-cooled copper baseplate of the crucible is continuously withdrawn. The shape and depth of the liquid pool in the crucible depend on feed rate and power input. Under these conditions, 100% melting cannot be ensured; some material can be trapped at the liquid/solid interface before melting. Figure 6 shows ingot density as a function of the raw material feed rate and power consumption in the plasma consolidation of titanium. Plasma consolidation processes can reduce manufacturing costs for the preparation of titanium VAR electrodes, which are usually plasma melted from pressed sponge or chip compacts. With plasma consolidation, metal chlorides can be removed from titanium sponge to sufficiently low levels. Tungsten carbide tool tips in titanium chips cannot be removed, and they eventually form high-density inclusions in the product. Therefore, the conventional x-ray inspection of scrap is still necessary. Low-density inclusions such as titanium nitride particles have not been found after plasma consolidation of contaminated titanium scrap and subsequent VAR remelting. Plasma consolidation has also been successfully applied to the processing of titanium aluminide (TiAl), a low-density alloy with excellent high-temperature service properties.
152 / Melting and Remelting Table 1 Characteristics and operating parameters of furnaces for plasma melting processes Type
Feed material
Plasma consolidation Titanium scrap and furnaces sponge Plasma consolidation Titanium scrap furnaces and sponge Plasma arc remelting Titanium, titanium furnaces alloys Plasma arc remelting Specialty steels, furnaces heat-resistant alloys
Plasma cold hearth melting furnaces Plasma cold hearth melting furnaces
Specialty steels, heat-resistant alloys Titanium scrap and sponge Titanium scrap and sponge, superalloys
Plasma cold crucible Titanium scrap melting and and sponge, casting furnaces superalloys
Fig. 5
Product form and size, mm (in.)
Ingot 355–430 (14–17) diam, 3000 (120) long; density >90% Ingot 700 (27.5) diam, 3800 (150) long; density 98% Ingot 125 (5) diam, 500 (20) long Ingot 150–200 (6–8) diam, 1200 (47) long; slabs 70 300 1200 (2.8 12 47) Ingots 650 (25.5) diam, 1500–2300 (60–90) long Slab 245 1125 2450 (9.6 44.3 96.5)
Specific power consumption, kW h/kg
Plasma gas
250–260 (550–570)
1.4–1.6
Argon
Atmospheric
Tungsten tip with external ignition bar; 6 90 kW
250–670 (550–1480)
0.89–2.4
Argon
Atmospheric
50–70 (110–155) 90–135 (200–300)
2.0–2.5
...
Atmospheric
1.5–2
...
Atmospheric
Hollow copper electrode with graphite liner; 600 kW Probably tungsten tip; 500 kW Probably tungsten tip; 500–1000 kW
400–900 (880–1985)
1.2–1.8
...
Atmospheric
Probably tungsten tip; 300–3600 kW
1–13 Pa (0.01–0.13 mbar)
Six hollow cathode torches; 700 kW on hearth; 1170 kW on crucible Hollow copper electrode, vortex stabilized; 200 kW on hearth, 100 kW on crucible Hollow copper electrode, vortex stabilized; 650 kW
Typical melt rate, kg/h (lb/h)
Typical furnace pressure
330 (730)
9
...
Ingot 195 (7.7) diam
90–227 (200–500)
1.3–3
Helium; gas recycling
Atmospheric
Casting ingots
380 (840)
1.3
Argon; gas recycling
50 kPa (500 mbar)
Plasma consolidation furnaces. (a) Multiple electrode. (b) Single electrode
Plasma Arc Remelting. An electrode that has already been melted by a primary melting process such as electric arc or induction melting can be plasma arc remelted into a water-cooled withdrawal crucible. The major objectives for plasma arc remelting are: To obtain directional solidification without
changing the chemical composition of the feed material To improve cleanliness by removal, size reduction, shape control, and redistribution of inclusions To lower gas impurity content To alloy with nitrogen Plasma arc remelting is used most frequently in the former Soviet Union countries to produce high-temperature alloys, bearing steels,
superalloys, and titanium alloys. Plasma arc remelting can be advantageous for the nitrogen alloying, deoxidation, or desulfurization of nonreactive alloys. An example of successful reactive melting is the nitrogen alloying of steel during plasma arc remelting with argon-nitrogen mixtures as the plasma gas. Figure 7 shows nitrogen concentrations in various metals as a function of nitrogen partial pressure in the gaseous phase after plasma arc remelting. The data indicate that excited molecules of nitrogen in the plasma react with the metal. In plasma cold hearth melting, bars, sponge, or scrap can be continuously fed into a water-cooled copper trough, usually called a hearth, and melted by means of plasma heating (Fig. 8). The liquid metal flows over the lip of the hearth into a withdrawal mold. The plasma cold hearth melting process is adapted from
Torch design principle and power
the electron-beam cold hearth refining process, which is a well-developed refining and casting method for superalloys. Plasma heat, instead of an electron beam, is employed to minimize selective evaporative losses that occur during electron-beam melting in the pressure range of 0.01 to 1 Pa (10 4 to 0.01 mbar, or 7.5 10 5 to 0.0075 torr). In the hearth, high-density inclusions from tungsten carbide tool tips will sink down to the skull. Low-density inclusions can be dissociated or dissolved in the base metal during interaction with the superheated melt and the plasma jet. Process optimization is necessary to prevent short-circuiting of the hearth by unmelted particles and to establish sufficient dwell time at high melting rates. The power distribution of one or more plasma torches required for smooth melting and fluid flow can be ensured by automated torch-moving devices (Fig. 9). In unconventional configurations, there is no mechanism provided for torch movement. Figure 8(b) shows a modern plasma cold hearth melting furnace concept with a doublewall, water-cooled, vacuum-tight furnace chamber and three plasma torches for the production of ingots 710 mm (28 in.) in diameter. Two torches serve the hearth, and the third the ingot mold. The torches use hollow water-cooled copper electrodes. The high gas consumption of this torch design means that an effective gas recycling system must be provided, especially if helium is used as the plasma gas. A key technology in the plasma melter is gas recycling systems that include reactive filters, oil-free high-capacity pump sets, and compressors. If necessary, the recycling system can be supplemented with cleaning systems for oxygen, carbon dioxide, nitrogen, hydrogen, and moisture. Helium gas recycling systems are
Plasma Melting and Heating / 153 now operating at 99.5% recycle efficiency (less than 0.5% of full gas flow). Plasma cold crucible casting is an alternative process to vacuum arc skull melting and casting or electron-beam skull melting and casting, and it can overcome some of the disadvantages of these processes. The disadvantages of vacuum arc skull melting in the melting and casting of titanium and superalloys are: Control of superheating is not possible;
increased power results in increased melting rate instead of superheating.
Expensive electrodes must be used instead
of the high-quality unconsolidated revert material that can be used for plasma cold crucible casting. For small batches of superalloys (up to 8 kg, or 18 lb), electron-beam skull casting is a well-established technique. Electron-beam melting and casting is rarely used for large castings of either superalloys or titanium alloys. This is because of the shallow melt pool in the crucible due to lack of stirring action and the potential for selective evaporation of alloying elements during melting.
The concept of plasma cold crucible melting is shown in Fig. 10(a). Solid material can be charged continuously into the water-cooled copper crucible, and additional material can be fed into the crucible during melting. The melt is poured into stationary molds or centrifugal casting molds situated in the mold chamber. The pouring stream can be superheated during casting or for production of atomized powder. An example of a cold crucible casting furnace (Fig. 10b) shows a horizontally installed mold chamber with the melt chamber on top of the mold chamber. In this example, a programmable torch-moving device is installed on top of the melt chamber. The pressure in the furnace during the melting operation can be preselected; pressures typically range from 30 to 50 kPa (300 to 500 mbar, or 225 to 375 torr). The melting operation can be started at every pressure in this range. The typical casting weight of this furnace is 27 kg (60 lb) of titanium with a 650 kW plasma torch. To achieve a deep molten pool, a high heat flow from the plasma torch to the melt must be established in combination with high stirring velocity in the melt to enhance the effective heat-transfer coefficient. The stirring action in the melt can be regulated by controlling current, gas flow, and torch-to-melt distance. Continuous feeding of the metal to be melted is required in order to achieve a deep pool. In contrast to vacuum arc skull melting, the crucible need not be inspected after each melt. The power densities in the plasma attachment area are adjusted to avoid damaging the crucible if it is inadvertently exposed to the plasma for short periods of time.
Fig. 6
Relationship among raw material feeding rate, ingot bulk density, and specific power consumption in the plasma consolidation of titanium. Source: Ref 1
Fig. 7
Nitrogen concentrations in various metals as a function of the partial pressure of nitrogen after plasma arc remelting. (a) Iron. (b) 16–25–6 stainless steel. (c) Austenitic stainless steel. Curves A and B, induction heating; minimum and maximum nitrogen concentrations, respectively. Curve C, plasma arc remelting. Source: Ref 2
154 / Melting and Remelting
Fig. 8
Plasma cold hearth melting. (a) Concept. (b) Configuration of system for recycling scrap
Fig. 9
Plasma torch-moving device. Courtesy of Leybold AG
REFERENCES 1. T. Yagima et al., Development of the Plasma Progressive Casting Process and Its Application for Titanium Melting, Titanium 1986: Products and Applications, Vol II, Proceedings of the Technical Program from the 1986 International Conference, Titanium Development Association, 1987, p 985–993 2. G.K. Bhat, Plasma Arc Remelting, Plasma Technology in Metallurgical Processing, J. Feinman, Ed., Iron and Steel Society, 1987, p 163–174
Fig. 10
Cold crucible plasma melting. (a) Melting, pouring, and superheating for powder atomization. (b) Industrialsized plasma cold crucible casting furnace with 650 kW torch power
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 155-159 DOI: 10.1361/asmhba0005198
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Crucible Furnaces George Giles, Morgan Molten Metal Systems
CRUCIBLES, in the form of well-fired clay pots, were present at the dawn of the metals industry, some 6000 years ago, and gave early metal workers the ability to melt metal in a durable container that could then be used to pour the molten material into a mold. Few, if any, metallurgical or industrial processes have a more distinguished history than the crucible melting of metals. Crucibles or earthenware smelting pots that were used to refine copper from King Solomon’s mines can still be found in that area of ancient Israel. Some pots had capacities of up to 0.40 m3 (14 ft3). The Wootz steels of India, the famous Damascus sword blades, and the equally famous steels of Toledo, Spain, were produced using crucible melting processes. At their most basic level of form and function, crucibles today (2008) are little changed from their early prototypes. However, in terms of materials, construction, thermal efficiency, and service life, modern 21st century crucibles (Fig. 1) are superior to those in use just a short time ago. Designed for a wide range of traditional applications and new processes, crucibles are formed using a wide variety of materials and processes that impart specific characteristics, capabilities, and limitations to each of the many types of crucibles available from crucible manufacturers. Crucibles may be as small as teacups or may hold several tons of metal.
Fig. 1
Large electric bale-out crucible furnace. Courtesy of Morgan Crucible
They may be fixed in place within a furnace structure or may be designed to be removed from the furnace for pouring at the end of each melt. Crucibles are used in gas- and oil-fired furnaces, electric resistance furnaces, and induction furnaces. They come with or without pouring spouts and in a wide variety of traditional and specialized shapes. There is no direct impingement of the flame on the metal, and the heat loss to the outside is limited by the refractory walls. However, the energy efficiency of the crucible furnace can be low (7 to 19%), with over 60% of the heat loss attributed to radiation (Ref 1). Without proper care, the life of the crucible furnace may be short. Temperature control may also be difficult.
Applications In the metals industry, most crucibles are used to melt, hold, and/or transfer nonferrous metals—commonly aluminum, zinc, and their alloys and copper and copper-base alloys—as well as precious metals. A crucible furnace is the least expensive melting method preferred for melting small volumes of nonferrous metals. It is also used for higher-temperature alloys, such as nickel bronze and cupronickel, and, to a lesser extent, for melting ferrous metals, such as gray iron. Crucible furnaces are popular in jobbing foundries and die casting shops because they are easy to tap and charge with another alloy. Foundry melting and handling operations with crucible equipment include furnace melting, pouring ladle equipment (either separately heated or as a ladle liner), and liners for electric induction melting furnaces. Crucible melting can also be used as a component of qualityadjustment equipment between melter and casting furnaces to alter the chemistry and temperature of the melt for control of such factors as porosity content prior to casting. Metal-processing operations include in-line filtration, fluxing, and hydrogen gas content control. Crucible Advantages. Crucible melting is a simple and flexible process. Crucible furnaces can generally be started or shut down at a moment’s notice. With their unique ability to
melt, hold, and transfer metal using a single vessel and to allow even incompatible alloy changes to be made simply by switching vessels, crucibles provide significant operational flexibility for various foundry applications. Even when fixed within the furnace structure, crucibles offer important advantages when compared to directly heated fuel-fired furnaces and to electric resistance or induction furnaces with rammed refractory linings. Advantages include: Small, compact size and relatively lower
acquisition cost are economical.
Smaller (and round) furnaces provide and
maintain good, even temperature profiles.
Open crucible furnaces, or even those with
pivoting lid covers (Fig. 2), allow easy access for ladling. Lower Metal Loss/Cleaner Metal. In fuelfired furnaces where the metal is heated directly by a flame produced by gas or oil burners, there are both high levels of metal loss and increased oxidation due to the combustion products generated. In fuel-fired crucible furnaces, however, the flame is isolated from the metal, and the metal is heated indirectly with heat conducted through the crucible. This indirect heating greatly reduces metal loss and metal oxidation, resulting in lower metal purchase costs, lower energy costs per pound of metal poured, and less slag or dross formation. Alloy Flexibility. Even when fixed in place in a furnace, the high-density interior surface of a crucible reduces the likelihood of cross contamination when alloy changes are made. Quick Replacement. Removing a worn or damaged crucible from a furnace and installing a new crucible is a faster, simpler, and less technically demanding process than removing and replacing a worn or damaged rammed refractory furnace lining. Also, because crucibles are preformed to exact material and dimensional specifications, there are few concerns about improperly compacted refractory or incomplete sintering. Matching Crucible and Application. Determining the required properties of an application is the most important step in selecting the proper crucible. Knowing the metal to be
156 / Melting and Remelting Glass: Glass is used to provide a semiperme-
able barrier to gases, controlling the ingress of oxygen and egress of combustion products. Typically, 10 to 20% of a crucible is glass. Aluminosilicates: Aluminosilicates are used to modify the refractory properties of the crucible or internal glass composition. Unfortunately, there is not a single crucible material recipe or manufacturing process that offers the highest level of every desirable characteristic for every application. Therefore, the best course is to prioritize the list of crucible properties most important for the application and review those requirements with the crucible provider to find the crucible product best suited for the specific needs.
Furnace Types
Fig. 2
Crucible furnace in die casting foundry. Courtesy of Morgan Crucible
melted, held, and/or transferred, however, is just the start in selecting the right crucible type for the application. The specific crucible properties that are important for a particular application must be considered. These properties include: Maximum temperature limits: It is important
that the highest temperature reached by the application does not exceed the maximum temperature limit for the crucible; otherwise, the crucible will fail. Tolerance of thermal shock: If the application results in rapid heating or cooling of the crucible, a crucible type must be selected that resists thermal shock. Impact resistance and mechanical strength: If the crucible in the application is used to transfer metal or if it is subject to the impact of dropped charge materials, a crucible with the strength and durability to survive these activities must be selected. Erosion/corrosion resistance: If the application produces slag, dross, or other by-products or uses fluxes that mechanically or chemically attack the inside surface of the crucible, then a crucible type must be selected that resists such attacks. Oxidation resistance: When exposed to air at high temperatures, graphite in a crucible will oxidize. If the application is likely to cause melt oxidation or excessively high temperatures (which exist in many foundry casting applications), a crucible should be selected with best oxidation resistance. Nonwettability: It is important in aluminum foundry application that the inside surface of the crucible resist the intrusion of molten metal into pores in the crucible material. Consequently, a crucible should be selected that possesses specific nonwetting agents.
Electrical properties: If the application
requires a crucible that heats inductively, select a crucible designed for use in an induction furnace and with a specific electrical resistivity that optimizes the coupling power of the crucible. Thermal conductivity: Not all crucibles offer the same level of heat transfer through the crucible and into the metal. To achieve the lowest energy costs, select the crucible with the highest level of thermal conductivity that meets all of the other requirements of the application. Crucibles have been called the first composites, and modern crucibles use an eclectic mix of materials to achieve the important characteristics listed here: Graphite: A foundry crucible contains 30 to
50% by weight of flake graphite, depending on the properties required. The carbon content of graphite imparts high thermal conductivity and nonwettability and, coupled with its platelike layered structure, provides high thermal shock resistance. This is critical to foundry applications where temperatures can change by several hundred degrees in seconds. Silicon carbide: A foundry crucible typically will contain 10 to 50% of silicon carbide, based on the desired characteristics. Silicon carbide provides excellent resistance to high-temperature erosion and corrosion and imparts a high level of thermal conductivity to the crucible. Silicon metal: Fine powders of silicon metal or alloys are used to increase strength and erosion resistance, to protect the graphite and carbon bonding matrix from oxygen, and to help ensure an even glaze coating.
There are many types of crucibles and shapes produced by crucible manufacturers for the foundry market. Most crucibles used in the metals industry for melting, holding, and/or transferring operations are preformed carbonbonded or ceramic-bonded vessels. Silicon carbide crucibles are a widely used type of carbonbonded crucibles, and clay-graphite crucibles are a good example of ceramic-bonded vessels. It is important to discern which type is appropriate for a given application. Table 1 lists the several shapes and the application that is generally appropriate for their use. The basic types of foundry crucible furnaces can be classified as stationary, tilting, or movable. Stationary furnaces use a crucible set inside a refractory-lined shell that has heating equipment attached. A lift-out crucible is shown in Fig. 3. Normally, molten metal is dipped out with hand ladles, or the crucible containing the molten metal is removed from the shell for the casting operation. By using tong and shank tools to lift and carry the bilge-shaped crucible to the mold, pouring of the casting is done directly from the crucible without the need for potentially damaging metal transfers. The maximum crucible size normally used in this manner holds 140 kg (300 lb) of aluminum or copper. The weight and the strength of the crucible limit this type of operation. Another variation is a pushout furnace (Fig. 4) with induction coils and mechanical lifts to raise and lower the crucibles within the coils. Some stationary crucible furnaces are equipped for hand-ladle dip-out pours. For most metals, it is imperative that the flame and flue gases do not contact the charge. Two methods of controlling this critical factor are illustrated in Fig. 5. The first is the side flue configuration used in an aluminum melter installation (Fig. 5a), and the second is the cover plate configuration used for a brass foundry furnace (Fig. 5b). The cover plate setup incorporates holes on the outside diameter of the crucible top. Also shown is the extensive use of highly insulating fiber refractory material. This material has a low heat
Crucible Furnaces / 157 Table 1
Application of specific crucible shapes
Shape
Advantages
Bilge
Reduced top diameter provides less metal surface area. This helps to conserve heat and cuts metal loss and oxidation. Provides large metal surface area to facilitate manual or automated baling out of metal. Wider lower profile provides more efficient heat transfer in electric resistance furnaces. This shape is also suitable for low-pressure furnaces where die casting stalks are used to feed a die above the metal bath. Provides large surface area to facilitate manual or automated baling out of metal. Narrow lower profile is advantageous in fuel-fired furnaces to provide space for burner flame to circulate. Large-capacity crucibles for bulk melting of metal. Metal is usually then transferred to a holding bale-out furnace. Parallel profile means crucible is equidistant from the induction coil, allowing even heating.
Basin
Bowl
Spouted basin
Cylinder
Furnace type
Application
Typical user
Lift-out (crucible is lifted out of furnace chamber for casting)
General melting work for small castings
“Jobbing” foundries
Bale-out (crucible remains fixed in furnace)
Holding and melt/holding
Die caster
Bale-out (crucible remains fixed in furnace)
Holding and melt/holding
Die caster
Fig. 3
Typical lift-out version of a stationary crucible furnace specifically adapted to the foundry melting of small quantities (<140 kg, or 300 lb) of copper alloys
Melting Basin tilting furnaces (crucible remains fixed in furnace; furnace is tilted to pour metal)
Die caster, ingot producer
Induction tilting furnaces Melting (crucible remains fixed in furnace)
Master alloy producer
Fig. 5
Stationary crucible furnaces equipped for handladle dip-out pours. (a) Side flue configuration. (b) Configuration with cover plate openings on outside diameter of crucible top
Fig. 4
Cross section of a double pushout stationary crucible furnace
content and serves to conserve fuel and to minimize the temperature override that often occurs with the heavier refractory furnace linings. The added refractory flame face is required to accommodate the higher melt temperature. Steel crucibles can also be used in stationary crucible furnaces, as shown in Fig. 6. The high
heat conductivity of the steel offers fuel and meltdown time savings. Steel and ductile iron crucibles are extensively used for magnesium and zinc melting. In these molten alloys, there is little attack of the crucible walls. On the other hand, aluminum alloys will attack the steel crucible walls by forming sludge
(combination of iron, manganese, and chromium), which eventually causes hot spots in the castings produced from this melt. However, there have been successful aluminum foundry uses of steel crucibles. This requires a thorough refractory coating on the crucible surfaces in contact with molten metal and careful cleaning of the rust that tends to accumulate in the firing chamber (for example, flame contact with the
158 / Melting and Remelting
Fig. 6
Cross section of a stationary fuel-fired furnace used for the open crucible melting of magnesium alloys
crucible causes rust and scaling). This can collect in the firing chamber. A special clean-out door is used to remove this residue. Both aluminum and magnesium can react explosively with iron oxide. The use of electrical resistance furnaces and steel crucibles reduces the seriousness of this problem, and long-term use of steel crucibles and resistance furnaces has been successful in the casting of aluminum alloys. Tilting furnaces incorporate a crucible that is supported in the shell such that both are free to rotate on an axis. This allows the molten metal to be poured into ladles or molds. The axis of the pour can be either at the center of gravity or at the lip of the furnace. This design simplifies the handling of molten metal in many foundries. Tilting crucible furnaces are designed to simplify metal removal and have been extensively used in numerous foundries for many years. Figure 7(a) illustrates a center-axis furnace; Fig. 7(b) shows a cross section of a lip-axis furnace. Tilting crucible furnaces are useful where castings require more metal than can be easily handled by stationary crucible furnaces. These units are more flexible and can easily handle production requirements that are too small for the efficient operation of larger hearth furnaces. Figure 8 illustrates the use of a crucible liner for an induction melting tilt furnace. Use of the crucible liner overcomes the problems that would occur as a result of the strong erosive currents existing at the crucible sidewalls. The low electrical conductivity of clay-graphite crucibles makes them ideal for use in such applications. A gradual buildup of molten metal in the heel and voids in the backup refractory, both caused by repeated tilting of the furnace, are detrimental to crucible life. Crucible liners, which provide a durable wear-resistant surface for molten metal, are often used in the brass, steel, and ductile iron foundry industries. This application for the foundry use of crucibles is similar to that for the induction furnace liners shown in Fig. 8.
Fig. 7
Two variations of a tilting crucible furnace. (a) Center axis. (b) Lip axis
Movable crucible furnaces are used in conjunction with automated pouring system methods. Methods include multiple crucible melters arranged on either a shuttle or a turntable, such that one furnace is moved into position for pouring while others are positioned for recharging, stabilizing, and metal treatment.
Design and Operation Proper Furnace Atmosphere. With few exceptions, a slightly oxidizing flame should be used. With aluminum alloys, this is a pale yellow color. For copper-base alloys, this will be tinged with green (due to the oxidation of the copper). A traditional test for the correct furnace atmosphere is to place a cold piece of zinc in the flame. A yellowish color indicates an oxidizing flame, and a black color indicates a reducing flame. The objective is to obtain a flame having a tinge of yellow, which indicates the presence of the desirable, slightly oxidizing atmosphere. Oxidizing conditions were maintained in old coke-fired crucible furnaces at the crucible walls by burning the coke in a 75 mm (3 in.) space around the crucible sidewall. The burning coke maintained a slightly oxidizing atmosphere at the walls and ensured good heat transfer. Burner Control. For efficient melting and optimal crucible life, it is necessary to have even heating of the crucible and to avoid hot and cold areas around the sides of the crucible bowl. Some furnace designers use two or four burners to minimize this problem. This will contribute to longer crucible life, better fuel economy, faster melting, and lower noise levels. Precise periodic adjustment of the burner(s) is essential to efficient crucible furnace operation.
Fig. 8
Cross section of a tilt furnace for the highfrequency induction melting of brass and bronze alloys. Crucible is of clay-graphite composition. Also shown are the locations of the molten metal buildup and the voids in the backup refractory, which shorten crucible life.
Special care is required to avoid overheating of the bath. Approximately 80 to 95% of the heat required to prepare a heat for the foundry is used to heat the charge to the melting temperature and then to convert the solid to a liquid. Very little heat is required to raise the temperature of the newly melted liquid to the required pouring temperature. This can cause problems. For rapid melting of the crucible contents, high heat inputs are normally used. As soon as the charge becomes entirely liquid, little or no additional heat is required. If the equipment or the operator does not immediately turn the heat down or off, the bath temperature can become far higher than required. For most metals, this causes quality-control problems. Currently, most foundries depend on their furnace operators to control this harmful variable. Burner Design. Because of the wide diversity in crucible applications using many different fuels, a summary of these data is well
Crucible Furnaces / 159 beyond the scope of this discussion. It should be noted that developments in the design and operation of burners have significantly improved this furnace component. Any new installations of significance probably should include recent developments in both burner designs and burner control units. Charging Techniques. The most efficient melting (in terms of heat transfer or efficient use of fuel) is achieved with faster melt times, and the highest metal quality is obtained when a heel of metal is used to start the melting. Here, the heat travels easily from the flame through the crucible wall and into the metal. A crucible that contains only a solid charge provides little contact between the crucible wall and the metal to be melted. As a result, there is a considerable time lapse between charging and the initial melting of the charge. For efficient melting, a good heel size is roughly 25% of the crucible capacity. A second point concerning charging is the need to keep a loose charge. When a cold charge is tightly jammed into the crucible, the expansion of this charge upon heating will often break or crack the crucible. The third point concerns the need for gentle charging of all heavy or sharp-cornered material. Crucibles are fragile and easily broken. Many crucibles have been broken by rough and careless charging with heavy and/or sharp ingot or scrap. Handling of Crucibles During Melting and Pouring. It is important to keep in mind that hot crucibles filled with molten metal are fragile and therefore deserve careful handling. Care should be taken to set them on sand or other nonconductive surfaces. Placing a hot crucible full of molten metal on a steel plate or damp floor can easily break or seriously weaken the bottom of the crucible. Crucible Care. With proper care, crucibles will provide long service life. However, like every system used in the processing of molten metal, crucibles must be handled properly and
must be operated within their design parameters. All crucibles may be damaged by: Excessive temperatures: Crucibles must not
be heated beyond their specified maximum temperatures. Excessive temperatures will cause crucible failure. Thermal shock: Structural damage may be caused by too rapid heating or cooling of the crucible. Uneven heating or frequent thermal cycling between cold and hot also can result in crucible damage, especially with frequent metal-level changes. Physical shock: Damage may be caused to the crucible by the impact of charge materials and/or, in the case of removable crucibles, by impacts during movement. Damage to crucible glazes: Crucible glazes serve as barriers against oxygen attack of the carbon material making up the crucible. These glazes may be damaged by any of the previously noted hazards, and damage to the glaze will lead to crucible failure. Chemical attack: Use of inappropriate or excessive metallurgical additives may lead to corrosion of crucible components, particularly at higher temperatures. Fluxing salts that contain fluorides are especially harmful.
Crucibles must be thoroughly inspected for shipping damage when received. If cracked, chipped, or otherwise damaged, they must not be installed. Crucibles should be stored inside in a warm and dry environment and must be completely dry before use. All clay-graphite crucibles must be properly annealed before being subjected to the first full heat. This will help the crucibles withstand thermal shock throughout their service life. All crucibles should be preheated gradually whenever entering use from a cold, that is, room-temperature, condition, following the crucible manufacturer’s specific recommendations. During operations, the crucible should be inspected
for damage after each heat, and slag or dross buildup inside the crucible must be carefully removed, avoiding damage to the crucible wall. Damaged or worn crucibles must be replaced immediately. Selecting the right crucible for the application; installing it properly; operating it within its design parameters; protecting it from physical, chemical, and thermal damage; and replacing it immediately when damaged or worn are best practices. Following these rules will enhance service life, operational efficiencies, and significant economies in crucible furnace operations. Cleaning of crucibles is impor tant in preventing premature failure. In many, if not most, metal-melting operations, slag or dross buildups develop at the metal surface line, and heavy, unmelted slush residue collects on the bottom. Both of these residues shorten crucible life by causing hot and weak wall areas. The poor and uneven heat transfer around the crucible walls causes hot spots and uneven stresses within the crucible walls. The metallic residues that will expand faster than the crucible upon heating can easily rupture the wall. At times, these residues expand and fracture the crucibles during the normal heating and cooling cycles. Cleanliness and careful workmanship on the melt room floor should be encouraged. ACKNOWLEDGMENT Portions of this article were from Casting, Volume 15, Metals Handbook, 9th ed., 1988. REFERENCE 1. Advanced Melting Technologies: Energy Saving Concepts and Opportunities for the Metal Casting Industry, BCS, Inc., Nov 2005
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 160-169 DOI: 10.1361/asmhba0005350
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Reverberatory and Stack Furnaces David White, Schaefer Furance Martin Reeves, Striko-Dynarad Paul Campbell and David Neff, Metaullics
REVERBERATORY FURNACES have a shallow hearth for melting metal, which is heated indirectly by burners mounted in the roof or in a sidewall of the furnace (Fig. 1). The flame used for melting the metal does not impinge on the metal surface itself but is reradiated off the walls or the roof of the furnace. The term reverberatory furnace is a throwback to the World War II era, when the most common fuels for hearth aluminum melting were coal or coke (Fig. 2). This necessitated locating the heat source or fire box at one end of the bath and the flue at the other end. The flue (stack) was designed to be strong enough to draw the flame over the bath. The roof was slanted to bounce or reverberate the flame (similar to sound echoes) off the ceiling, down across the metal to be melted in the molten bath, and then onto the flue. A primary objective was to heat the refractory walls and ceiling and then use these surfaces to radiate heat to melt the charge. This same flame pattern and the method of using a flame to heat the bath were implemented when
the fuel used was changed from coke to oil and gas. This older concept of aluminum metal melting is now uncommon but may still be used in some foundries or smelters. The modern method is to use roof burners or side burners directed toward the metal surface. In effect, the objective is to heat the metal, not the furnace, for efficient operation. Gas- or electric-heated burners are used in reverberatory furnaces. Aluminum foundry reverberatory furnaces range from small electric or gas-fired vessels of just a few tons volume to upward of 40 to 50 tons. Secondary smelters that produce ingot or supply liquid metal to foundries operate even larger furnaces. Gas reverberatory furnaces are commonly used in large nonferrous casting foundries because they provide a large reservoir of molten metal, ensuring a steady and reliable supply of molten metal to the shop. Small gas or electric everbretatory furnaces (Fig. 3) are also used for melting. Units are also used for holding furnaces, although typically the term reverbretatory furnace is used for melting
units—not holding furnaces. Practical sizes range from 1 to 45 Mt (2000 to 100,000 lb). The advantages of reverberatory furnaces include high-volume processing rates and low maintenance costs. Many foundries and die casters have used reverberatory furnaces for many years. Large- and medium-sized shops use them as central melters to supply multiple operations located within the shop. By using double-end clean-out and preheat hearths, these furnaces are some of the easiest to clean and are much more efficient than just a few years ago. Advances in insulating back-up linings have made these furnaces cooler to work around and more energy efficient than ever before. Many of the newer furnaces can reach an energy efficiency capability of 1250 Btu/lb of metal melted. Typical capacity of the furnace can range widely, from 1 to 75,000 tons. The energy efficiency of reverberatory furnaces is in the range of approximately 20 to 30% (Ref 2). Energy is mainly lost through the hot flue gases. In addition, as the molten
Fig. 2
Fig. 1
Basic components of a reverberatory furnace. Source: Ref 1
Early version of a coal-fired melter hearth furnace used to produce aluminum castings. (a) Plan view. (b) Vertical cross section
Reverberatory and Stack Furnaces / 161
Fig. 5
Fig. 3
Aluminum reverberotory furnace (a) 180 kg (400 lb) per hour electric reverberotory furnace, (b) small gas reverberotory furnace, Courtesy of Schaefer Furnace
Table 1 Melt loss and themal efficiency of electric and gas reverberatory furnaces Melt
Aluminum Aluminum Zinc Zinc
Type
Melt loss, %
Thermal efficiency, %
Electric Gas Electric Gas
1–2 3–5 2–3 4–7
59–76(a) 30–45 59–76(a) 32–40
Major components of a stack melter
metal comes in contact with the furnace gases, it forms slag/dross. The melt loss rate is 3 to 5% in aluminum gas reverberatory operations (Table 1). Improvements in burner technology, fuel/air ratio control, insulating refractories, and temperature control have contributed to a slow but steady improvement in the efficiency of reverberatory furnaces. Efficiency is also obtained in furnaces with preheating hearths (Fig. 4). Fuel savings of approximately 15% per year are possible by allowing a sow or ingot to sweat on the hearth as the internal temperature is raised to approximately 480 C (900 F). There is no moisture left in the sow or ingot, thus making it safer to charge. The stored thermal energy (Btus) provides further energy savings in maintaining the melt.
(a) The primary energy efficiencies of these furnaces are much lower (approximately one-third) than the listed efficiencies due to the use of electricity, which involves sizeable energy losses during generation and transmission. Source: Ref 2
Furnace Types
Fig. 4
In most large facilities, the conventional means for melting and holding aluminum is a gas-fired reverberatory furnace. In this arrangement, radiation from the burners, roof, and sidewalls heats the bath and solid charge. Forced convection through stirring or pumping is also needed to overcome a number of process limitations. Without forced circulation, the concentration of heat at the bath surface readily produces undesirably high temperatures. The high heat flux generated by direct radiation at the bath surface is typically much greater than conductive and natural convective forces, which distribute the heat through the bath. In turn, surface oxidation and dross formation ensue. There are two types of gas reverberatory furnace, based on whether the sows, ingots, and revert materials are preheated before melting or not. A dry hearth reverberatory charges and
Preheat hearth
preheats sows, ingots, and revert materials on a slope between the charging door and the molten metal bath, removing any contamination or water residue prior to melting, which prevents splashing or explosions. A wet bath reverberatory directly dumps the sows, ingots, and reverts into the molten metal bath. The wet bath system is beneficial for melting scrap types with large surface-to-volume ratios, which may oxidize excessively unless melted quickly. Another type is a side-well reverberatory furnace that consists of a number of burners firing inside the hearth, against the furnace hot wall or door. A charging well and a pump well are attached to the furnace hot wall on the outside of the furnace. Both wells are connected to each other and with the furnace hearth by arches, which permit aluminum circulation between the furnace chambers. During the charging period of the furnace cycle, molten aluminum circulates through the charge well where it melts the scrap, then circulates back into the hearth for reheating. Charging continues until the molten bath reaches the desired level. Subsequent steps in the smelting process may include degassing and alloying. Improvements in energy efficiency have also been made with the use of stack furnaces (Fig. 5), which can be considered as a modified reverberatory furnace with improved efficiency by better sealing of the furnace and the use of the flue gases to preheat the charge materials. Stack furnaces are more common than reverberatory furnaces for melting aluminum in Europe. In the United States, 95% of aluminum is melted in gas reverberatory furnaces, which operate with lower energy efficiency but have lower capital cost and provide easier operation and maintenance than stack melters (Ref 2). Dry Hearth Reverberatory Furnace. In a dry hearth furnace, the charge of solid metal
162 / Melting and Remelting is positioned on a sloping hearth above the level of the molten metal so that the charge is completely enveloped by the hot gases. Heat is rapidly absorbed by the solid charge, which melts and subsequently drains from the sloping hearth into the wet holding basin or chamber. Dry hearth furnaces are typically used where castings and/or returns contain nonaluminum inserts or in sand casting operations where casting scrap containing sand must be reclaimed. These castings can be melted on the hearth, scraping the sand or insert off the hearth and keeping them out of the molten furnace bath. This keeps the furnace cleaner and the aluminum from being contaminated with iron (in the case of cast iron inserts). The advantages of the dry hearth furnace include:
the metal is transferred to a holding furnace(s) that feeds the casting line(s). These furnaces are designed to use radiant heat (a highly efficient method of transferring heat into aluminum), where the heat source (burners) is located 46 to 61 cm (18 to 24 in.) off the bath for best practices. The closer the heat source is to the aluminum, the faster the transfer of Btus. The temperature differential between the thermal head and the metal bath provides the driving force for heat transfer. The advantages of the wet bath furnace include: The up-front cost is lower than high-headroom
Safety: Everything melts on the hearth in the
atmosphere of the furnace and not under the bath. Such preheating avoids the potential for moisture entrapment and subsequent explosive force in and beneath the bath surface. The dry hearth allows the saving of inserts and avoids quality problems with undesirable impurities being incorporated into the melt. Of course, there are other potential sources of impurities that must be considered. The dry hearth holding chamber is typically smaller, approximately a 4 to 1 hold-to-melt ratio. The furnace holds four times what it melts or less, so the furnace is smaller. If the flue is maintained correctly, very good temperature control can be achieved in the hold zone. The disadvantages of the dry hearth furnace include: Typically, these furnaces have two flues: one
reverberatory furnaces and shaft melters. Efficiencies can be achieved on average of 1500 Btu/lb of metal melted. Everything is melted under the surface of the bath for greatly reduced metal melt loss (2 to 5% in gas-fired furnaces and less than 1% with electric furnaces). Many options are available: preheat hearths for sows and ingots, circulation wells for scrap, vortexing for chips, flue pressurization, well space for continuous degassing and filtration, and so on. With circulation and preheating, efficiency can be increased to 1250 Btu/lb of metal melted. By adding radiant roof regenerative burners, energy efficiencies can be further improved to 1000 Btu/lb of metal melted.
The disadvantages of the wet bath furnace include: These furnaces have a higher hold-to-metal
ratio and therefore are larger (higher capital cost). With a typical hold-to-melt ratio of 8 to 1, a minimum of eight times the solid charge weight must be maintained to help
dilute the cold metal charge and not cause excessive bath temperature reduction. Larger furnaces require extra floor space. Sows and ingots must be preheated before loading in the furnace, whether on the charge well wear plates (i.e., furnace sill) or on a preheat hearth. Electric wet bath reverberatories must be charged evenly or they will lose temperature and take time to recover. Larger electricpowered melting furnaces are extremely expensive and not as efficient as large gasfired furnaces. Low headroom requires care when cleaning. Sows or ingot bundles cannot be double stacked on hearths because of lower headroom.
Stack Melters. For aluminum metal casting, there are substantial opportunities for energy savings in replacing reverberatory furnaces with stack melters (also referred to as shaft melters). Stack melters improve efficiency by better sealing of the furnace and use of the flue gases to preheat the charge materials. The charge materials slide down the shaft and reach the melting zone, where they are melted by the burners, and the molten metal flows down to the holding area (Fig. 5). The hot exhaust gases from the melting zone flow through the shaft to preheat the incoming charge, improving the energy efficiency of the stack furnace by 40 to 50% (Ref 2). A typical example of a shaft melter furnace in a foundry is shown in Fig. 6. One of the central advantages of the stack melter is the physical separation of the charging, preheating, melting, and holding areas of the furnace. Advanced burner controls can augment the advantages of this kind of furnace. The stack furnace also can be coupled with
over the hearth and one over the hold chamber. Two separately controlled burner systems are required for optimal operation. High metal melt loss (possibly 7 to 12%) can result due to melting in the atmosphere, which creates a very high oxidation rate at these temperatures (700 to 760 C, or 1300 to 1400 F). When melting light-gage scrap and small foundry ingot, the metal melt loss is high and the thermal efficiency is low (1800 Btu/lb or higher). When melting larger sows (i.e., typically 540 kg, or 1200 lb), the efficiency is better and approaches the 1500 Btu/lb range. Metal melt loss is lower as the volume-to-surface area of the charge material increases. Wet Bath Reverberatory Furnaces. In a wet bath furnace, the products of combustion are in direct contact with the top of the molten bath, and heat transfer is achieved by a combination of convection and radiation. Wet bath reverberatory furnaces generally are used where there is just one alloy being melted or minimal alloy changes. Melting can be performed in a central furnace, sometimes called a break-down furnace, and then
Fig. 6
Stack melter employed in foundry practice. Courtesy of Striko-Dynarad Corporation
Reverberatory and Stack Furnaces / 163 other melting systems to improve efficiency. For example, it can be used with a crucible furnace, reducing its melt losses since the crucible melts by indirect heat to the metal through the wall of the furnace. The disadvantage of the stack furnace is that, due to the charging mechanism, the height of the stack furnace must be more than 6 m (20 ft) (Ref 3). In addition, as the furnace is charged from the top of the shaft, the refractories at the bottom of the charging door are subjected to repeated impact and wear. This requires additional maintenance. The shaft construction and melting design are at the heart of the advantages offered. The choked shaft design allows the consistent preheating of the charge material prior to melting. The hot exhaust gases from the holding bath pass through the melt zone and from there into the preheating zone before exhausting from the furnace. The additional energy required for melting comes from burners located in the melt zone and focused into the area at the base of the preheating area. Exhaust gases from these burners also exit through the shaft and therefore ensure a greater degree of preheating as well as greater speed and efficiency of melting. Energy Efficiency. Compared to other furnace designs, stack furnaces can result in substantial savings from lower fuel consumption and consequently lower melting costs. In the practical foundry operation, for example, with the use of 50% ingot and 50% bulky return material and a bath temperature of 720 C (1330 F), a power requirement is only 600 kW per hour per ton. This corresponds to a thermal efficiency of over 50%. Improved Melt Cleanliness, Quality, and Yield. By the separation of melting and preheating zones, the production of clean, high-quality metal can be achieved with stack melters. Moisture and other contaminants are evaporated or burnt off in the shaft by preheating with exhaust gases, ensuring that only dry material reaches the melting chamber. The primary source of contamination is moisture from storage of the scrap or ingots; the result of this is an increase in the hydrogen content of the metal, with its consequent quality problems in the cast parts. A further advantage is the avoidance of excessive contact between the melt and atmospheric oxygen, thereby reducing the tendency to form oxides. In order to minimize this danger, the burners in the melt chamber are arranged to avoid direct metal impingement. This allows melting on the hearth only and allows the metal to flow quickly but without turbulence into the holding bath without superheat. This ensures an optimal value for metal yield; typical values lie between 0.8 and 1.5%. A third measure for the maintenance of metal quality is the avoidance of contamination by impurities entering the metal holding bath. The separation of melting and holding ensures that contaminated charge material is not charged directly into the bath. In a crucible furnace or wet bath melter, what goes in stays in and must be removed by additional
metal-cleaning operations. Cleanliness is also assured by allowing the metal to run laterally into the holding bath. In this way, dross and other impurities remain on the melt hearth, to be removed during cleaning. An additional advantage from the separation of melting and holding is an almost continuous availability of molten metal, on demand, with a very consistent temperature (to +/5 C, or 9 F), which itself represents an important feature of quality. Melting continues while the holding bath level remains below maximum. The holding bath also offers the foundryman a source of molten metal at a consistent and correct temperature without fluctuations due to melting. Automated charging (Fig. 7) also contributes to this factor by ensuring that material is always available to melt when required. This is done using an automated lift-and-tilt system. Stack melters can also be combined with more conventional sow-melting technology to improve efficiencies and provide more options for charge materials. Various levels and forms of automation are available to cater to the different needs of the foundry and material supply. This can be by conveyor feed and automatic bin loading, allowing up to several hours, depending on melting rate and metal demand.
Reverberatory Furnace Practice Radiant Heating Factors. Radiant roof reverberatory furnaces rely on three major factors to make the process more efficient: Distance from the heat source (the burners)
to the load (the aluminum)
Hold-to-melt ratio Even charging of the load
One major design criterion for radiant roof melting is based on the Stephan-Boltzmann law of radiant heat, which states that the rate of heat transfer increases the closer an object is to a heat source and the greater the temperature differential. By using low-velocity, flat-flame, highthermal-release burners that heat the refractory block construction to 1400 C (2600 F), a huge temperature differential is obtained between the aluminum and the heat source. In this type of furnace, efficient heat transfer into the aluminum is principally by radiation. The greatest efficiency is obtained when the temperature differential is the greatest between the radiant source (the burner blocks, not the furnace upper chamber) and the receiving medium (the molten aluminum bath). Heat transfer is also more efficient when the radiant source is closer to its receiving medium. When the radiant heat sources are closer, more burners are needed to obtain coverage of the medium. Most reverberatory, radiantly heated melting furnaces will employ practices in aluminum melting wherein the metal charged is either all solid (ingot, foundry scrap, gates, runners, etc.) or a combination of solids plus liquid
Fig. 7
Charge for stack melter
(metal delivered in over-the-road ladles). Consequently, a portion of the charge will be cold (solid) metal. The hold-to-melt ratio is critical as a dilution factor for the cold metal being charged into the furnace. The higher that ratio, the less temperature differential is obtained when the furnace is evenly charged. Either radiant roof or highheadroom side-fired melters should be charged at ¼ of their hourly rated capacity every 15 min. This will ensure that an even temperature is maintained in the bath at +/10 during the melting process. Charging. Overcharging of these furnaces can create sludge in the furnace (if the metal temperature drops below the sludge point of the alloy), which will result in poor-quality metal for casting. Such overcharging and causing greater temperature swing in the melting and already molten bath can also result in additional dross formation at the melt surface, due to increased exposure time (i.e., the overall melt cycle) and the absolute temperature at the bath surface as it continues to absorb heat during recovery of the bath temperature. Typically, the radiant roof approach of melting everything under the surface of the bath greatly reduces metal melt loss (2 to 3% in gas furnaces and less than 1% in electric furnaces) and increases melt efficiencies for melting small, lightweight scrap. Recuperation. While there is little doubt that there is a great deal of savings possible in recouping heat lost out of a flue, some care is needed on how far this can be taken. For example, keeping the recuperation temperature below 370 C (700 F) can substantially reduce the system cost. Below 370 C (700 F), insulated mild steel manifolds can be employed as opposed to more costly stainless steel. Even at 370 C (700 F) preheated combustion air, a 25% savings in fuel usage is possible. An example of the cost trade-off is based on numbers produced by a major burner manufacturer at $10.00/106 Btus on an 1800 kg/h (4000 lb/h) radiant roof melter. One can expect to pay approximately $85,000.00 more for a burner system using an installed recuperator and save approximately $54,750.00 per year, or have a
164 / Melting and Remelting 17.4 month payback on this investment differential. Actual costs may vary due to the age of the furnace and components or the type of furnace and the number of burners. Other factors to consider in evaluating recuperation are: The flue gases should be dampered to be
able to bypass the exchanger when the furnaces are fluxed. Otherwise, excessive wear and maintenance on the heat exchanger will result. The nozzles in the burner should be changed to stainless steel to handle the additional temperature of the preheated combustion air. Older furnaces should consider changing to an ultralow-NOx burner. As the combustion air temperature increases, the NOx emissions go up. Regenerative burners enhance melting efficiency by providing a steady supply of hightemperature preheated combustion area. These burners contain dual firing ports, reversing firing as the air is drawn over heat-storage media in each port. In most cases, initial first-cost of recuperation provides a faster payback than does regeneration on smaller (2700 kg/h, or 6000 lb/h, or less) furnaces. This assumes a new furnace. For larger furnaces, the regenerative burners really begin to pay off. Payback varies depending on the type and age of the furnace in which regenerative burners are installed. Certainly, if the furnace is melting metal at a high end (1800 to 2000 Btus/lb of metal melted), then the return on investment becomes more lucrative. Regenerative burners become cost-efficient with high natural gas prices. As a further example, buying a 3200 kg/h (7000 lb/h) melter with regenerative burners to replace a less energyefficient high-headroom furnace that melts at 2000 Btu/lb of metal melted would result in a savings of over $452,000.00/year in fuel cost. For a melter of this size, the return on investment payback can be two years or less. (These figures are based on 3200 kg/h, or 7000 lb/h, for 20 h of melting/day, 340 days/year, and $10.00/1000 ft3 of gas.) Reducing Fixed-Heat Losses. Highly insulated furnaces with substantial linings help provide for the lowest fixed-heat loss. Fixed-heat loss is the amount of energy being absorbed into and passing through the lining of the furnace and out the doors and open wells on the furnace. An example of fixed-heat loss is: If the sidewall casing temperature of a given
furnace is 96 C (205 F), then 300 Btus/ft2/h are lost from that casing. By simply reducing the temperature to 64 C (148 F), 152 Btus/ft2/h are saved. The older refractory configuration concept taught that the freeze plane (the point where the molten metal would solidify if it leaked through the lining) had to be in the hot face or first layer of refractory material. With the
availability of newer, nonwetting insulating materials, the freeze plane can now be designed to be further back in the thickness of the lining. This gives much more flexibility to highly engineer this lining to not only contain the metal but actually save energy. Melting furnace casing temperatures are being designed to 50 to 60 C (125 to 135 F) instead of the old 75 to 85 C (165 to 185 F) of just a few years ago. New superinsulating materials can be used in the roof and floor of the furnaces to further reduce energy consumption. Some furnaces need to have individual burner controls, just to keep the furnaces from overheating over long weekends, because of this thermos-bottle approach to highly insulated linings. The fixed-heat loss has been reduced to the point that the burners on the furnace at low fire will overcome the heat losses and actually raise the temperature if they are not turned off. Covering up open wells will reduce heat loss substantially. An open well of 700 C (1300 F) aluminum with a 0.5 emissivity of the bath surface in the well will lose 8214 Btus/ft2/h. If the average charge well is 2.4 m long by 1 m wide (8 ft long by 3 ft wide) (2.2 m2, or 24 ft2), then 197,136 Btus/h are being lost off of that well. That corresponds to 5.6 m3 (197 ft3 of gas per hour, or 134 m3 (4731 ft3) a day by 340 days a year is 45,560 m3 (1,608,540 ft3) of gas lost. That is $16,085.40 worth of natural gas at $10.00/therm lost a year. Improving and Maintaining Furnace Efficiencies. Burners are not stoichiometric. The combustion system must be set up properly, or more dross will be generated. Typically, most natural gas-fired burners are set up on a 10 to 1 air-to-gas ratio. At this ratio, 95% of the oxygen in the burner system is consumed. However, if a furnace is set up with ratios in excess of this, dross increases in the furnace. On the other hand, a furnace that is set up fuel-rich will produce a colder flame and make the furnace work harder to get the job done. Keep the burner in ratio. A qualified technician should check the combustion system for the proper set-up. The burner system should be firing stoichiometric, or in ratio, air to gas. Typically in reverberatory furnaces, this is a 10 to 1 air-to-gas mixture. By analyzing the flue gases, a technician can determine if indeed the system is running at peak efficiency. This can also be done with a manometer to test pressure (inches of water column) of gas and air, if the technician knows the burner manufacturer’s recommended set-up pressures. Fuel-rich burners produce a colder flame, and some of the gas will ignite when it reaches the atmosphere at the flue level. This results in a wall of flame coming out of the flue. If the burner is set up too lean or has excess air, a colder flame results, together with an increase in oxide formation and dross in the furnace. This may cause the burners to work harder to produce the same Btus, resulting in lower energy efficiency and therefore higher costs of production.
In addition to burner ratio, improving and maintaining furnace efficiencies can be achieved in several ways, as summarized subsequently. Remove all fans that blow on the furnaces. This will result in less heat loss from the casing where the fan is hitting the furnace. Instead, use individual cooling stations that bring the air in from overhead to a specific spot where the operator can stand to cool off. Install well covers (Fig. 8) whenever the furnace is idle for longer than a half-hour. Install a half-flue cover over the flue during long idle conditions. Warning: The flue cover must be removed before turning the furnace to high fire. At low fire, the flue opening is still sized for 100% output of the burners. Without automatic flue pressurization, the holding efficiencies will drop considerably. The furnace may not need to be on high fire as much. Charging. Make sure the furnace is used at full capacity and charged ¼ of its hourly rated capacity every 15 min. Overcharging the furnaces causes dramatic swings in temperature and increased sludge buildup on the floor of the furnace. The more sludge on the floor, the less likely the furnace will melt at its rated capacity. The furnace may not melt anywhere close to the Btus/lb of metal melted that it was designed to do. Reduce molten metal temperature over weekends by 11 C (20 F) or more if feasible. Clean furnaces during idle shifts or idle times, but clean them daily! This will cut down on the amount of oxides growing in the furnace. Oxide is dense and absorbs additional Btus from the metal. Do not overflux the melters. Overfluxing of the walls can break down the binder in the refractory and cause premature erosion of the belly-band area. Clean or change combustion air filters regularly. Hold a dollar bill 50 mm (2 in.) away from the filter. If the bill is not drawn to the filter rapidly, it is time to clean or change the filter on the blower. Double-check and monitor combustion set-up for maximum burner performance (requires a manometer). If a wall of flame is coming out of the flue, a fuel-rich condition probably exists inside the furnace. The raw gas that does not have enough air in the furnace for combustion
Fig. 8
Pneumatic well cover
Reverberatory and Stack Furnaces / 165 ignites when it reaches the atmosphere just outside the flue, because of all the free air available to it at the flue and hood junction. Overpressurization of the furnace can also cause this condition. Preheat Ingots. In reverberatory (radiant roof low-headroom or high-headroom) furnaces, preheat all ingots on a charge well (wear plates) or on a hearth before putting them into the bath. Preheat all sows on a hearth before pushing them into the bath. When a sow or ingot begins to sweat, the internal temperature is close to 480 C (900 F). There is no moisture left in the aluminum, and it can be safely charged under the bath of aluminum, where the stored Btus in the metal help finish the melting process. Ingots left on a well wear plate to preheat will reach 120 C (250 F) after 20 min. This is sufficient to load into the bath. Some bubbling may occur from moisture left in the very center of the ingot. Metal delivery to the holding furnaces should be every 20 to 30 min; less frequently will mean overcharging the melters and drawing the metal down too low in the holders. The closer the metal is to the heat source, the faster the transfer of Btus. A loss of 10% in efficiencies could take place by drawing the holders that low. Ensuring that the metal tenders put back what they take out in the melters results in a difference in metal temperature stability. Remove dross buildup whenever it becomes thicker than 13 mm (½ in.), or the dross will begin to act like an insulator and make the furnace work harder to do the same job. A completely clean bath surface will reflect the heat, so it should not be cleaned any more frequently than once per shift. If a furnace creates a dross layer over 13 mm (½ in.) in less than 6 to 8 h, then the furnace is creating more dross than normal. Electric Pump Motors. To further conserve energy, the pump should have an electric motor rather than an air motor. Plant air is the most expensive energy source in the plant. A small blower keeps the pump cool. Treat the cause of dross buildup, not the symptom. Investigate why the furnace is generating dross. Typical reasons are: Air leaking in and around the doors Drawing the furnace down below
the submerged arches or door blades Burners that have a flame that touches the bath or scrubs the solid metal on a hearth Too large of a flue opening for the Btus connected, causing the furnace to draw in more air Keep the burn in ratio!
Molten Metal Circulation In the melting of aluminum and its alloys, forced convectional heat transfer within the melt often plays an important role. There is a considerable latent heat of fusion (95 cal/g) associated with taking aluminum at its melting
point from a solid to a liquid. Forced convection assists this transition, particularly for larger volumes of metal, by better mixing of the solid metal with already molten metal, thus reducing time and energy input into the melting process. Other nonferrous alloys, such as zinc, magnesium, and copper, could also benefit from molten metal circulation, but pumping systems for circulation are not usually employed. Mechanical stirrers are effective, particularly for zinc, due in part to the lower heat of fusion of zinc and the smaller melt sizes usually used. In reverberatory furnaces, forced convection enables the radiant heat energy absorbed by the bath surface to move, that is, to be transmitted throughout the bath depth, so that a steady heat flux is developed and continuing heat transfer to the surface occurs by radiation. Without such forced convection, the surface layer of the molten metal becomes heat saturated, and the surface temperature rises. Ultimately, excessive surface temperature will result in oxidation and dross buildup, which reduces energy efficiency. Forced convection is produced by the circulation of the molten metal bath. Current methods to achieve circulation or metal movement include manual stirring, mechanical pumps, electromagnetic pumps, induction stirrers, porous plugs, jet pumps, and towmotor/boom agitation (see the section “Molten Aluminum Circulation Methods” in this article). Manual stirring is only practical in smaller furnaces (in small pot furnaces, crucibles, and so on). Forced circulation is usually best achieved in large furnaces (2000 to 110,000 kg, or 5000 to 250,000 lb, capacity) by using a submerged discharge molten metal pump. Several types of molten metals pumps have appeared over the years (such as electromagnetic and pressure pumps), but the centrifugal pump has been a common type. These pumps are usually installed in the pump wells and discharge into the charge well of reverberatory furnaces for remelting in large aluminum foundries, die casting operations, secondary smelters producing foundry ingot, and in mill product casting houses. Typical installations are illustrated in Fig. 9 (a and b). A summary of the benefits of forced circulation in large furnaces employed in aluminum foundry practice, as well as a comparison of the different methods of circulation, is in the following sections.
much greater than conductive and natural convective forces, which distribute the heat through the bath. In turn, surface oxidation and dross formation ensue. Figure 10 (Ref 4) provides the oxidation rate of molten aluminum as a function of temperature. Because aluminum oxide is thermally insulative, the variation between surface and bath temperatures increases with the growth of the oxide layer. Temperature differentials in a 915 mm (3 ft) deep bath with a clean surface and without forced convection can typically reach 50 to 85 C (90 to 150 F). Significant melt loss is incurred due to ever-increasing dross formation under these conditions. Without effective heat transfer to the metal, other quality and furnace performance factors may be less than optimal.
Fig. 9
Molten metal circulation pumps. (a) Typical pump installation in a sidewell furnace. (b) Circulation well and pump
Benefits of Forced Circulation The benefits of circulating molten metal in reverberatory furnaces are well documented, especially for aluminum. Forced convection through stirring or pumping overcomes a number of process limitations of reverberatory furnaces for melting and holding. Without forced circulation, the concentration of heat at the bath surface readily produces undesirably high temperatures. The high heat flux generated by direct radiation at the bath surface is typically
Fig. 10
Oxidation of aluminum versus temperature
166 / Melting and Remelting The result of forced convection is to reduce metal temperature stratification in the bath, improve melting rates, enhance metal quality through alloy homogeneity, reduce energy consumption as a result of improved heat transfer, extend refractory life, reduce sludge (dross) formation and therefore lower furnace-tending requirements, and accelerate alloy dissolution rates. Operational benefits of more homogeneous bath temperature of the entire bath of metal (top to bottom, inside and out in the well) include: Quicker melting with reduction of melt
cycle times by 10 to 50% Reduce energy consumption by 15 to 25% Extend refractory life at least 25% Improve recovery rates by a minimum of 0.25% With these major factors at stake, it is no wonder that a wide variety of techniques have been developed to provide forced circulation. While many production parameters may be difficult to fully quantify, forced circulation clearly improves many key aspects that affect total quality. These benefits of circulation are influenced by a variety of production/operation factors, as briefly described here. The next section, “Molten Aluminum Circulation Methods,” in this article also summarizes the performance of currently applied circulation technologies in large-furnace foundry practice. The review covers six methods that have evolved to some degree of acceptance: mechanical pumps, electromagnetic pumps, induction stirrers, Alcan jet stirrer, porous plugs, and towmotor/boom. With each technique, the production manager must assess the accompanying benefits, characteristics, and drawbacks. Reduced Temperature Stratification. Bath temperature is a key factor that impacts overall product quality. Forced circulation improves bath product quality by ensuring the bath temperature will be consistent throughout the furnace. Typically, reductions in temperature differentials for a 915 mm (3 ft) bath are from 50 to 85 C (90 to 150 F) in an uncirculated furnace to 3 to 7 C (5 to 13 F) for stirred furnaces. Circulation ensures that the temperature of the metal exiting the furnace is virtually the same as that indicated by the furnace thermocouple. Production parameters can be set and more easily controlled when the metal delivered is at a constant temperature. Additionally, there will be less superheat of the metal required because there will be less variance in the furnace temperature. Increased Melt Rate. Constant circulation of the furnace will increase the melt rate of the furnace as well as improve the thermal profile. The high-velocity stream from a circulation pump will increase the melt rate of large submerged solids in the furnace. Insulating boundary layers that are formed around large solids as they melt are swept away by the hot metal stream. Such a device equipped with gas injection can provide bath fluxing and degassing during the melting cycle, eliminating
or greatly reducing the time required at the end of the cycle for metal treatment. The same holds true for temperature control. Continuous circulation provides excellent temperature control, so temperature adjustments are usually not required prior to transfer. Cunard, using a direct-charge furnace in a billet casting process, reported in a study over a five month time period that furnace cycle times increased between 8.1 and 20.6% when the pump was not in operation (Ref 5). He estimated that using a pump produced a conservative value of 10% cycle-time savings. Charging time reductions of 25 to 50% have often been observed in sidewell furnaces when a pump is used. Savings may be obtained by increasing output with existing equipment or by maintaining production levels with fewer furnaces. Manufacturers of foundry reverberatory furnaces estimate that the melting capacity of a given furnace will increase by 10 to 15% when forced circulation is added to an uncirculated furnace. Alloy Homogeneity. One method for stirring a furnace was developed by Alcan during the 1980s. They attached a large pressure vessel to the side of the furnace. The pressure vessel communicates with the furnace through an opening below the metal surface. When a vacuum is applied to the vessel, metal is drawn in from the furnace through the opening. The vessel is then pressurized, forcing metal back into the furnace through the opening. The vacuum is reapplied to the vessel, and the cycle repeats. The results of an Alcan jet stirrer (AJS) installed in a top-charge 70 ton round melter have been documented (Ref 6). Figure 11 schematically represents the main elements of the AJS, while Fig. 12 shows a typical adaptation to a top-charge round melter. The results were obtained during the preparation of an AA4000series alloy with a high silicon content. The silicon specification for this alloy is 7.15 to 7.45%. The silicon content was monitored in the trough during the cast and compared with samples taken from the furnace at the end of the batch casting process. A significant difference was observed between the two samples for the same batch. For eight runs without stirring, the average silicon composition in the furnace at the end of the cast was 7.28%. In contrast, the average for trough samples during the cast was 7.22%, with a standard deviation of 0.08%. In fact, two ingot batches were outside of specification. In five runs where stirring was used, no significant difference between furnace and trough samples was observed. All ingots were within specification, and the standard deviation for the cast samples dropped to 0.0%. In foundry applications, similar results can be expected when the furnace is circulated. Without forced circulation, alloy additions must be stirred in mechanically to ensure uniformity. Since mechanical stirring is done over a relatively short period of time, it is difficult to ensure that the alloy is mixed uniformly
Fig. 11
Schematic of the main elements of the Alcan jet stirrer
Fig. 12
Typical Alcan jet stirrer installation in a top-charge melter
throughout the furnace. Some additions will also begin to sink to the bottom or float to the surface if not continually stirred. When a stirring device such as a mechanical pump is used, alloy elements are constantly mixed with the rest of the metal in the furnace. In a typical foundry application, the volume of metal in the furnace passes through the pump every 15 min. Since the pump is operated 24 h per day, samples taken from any point in the furnace will show uniform chemistry. One other consideration with foundry alloys is the formation of sludge. This is an Fe-CrMn compound that precipitates out of the metal if the temperature drops below a critical point. This is discussed in detail by Jorstad (Ref 7). With a temperature differential of 50 to 85 C (90 to 150 F) between the top and bottom of the bath, it is easy to see how the metal at the bottom of the bath can cool below the critical point if the furnace temperature is monitored and controlled by a thermocouple located near the top of the bath. As previously discussed, temperature is uniform throughout the bath when the furnace is constantly stirred. This allows for accurate temperature control and the avoidance of sludge formation. Reduced Energy Consumption. The rate at which energy is transferred from the burners
Reverberatory and Stack Furnaces / 167 and refractories to the metal is referred to as the heat flux (q). Heat flux in a reverberatory furnace is determined by the following equation (Ref 8): q ¼ Ah½ðt1 Þ4 ðt2 Þ4 where A is the surface area of the bath inside the furnace, h is the heat-transfer coefficient of the area inside the furnace above the bath, t1 is the temperature of the heat source or refractory walls and ceiling, and t2 is the temperature of the bath surface. From this equation, it can be seen that the rate at which energy is transferred into the metal is dependent on the difference between the surface temperature of the metal and the temperature of the refractories, to the fourth power. Constantly stirring the furnace allows heat to be removed from the surface to the rest of the bath by convection. As heat flows more efficiently away from the surface of the bath, the bath surface temperature is reduced. As the surface temperature drops, the temperature differential between the refractories and the metal increases, which translates into much higher heat-transfer rates. This subject is covered in more detail by Campbell (Ref 9). Since stirring occurs during the melting process and with the doors closed, energy that would have been lost out the door when the bath is stirred manually is saved. Additional savings are possible from the elimination of throughthe-door fluxing operations if the circulation pump is equipped with gas injection capability. Actual furnace efficiency improvements vary from operation to operation. Experience has shown that a minimum of a 15% improvement can be expected. A 25 to 30% improvement is possible, depending on furnace and operationspecific variables. In the study done at VAW in Ellenville, New York (Ref 5), savings ranged from 10.3 to 20.2% over a five-month period. Longer Refractory Life. By reducing the metal surface temperature, more heat is transferred to the metal and less is absorbed by the refractories. Roof and stack temperatures are usually reduced by 80 to 115 C (145 to 205 F). These lower operating temperatures translate into longer refractory life. Lower temperatures will also result in less corundum formation at the metal line. Less buildup means less need for mechanical cleaning and therefore less opportunity for damage from cleaning tools. While specific information on savings from improved refractory life is not widely shared, dramatic improvements have been reported by some companies using circulation pumps (Ref 10). Decreased Dross Formation. Another aspect that should not be ignored is melt loss. Oxidation increases with temperature (Fig. 10); at temperatures above 775 C (1425 F), the rate at which aluminum oxidizes increases quickly. Reducing the bath surface temperature can significantly reduce the rate of dross formation. Minimizing surface agitation with submerged circulation removes another source of dross formation.
Many foundries do not pay enough attention to melt loss that results from dross formation. Every pound of aluminum that is lost to oxidation in the form of dross is lost to the manufacturing process and must be replaced. While not widely documented, the typical dross reduction in the hearth when using a centrifugal pump for forced circulation is approximately 0.5%. Using a conservative value of 0.25%, a foundry that melts and casts 454 MT (1 million 1b) of aluminum a month would save the replacement cost for 1.1 MT (2500 1b) of metal every month. In another study, a large die casting company augmented its production facility with centrifugal pumps (Ref 11). They estimated a 98% recovery of charged material during the first six months of pump use versus a 96% recovery prior to its use. Easier Alloy Adjustments. The benefits associated with alloy additions can manifest themselves in three areas: alloy addition dissolution, treatment during melting, and improved melt quality before downstream treatments. By continually providing forced circulation, alloy additions and adjustments are easier to control, since the bath will remain homogeneous. Silicon and other alloy additions go into solution much faster when exposed to the discharge stream of a pump. When heavy chlorination is required due to scrap charge composition and desired alloy chemistry, a gas injection pump is an extremely effective tool (Ref 12). The dispersion and subsequent reaction of chlorine gas as it exits the pump in the molten metal stream creates a very effective kinetic environment for removal of magnesium from the melt. Figure 13 provides a typical representation of demagging results in the production of foundry ingot in a 68 MT (150,000 lb) furnace with a 34 MT (75,000 lb) heel from the previous heat. The magnesium content was decreased consistently and effectively during the normal charge cycle.
Fig. 13
Demagging results using a gas injection pump
Results from tests conducted at Golden Aluminum in 1995 show that with a mixture of 95% Ar and 5% Cl, a gas injection pump can significantly reduce sodium and calcium levels without affecting magnesium content (Ref 13). The tests also showed an improvement in porous disc filtration analysis results and a reduction in gas levels while the gas injection pump was in operation. By removing inclusions and hydrogen in the melting furnace, the load is reduced on downstream operations such as degassers and filters.
Molten Aluminum Circulation Methods As noted, methods of molten metal circulation include manual stirring, mechanical pumps, electromagnetic pumps, induction stirrers, porous plugs, jet pumps, and towmotor/ boom agitation. These technologies have evolved to facilitate efficient and consistent molten aluminum processing in melting and holding furnaces as the commercial importance of aluminum continues to grow. The features, limitations, and costs of each device must be considered with the requirements of the operation to determine the best solution. While all of these devices impart motion to the molten aluminum bath, the effectiveness of heat and mass transfer, capital investment, running expense, maintenance issues, ease of installation, and capabilities for alloy modification vary significantly. One of the most important factors in today’s (2008) competitive environment is product quality. While the economic benefits of quality can be difficult to quantify for a specific operation, the benefits of improved productivity, reduced cycle time, and better thermodynamics can be readily translated into economic savings. Many papers have been published on the application and performance of mechanical circulation and gas injection pumps, as provided
168 / Melting and Remelting in the references. These are typically centrifugal pumps with a rotating impeller in a volute casing. Materials of construction consist of graphite and/or ceramics so that they can survive in contact with the molten metal. They require an external sidewell (Fig. 9a) on the furnace. Mechanical pumps offer the most flexibility, low capital cost, and moderate maintenance requirements. On the other hand, they are not suitable for every application. The operating principle of an induction stirrer and an electromagnetic pump is the same as that of a linear motor (Ref 14). An arrangement of a bottom-mounted induction stirrer with a furnace is provided in Fig. 14. The melt is stirred by the force created by the inductor attached to the bottom of the furnace. A variation of this mounts the stirrer on the side of the furnace. Electromagnetic stirrers typically have the highest capital and energy costs, but with no metal contact, they tend to be the most reliable of all the systems considered here. In contrast, the electromagnetic pump moves molten metal through a tube within a coil by the generation of an electromagnetic force created by an applied magnetic field and induced eddy currents (Ref 15). Figure 15 depicts the plan view of an electromagnetic pump in a furnace with charge and dross wells. Electromagnetic pumps offer many of the advantages of a mechanical pump, generally require less maintenance, and can be used in applications not suitable for mechanical pumps. Capital costs
Fig. 14
Fig. 15
Schematic side view of an induction stirrer
Plan view of an electromagnetic pump, charging well, drossing well, and furnace
and energy costs are higher than mechanical pumps. The operating principle of the AJS is based on the repetition of two phases: a vacuum/accumulation phase followed by the pressurization/ expulsion phase. Metal is drawn into the unit through an opening with the furnace during the vacuum phase, then forced out through the same opening during the pressurization phase. For several years, Reynolds Metals used porous plugs in melting and holding furnaces at the alloys plant in Muscle Shoals, Alabama (Ref 16). This technology uses porous refractory plugs that are permanently installed in the furnace floor. Gas, usually nitrogen or argon, passes through porous refractory and produces a plume of fine bubbles over each plug. This gas plume creates localized circulation over each plug. Finally, a towmotor/boom system is used in some foundries and is considered to provide circulation within the furnace hearth. With this method, the door is opened, and the towmotor operator works a boom-and-rake arrangement back and forth to stir and skim the molten melt. Maintenance Comparisons. As a general rule, mechanical pumps will require more maintenance than electromagnetic pumps and stirrers. Mechanical pumps have seen significant improvements in designs and materials of construction in recent years, and reliability has improved to a point where at least 1 manufacturer now offers a 1 year warranty on the major components. Electromagnetic pumps and stirrers have no moving parts and therefore require little maintenance (Ref 17). This is somewhat offset by higher capital and energy costs. Porous plugs require no maintenance. Life and reliability are determined by installation and operation techniques. Major refractory repairs are required when they need to be replaced. Flow Rate/Effectiveness Comparison. As described in the first section of this article, there are many benefits of strong, forced circulation. Mechanical pumps, electromagnetic pumps, and induction stirrers receive high marks for delivering these benefits to the production process. Porous plugs, which are probably better categorized as a treatment technique rather than forced circulation, rate as moderate in overall performance. The towmotor/boom system is a “get-what-you-pay-for” arrangement. Not only is the circulation poor and intermittent, there are quite a number of negative factors that are associated with this technique. Since the furnace doors must be opened for access, there is a high heat loss. With this type of thermal cycling and exposure, the refractory life is significantly reduced. While this method is inexpensive from a capital standpoint, the increased dross formation and associated metal losses can be a large operating loss. Finally, safety is an issue. Certainly, the towmotor operator is at risk in this environment; it is often a topic of lively discussion at aluminum
meetings. Additionally, the back-and-forth action that is required presents a hazard in an aisle that is common to both plant personnel and equipment traffic. Alloy Modification. In today’s (2008) demanding production environment, concurrent processes are often added to classical charging, melting, and homogenization schemes to reduce overall production time. These processes include chlorine injection to remove magnesium from the melt, gas injection to remove hydrogen and increase inclusion flotation, and solids addition for alloy modification. A mechanical pump structured as a gas injection model has documented performance for all three key needs. An electromagnetic pump can be used for alloy modifications by the addition of solid materials, but it has not been shown to be effective for gas injection. In contrast, the induction stirrer does not provide any of these benefits as it is currently configured. Porous plugs are excellent for hydrogen and inclusion reduction. They are not employed for demagging, and their effectiveness for alloy additions is not documented. Towmotor/booms are often used to stir in alloy additions, but they cannot effectively improve operations over any of the other stirring methods. Ease of Installation/Retrofit. The application of a mechanical pump requires an open well and a residual molten metal heel during normal furnace cycles. A floor-mounted induction stirrer system is usually only considered for new construction, since the installation requires a basement area to house the inductor, and the furnace floor must be stainless steel. Side-mounted induction stirrers can be added to a furnace fairly easily. Electromagnetic pumps can be adapted to most furnaces. The furnace refractory and steel shell must be modified so that the equipment can be bolted to the side of the furnace. In order to install porous plugs, a major furnace reline is required. The towmotor with boom may be used in all instances. REFERENCES 1. “Refractories for Industrial Processing: Energy Reduction Opportunities,” Oak Ridge National Laboratory, Dec 2004 2. Advanced Melting Technologies: Energy Saving Concepts and Opportunities for the Metal Casting Industry, BCS, Inc., Nov 2005 3. J.F. Schifo and J.T. Radia, Theoretical/Best Practice Energy Use in Metalcasting Operations, U.S. Department of Energy Industrial Technologies Program and Keramida Environmental, Inc., May 2004, p 36 4. D.V. Neff, Molten Aluminum in Motion, Light, Metals, TMS, 1993, p 197–211 5. T. Cunard, Application of Graphite Pump to Direct Charged Aluminum Melting Furnace, Aluminum Industry Energy Conservation Workshop X, Oct 20–21, 1987, Aluminum Association, p 159–185
Reverberatory and Stack Furnaces / 169 6. M.A. Thibault, F. Tremblay, and J.C. Pomerleau, Molten Metal Stirring: The Alcan Jet Stirrer, Light Metals, TMS, 1991, p 1005–1011 7. J. Jorstad, Understanding Sludge, Die Casting Engineer, NADCA, 1986 8. M.C. Mangalick, Reduced Fuel Consumption Through Mechanized Molten Metal Circulation, Soc. Die Cast. Eng., Vol 17 (No. 5), Sept–Oct 1973, p 22–31 9. P.S. Campbell, Use of Molten Metal Pumps to Reduce Energy Consumption in Reverberatory Furnaces, Aluminum Industry Energy Conservation Workshop XI, Nov 1–2, 1990, Aluminum Association, p 537–546
10. T. Thornton, C. Hammond, J. van Linden, P. Campbell, and C. Vild, “Improved UBC Melting Through Advanced Processing,” 136th Annual Meeting and Exhibition, TMS, 2007 11. L-35 gas injection customer, private communication 12. D.V. Neff, Use of Gas Injection Pumps in Secondary Aluminum Metal Refining, Recycle and Secondary Recovery of Metals, Dec 1–4, 1985, Metallurgical Society of AIME, p 73–95 13. R.S. Henderson et al., “Quantification of Molten Metal Quality Improvements Using an L-Series Gas Injection Pump,” Third Int. Symposium on Recycling of Metals
14. 15.
16.
17.
and Engineered Materials, Nov 12–16, 1995, TMS C. Clark et al., A Molten Metal Pump Installation at New Zealand Aluminum Smelters Limited, TMS, 1988 R.M. Jones, A New Electromagnetic Metal Pumping and Stirring Technology, Fourth Australian-Asian-Pacific Course and Conference, TMS, 1995, p 275–289 J.R. Guttery and W. Evans, A New Generation of Fluxing in Aluminum Melting and Holding Furnaces, Light Metals, TMS, 1994, p 921–927 A. Katyal, Metal Stirring Using an Electromagnetic Pump, Optimization ’95 Conference, Proc. BCIRA, Vol 8, 1995, p 1–20
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 173-184 DOI: 10.1361/asmhba0005352
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Transfer and Treatment of Molten Metal—An Introduction MOLTEN METALS are processed and handled in various ways, depending on the alloy and melting practice. The very nature of melting is a way of combining elements for the purpose of alloying based on the laws of physical chemistry (see the article “Chemical Thermodynamics and Kinetics” in this Volume). However, the reactivity of metals and the high temperatures of melting also introduce the deleterious effects of dissolved gases and inclusions in molten metals, as described in the articles “Gases in Metals” and “Inclusion-Forming Reactions” in this Volume. Molten-Metal Treatments. Of primary importance during melting and the handling of molten metal is the mitigation and reduction of dissolved gases and inclusions in the melt. This objective may be accomplished in different ways, depending on the base metal, alloying constituents, and residual impurities. Fluxing of molten metal is often the first step in obtaining a clean melt. Fluxing is a term commonly used in foundries and melting departments that refers to molten-metal treatments from the addition of reactive chemical compounds or inert gases. The various functions of molten-metal fluxing additives, either singly or in combination, include:
To render the high-melting refractory oxide
impurities or gangue fusible and separable from the molten metal To provide a medium with which the impurity elements or compounds would combine in preference to the metal In steelmaking, the primary basic fluxes are either limestone (CaCO3) or dolomite ((Ca, Mg)CO3), depending on the other slag constituents present and the required ability of the slag to remove sulfur. Limestone is preferred for large sulfur removal. The slag-forming compounds in steel melting can be classified as either basic or acidic (see the article “Electric Arc Furnace Melting” in this Volume). Likewise, secondary refinement of steel melts may include various fluxing techniques in converter and ladle metallurgy (see the article “Steel Melt Processing” in this Volume). Fluxing is also commonly done to some extent with virtually all nonferrous mill and foundry melting operations. In particular, fluxing is used extensively in controlling dross formation and removing oxides from aluminum melts and melting furances. The basic types of aluminum fluxes are: Cover fluxes designed to barrier “blanket”
on the surface of molten aluminum
the article “Casting of Copper and Copper Alloys” in this Volume) Fluxing of zinc alloys with chloride-containing fluxes that form fluid slag covers, which can be used to minimize melt loss if they are carefully skimmed from the melt before pouring (see the article “Casting of Zinc Alloys” in this Volume) Degassing. Removal and reduction of dissolved gases in molten metal is typically accomplished with injection of inert gases in the melt (see the article “Degassing” in this Volume). In addition to dissolved gases, other sources of gas porosity include various mold-metal reactions (see the article “Gases in Metals” in this Volume). Molten-Metal Filtration. In addition to fluxing, inclusions in molten metal can also be controlled with various types of filters, such as screens or foam. Filters are commonly used in both aluminum melting furnaces and in-mold systems (see the article “Molten-Metal Filtration” in this Volume). Filters are also used in steel foundries. Molten-metal handling systems are also important components in the effective transfer, holding, and pouring of melts. After the melting process, molten metal may be handled and processed in various systems that may include:
Removing inclusions, impurities, or dissolved
Drossing fluxes with wetting action that
Launders, which are channels for transport-
gases from the melt Preventing excessive formation of solidphase inclusions Preventing and/or removing oxide buildup on furnace walls Injecting or distributing alloying constituents
promotes coalescence of aluminum Melt-cleaning fluxes designed to remove aluminum oxides from the melt Wall-cleaning fluxes specifically designed for the softening and removal of excessive aluminum oxide buildup that occurs on melting furnace walls
Thus, fluxes are commonly used to some extent with virtually all molten-metal operations in both the foundry and in the production of mill products. Fluxing can occur throughout the various stages of melting and handling of molten metal—from the melting or holding furnaces to the transfer ladles. Molten-metal fluxing agents may be chemically active solids, inert gases, or a chemically active gas. The type of flux depends, of course, on the composition of the melt and refractories. In iron and steelmaking, for example, the function of fluxes is twofold:
These fluxing treatments of molten aluminum are described in more detail in the article “Aluminum Fluxes and Fluxing Practice” in this Volume. Fluxing of other nonferrous alloys includes:
ing molten metal from melting units to tundishes, ladles, or molds Tundishes, which are intermediate holding vessels that have the ability to regulate molten-metal flow using slide gates, sliding nozzles, or a stopper Holding furnaces or transport crucibles that maintain a melt (or superheat) over a period of time Molten-metal transfer pumps Teeming ladles Dosing furnaces for delivery of controlled quantities (doses) of molten metal to the mold or die
Fluxing of magnesium alloys using salt
fluxes or inert gas as a cover (see the article “Magnesium and Magnesium Alloy Castings” in this Volume). Fluxing of copper alloys to remove gas or prevent its absorption into the melt, to reduce metal loss, and to remove specific impurities and nonmetallic inclusions (see
This article describes these typical systems for transporting, holding, or delivering molten metal to the mold/die system. Specific moltenmetal treatments and refractories are covered in other articles, as noted in the preceding overview on fluxing.
174 / Molten Metal Processing and Handling
Launders Launders are refractory-lined steel troughs or channels for transporting molten metal. The refractory may be brick or, if the launder size is not excessive, cast and sintered as single pieces of refractory. Launders often incorporate projections (dams and weirs) into the line of flow of the molten metal to help prevent the transfer of slag into the molds during pouring. Dams are the projections rising from the bottom of a launder. Weirs are projections coming down into the top of the pour stream. Launders are used in the tapping of blast furnaces (several hundred in 30 min) and cupolas (1=10 that of blast furnace tapping). Launders are also used in smaller furnaces, where melt is transferred from one chamber to another (see, for example, the article “Vacuum Induction Melting” in this Volume). With systems opertaing under atmospheric conditions, the launder is usually covered so the metal is not
Fig. 1
Fig. 2
Electric low-energy holding furnace with laundered inlet. Courtesy of Schaefer Furnace Inc.
directly exposed to the atmosphere. The covers are heavily insulated and are hinged to permit inspection of any portion of the launder. The trough is lined with heavy insulation having low thermal conductivity. This insulation is also nonwettable by the molten metal. Metal conveyance free from turbulence throughout the launder system is important, so the metal moves smoothly and continuously usually under a protective layer of stationary metal oxides. Most importantly, metal temperature fluctuations are thus significantly reduced, and manual work is lessened. The size of launders varies, of course, depending on the scale of the production operation. Small holding furnaces may have an integrated launder inlet (Fig. 1). A much larger launder unit is shown in Fig. 2 with its connection to a tilting aluminum furnace. In this case, the tilting furnace is connected to the launder with two cast iron parts. These cast iron parts need maintenance after each cast in order to keep them working properly. If they start to leak, it takes quite some time to prepare them or replace the cast iron pieces before the next cast. Periodic skimming of the launder is also another maintenance task. Flexible launder joints (Fig. 3) are also used for the transfer of molten aluminum. They are very lightweight and easy to install. One-layer launder joints are used only once because of possible leakage. Flexible launder joints also include multiple-layer designs that can extend the load-carrying capacity and be used multiple times. The thickness and density of the blanket and the hot-face cloth can be adjusted according to the operating temperature, possible chemical reaction, and whether or not there is any flame contact.
Transfer area from a tilting aluminum furnace to a launder. Courtesy of AGN Nachrodt
Tundishes The tundish is an intermediate holding vessel used in steel continuous casting systems or in the handling of other metals such as superalloys. The metallurgical role of a tundish is to facilitate separation of inclusions and slag, which floats with sufficient residence time of the metals in the tundish. Tundishes also perform the function of a metal reservoir that allows unabated casting during ladle changeovers for sequence casting of heats. In this case, it is imperative that the ladle change be done in the shortest possible time. The tundish is a ceramic bricked or cast tub into which the metal flows after leaving a teeming ladle. Some systems are designed to pour directly from the melting crucible into the tundish. A typical tundish design has significant depth and provides a low-velocity flow path from the point of entry of the metal to the bottom installed nozzle through which the metal exits into the molds or feeding system. Tundishes are preheated prior to metal pouring and are often covered to minimize radiation heat losses. Tundishes may also be heated. For highly deoxidized steels, they are protected with an argon gas cover to reduce reoxidation and inclusion formation. An artificial slag is usually added to absorb inclusions that are floating out. Metal is poured away from the nozzle location directly on a wear resistant pad. A constant metal height above the tundish nozzle is maintained to discharge the metal at a constant rate and, in turn, to maintain constant casting speeds. The height of metal in the tundish is kept constant by regulating the metal pour rate from the ladle. The height of the liquid head in the tundish makes the flow (pour rate) out of the tundish nozzle a fixed rate. For specialquality steels or other melts with stringent surface and cleanliness requirements, shrouding is commonly employed to protect the steel from air, particularly for aluminum-killed grades, which have a potential of forming alumina inclusions. The shrouding systems include inert gas shrouding and refractory-tube shrouding. The concept of a refractory-tube shrouding
Fig. 3
Flexible launder joint in aluminum unit. Courtesy of Alunorf and Ceranex
Transfer and Treatment of Molten Metal—An Introduction / 175
Fig. 4
Schematic of a refractory-tube shrouding system for minimizing oxidation during pouring. Courtesy of Metaullics Systems Division, Pyrotek, Inc.
system is shown in Fig. 4 for ladle-to-tundish and tundish-to-mold pouring. Refractory-tube shrouds for steel are made of fused silica or alumina graphite. Nominal residence time of the metal in the tundish is important, because this controls the extent to which entrained slag will float to the top of the tundish and be removed from the pour. This has led to a trend toward tundishes with larger volumes and thus longer residence times for a given throughput. Flowcontrol devices, such as refractory dams and weirs, are attached to the tundish to distribute the metal flow, to minimize turbulence, and to eliminate dead flow zones (Ref 1). Refractory baffle walls in tundishes are also used to help direct the flow of liquid, so inclusions have an easier chance to float out at a later point. The holes in the baffle walls are typically approximately 50 mm (2 in.) and do not catch or trap inclusions (as in the case of molten-metal filters for inclusions in aluminum melts). Dams and weirs also are used in the tundish to collect slag particles. Bottom-projecting dams direct metal flow to the top of the tundish as a flotation aid for the slag particles. Flow modifiers have no influence on the very small inclusions collected by ternary refining, but they can help clean the steel of midsized inclusions in the 50 to 200 mm (2 to 8 mils) range. Flow controls are not needed for larger inclusions (which have higher Stokes rising velocities). The refractory lining in a tundish includes a service lining and a back-up safety lining. Refractories in the tundish are subject to corrosive wear, slag penetration, and thermal shock. Until recently, refractory wear in a tundish was usually of secondary importance to the ease of installation and tearout, as well as metallurgical considerations of the steel. During deskulling, the hot-face service lining is usually lined with a thin, 2.5 to 7.5 cm (1 to 3 in.), coating layer of olivine (magnesium iron silicate), MgO, or a mixture of these. This coating is
usually a slurry spray material or a dry vibratable powder that is installed behind a steel form that is heated to set a phenolic resin binder. With the drive for longer and longer cast sequences, the emphasis of tundish linings is changing toward more corrosion-resistant coatings that last longer and are cost-effective but that also maintain good deskulling properties. In steel continuous casting, the tundish continues to play a critical role in postfurnace steelmaking practice (along with steel ladle metallurgy treatments, as described in the article “Steel Melt Processing” in this Volume). Properly designed tundish flow modifiers provide substantial benefits in terms of both quality and productivity, and extensive work in physical and mathematical modeling has been done to understand and improve the fluid flow pattern in the tundish (Ref 2–4). Mathematical prediction of flow pattern has become more widely used in analyzing turbulence. Molten-Steel Filters. For critical applications of steel, the presence of inclusions with a diameter greater than approximately 50 mm (2 mils) needs to be prevented. There was a move in the industry in the 1990s to use filters (foams or screens) for such applications to help eliminate inclusions under approximately 50 mm (2 mils) in size, which are not susceptible to flow modifiers. However, this use of filters in tundishes for steel continuous casting did not prove effective, because the large number of inclusions would quickly clog the filter.
Molten-Metal Pumps Molten metals have traditionally been transferred from the melting vessel to a mold or casting unit by tapout or lip pouring (essentially gravity flow systems) or by using an intermediate ladle. Molten-metal transfer/recirculation pumps are now commonly used with aluminum, magnesium, zinc, and lead alloys. The pumps may be a mechanical device or a device operating on the principle of electromagnetic pumping. Mechanical pumps are typically centrifugal pumps with a rotating impeller in a volute casing. They can be used as either transfer pumps or as recirculation pumps in melting or holding units. Mechanical pumps are also used for gas injection. Electromagnetic pumps are also widely used in industry, because they have no moving parts, so routine maintenance is minimized.
Mechanical Pumps Mechanical pumps for molten-metal recirculation and transfer (Fig. 5) traditionally have graphite posts, which must be rebuilt approximately three times per year when they become oxidized or broken. Newer techonolgy includes molten-metal pumps with ceramic posts (Fig. 5) that are held in compression from a tension rod made of high-temperature alloy steel.
By replacing the posts with a stronger material that will not oxidize, the life of the pump has been extended, with fewer rebuilds. In this design, fragile ceramic sleeves are no longer needed to protect graphite posts from oxidation and abrasion. Ceramic sleeves can also hide post oxidation, which could result in unexpected failures. The ceramic posts held in compression are stronger than graphite; if they do break, the base is still supported by the steel rods. Mechanical recirculation pumps are used to improve heat transfer by forced convection. This is particularly important in aluminum reverberatory furnaces (see the article “Reverberatory and Stack Furnaces” in this Volume). In reverberatory furnaces, forced convection keeps the entire melt at a more uniform temperature, so that the burners above the melt transfer more radiant heat to the surface of the melt. If the surface temperature becomes hotter than the lower parts of the bath, then the heat transfer for melting become less efficient in reverberatory furnaces. Other nonferrous alloys, such as zinc, magnesium, and copper, also benefit from molten-metal circulation, but pumping systems for circulation are not usually employed. Mechanical stirrers are effective, particularly for zinc, due in part to the lower heat of fusion of zinc and the smaller melt sizes usually used. Mechanical transfer pumps are now commonly used with aluminum, magnesium, zinc, and lead alloys. A mechanical pumping system has several advantages. Pressurized flows permit larger volumes of metal to be transferred per unit time, with resultant productivity gains. The ability to move greater volumes of metal in shorter times also means that less superheat must be put into the metal to accommodate temperature losses during transfer. Thus, shorter overall furnace melting cycles are possible. Because most metal-transfer pumping systems have submerged intakes, dross entrainment from the surface is avoided, providing better metal quality. Transfer pumping per se does not produce additional oxidation or dross losses, because molten nonferrous metals are always subjected to turbulent flows (high Reynolds numbers), even with gravity flow tapout systems. Transfer pumps can often be used with closed piping systems to provide transfers surmounting physical obstacles (over other equipment, for example) or overcoming head and height limitations when gravity flow is limited or not possible. Melting and casting facilities can therefore be mixed and matched and need not be directly in line with one another. Transfer pumps are now commonly used in many segments of the aluminum casting industry to move molten metal from melting furnace to holding furnace, to feed or empty auxiliary refining vessels (used for alloying, degassing, and filtration), to empty and fill transfer ladles, to directly pour castings and fill molds, or to feed casting machines in the wrought industry. In the magnesium, zinc, and lead metal casting industries, mechanical pumps transfer metal
176 / Molten Metal Processing and Handling
Fig. 5
Fig. 6
Molten-metal pumps with ceramic posts and steel tension rods. (a) Recirculation pump. (b) Transfer pump. (c) Pump elements. Courtesy of Metaullics Inc.
Schematic of metallic centrifugal pump used to transfer magnesium
between furnaces, fill ingot and pig molds, feed die casting machines, cast electrogalvanizing anodes, and empty galvanizing kettles when repairs or desludging is required. Molten copper alloys are also transferred by pressure pump systems. Of the various kinds of mechanical pumping systems, a positive-displacement pressure pump is often quite useful for small furnaces or small flows (<45 kg, or 100 lb). When actuated, external air pressure displaces the volume of air in the cylinder, which forces an equivalent volume of liquid metal through the delivery tube. Another type of pump is the siphon pump. Both of these pumps are quite capable of filling small molds where the accuracy of fill demanded is moderate (several percent variation) and where pressure head requirements are minimal (0.3 to 0.6 m, or 1 to 2 ft, maximum). When larger volumes of metal, more sustained flow, or higher head is required, a centrifugal pump can be used. Flow is directed by an impeller up through a riser tube. Operated either by air or electric drive, these pumps are
typically constructed of graphite and are easily operated in molten aluminum, magnesium, and zinc alloys. It is important to keep the riser tube open, using insulation and careful external heating if necessary, to prevent freezeups, especially if use is intermittent. Another type of centrifugal mechanical pump often used with magnesium alloys is shown in Fig. 6. Pump construction is usually cast iron, low-carbon steel, or 400-series stainless steel. Nickel impurities in metal pump components should be specified to be below 0.20% to avoid the dissolution of nickel in the molten magnesium and the subsequent reduced corrosion resistance of magnesium castings. Centrifugal pumping systems for zinc are usually constructed of graphite, but metallic centrifugal pumps are often used as well. Pumping systems for lead are usually constructed with cast iron because little corrosion is experienced. Cast iron can also be used to pump zinc as long as the temperature is kept below 480 C (900 F), above which molten zinc begins to dissolve iron. Austenitic stainless
Transfer and Treatment of Molten Metal—An Introduction / 177 steels can be used as pump materials with aluminum-containing zinc alloys (0.2 to 3.5% A1). Stainless steel must not be used, however, with pure zinc, because liquid metal embrittlement and intergranular attack occur quite rapidly. With zinc-aluminum galvanizing alloys (40 to 60% Al), the high operating temperature (595 to 650 C, or 1100 to 1200 F) is in the region where austenitic stainless steels become sensitized to greater intergranular carbide precipitation and subsequent accelerated corrosion. Stainless steel pumps should therefore not be used for this application. A useful but expensive alternative for the latter two applications is an Mo-30W or Mo-40W pump shaft, which provides excellent corrosion resistance but is fairly brittle at high temperatures.
Electromagnetic Pumps Electromagnetic pumps have been developed for circulation of molten metal and controlled mold-filling capabilities. This type of pump is constructed from ceramic materials and has a linear motor that induces pumping on the basis of the left-hand rule of electromagnetic field theory, using the molten metal as the secondary field. It is particularly useful for metals with good electrical conductivity. One specific advantage of this type of pump is that there are no moving parts. However, the moltenmetal channel opening must be kept very clean and free of oxide buildup or dross; otherwise, the field strength and resulting pumping capacity can be greatly reduced. Electromagnetic Pumps for Furnace Circulation and Scrap Recycling Technology. Electromagnetic (EM) pumping systems are used for application of forced circulation in both foundry and casthouse aluminum melting. The EM pumps are supplied to suit the furnace size and application, with mass flow capability from 1 metric tonne per minute to greater than 20 metric tonnes per minute. All pumps are operated at main voltage and frequency and are fully controllable for both speed and direction. The consumable pump tube is replaced approximately every 6 months, especially with mixed-charge scrap melting, although with certain applications, up to 18 months of continuous operation are possible. An EM pump is especially suitable for those melting furnaces that may not have sidewells, or furnaces that are not stationary but rather tilt-pour. With the consolidation in the foundry industry, larger, volume-oriented castings producers are increasing their central melting furnace sizes well above 10 tons, making the application of an EM pumping system to circulate the furnace a viable option. In terms of operation, an EM pump is connected via insulated tubes to the furnace charge chamber and an external refractory-lined charge well, where a unique internal bowl shape develops a vortex with, low-turbulence scrap submergence (LOTUSSE Metaullics
Systems) in the molten-metal stream. The vortex creates an ideal entry point for high-surface-area, low-density scrap and alloying materials that need rapid submergence to avoid oxidation with direct flame impingement. These include machining chips, borings, turnings, flashings, used beverage containers, and so on. While many foundries would only melt a small portion of such materials, cast wheel producers as well as producers of castings with inhouse extensive machining produce a sufficient quantity of such light-gage materials (machining chips) to make the EM pump/LOTUSS scrap recycling technology quite feasible and economical. Indeed, the same LOTUSS scrap recycling technology can be employed with mechanical pumps as well. The choice of which system to employ—EM pumping or mechanical pumping—will depend on many specific user’s economical and process factors. Metal flow capability and electrical requirements of typical EM pumping parameters are: Metal flow rate, t/min
Kilovolt amperes (VA) rating
8–10 8–10
110 110
15
130
8–10
110
15
130
20 12–14
150 110
Features
100 mm (4 in.) pump 100 mm (4 in.) pump and redesign tie bars 100 mm (4 in.) pump and laminations 150 mm (6 in.) pump and redesign tie bars 150 mm (6 in.) pump and laminations 150 mm (6 in.) pump, 9 poles 200 mm (8 in.) standard pump
Use of a either a centrifugal or electromagnetic metal pumping system has the following benefits: Increased metal yield (typically for dry
Alloying materials in foundry melting are prone to problems with dissolution, incomplete melting, and less-than-optimum recovery. The combination of rapid-submergence charging capability, coupled with the forced convectional circulation provided by the pumping action, helps to ensure quick dissolution and higher recovery than would otherwise be possible. Pure copper, silicon, or aluminum master alloys can be processed by this technology. Smaller lump forms, granules, and small ingot bars are preferable to powdered additions. Recent advances (Ref 5) in EM pumping technology have integrated several other functions into the heretofore single application of forced circulation. The EM pumps are now capable of providing transfer of metal out of the closed system into a transfer ladle, launder, or a holding furnace, and with suitable gas injection systems, degassing is also possible. Additionally, such systems can be outfitted with noncontact sensor devices (video camera, level sensor, etc). Mold-Filling EM Pumps. Electromagnetic pumps are easily used for controlled mold filling in such applications as permanent mold casting, large die casting, low-pressure die casting, and perhaps small ingot casting. They offer exceptional control of metal flow throughout mold filling, and turbulence is minimal since there are no moving parts. This puts the highest-quality metal into the mold. Other benefits are: Metal filtration is easily performed using a
chips, 98%)
compared to a static bath)
yield; silicon, +95% yield)
Increased melt rate (typically 10 to 20% High-efficiency alloying (magnesium, +95% Reduced energy use (equivalent to increased
output)
Chemical and thermal homogeneity within
the furnace (þ /5 C, or 9 F, within the furnace) Reduced dross formation (depending on scrap types)
With the LOTUSSE scrap recycling system, many different materials are possible to be charged into the special-shaped bowl. However, it is not possible to charge large sows or large scrap castings through the system. The following are typical smaller, lighter-gage aluminum materials that are suitable for processing in this vortex scrap charging system: Machining chips, turnings, borings (which are
most prevalent in integrated foundry operations)
Shredded sheet scrap Extrusions
Large fragmented castings Small castings Scalper chips Used beverage containers Aluminum chips
filter over the pump inlet, filtering all metal to the casting. The response to a desired change in flow rate is very fast. Plant safety is increased due to automation— no need for hand ladling. The alloy in the furnace can be tested at any time, and additions made without interfering with production. In low-pressure casting, the furnace is not a closed pressure vessel. The temperature of the metal delivered by a pump is very close to that of the moltenmetal bath, so superheating of the furnace is not required.
Types of EM Pumps. The types of EM molten-metal pumps include:
Single-phase conduction pumps Jumping ring pumps Multiphase linear pumps Pinch-effect pumps Direct current conduction pumps
Single-phase EM pumps (Fig. 7) can be applied to almost any metal casting process. They are particularly suited to low-pressure casting,
178 / Molten Metal Processing and Handling
Fig. 7 Example of two sizes of single-phase conduction pumps with metal filters over the bottom inlet where they fill the mold from below while exhibiting very good control over the rate of metal delivery, along with the ability to almost instantly change the flow rate. Repeatability is generally excellent. It is also very easy to hold the metal just below the mold between castings to minimize oxide generation and to reduce overall cycle time. Other uses include squeeze casting through the side of a vertical shot sleeve, just above the top of the shot tip, to eliminate dross from top-filling; high-pressure die casting; bottom-fill lost foam; and all kinds of gravity casting. Single-Phase Conduction Pump. The singlephase conduction pump uses alternating current applied to a coil of many turns of wire that is wound around a laminated ferrous core, which constitutes the primary winding of a transformer. Around the same laminated core is a ceramic material that is compatible with aluminum, and this ceramic has an internal passage into which molten aluminum flows from the metal bath. This passage forms a complete circuit (loop) around the laminated core; hence, the molten aluminum therein forms a one-turn shorted secondary winding on the transformer. By virtue of this single turn, a high-amperage flow is induced into the molten aluminum. A portion of the loop is flattened to allow for close spacing of the magnet pole pieces and to make the current density in that portion of the loop relatively uniform throughout the cross section. This is the area (pumping vein) in which the pumping force will be produced. Two pole pieces are placed as closely as possible to the broad face of the pumping vein, and additional electrical coils are around the laminated core that forms the magnet pole pieces. This creates an alternating magnetic field across the pumping vein. Since the transformer primary and the magnetic field coils are powered from the same alternating current source, both the current flow in the metal and the applied magnetic field alternate polarities together. If either the current in the metal or
the magnetic field were static while the other one alternated, the metal would experience an alternating force, first in one direction, then in the other. Since the magnetic field and the current in the metal are both alternating together, the force on the metal simply pulsates from zero to maximum at twice the applied line frequency. These pressure pulsations may prove beneficial when pumping thixotropic materials such as metal-matrix composites, because the pulsations introduce shear into the moving metal, thus greatly increasing the fluidity, which helps the metal to flow into mold details. Since the electrical current is induced directly into the molten metal, there is no need for electrodes in contact with the metal, also eliminating the need for heavy cables and very large direct current power supplies. This is the type of pump most commonly used for mold-filling applications in aluminum foundries, because of its ease of use, reliability, and repeatability. Jumping Ring Pumps. If a coil of wire is wound around a straight ferromagnetic bar or rod, and a ring of any conductive metal is placed near it over the same bar, with alternating current applied to the first (primary) coil, the ring is very strongly repelled from the coil. This is called the jumping ring pump. These pumps do not generate much force in the molten metal and are most often used for stirring metal in a furnace. Multiphase Linear Pump. If a multiphase electric motor stator (the stationary part with the electrical windings) were cut open and rolled out so the windings are all in a line, when power is applied, there would be a moving magnetic field. If a conductive material is placed above this stator, electrical current would be induced into that material. This would also produce corresponding magnetic fields, which would interact with the moving magnetic field, and the interaction between these magnetic fields would produce a force on the conductive material, tending to move it in the direction of the moving magnetic field. As described, this would be a very inefficient pump due to poor control over the magnetic field. There are two common ways to improve this. One is to place a plain pole piece opposite the stator while keeping the molten metal between the stator and the coil assembly. The second improvement is that the stator could be rolled to form a long tube, with the moving magnetic field traveling down the length of the tube. This type of pump is typically used for high-volume metal transfer and stirring of metal in large furnaces. Pinch-Effect Pumps. Another form of pump induces current into a cylinder of molten metal that has a tapered ceramic mandrel in the center. Since the metal is trying to collapse on itself, it is forced toward the narrower end of the taper, which is also the metal discharge area. These pinch pumps typically use a highfrequency coil to improve the efficiency of the induction, and they also typically use water cooling to keep the coil from overheating.
Direct current conduction pumps use electrodes to pass a large direct current into the fluid to be pumped, and a permanent magnet supplies the magnetic field. These pumps are not suitable for use with aluminum, since they would require electrodes to be in contact with the molten metal, and there is no suitable material available for this purpose. Additionally, the amperage required is very high, which would need very heavy cables and a direct current power supply capable of supplying as much as 10,000 A.
Pouring and Dosing As noted, molten metals traditionally have been transferred to a mold or casting unit by tapout or lip pouring by using an intermediate ladle. Traditional metal transfer from the molten-metal supply to the mold is done with open-top ladles ranging from hand held to multiton, crane-supported, mechanized tilt ladles. Underpouring from a ladle, either through the incorporation of a teapot spout or bottom stopper rod and nozzle, reduces slag inclusions in castings (Fig. 8). In addition to gravity-feed pouring, a number of other methods are also used to fill molds. For example, molten-metal pumps (as briefly described in the preceding section of this article) are used. Other methods include the use of dosing furnaces in die casting operations (e.g., see the article “Automation in High-Pressure Die Casting” in this Volume). Dosing furnaces are replacing ladle furnaces and are described in the next section of this article. Another method is low-pressure filling systems for sand casting of heavy metals. In lowpressure pouring (Fig. 9), the molten metal is pushed into the open mold from below. Compared with gravity casting, the low-pressure method offers a number of important advantages: Reliable, simple, and tight mold-to-nozzle
docking action
Adaptable pressure buildup for active control
of mold-filling characteristics
High dosing accuracy based on pressure and
(where applicable) level control
Essentially laminar mold-filling flow Cleaner melt supply
Finally, automated methods encompass a wide range of mechanized methods and specialized equipment. This includes mechanized ladles and pouring robots that are progarmmed to transfer measured amounts of metal from a storage vessel or a controlled pouring unit to the mold. Pumps and dosing furnaces are also suited to integrated automation. Benefits of automated pouring include: Increased productivity due to the speed and
consistency of pours from one mold to the next
Transfer and Treatment of Molten Metal—An Introduction / 179
Fig. 8
Four types of pouring ladles. (a) Open lip-pour ladle. (b) Teapot ladle. (c) and (d) Bottom-pour ladles
Improved process control with tracking of
production and inventory information. In some cases, automatic pouring equipment also allows a reduction in pouring temperatures with significant savings in energy and refractory costs. Improved working conditions in terms of health and safety, environmental, and operational factors
Pouring Methods
Fig. 9
Low-pressure mold filling. (a) Bottom fill. (b) Fill from top of the mold
Increased yield due to fewer overpours,
spills, and allowance for smaller pour basins. Consistent, controllable pour rates also may allow for the design of less complex gating and riser systems as well as for simultaneous multiple-pour streams in the same mold. Reduced scrap costs due to reduction of inclusions in castings and the reduction of pouring failures such as underpours, interrupted pours, and irregular pour rates Improved quality control with better temperature control within the pouring vessel, more consistent metallurgy, and automatic feeding of alloy and inoculant additions into the pouring ladle or metal pour stream
The two key variables in pouring are pouring rate and duration. The optimum pour rate is the maximum rate at which a mold will accept metal, and the optimum duration is the time from the start of the pour until the mold will no longer accept metal. Other important considerations in automatic pouring systems include the relative positioning of the mold and pour spout or nozzle and the sequencing of the pour. In manual pouring, the worker compensates for variables in the mold, ladle, and metal while pouring to produce the best results for each pour. Duplication of this effort is the challenge of automatic pouring. Optimum pouring time is based on the weight and shape of the casting, the temperature of the molten metal and its solidification characteristics, and the heat-transfer characteristics and thermal stability of the mold. Most foundries may also pour many different castings from one heat, or even from one ladle. Therefore, these foundries do not attempt to control pouring time, but they do control the speed at which molten steel enters the mold cavity. This control is achieved through the design of the gating system. Pouring Ladles. The three basic types of pouring ladles are: Bottom pour Teapot Lip pour
Ladle capacity normally ranges from 45 kg to 36 Mg (100 lb to 40 tons), although ladles having much larger capacities are available.
Ladle Linings. Ladles with capacities above approximately 360 kg (800 lb) are lined with layers of firebrick covered with a thick refractory coating. Smaller ladles are protected by monolithic linings prepared by ramming or by pouring and vibrating refractory mixtures around forms. For molten steel, the optimum refractory for this purpose is high-purity alumina, but because of the high cost of alumina, alternative materials such as olivine and zircon are more widely used. The service lives of these monolithic linings, which are most commonly used for small ladles, vary considerably, depending on the refractory mixture and the type of service. Thermal expansion and contraction can cause hairline cracks, into which hot metal enters. This results in erosion of the lining. Erosion not only shortens the life of the lining but also contributes to excessive dirt in the steel. Disposable fiber linings that do not require preheating are gaining acceptance. Preheating the ladle helps to prolong the life of the ladle lining. Heating the lining to a temperature close to that of molten steel serves to lessen the thermal shock to the lining when the molten steel contacts it. Common practice is to preheat ladle linings to 815 to 1095 C (1500 to 2000 F) with oil-fired or gas-fired torches. Ladles lined with high-purity alumina can be preheated to temperatures as high as 1370 C (2500 F). Bottom-pour ladles have an opening in the bottom that is fitted with a refractory nozzle. Stopper rods (Fig. 10a) were commonly used to control flow from the unit, but more typically a valve-type mechanism (Fig. 10b) is often used. Bottom pouring is preferred for pouring large castings from large ladles, because it is difficult to tip a large ladle and still control the stream of molten steel. Also, the bottom-pour ladle delivers cleaner metal to the mold. Inclusions, pieces of ladle lining, and slag float to the top of the ladle; thus, bottom pouring greatly reduces the risk of passing nonmetallic particles into the mold cavity. Pouring ladles may also include options for melt treatments or shrouding. For example, Fig. 11 shows a schematic for a bottom-pour
180 / Molten Metal Processing and Handling
Fig. 13 Fig. 10
Bottom-pour mechanisms (a) stopper rod (b) sliding value
Fig. 12
Fig. 11
Bottom-pour teeming with argon insertion ring. The term teeming typically applies to the pouring of molten metal from a ladle into ingot molds.
mold setup with an argon insertion ring that shrouds the pour from oxygen and nitrogen pickup from the atmosphere. Mold powders or fluxes are added to the advancing metal front to improve surface quality by providing thermal insulation from the mold surface. At the end of the pour, exothermic compounds may be added to the top of the mold to insulate the top of the solidifying ingot and provide a degree of hot
Typical teapot ladle used to pour small- to medium-sized castings
topping, where the molten metal will fill the shrinkage occurring upon solidification lower in the ingot. Argon-oxygen decarburized electrodes also may be top poured, but the need to continually close the ladle tap, transfer the ladle, and reinitiate argon shrouding produces variable conditions from ingot to ingot. Should difficulty in closing the nozzle be experienced during top pour, then a safety problem arises. Thus, bottom pouring is the more generally used process. On the other hand, it is impractical to pour molten metal into small molds from a large bottom-pour ladle. The pressure head created by the metal remaining in the ladle delivers the molten metal too fast. Also, the time required to fill a small mold is short, thus requiring that a large bottom-pour ladle be opened and closed many times in order to empty it. This may cause the ladle to leak, although special nozzles have been developed to minimize leakage. Despite the fact that the size of bottom-pour
Schematic of automatic mechanized ladle
pouring
with
ladles could be scaled down for pouring smaller castings, this is unnecessary because of the almost equal ability of the teapot ladle to deliver clean metal. The teapot ladle incorporates a ceramic wall, or baffle, that separates the bowl of the ladle from the spout. The baffle extends almost four-fifths of the distance to the bottom of the ladle (Fig. 12). As the ladle is tipped, hot metal flows from the bottom of the ladle up the spout and over the lip. Because the metal is taken from near the bottom of the ladle, it is free of slag and pieces of eroded refractory. The teapot design is feasible in various sizes, generally covering the entire range of casting sizes that are below the minimum size for which the bottom-pour ladle is used. Lip-pour ladles resemble the teapot type but have no baffles to hold back the slag. Because the hot metal is not taken from the bottom of the ladle, this type of ladle pours a more contaminated metal and is seldom used to pour high-alloy castings. Nevertheless, it is widely used as a tapping ladle (at the melting furnace) and as a transfer ladle to feed smaller ladles of the teapot type. Methods of Automatic Pouring. There are two basic methods of automatic pouring: mechanized ladle pouring and direct pouring. Mechanized ladles generally have capacities matched to single mold requirements and include load cell weighing systems and controlled tilt rates (Fig. 13). In the case of direct pouring, metal is metered directly into the mold from the pouring vessel, with the appropriate controls to ensure the desired results. The use of a mechanized ladle allows for inladle alloying and inoculation before pouring. Dripping of molten metal onto the mold is eliminated when a mechanized ladle is used in combination with a stopper rod and nozzle
Transfer and Treatment of Molten Metal—An Introduction / 181
Fig. 14
Fig. 15
Automatic pouring into molds on continuously moving line. (a) Pouring into flaskless molds on an indexing line from bottom-pour unit with stopper rod. (b) Synchronized pouring on a continuously moving line with mechanized ladles capable of x-axis and y-axis travel and synchronization with the pouring line.
Rotary mechanical pourer with molten metal level detector and control
pouring unit. This is very important when pouring lost foam molds. Ladle tilting about an axis in line with the pouring lip can reduce moltenmetal stream travel during pouring. The capacities of mechanized pouring ladles are based on anticipated metal requirements. The dispensing of a metal stream from a ladle at a specific point requires that the nozzle or pouring lip not move, relative to the mold,
during pouring. In semiautomatic pouring units, bottom-pour ladles may be mounted above indexing molding lines to pour into molds indexed into position beneath the ladles (Fig. 14a). Synchronized pouring on a continuously moving mold line (Fig. 14b) may also require the use of a pouring vessel to measure predetermined weights of metal and either one or two (depending on line speed) mechanized
pouring ladles. Another approach is a mechanized rotary pourer with optical or laser monitoring for automatic control of melt level (Fig. 15). Direct bottom-pour ladles use either slide gates or stopper rods and nozzles to control and stop the flow of metal to the mold. Slide gates and nozzles are most commonly used to control flow rates on bottom-pour ladles for continuous and semicontinuous strand casting operations. Slide gates (Fig. 10b) offer more reliable shutoff than stopper rod valves; however, service life of the slide gate can be shorter than that of the stopper rod. Stopper rods and nozzles offer more control of pour rates than slide gate valves and are less expensive. For these reasons, bottom-mounted stopper rods and nozzles are more common in foundry pouring operations. Controls of Pouring. Satisfactory pouring results can be obtained in some automatic pouring applications without complex electronic control systems. Some requirements can be met by establishing a flow rate and simply timing the pour. A slightly more complex system includes dual pour rates and pour times. Simple infrared photosensors designed to detect the presence of molten metal can be used to end the pour cycle. If the exact weight of metal required for a mold is known and a mechanized ladle is weighed while filling, at least one variable is under control. If the movements of the operator controlling the pouring of a mechanized ladle are measured and then stored in a programmable controller, these movements and the results can be duplicated under automatic control. The programming of a controller by duplication of remote control is referred to as a teach-in system. The program is referred to as the pouring profile. Weighing of a bottom-pour ladle and pouring by weight can also be effective in some cases, although the relatively small capacity of the
182 / Molten Metal Processing and Handling ladle results in substantial variations in metal head pressure above the stopper and subsequent variations in pour rate. Increasing the capacity of the pouring vessel, facilitated by electric heating, reduces the variation in head pressure. The incorporation of a process control system to compensate for reduced head pressure through stopper positioning by feedback of weight, anticipated weight, or actual bath height is effective in some cases. If the rate of metal flow through a nozzle is constant for a constant head pressure, the repeated variations in the position of a stopper rod above the nozzle will yield repeatable and controlled pouring rates. A teach-in system is used to program the pouring profile. If slag buildup on the nozzle and variations in metal head pressure cause flow variations, the system pouring profile program must be corrected by a process control feedback loop. Feedback can be provided by a timing and metal-presence sensing system that checks the rate of mold fill or by a more sophisticated control system for mold-level measurement. Mold-level measurement and control systems fall into two categories: laser systems and vision systems. Both of these electronic process control systems monitor the level of metal in the pour cup and control the position of the stopper rod to maintain a nearly constant level throughout the pour. Optical sensors provide input to help regulate the electromechanical stopper drive in such a way that the liquid level in the filling head is kept at a constant level. This type of feedback system most closely duplicates the efforts of a skilled operator. Camera or laser level sensors are used depending on production factors. Manufacturers who mainly use a sprue cup diameter of 85 mm (3.3 in.) or larger will usually select the camera-based solution, whereas the laserbased solution will be used particularly where the focus is on minimizing the circulation material and a sprue cup diameter smaller than 80 mm (3.1 in.) is employed. The laser reflection angular detection system (Fig. 16a) determines the position of a lowpower laser beam reflected from the surface of the molten metal in the pouring basin. The laser beam is directed at the pouring basin from an angle. The distance to the molten-metal level will affect the incident and reflected angle and therefore the position of the beam on a sensing unit. The comparative value of the beam position on the sensor is used to control the position of the stopper rod; thus, the level of the metal in the pouring basin is controlled. A vision system uses a video camera to observe the pouring process (Fig. 16b). The video signal is displayed on a monitor for the operator and processed electronically in a manner that separates the brightness levels of the molten metal from the other portions of the picture. The molten-metal stream is electronically blanked from the composite picture of the metal in the pouring basin and the area of metal fill compared to the actual area of the pouring
Fig. 16
Optical sensors for pouring control. (a) Laser sensor. (b) Vision sensor
basin. The comparative value is used to control the position of the stopper rod, and the level of the metal in the pouring basin is controlled. The most sophisticated pouring systems may incorporate more than one set of controls. The use of integrated control schemes allows the system backup, improved safety, and increased flexibility that are required for pouring a variety of castings. For example, the benefits of moldlevel measurement and control are enhanced when used in combination with a teach-in duplicated pouring profile in some pouring applications. Pouring Furnaces. Radiant element, electric resistance-heated pouring furnaces are used in both nonferrous and ferrous pouring applications. For nonferrous, low-temperature applications, metal or silicon carbide heating elements are used. For ferrous high-temperature applications, graphite rod heating elements are used. Radiant-heated furnaces are typically bottompour units with stopper rods and nozzles mounted in the furnace bottom or in a detachable pour box. The furnace can be tilted to drain the pour box and to contain the molten metal beneath the radiant heat source. Crucible-type electric resistance-heated and coreless induction-heated furnaces are most commonly used in nonferrous pouring applications. These applications vary from the simple dip-and-pour mechanized ladle systems to the underpour stream discharge, axial-pour, and trunion-pour furnaces. Underpour stream
discharge in aluminum pouring for continuous and semicontinuous casting processes uses the oxide skin on the metal surface to protect the metal stream from further oxidation while transferring metal from the melting furnace to the automatic pouring ladle. Channel-type electric induction-heated pouring furnaces are sealed and designed to use gas pressure to discharge metal from the furnace vessel. Pouring can be accomplished over the lip of a pour spout or through a nozzle, with or without a stopper rod, in a pour box mounted on the furnace vessel. Vessel gas pressure can be varied to control the rate of pouring or to control the level of metal over the pouring nozzle (Fig. 9b). Control of the level of metal in the pour box establishes a constant static head pressure above the nozzle and ensures consistent pour rates. Venting of the pressure in the vessel causes the metal in the pour box to drain back into the furnace. Inert gas can be used as the pressure medium in applications where the metal is particularly susceptible to oxidation.
Dosing Furnaces Dosing furnaces are replacing ladle furnaces in many applications, especially in aluminum high-pressure die casting (see the article “Automation in High-Pressure Die Casting” in this Volume). Aluminum castings traditionally have been produced by either direct pouring from the melting furnace into the mold, or from a
Transfer and Treatment of Molten Metal—An Introduction / 183 transfer ladle in which further treatment has been performed just prior to pouring the mold, or from a holding furnace dip-well from which the casting is poured either by automatic robotic ladle or by manual means. Increasingly, automated and controlled direct pouring is becoming popular, especially in high-pressure die casting operations where the dosing furnace technology can be meshed with the automated die casting machine controls. A number of different units exist commercially that are comprised of holding furnaces with discharge capabilities directly into the mold cavity. These devices provide a very precise dosage of molten metal and are thus known as dosing furnaces. The dosing operation (Fig. 17) begins with the introduction of compressed air into the pressure-sealed furnace, thereby causing the ascension of the melt in the riser tube. With further increase in the pressure, the melt is discharged from the riser tube and flows via the dosing channel into the shot sleeve of the pressure die casting machine. Flow of the melt is stopped by rapid discharge of the compressed air, the molten metal returns to the level of that in the dosing furnace, and the die casting machine can now execute the casting operation. There has been constant further development in dosing with the dosing furnace to fulfill casting demands. The following describes several of the key technologies that represent the current state of the art. The following demands are placed on the overall process: High and reproducible dosing accuracy Short main time and minimum waiting time No loss of time with fluctuations in the cycle
cycle times without further adjustments, thereby excluding human influence. It is only necessary to enter the melt temperature and dosing weight. The dosing furnace automatically synchronizes itself with the die casting machine, thus always resulting in an optimized dosing cycle. Detection of Melt Level. It has been and still is necessary to be able to detect the level of the melt in front of the discharge edge at the rising tube. This was previously achieved by an electrode that generated an electric signal on contact with the surface of the melt (Fig. 18a). In order to avoid deposits of aluminum, this electrode had to be rapidly swung out after contact by means of a pneumatically operated device. Heat, oil, spraying agents, and sometimes open fire often led to failures in this component. A pneumatic scanning system (Fig. 18b) today (2008) replaces the electrode for detection of the melt level. Pneumatic scanning employs no moving components, requires no adjustment, and is especially advantageous with regard to high operational reliability and the absence of maintenance in a particularly harsh environment. The scanning accuracy is a precondition for exact dosing and is very well fulfilled by this system. The high degree of reliability in a pneumatic scanning system allows a level-scanning process known as the “top-stop-position” that has been used in the dosing furnace technology since approximately 2005. In this process, the height of the melt in the riser after dosing is not reduced to the level of the melt in the furnace but is held a little below the discharge
edge (Fig. 19). The furnace is thus always ready for dosing, so that there is no delay in discharge of the melt on receipt of the signal from the die casting machine. This compensates for fluctuations in the cycle of the die casting machine, and there is no loss of time through nonsynchronized process sequences. As a rule, this system achieves up to an 8% savings in the cycle time. Because the dosing furnace is a closed vessel that cannot be looked into, the filling level previously had to be determined by a weight load cell. This method was inaccurate since external influences on the furnace were also detected and could falsify the reading. With the development of new technology, the level in the furnace is precisely detected during the dosing process by means of the dosing contact and the system, so that the result is free from external influences. Level scanning is used to ensure that the furnace also remains ready for dosing while it is being refilled with molten metal (Fig. 20). Level scanning is started after opening of the filling cover, the level and the required amount of refill being semicontinuously detected in short cycles with indication of the current level. Furnace Recharging and Melt Holding. The quality of the melt in the transport ladle is only conditionally decisive for the casting quality. The decisive factor is the quality of the melt in the holding or dosing furnace and, finally, in the mold itself or the shot sleeve of the pressure die casting machine. If the quality of the melt in the shot sleeve is to be maintained or increased, closer consideration must
Fig. 18
Fig. 19
time Controlling and Monitoring the Dosing Operation. A digital control unit provides for real-time pressure control in the dosing furnace, and process information is monitored on a computer display with a large-surface keyboard for the operator. Precision hardware ensures adherence to dosage accuracy. The dosing program for the cold chamber pressure die casting machine optimizes the dosing to the shortest
Fig. 17
Dosing furnace
Level detection in dosing furnace. (a) Contact electrode. (b) Pneumatic scanner
Dosing sequence. (a) Melt hold, ready. (b) Melt flow (casting). (c) Melt again held, ready
184 / Molten Metal Processing and Handling
Fig. 20
Detection of melt level by level scanning
Fig. 21
Fig. 22
Recharging a dosing furnace
Dosing furnace with porous plugs
be given to the recharging operations and melt holding in the dosing furnace. Each filling or refilling operation is associated with a reduction in the quality of the melt. A correctly implemented filling of the dosing furnace can thus also substantially restrict the loss in quality. Figure 21 shows a correctly implemented filling operation. The filling funnel is kept full, comparable with the filling of a die in gravity die casting. This results in a lower formation of oxides, which remain on the surface, do not enter the dosing furnace, and can be skimmed off at the end of the filling operation. Holding of an aluminum melt in the dosing furnace allows for settling of inclusions to a major extent. Measurements confirm that the density index or other quality parameters usually improve during quiescent holding periods. The quality of the aluminum melt thus continuously improves with the holding time, and the melt quality is not negatively influenced by the dosing process itself with dry compressed
air (protection by the oxide skin). One option to provide further improvement in melt quality within the dosing furnace is through the installation of porous plugs (Fig. 22). Introduction of inert gas, preferably argon, at the lowest point in the furnace enables reliable maintenance of a hydrogen content in the furnace to acceptable low levels. ACKNOWLEDGMENT The editors extend thanks to the following for contributions to this article: Edward S. Szekeres, Casting Consultants
Incorporated, and Robert D. Pehlke, University of Michigan, on tundishes in continuous casting of steel Martin Reeves, Striko-Dynarad, on dosing furnaces Lee Gouwens, CMI Novacast, on electromagnetic mold-filling pumps Dave Neff and Paul Campbell, Metaullics Systems Division, Pyrotek, Inc.
REFERENCES 1. I.V. Samarasekera and J.K. Brimacombe, The Continuous Casting Mould, Continuous Casting, Vol 2, Heat Flow, Solidification and Crack Formation, ISS of AIME, 1984, p 33–44 2. Y. Sahai and Toshihiko, Melt Flow Characterization in Continuous Casting Tundish, ISIJ Int., Vol 36 (No. 6), 1996 3. M.M. Collur, D.B. Love, and B.V. Patil, “Use of Flow Modifiers to Improve Performance of a Tundish,” ISS Steelmaking Conference, April 1997 4. M.L. Lowry and Y. Sahai, Investigation of Steel Flow in a Continuous Casting Tundish with Multiple Baffles Using Mathematical Models and Tracer Studies, 1989 Steelmaking Conference Proceedings, p 71–79 5. A.M. Peel, Electromagnetic Pumps and Stirring Solutions for the Aluminum Industry, Alum. Times, Vol 7, 2005, p 32–33
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 185-193 DOI: 10.1361/asmhba0005353
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Degassing David V. Neff, Pyrotek, Inc.
GAS POROSITY is a major factor in the quality and reliability of castings, and there are various sources of gas porosity in castings. One source is the reaction between the casting mold and the molten metal. For example, the moisture in sand molds can react with the elements in the cast iron (carbon, silicon, aluminum, or iron itself) to form dissolved gases such as hydrogen or carbon monoxides (see the article “Gases in Metals” in this Volume). In addition to reactions with the mold wall, porosity also occurs from entrapped air during filling, centerline shrinkage that occurs during the final solidification, and blow holes from unvented cores (Ref 1). The other major cause of gas porosity in castings is the evolution of dissolved gases from melting, and dross or slag containing gas porosity. For example, liquid cast iron may have dissolved hydrogen and nitrogen. The solubility of these gases in the solid may be less than that in the liquid, and the gases may therefore be evolved during solidification. Whether or not this will occur depends on the amount of hydrogen and nitrogen present, the alloy being cast, chemical kinetics, and the surface tension of the alloy. Similarly, hydrogen can dissolve in liquid aluminum, but its solubility is much lower in solid aluminum and can cause porosity. For copper, the evolution of hydrogen, water vapor, or carbon monoxide could cause gas porosity. Gases can be a major problem in the casting of iron, aluminum, magnesium, and copper. Dissolved hydrogen is a particular concern because of its solubility in these metals (see the article “Gases in Metals”). The most common method for removing hydrogen from aluminum, copper, and magnesium is inert gas flushing, as described in this article. Inert gas flushing can also be used for removing hydrogen and nitrogen from cast iron (see the article “Gases in Metals”). Degassing of steels is described in the article “Steel Melt Processing” in this Volume.
Aluminum Foundry Degassing Degassing of aluminum alloy melts is a critical melt treatment technology necessary to provide readiness of the melt for casting into many
applications (Ref 1–8). Hydrogen is the only gas that is soluble in aluminum, as monatomic hydrogen. Figure 1 exhibits the hydrogen solubility of typical aluminum-silicon foundry alloys. Clearly, a substantial amount of hydrogen can be present in the liquid state, but as the melt solidifies, there is a substantial reduction in the amount of hydrogen that can exist in the solid state. How this change takes place will determine whether porosity will form. If not removed or reduced to an acceptable level determined by prior process and quality studies, the gas dissolved in molten aluminum, if not treated, can result in detrimental properties to castings. Figure 2 is a classic example of gas porosity developed during solidification in an aluminum alloy casting. Loss in mechanical properties, such as tensile strength, elongation, fatigue strength, and fracture toughness, are common results. In addition, interconnected porosity can result in “leakers” during pressure-tightness testing and blistering upon subsequent heat treatment as the pores grow and open up; this is especially a problem on the surface when surface finish/cosmetic properties are required. However, gas can arise not only from entrapped hydrogen dissolved during the melting process and evolving as gas during the liquid-solid transition, but also gas can result from aspirated air or gas evolution from molding materials (i.e., cores, binders, etc.). Thus, it is not totally clear in Fig. 2 that the porosity is due to evolution of hydrogen gas. Observable porosity may also very well be caused by lack of feeding, that is, shrinkage, during solidification. In fact, Fig. 2(a) depicts porosity that follows very well the interdendritic structure and hence is very likely shrinkage, not gas. In Fig. 2(b), the rounded porosity would suggest gas evolution during final stages of solidification and hence could very well be caused by hydrogen, However, further microscopic investigation may be necessary, since this porosity may also be caused by entrapped air or lubricant decomposition. In actual practice, it is therefore best to degas the melt to an acceptable level before casting. A typical target value, below which hydrogen-caused porosity is not normally a problem in the final casting, is that the degassing process should achieve a
level of 0.15 mL H2/100 g Al, the usual standard of measurement in a quantified test such as Alscan (Alcan Aluminum Ltd.). If one is instead using the reduced-pressure test as a measure of minimal porosity potential, it is not possible to accurately predict lack of hydrogen-caused porosity by this means alone. However, reaching a high percentage value in the reduced-pressure test related to the specific alloy theoretical density is a good measure, for instance, 2.62 in A356, whose theoretical density is 2.68 g/cm3.
Hydrogen Sources Given the propensity of aluminum to readily absorb hydrogen, there are many ways that hydrogen can and will always be present in a melt to a certain degree. Fuel-fired furnaces (natural gas, oil) have hydrogen available from fossil fuel decomposition. Water vapor is almost always present in the atmosphere, even at high temperatures directly in contact with the melting furnace. Indeed, high humidity conditions exacerbate not only hydrogen dissolution in the melt but also can retard the ability of any melt
Fig. 1
Hydrogen solubility in aluminum casting alloys. Source: Ref 2
186 / Molten Metal Processing and Handling
Fig. 3
Fig. 2
Typical porosity configurations in aluminum-silicon casting alloys. (a) Shrinkage pore found within a casting. (b) Gas pore in an Al-8% Si alloy. (c) Microporosity (gas plus shrinkage). (d) Microporosity (gas plus shrinkage). Source: Ref 2
treatment to remove hydrogen, because a degassed melt can then readily redissolve additional hydrogen. Ingot, additives, and fluxes are other common sources of hydrogen. In the case of fluxes, fluxing compounds are usually mixtures of various salts, including halides, and become very hygroscopic when exposed to air. In addition, the tools used to add and work in fluxes—rakes, spoons, and so on—will often retain flux and oxide debris after use, which will then adsorb moisture from the air as they cool down. Reuse of these tools, without intermediate cleaning, not only poses a potential safety hazard but also can reintroduce hydrogen. Scrap castings, machine turnings and borings, die cast trim press scrap, gates and risers, and sand and other nonmetallic molding material debris are additional sources of hydrogen.
Aluminum Degassing There are many means available to minimize hydrogen pickup or to degas a melt. Noting the increased solubility with temperature, as depicted in Fig. 1, merely minimizing the melting or holding times and temperatures will lessen the hydrogen absorption. Reducing the temperature will result in natural outgassing, given time, which is not always available. Therefore, additional degassing measures are often needed.
The basic principle of practical foundry degassing is that any gaseous component has the ability of collecting hydrogen gas, since all gases are insoluble in aluminum, with the exception of hydrogen, which exists as a single atom. Thus, a purge or process gas can function as in Fig. 3, depicting gas fluxing, which not only collects the hydrogen but also removes particulate matter, that is, inclusions, by flotation. The first application of this principle was mainly through the decomposition of a solid product that reacted in molten aluminum. Historically, the degassing material commonly employed was a hexachloroethane tablet (C2Cl6). When submerged and plunged into the melt, the hexachloroethane decomposed into its respective components. The chlorine component immediately forms a metastable aluminum chloride gaseous compound, which is insoluble in the melt and served as the purge gas. With any purge gas, the monatomic hydrogen diffuses into the purge gas bubble, combines with another hydrogen atom to form molecular hydrogen, and the bubble rises to the surface, where both gaseous species are released. The fluxing components in the hexachloroethane tablet, which are often halide salts, served not only to bind the tablet, but these salt components do aid in wetting of the gas bubble in the aluminum, reducing the interfacial energy
Schematic of gas fluxing in molten aluminum
and allowing easier transmittal of the hydrogen atoms into the collector aluminum chloride gas bubble. This use of hexachloroethane tablets has given way to other degas methodologies, in most instances because of environmental concerns and more efficient technologies. However, some foundries still employ this technique, especially on small crucible furnaces. The second most popular degassing technique in historical use—and still used today (2008)—is a simple flux wand or static lance. This can be a steel pipe or, more commonly, a graphite lance. A purge gas such as nitrogen or argon is injected under low pressure (less than 207 kPa, or 30 psi) into the melt. The insoluble nitrogen or argon gas bubble then rises within the melt, and the hydrogen atoms diffuse to this process gas bubble and form a hydrogen molecule. When the bubble reaches the surface of the melt, the hydrogen is released. If the relative humidity in the atmosphere is particularly high, this will be a dynamic process, because new hydrogen may become absorbed into the melt. Fundamentally, the lance degassing technique is relatively inefficient, as is seen later in this section, in relation to newer technologies. The reason for this is the relatively large gas bubble that emanates from the bottom of the lance and then agglomerates and rises quickly to the surface. The rate of rise increases with bubble size, and the bubble accelerates as it climbs toward the surface. Fig. 4(a) depicts a water model showing that a static lance only influences a small volume of fluid surrounding the lance, unless it is manually moved about. While this is possible in relatively small-diameter melting and holding vessels, it is impractical in larger ones. In Fig. 4(b), the use of either a disperser on a static lance or a rotor on a rotating shaft achieves a much finer bubble size, more dispersion, and better mixing of the gas bubbles and the liquid throughout a greater vessel volume. Visually, when performing this evaluation, one would see that the rate of rise of these smaller bubbles is measurably slower than with the larger gas bubbles in the static lance. This greater residence time increases the probability of reaction between the purge gas bubble and the hydrogen, thus increasing the efficiency. The efficiency of a degassing process is directly
Degassing / 187
Fig. 4
Water model with (a) static lance and (b) disperser or rotating shaft/rotor
Fig. 5
Fig. 6
to manipulate the gas injection into the system, and the shaft/rotor assembly itself. Equipment can be as simple as a well-mounted drive unit or more complex, with a fixed-mast drive-up ladle treatment station or a mobile unit that can be moved from furnace to furnace. Many manufacturers offer these various equipment designs, together with varying degrees of automated controls, and options to inject flux (see the article “Aluminum Fluxes and Fluxing Practice” in this Volume). For the operating foundry person, several important variables must be considered in developing the degassing process in the foundry to meet specific casting requirements. The parameters that must be integrated include:
Relative degassing efficiencies. Source: Ref 4
influenced by the bubble size, as shown in Fig. 5. A classic depiction of the fundamental difference in degassing efficiency between a static lance, a rotating injector, and a porous plug (i.e., a fine pore disperser device) is shown in Fig. 6 (Ref 4). It is clear that high hydrogenremoval efficiencies are possible with a disperser, but a rotating injector with an impeller achieves a much faster rate of reduction. Rotary Impeller Degassing. This technique has been borrowed from the chemical process industry and greatly improves mixing capability between the aluminum and the process gas. This technique was introduced to the aluminum
Degassing efficiency as a function of purge gas bubble size. Source: Ref 5
foundry industry in the mid-1980s and has become the preferred technology of choice for use in degassing wells in melting furnaces, transfer ladles, and in continuous-flow launder systems. In this technique, the purge gas is injected through a bored shaft, usually made of graphite. The incipient gas volume is then fractured into many small bubbles, that is, bubble shear, by the rotor attached to the bottom of the shaft, as the gas leaves the shaft. Many different rotor designs, sizes, and configurations are now employed throughout the industry. The basic components of a rotor degassing system include a mechanical drive unit to rotate the shaft/rotor assembly, a gas source, controls
Initial hydrogen level versus desired final
hydrogen level as determined by evaluation. This is a dynamic, ever-changing process where the initial hydrogen level is influenced by charge materials, melting practice, and relative humidity. Vessel (well or transfer ladle) size and volume. In a continuous-flow launder process, the metal volume residence time must be considered. Available time in a transfer ladle for melt treatment. The metal must then be delivered to the casting station(s) without excessive temperature loss, considering the relationship between rotor type and revolutions per
188 / Molten Metal Processing and Handling minute (rpm), gas volume being injected, avoidance of surface effects (vortexing, splashing), and time necessary and available to achieve the desired degassing results. The interplay among these variables must be determined on a case-by-case basis by the individual foundry and often on the specific casting(s) to achieve the optimum combination of process and equipment parameters. In general, the optimum result desired is to achieve the necessary degassing specification in as short a time as possible, at the lowest cost, and without excessive turbulence. As a general rule, most well-developed foundry degassing practices should achieve the desired degassing level in 5 min when treating a vessel with 455 kg (1000 lb) volume and when optimizing the foregoing parameters of rotor type, rpm, and purge gas flow rate. Selecting a Process Gas. Since all gases except hydrogen are insoluble in molten aluminum, there are theoretically many choices of which gas to use. Nitrogen is the most commonly employed and is the least expensive. Nitrogen creates a wet dross; that is, the dross that is formed is high in metallic aluminum content. Furthermore, since the density of nitrogen is comparable to air overall, there is a tendency to pick up moisture in nitrogen lines as fittings deteriorate and humid air is aspirated into the delivery lines. For this reason, all nitrogen sources should include an in-line dryer. Nitrogen purity should be at least 99.99% for any critical casting applications. Argon, while significantly more expensive than nitrogen, produces a less metal-rich dross. Argon is more inert, and, being heavier, provides a protective cover over the melt during degassing, precluding not only hydrogen absorption but further oxidation as well. It is also easier to keep the argon supply drier and cleaner as hoses and fittings deteriorate. While argon is the preferred process gas for in-line degassing treatment in the production of wrought alloys, it is not extensively used in foundry casting because of the higher cost. Active Halogens. Historically, a small percentage of chlorine (5 to 10%) was used with an inert gas when degassing with a static lance. The chlorine enhanced the reaction between the melt and the hydrogen atoms because it lowered the interfacial tension at the purge gas bubble surface (through the formation of aluminum chloride, a metastable gaseous phase) and facilitated the diffusion of the hydrogen atom into the purge gas bubble. Several commercial mixtures of halogen gases—Freon (CCl2F2) (Dupont) being the most popular—were employed in the past. With the growing concern and difficulties with the environmental impact of chlorine use, sulfur hexafluoride (SF6) gained prominence 20 years ago as a replacement for chlorine. However, subsequent experience showed that SF6 also posed environmental problems with ozone depletion, and the gas is
very corrosive, even in small amounts, to foundry metal structures. Fortunately, with the advent of rotor degassing and the resultant better kinetics of reaction and mixing, it has been found that chlorine does not enhance the degassing rate or efficiency when the purge/process gas bubbles are sufficiently small, as they are with rotor impeller degassing. Rather, when chlorine is used, there is better wetting of the nonmetallic inclusions during the degassing process, so that metal cleanliness is improved, and the burden on any subsequent particulate removal process is lessened, which in turn creates a better-quality cast product. Other Considerations for the Degassing Process. There are other issues that must be taken into account during the degassing process. Temperature. It is well understood that hydrogen absorption in the melt increases with increasing temperature. Conversely, hydrogen solubility and partial pressure of hydrogen gas in the melt decrease rapidly as the melt temperature is decreased. It follows that minimizing the temperature will result in lower hydrogen value. Given sufficient time, allowing the temperature to drop will result in natural outgassing. However, the injection of a purge gas at normal standard temperature and pressure creates an endothermic reaction, and heat will therefore be lost at an accelerated rate. A general ruleof-thumb is that degassing a metal volume of 455 kg (1000 lb) in an unheated refractory vessel for 5 to 10 min will result in a temperature drop of nearly 28 C (50 F). Whatever the actual magnitude of the temperature drop due to gas injection/treatment time, this reduction must be accommodated in relation to the necessary temperature that must be maintained as the treated melt is transferred into the mold or into the casting station(s). A degassing process should nevertheless be carried out at the lowest practical temperature possible. Grain Refining. Most aluminum melts in sand, permanent mold, and low-pressure foundry production require grain refinement to achieve the best mechanical properties in the finished casting. In most instances, grain refiner is added during the degassing process, because the increased mixing and agitation facilitates the dissolution and reaction of the grain refiner itself and provides more uniform dispersion throughout the melt or treated volume. Modification. Historically, the element sodium was used to provide modification of the eutectic silicon component in aluminum casting alloys—a necessary practice, together with grain refining, to produce desired mechanical properties. In current practice, sodium has given way to a strontium addition for many reasons, including safety in handling and in the environment, better consistency, and reduced fade, which is the loss of modification with time. Because both sodium and strontium are prone to oxidation (although strontium vapor pressure is substantially lower than sodium),
and the degassing mixing creates more convection during treatment, degassing should take place before modification, so as not to lose the benefits of the dissolved modifier due to oxidation. If chlorine is used, even in small quantities, in the degassing process, both sodium and strontium will react chemically to form solid chloride reaction products. Dross Formation and Skimming. The action of the purge gas injection process by whatever means—lance, rotor, or porous plug/disperser—will result in flotation of debris and some degree of particulate inclusion content within the melt. Naturally, the degassing process creates dross. This must be thoroughly skimmed off the surface to the greatest practical extent possible to avoid transfer into the casting mold or casting station. Determining Hydrogen Content. To control hydrogen content, one must be able to measure the hydrogen content in the melt and the subsequent result of any hydrogen-removal treatment. Laboratory evaluations are possible using vacuum fusion (Leco technique) and subvacuum fusion on previously solidified test specimens. Obviously, laboratory measurements cannot be used to make real-time decisions but rather are useful in process development. Real-time determinations are limited to the simple reduced-pressure test (RPT) and its variants, and in situ techniques, such as Alscan, continuous hydrogen analysis by pressure evaluation, and others. The RPTs are the simplest, but these tests are a measure of molten metal cleanliness and not absolute hydrogen content per se. The presence of inclusions in the melt during an RPT evaluation will create exaggerated porosity nucleated during solidification and a resultant lower density value. The Alscan test and others of this type do generate a real-time result through the principles of recirculating gas of a known standard hydrogen content and the thermal conductivity of dilute solutions.
Degassing of Magnesium Magnesium behaves similarly to aluminum in its ability to dissolve hydrogen easily in the molten state (Fig. 7). It too will suffer from porosity when cast if the hydrogen is not removed or naturally expelled during solidification. Aluminum and zinc, which are typical alloying elements in magnesium alloys, lower this solubility somewhat. Figure 8 indicates that much more gas is evolved from magnesium alloys containing zinc and aluminum than from pure magnesium upon solidification. Gas porosity occurs in solidified castings when the residual hydrogen exceeds 15 mL/100 g. Zirconium is also a major constituent of certain magnesium alloys and has a marked effect on hydrogen solubility. Hydrogen solubility decreases as the zirconium content of the alloy increases (Ref 9). Because of this low solubility even in the liquid state, hydrogen porosity is minimal in castings of magnesium-zirconium
Degassing / 189 alloys. However, excess hydrogen in the melt readily precipitates as zirconium hydride, thus depleting the alloy of zirconium in solid solution. It is therefore advisable to minimize hydrogen absorption into the melt and/or to employ degassing methods even with this alloy. Sources of gas in magnesium alloy melts are the same as those for aluminum—atmospheric conditions, fluxes, tools, and so on. Flux covers or antiburning agents do not necessarily preclude hydrogen absorption during melting. Like aluminum alloys, molten magnesium metal and alloys can be degassed with hexachloroethane, chlorine, or nitrogen. Hexachloroethane has the added advantage of supplying carbon through the decomposition of the hexachloroethane, with the carbon serving as a grain refiner (see the article “Magnesium and Magnesium Alloy Castings” in this Volume). Chlorine
or chlorine/nitrogen gas purging is optimally performed at temperatures of 725 to 750 C (1335 to 1380 F). In this range, liquid MgCl2 will also be formed, which will retard burning on the surface. Liquid MgCl2 wets the oxides and the magnesium melt surface, facilitating hydrogen escape. Pure chlorine provides for more liquid MgCl2 formation than either gas mixtures or hexachloroethane and therefore creates better degassing. Fluxes higher in MgCl2 that are used as covers during degassing also aid in hydrogen desorption at the flux/metal interface and in subsequent expulsion to the atmosphere. The degassing of magnesium alloys in crucible or pot-type furnaces should be carried out with clean, dry lances, preferably graphite, with thorough mixing for 5 to 10 min. Pressures and volumes should not exceed those required to achieve a mild, gentle roll or small bubble (25 mm, or 1 in., in diameter) on the melt surface. Rare-earth-containing magnesium alloys should only be degassed when required with nitrogen or argon, because chlorination will readily remove these expensive alloying elements.
Degassing of Copper Alloys In the melting and casting of many copper alloys, hydrogen gas absorption can occur because of the generous solubility of hydrogen in the liquid state of these alloys. The solid solubility of hydrogen is much lower; therefore, the gas must be rejected appropriately before or during the casting and solidification process to avoid the formation of gas porosity and related defects (excessive shrink, pinholing, blow holes, and blistering). These defects are almost always detrimental to mechanical and physical properties, performance, and appearance. Different copper
alloys and alloy systems have varying tendencies toward gas absorption and subsequent problems. Other gases, particularly oxygen, can cause similar problems (see the article “Copper and Copper Alloy Castings” in this Volume).
Gases in Copper The type and amount of gas absorbed by a copper alloy melt and retained in a casting depend on a number of conditions, such as melt temperature, raw materials, atmosphere, pouring conditions, and mold materials. Table 1 lists the gases that can be found in a number of copper alloy systems. Hydrogen is the most obvious gas to be considered in copper alloys. Figure 9 (Ref 11) shows the solubility of hydrogen in molten copper, tin, and copper-tin alloys. As in aluminum and magnesium alloys, solubility is reduced in the solid state; therefore, the hydrogen must be removed prior to casting or rejected in a controlled manner during solidification. Alloying elements have varying effects on hydrogen solubility (Fig. 10). There are many potential sources of hydrogen when melting copper alloys. For the most part, these are identical to the sources of hydrogen in aluminum melting. Oxygen also presents a potential problem in most copper alloys, but oxygen alone does not cause problems in copper alloys. In the absence of hydrogen, oxygen alone may not cause problems, because it has limited solubility in the melt. However, it forms a completely miscible liquid phase with the copper in the form of cuprous oxide (Fig. 11). During solidification, the combination of cuprous oxide and hydrogen can give rise to casting porosity resulting from the steam reaction (discussed later). Sulfur gases have significance in primary copper through the smelting of sulfide ores
Fig. 7
Solubility of hydrogen in molten magnesium at 101 kPa (760 torr). Source: Ref 9
Fig. 8
Effect of alloying elements on hydrogen solubility in liquid magnesium. AC, as-cast; HT, heat treated. Source: Ref 9
190 / Molten Metal Processing and Handling Table 1 Summary of gases found in copper alloys Alloy family
Pure copper Copper-tin-leadzinc alloys Aluminum bronzes Silicon brasses and bronzes Copper-nickels
Gases present
Water vapor, hydrogen Water vapor, hydrogen, CO Water vapor, hydrogen, CO Water vapor, hydrogen Water vapor, hydrogen, CO
Remarks
Approximate hydrogen/water vapor ratio of 1. Higher purity increases the amount of water vapor and lowers hydrogen. Lead does not affect the gases present. Higher tin lowers total gas content. Increased zinc increases the amount of hydrogen, with a loss in water vapor. The presence of 5 wt% Ni in alloy C95800 causes CO to occur rather than water vapor. Lower aluminum leads to higher total gas contents. Approximate hydrogen/water vapor ratio of 0.5. Increased zinc decreases hydrogen and increases water vapor. All three gases are present up to 4 wt% Ni, after which only CO and hydrogen are present. Hydrogen increases with increasing nickel up to 10 wt% Ni but is decreased at 30 wt% Ni.
Source: Ref 10
Fig. 11
Fig. 9
Solubility of hydrogen at 1 atm (760 mm Hg) in pure copper, pure tin, and copper-tin alloys. Source: Ref 11
and in the remelting of mill product scrap containing sulfur-bearing lubricants. Sulfur dioxide is the most probable gaseous product. Foundry alloys and foundry processing usually do not experience sulfur-related problems unless high-sulfur fossil fuels are used for melting. Carbon can be a problem, especially with the nickel-bearing alloys. The nickel alloys have extensive solubility for both carbon and hydrogen. Carbon can be deliberately added to the melt, along with an oxygen-bearing material such as nickel oxide. The two components will react to produce a carbon boil, that is, the formation of CO bubbles, which collect hydrogen as they rise through the melt (Ref 13). If not fully removed, however, the residual CO can create gas porosity during solidification. Nitrogen does not appear to be detrimental or to have much solubility in most copper alloys. However, there is some evidence to suggest that nitrogen porosity can be a problem in cast copper-nickel alloys (Ref 14).
Fig. 10
Effect of alloying elements on the solubility of hydrogen in copper. Source: Ref 12
Water vapor can exist as a discrete gaseous entity in copper alloys (Ref 10, 15, 16). In many copper alloys, the most important gas causing porosity is water vapor (Ref 11). Water itself is not dissolved, but dissolved atomic oxygen and hydrogen in the melt combine to form water until the oxygen becomes depleted; hydrogen is produced as a separate species. Water vapor is evolved from solidifying copper alloys, which always have some residual dissolved oxygen. Thus, gas porosity can occur even with low hydrogen content when oxygen is present. Gas porosity occurs when the gas pressure exceeds 1 atm (760 mm Hg). To prevent such porosity, the solid solubility limits must not be exceeded, and the sum of individual gas partial pressures must be less than 1 atm (760 mm Hg) through the solidification process (Ref 11). Testing for Gases. There are essentially three ways to determine the presence of gas in a copper alloy melt. The easiest and simplest is a chill test on a fracture specimen. In this
Copper-oxygen phase diagram. Source: Ref 12
method, a standard test bar is poured and allowed to solidify. The appearance of the fractured test bar is then related to metal quality standards (Ref 17). The second method is an RPT similar to the Straube-Pfeiffer test used for aluminum alloys. In recent years, considerable development and definition have been given to this test as adapted for copper alloys, because the different classes of copper alloys have different solidification characteristics that affect the response to the test (Ref 10, 18–20). The objective of the test is to use a reduced pressure only slightly less than the dissolved gas pressure so that the surface of the sample mushrooms but does not fracture (Fig. 12). The test apparatus is shown schematically in Fig. 13. The key to establishing controllable results, that is, unfractured but mushroomed test sample surfaces, is to provide wave-front freezing. This is accomplished by proper selection of vessel materials to achieve the correct solidification characteristics. Because the freezing ranges of copper alloys vary from short to wide, different materials must be used. The various material combinations that have proved successful are described in Ref 10.
Degassing Methods There are several kinds of treatments which may be employed for the improvement of melt quality in copper alloys, such as oxidationdeoxidation practice, zinc flaring, inert gas fluxing, and the use of solid degassing fluxes. Metallurgical principles for common copper alloy classes are discussed in detail in Ref. 21. Oxidation-Deoxidation Practice. The steam reaction previously mentioned is a result of both hydrogen and cuprous oxide being present in the melt. These constituents react to form steam, resulting in blow holes during solidification as the copper cools: 2H þ Cu2 O ! H2 OðsteamÞ þ 2Cu The proper deoxidation of a copper alloy melt will generally prevent this steam reaction,
Degassing / 191
Fig. 12
Effect of pressure on the appearance of copper alloy reduced-pressure test samples containing the same amount of gas. (a) Pressure of 7 kPa (55 torr) results in surface shrinkage. (b) At 6.5 kPa (50 torr), a single bubble forms. (c) Boiling and porosity occur at 6 kPa (45 torr).
Fig. 13
Fig. 14
Schematic of the reduced-pressure test apparatus used to assess amounts of dissolved gas in copper alloys
Hydrogen/oxygen equilibrium in molten copper. Source: Ref 10
Fig. 15
Influence of zinc content on boiling point or vapor pressure in copper alloys. Source: Ref 12
although excessive hydrogen alone can still cause gas porosity if it is not expelled from the casting before the skin is completely solidified. Fortunately, there is a mutual relationship between hydrogen and oxygen solubility in molten copper (Fig. 14). Steam is formed above the line denoting equilibrium concentration, but not below. Consequently, as the oxygen content is raised, the capacity for hydrogen absorption decreases. Therefore, it is useful to provide excess oxygen during melting to preclude hydrogen entry and then to remove the oxygen by a deoxidation process to prevent further steam reaction during solidification. This relationship gave rise to the Pell-Walpole oxidation-deoxidation practice for limiting hydrogen, especially when fossil-fuel-fired furnaces were used for melting. The melt was deliberately oxidized using oxygen-bearing granular fluxes or briquetted tablets to preclude hydrogen absorption during melting. The melt was subsequently deoxidized to eliminate any steam reaction with additional hydrogen absorbed during pouring and casting. Zinc Flaring. The term zinc flare applies to those alloys containing at least 20% Zn. At this level or above, the boiling or vaporization point of copper-zinc alloys is close to the usual pouring temperature, as shown in Fig. 15 (Ref 12). Zinc has a very high vapor pressure, which precludes hydrogen entry into the melt. Zinc may also act as a vapor purge, removing hydrogen already in solution. Furthermore, the oxide of zinc is less dense than that of copper, thus providing a more tenacious, cohesive, and protective oxide skin on the melt surface and increasing resistance to hydrogen diffusion. In zinc flaring, the melt temperature is deliberately raised to permit greater zinc vapor formation. Because some zinc loss may occur, zinc may have to be replenished to maintain the correct composition. Inert Gas Fluxing. With gas fluxing, an inert collector or sparger gas such as argon or nitrogen is injected into the melt with a graphite fluxing tube. The bubbling action collects the hydrogen gas that diffuses to the bubble surface, and the hydrogen is removed as the inert gas bubbles rise to the melt surface. The reaction efficiency depends on the gas volume, the depth to which the fluxing tube or lance is plunged, and the size of the collector gas bubble generated. Again, finer bubble sizes have higher surface-area-to-volume ratios and therefore provide better reaction efficiencies. Figure 16 depicts the amount of purge gas necessary to degas a 450 kg (1000 lb) copper melt. Figure 17 shows the response of an aluminum bronze alloy to nitrogen gas purging. The curve for oxygen purging is also shown. Solid Degassing Fluxes. Other materials can be used to provide inert gas purging. A tabletted granular flux such as calcium carbonate (CaCO3), which liberates CO2 as the collector gas upon heating, has been successful in degassing a wide variety of copper alloys. This type of degassing is simpler than nitrogen but must
192 / Molten Metal Processing and Handling
Fig. 16
Amount of purge gas required to degas a 450 kg (1000 lb) copper melt. Source: Ref 20
Fig. 17
Fig. 18
Comparison of the effectiveness of solid degassing flux versus nitrogen purging. Source: Ref 21
be kept dry and plunged as deep as possible into the melt with a clean, dry plunging rod, preferably made of graphite. This type of degasser may also have the advantage of forming the collector gas by chemical reaction in situ. This results in an inherently smaller initial bubble size than injected gas purging for better reaction efficiency. Such an advantage can be inferred from Fig. 18, although after 5 min of degassing, the end results are the same. Other solid degassers in the form of refractory metals or intermetallic compounds, such as calcium-manganese-silicon, nickel-titanium, titanium, and lithium, are effective in eliminating porosity due to nitrogen or hydrogen by their ability to form stable nitrides and hydrides. Again, maintaining dry ingredients, deep plunging, and stirring or mixing will enhance their effectiveness. For best results and optimum reaction efficiency, such degassers should be wrapped or encapsulated in copper materials to provide controlled melting when plunged. Alternatively, copper master alloys containing these elements are available.
Effect of nitrogen flow rate during purging on the residual gas pressure remaining in an aluminum bronze melt. The cure for oxygen purging is also shown. Source: Ref 19, 22
For both inert gas fluxing and solid degassing additions, the sparging gas reaction should not be so violent as to splash metal and create an opportunity for gas reabsorption. Furthermore, a melt can be overdegassed; an optimum amount of residual gas remaining in the melt helps to counter localized shrinkage in longfreezing-range alloys such as leaded tin bronzes. Vacuum degassing is not generally applied to copper alloys, although it can be very effective. However, the cost of the equipment necessary is relatively high, and there may be a substantial loss of more volatile elements having a high vapor pressure, such as zinc and lead. Furthermore, significant superheat (up to 150 C, or 300 F) may be required to accommodate the temperature drop during the degassing treatment, further aggravating the vapor losses. Auxiliary Degassing Methods. There are other ways to degas a copper alloy melt than using a specific treatment. The technique of lowering the melt superheat temperature, if possible, and holding the melt (in a dry, minimal gas environment) provides outgassing simply by lowering the equilibrium liquid-state solubility of the gas. During casting, the use of chills to provide directional solidification, particularly for the long-freezing-range alloys such as the red brasses and tin bronzes, results in less tendency for gas porosity. In the mold-metal reaction, previously degassed and deoxidized metal containing excess phosphorus or lithium deoxidants can react with green sand containing moisture. Hydrogen is liberated and absorbed as the metal comes into contact with the sand. This can be
minimized by correct deoxidant additions. In addition, finer facing sands or mold coatings (inert or reactive, such as sodium silicate or magnesia) can be used; these confine any reaction to the mold/metal interface area and retard hydrogen penetration into the solidifying skin of the cast metal.
REFERENCES 1. R. Monroe, Porosity in Castings, Trans. AFS, Vol 113, 2005, p 519–546 2. J.E. Gruzleski and B.M. Closset, The Treatment of Liquid Aluminum-Silicon Alloys, AFS, 1990 3. D. Apelian, Clean Metal Processing of Aluminum, Proceedings, Materials Solutions Conference ’98 on Aluminum Casting Technology, ASM International, 1998 4. G. Sigworth and T. Engh, Chemical and Kinetic Factors Related to Hydrogen Removal from Aluminum, Metall. Trans., Vol 138, 1982, p 447 5. G. Sigworth, A Scientific Basis for Degassing Aluminum, Trans. AFS, Vol 95, 1987, p 73 6. D.V. Neff, Understanding Aluminum Degassing, Mod. Cast., May 2002 7. D. Apelian and M.M. Makhlouf, Ed., High Integrity Aluminum Die Casting, NADCA, April 2004 8. M.M. Makhlouf, L. Wang, and D. Apelian, Measurement and Removal of Hydrogen in Aluminum Alloys, AFS Special Report, 1998 9. E.F. Emley, Principles of Magnesium Technology, Pergamon Press, 1966
Degassing / 193 10. P.K. Trojan, T.R. Ostrom, and R.A. Flinn, Melt Control Variables in Copper Base Alloys, Trans. AFS, 1982, p 729; “High Conductivity Copper Alloys,” Bulletin 43, Foseco, Inc., 1972 11. Casting Copper-Base Alloys, 2nd ed., American Foundry Society, 2007, p 128–131 12. R.J. Kissling and J.F. Wallace, Gases in Copper Base Alloys, Foundry, 1962, 1963 13. M.P. Renatv, C.M. Andres, G.P. Douglas, and R.A. Flinn, Solubility Relations in Liquid Copper-Nickel-Carbon-Oxygen Alloys, Trans. AFS, 1976, p 641
14. B.N. Ames and N.A. Kahn, Gas Absorption and Degasification of Cast Monel, Trans. AFS, 1947, p 558 15. T.R. Ostrom, P.K. Trojan, and R.A. Flinn, Gas Evolution from Copper Alloys—A Summary, Trans. AFS, 1981, p 731 16. T.R. Ostrom, P.K. Trojan, and R.A. Flinn, The Effects of Tin, Aluminum, Nickel, and Iron on Dissolved Gases in Molten Copper Alloys, Trans. AFS, 1980, p 437 17. Casting Copper Base Alloys, American Foundrymen’s Society, 1984 18. T.R. Ostrom, R.A. Flinn, and P.K. Trojan, Gas Content of Copper Alloy Melts—Test
19. 20. 21. 22.
Equipment and Field Test Results, Trans. AFS, 1977, p 357 T.R. Ostrom, P.K. Trojan, and R.A. Flinn, Dissolved Gases in Commercial Copper Base Alloys, Trans. AFS, 1975, p 485 R.J. Cooksey and R.W. Ruddle, New Techniques in Degassing Copper Alloys, Trans. AFS, Vol 68, 1960 D.V. Neff, Fluxing, Degassing, Deoxidation of Copper Base Alloys, Trans. AFS, 1989, p 439 P.K. Trojan, S. Suga, and R.A. Flinn, Influence of Gas Porosity on Mechanical Properties of Aluminum Bronze, Trans. AFS, 1973, p 552
ASM Handbook, Volume 15: Casting ASM Handbook Committee, p 194-205 DOI: 10.1361/asmhba0005351
Copyright © 2008 ASM International® All rights reserved. www.asminternational.org
Molten-Metal Filtration D.V. Neff, Pyrotek Inc. Robert P. Pischel, FOSECO Metallurgical Inc.
FILTRATION has become a primary means of removing solid impurities from nonferrous molten metal. Aluminum alloys, especially with their tendency toward oxidation, often contain impurities (inclusions) that are deleterious to physical, mechanical, electrical, and aesthetic properties. In the handling of molten aluminum, it is fairly common to use filters as a part of the melting unit (in-furnace filtration) and in the gating and/or riser system (in-mold filtration). Zinc and copper alloy melts also are filtered but to a much lesser extent than aluminum. In almost all instances of metal casting, the presence of dross, inclusions, or particulate impurities detracts from casting performance and can be a major contributor to scrap castings and lost productivity and profitability for the foundry. For example, the presence of dross, inclusions, or particulate impurities detracts from the natural fluidity of any melt (Fig. 1) (Ref 1). Mechanical properties of the cast product can also be degraded, particularly in elongation/ductility and fatigue strength, which are most sensitive to the presence of inclusions. Inclusions also are harder than the surrounding matrix and give rise to tool chatter, tool wear, and ultimately, tool breakage during machining operations. In addition, inclusions may result in poor continuity, such as creating leakers in
pressuretight castings by nucleating porosity formation during solidification or by causing obvious discontinuities and blemishes in cosmetic surfacefinishing operations such as powder coating, plating, painting, polishing, and anodizing. Thus, inclusion-related defects should be minimized to the greatest extent possible to prevent scrap castings, lost productivity, and lower profitability. This article describes the methods of in-furnace and in-mold filtration, with emphasis on the filtration of molten aluminum. Moltenmetal filtration is an ongoing area of interest in aluminum foundry practice (Ref 2–6). For example, the following table comprises typical quality improvements in several aluminum foundries using in-furnace filtration: Foundry
Al alloy
Result from in-furnace filtering
380 319 380 319 380 518
Hardspots eliminated Leakers reduced 75% Tool life extended 600% Scrap reduced 95% Tool breakage reduced 95% Scrap eliminated
A B C D E F
Sources of Inclusions Undissolved particulate contaminants (inclusions) include slag (ferrous), dross (nonferrous), nonmetallic impurities (oxides, sulfides, nitrides), or insoluble intermetallic precipitates (see the article “Inclusion-Forming Reactions” in this Volume). Inclusions arise from many sources. Charge materials can contain nonmetallic inclusions from unfiltered primary or secondary ingot. Scrap materials may contain molding sand, debris from back-charged gates and runners, metalworking lubricants, paint and coatings, plastics, or unsorted and unwanted heavy-metal contaminant oxides (iron, manganese, chromium, and so on). Furnace tools, such as skimmers, puddlers, and rakes, can also add solidified oxide skins back into the melt if they are not cleaned before use. Factors that influence the formation of inclusions include: Excessive
Fig. 1
Effect of metal cleanliness on fluidity of an aluminum die casting alloy. Source: Ref 1
temperatures during melting accelerate the oxidation of aluminum and alloying elements. Where fluxes are used,
incomplete skimming or separation of fluxing salts often results in contamination and salt-wetted oxide inclusions. Grain-refiner additions can be a source of inclusions. Intermetallic compound precipitation may also contribute to inclusions. Iron, chromium, and manganese form intermetallics with aluminum, and the formation of intermetallic sludge is a common problem in die casting alloys. Refractory inclusions from the erosion and wear of melting and holding vessels may add alumina, spinel, silica, and silicon carbide inclusions to nonferrous alloy melts. Aluminum Inclusions. Inclusions in molten aluminum foundry processing arise from several sources. These inclusions exist as oxides present as skins on charge materials; from oxidation of the melt itself; from erosion of refractory vessels used for melting, transfer, and holding of molten aluminum; from additives such as master alloys; and grain refiners and modifiers as well. The various types of solidphase inclusions in molten aluminum alloys include:
Oxides (Al2O3, MgO) Spinels (Mg2AlO4) Borides (TiB2, VB2, ZrB2) Carbides (Al3C4, TiC) Intermetallics (MnAl3, FeAl3) Nitrides (AlN) Exogenous refractory inclusions (oxides and/ or carbides of iron, silicon, aluminum, etc.)
The size and shape of the inclusions vary from globular (oxides and borides) to stringerlike (oxides and spinels). The latter are particularly insidious in their effect on castability and the properties of the casting. Aluminum oxide is present in virtually all aspects of aluminum alloy casting production, because aluminum has a high oxidation potential. Prime metal in either liquid or solid form, alloy additions, and so on will always have an aluminum oxide skin. Melting in nearly all processes is done in air, thus providing another source for aspirated air to enter, especially in closed or even
Molten-Metal Filtration / 195 flue-operated larger furnaces that are not maintained under positive pressure. Foundry returns— an essential part of the foundry economic charge mix—also contain not only aluminum oxide skin but also can contribute sand debris, machining lubricant residues, die casting release agents, plunger lube, and other fluids that may contain some level of solids content. In addition, alloy and metallic impurity components in excess of the solubility limit form intermetallic compounds that do not dissolve during processing and hence can become inclusions in the solidified metal. In high-pressure die casting alloys and most aluminum-silicon shape casting alloys, an intermetallic compound known as sludge can form with sufficient constituent concentration (iron, manganese, chromium) and low temperature to form such an intermetallic. Additionally, segregated heavy metals (zinc, copper, iron, manganese) can also form a segregated agglomerate in the melting furnace, which has the potential to be transferred out and into a casting. During fluxing operations in a melting furnace or transfer ladle, residual flux inclusions may not become fully skimmed or removed from the surface nor have time to either float out or sediment. If these flux inclusions do become entrapped within the casting, they create sources for galvanic corrosion to take place. Typical solid-phase inclusions that may be found in aluminum casting alloys are listed in Table 1. Oxide-based inclusions are varied and will exhibit a spectrum of densities—some with specific gravity lower than the melt (nominally 2.4 g/cm3, or 0.087 1b/in.3) and some higher than the melt. Given sufficient time, separation of the solid particulate from the liquid is possible to a degree, based on the density difference. Thus, lighter particles could float to the surface; heavier particles will sink or sediment to the bottom of a melting or casting vessel or furnace. As Fig. 2 demonstrates, however, this phenomenon is time-dependent as well as size-dependent. In casting alloys, a discrete particulate soon becomes an agglomerate. Surface tension/interfacial energies usually dictate that complexes of inclusions have a bulk density nearer that of the melt, and hence, separation purely by gravity means becomes retarded. Liquid-phase inclusions can occur in molten aluminum alloys. The most prevalent is MgCl2 arising from the chlorination of magnesium-containing melts. These inclusions are not generally removed to a great extent by conventional filtration techniques and therefore must be avoided. Preventing excessive and inefficient chlorination of the melts during degassing and thorough skimming at least help to avoid such inclusions. The nature of these and other inclusions in aluminum alloys is discussed in more detail in Ref 7 and 9 to 12. Inclusions in Magnesium Alloys. The inclusions found in magnesium melts are similar in many cases to those found in aluminum—oxides, spinels, furnace refractory particles, and so on. In addition, alloying elements (zirconium and the rare earths) lead to oxides of these elements
as well as intermetallic compound precipitates. If the fluxing process is used to clean magnesium, fluxing salt inclusions may also be present. Inclusions in Copper Alloys. In addition to the usual inclusions arising from oxides, fluxing salts, and intermetallics, copper oxide inclusions and phosphorus pentoxide (from deoxidation) may be present in copper alloys if the melt is not allowed to settle or if it is inadequately skimmed before pouring and casting. Inclusions in Zinc Alloys. In the zinc foundry and die casting alloys containing aluminum, nonmetallic oxide inclusions are
usually less important than intermetallic inclusions. Iron-zinc and iron-aluminum intermetallic compounds may also be present, especially above 480 C (900 F) in melting in iron-base vessels. These particles may be small crystals less than 5 mm (200 min.) in size or may agglomerate into millimeter-sized aggregates. Other heavy metals present, such as lead, cadmium, nickel, manganese, and chromium, can also lead to intermetallic compound precipitation (for example, Mn2Al5, Ni2Al3, and CrAl4). These compounds (often called aluminides) can form when the concentration of impurity exceeds
Table 1 Common types of inclusions in aluminum foundry melts Inclusion type
Chemical formula
Density Form
g/cm3
Size range lb/in.3
mm
min.
8–1200 400–2 105 4–200 400–2 105 4–200 400–2 105 20–200
Oxides Alumina
Al2O3
Particles, skins
3.97
0.143
Magnesia
MgO
Particles, skins
3.58
0.129
Spinel
MgAl2O4
Particles, skins
3.60
0.130
Silica
SiO2
Particles
2.66
0.096
0.2–30 (particles) 10–5000 (skins) 0.1–5 (particles) 10–5000 (skins) 0.1–5 (particles) 10–5000 (skins) 0.5–5
Salts Chlorides Fluorides
Varies Varies
Particles Particles
1.98–2.16 1.98–2.16
0.072–0.078 0.072–0.078
0.1–5 0.1–5
4–200 4–200
Carbides Aluminum carbide Silicon carbide
Al4C3 SiC
Particles Particles
2.36 3.22
0.085 0.116
0.5–25 0.5–25
20–1000 20–1000
Nitrides Aluminum nitride
AlN
Particles, skins
3.26
0.117
10–50
400–2000
Borides Titanium boride
TiB2
4.50
0.163
1–30
40–1200
Aluminum boride
AlB2
Agglomerated particles Particles
3.19
0.115
Sludge Various
Al(FeMnCr)Si
Particles
>4.0
>0.145
Source: Ref 7, 8
Fig. 2
Sedimentation/flotation of inclusions in aluminum melts
0.1–3
4–120 ...
...
196 / Molten Metal Processing and Handling the liquid solubility level of the element, which is often quite low (<0.02% for manganese, silicon, chromium, and nickel in die casting alloys). The intermetallic particles can be lighter than the base alloy and therefore rise to the top dross. Fluxing salt inclusions and exogenous refractory particles can also be present, as with the other nonferrous alloy systems. Oxide inclusions in zinc are somewhat limited because of the high positive vapor pressure of zinc. However, alloys containing aluminum can be subject to greater oxidation of the aluminum constituent. Oxide inclusions can also result during melting or casting when excessive turbulence is encountered; thus, mechanical stirring or molten metal pumping should be controlled. Other nonmetallic inclusions arise from refractory vessel erosion and include silica, alumina, and silicon carbide. There is an appreciable density difference between these oxides and the base melt, unlike in aluminum; for this reason, gravity separation and thorough removal are substantially easier.
Removal of Inclusions There are three basic methods of mechanically removing or separating inclusions from molten metal: Sedimentation Flotation Positive filtration
Fig. 3
Two modes of positive filtration. (a) Depth filtration. (b) Cake filtration
The first two processes—sedimentation and flotation—are based on density differences between the liquid and solid phases. Actual separation capability is based not only on the density differences but also particle size. For example, rotary degassing is a method that removes inclusions by flotation. It is usually effective in removing inclusions larger than 40 mm (1600 min.) in size. Figure 2 is an application of Stokes’s law governing phase separation as applied to inclusions in molten aluminum. The greater the density difference from molten aluminum and the larger the particle size, the faster the rate of rise (flotation to the melt surface) or fall (sedimentation to the bottom of the furnace or holding vessel). Thus, in many metal casting operations, a holding or quiescent period within a furnace will result in much of the inclusion content being able to be removed by skimming the surface, or avoidance of the sedimented inclusions by careful pouring so as not to entrain the inclusions within the pouring stream or into the transfer ladle or casting mold. However, two factors combine to render such gravity-related processes less than efficient—at least for light nonferrous metals such as aluminum. First, there is a high surface tension/interfacial surface energy between most oxide inclusions in aluminum and the melt, which makes actual separation difficult. Secondly, many oxide inclusions, even small ones, can be detrimental to casting properties if concentrated into the casting during solidification. It can be seen that small-sized inclusions have very slow rates of separation. Additionally, many inclusions and/or inclusion complexes will have bulk
Fig. 4
densities that are very similar to the density of the melt itself and thus will also not separate easily or quickly. Consequently, positive filtration will need to be employed to remove these. Positive filtration involves the passing of molten metal through a porous device (a filter) in which the inclusions contained in the flowing metal are trapped or captured by one or more filtration mechanisms. With any filter, there are two mechanisms of positive filtration (Fig. 3): Cake filtration Depth-mode (or deep-bed) filtration
Cake filtration occurs when there is a buildup of particulate (filtrate) on the inlet surface of a filter. Particles that are larger sized than the inlet openings on a filter will naturally accumulate at the surface to form a cake. In most instances, however, inclusion particles are considerably smaller than the filter pore openings, but the filter body itself connecting the pores serves to slow the molten-metal flow through the filter, such that the particles begin to adhere to the surface. Fortunately, even though these inclusions begin to build up, the accumulation remains quite lacy and permeable to continued metal flow, up to an extent. Depth-mode filtration occurs within the body of a filter. In this mode, inclusion particles flowing with the liquid are subject to changes in velocity and direction created by the internal complexity or tortuosity of the flow path within the filter. The inclusions become captured or adhere to the internal surfaces of the filter by several possible mechanisms (Fig. 4) of diffusion (important for inclusions less than 1 mm, or
Transport and capture mechanisms of depth filtration. (a) Diffusion. (b) Direct interception. (c) Inertia or sedimentation. (d) Fluid dynamics effects on particulate transport
Molten-Metal Filtration / 197 40 min.), direct interception, sedimentation, and fluid dynamics. Filtration is a matter of probability, and so, the greater the flow path through the filter, the greater the opportunity for an inclusion to be subject to changes in direction and velocity and come into contact with the filter internal surface for adhesion and capture. Depth filtration is the filtering mode most prevalent in bed filtration. Depth filtration is capable of removing particles much smaller than 30 mm (1200 min.) as they are carried in a molten-metal flow stream through a very tortuous path in a deep-bed filter (Fig. 5), which is commonly used in mill product casthouses. The tortuosity provides increasing probability for a two-step particle-capture mechanism involving transport from the liquid metal flow and attachment to the filter surface itself. Bed filters composed of tabular Al2O3 are quite common in the casting of mill products, where up to 1800 Mg (2000 tons) of a single alloy may be filtered in-line between holding furnace and casting station before replacement is necessary. Bed filters provide the best source of fine particle filtration, but their size, expense, and single-alloy use preclude their economic use in multiple-alloy wrought casting (ingot) or foundry casting operations. Bed filters are capable of providing nearly the highest filtration efficiency attainable, but because the bed is not rigid, movement and flow channeling can occur, which can dump inclusions.
Filter Characteristics Molten-metal filters obviously require hightemperature integrity in terms of strength, refractoriness, thermal shock resistance, and corrosion resistance. The filters may be placed directly in either the melting or casting furnace (or sometimes both) in the cases of in-furnace applications, or the filters may be in the running system of the casting mold (in-mold filters). In either application, it is important to understand the basic characteristics of the filters. The mechanical removal of undissolved particulates from molten metal follows the same
Fig. 5
basic physics as the filtration of particulate from any liquid. The basic principles of filtration are: Larger inclusions are more easily removed
from melts than smaller ones.
Lower flow rates of metal through a filter
result in greater efficiency of inclusions removal. Filters with larger inlet surface areas have greater inclusion-holding capability and longer filter life, for a given incoming inclusion concentration. Filter bodies with greater internal tortuosity or filter depth provide a higher probability of inclusion capture. However, some basic misunderstandings persist in understanding the characteristics of filters. One misconception is that filters can catch everything unwanted in the melt. As a basic physical principle, mechanical filtration can only remove the insoluble particulates in a liquid. Mechanical filtration cannot remove contaminants that are liquid to semiliquid themselves or dissolved constituents such as hydrogen gas, various low-melting fluxes, or alloying eutectic components such as silicon. In addition, it is critical to understand that reactive molten metals (especially aluminum) react with the environment for the entire time of its molten state. Aluminum is particularly reactive with the atmosphere and reacts easily with oxygen, forming solid particulate aluminum oxide, which is filterable, and hydrogen, which is soluble and therefore not filterable. This includes the time spent traveling in the runner system and filling the mold postfilter placement. Hence, some particulate contamination may still occur if the molten aluminum is allowed to travel in a turbulent state postfiltration. This would be especially true in sand molds where lower compaction, incomplete bonding, and turbulence of metal flow will create the possibility of sand erosion. Secondly, there are significant differences in the mechanism of various types of media described as a filter. The two most common
Example of one type of bed filter used for in-line processing
media found in today’s (2008) aluminum foundries are screens and refractory foam filters (Fig. 6, 7). Screens, for all intent and purpose, are two-dimensional structures with evenly spaced round or square holes. Regardless of composition, that is, metal or fiberglass being the most common, they function primarily as sieves. Sieves can only remove particulate larger than the opening in the sieve or, in this case, the screen. Foam filters, on the other hand, have two completely different and distinct modes of filtration that allow them to act as true filters rather than just simple sieves. Because foam filters are three-dimensional and have a random pattern of openings, they exhibit two completely different behaviors. On the surface, they behave similarly to screens in that particulate larger than the openings of the pores of the foam will be trapped on the inlet side of the foam. This is often referred to in the literature as cake filtration, as previously noted. However, the metal must continue to pass through the pores and cells that make up the internal honeycomb structure of the filter. Within the structure of the filter, due to rapid changes in pressure and metal velocity, particulate much smaller than the actual pores and cells themselves may become trapped. This is commonly referred to as deep-bed or depthmode filtration. It is important to understand that typically, sieving or cake filtration can be nearly 100% under normal conditions for large particulates.
Fig. 6
In-mold ceramic foam filters
Fig. 7
In-mold screens
198 / Molten Metal Processing and Handling Unless excessive force is applied, all particulates greater than the opening size of the holes cannot pass through the holes. On the other hand, deep-bed filtration can only function as a percentage-reduction filtration mechanism dependent on many variables: metal fluidity, velocity, foam filter porosity, foam filter thickness, particle size of the contaminant to be trapped, and so on. That is, some percentage of extremely small particulate will be able to escape even the best foam filter. It therefore follows that the choice of filter medium is based at least in part on the requirement for casting quality. For those castings where very small particulates do not contribute to quality issues, screens may be adequate. However, as the quality requirement of the casting increases, a point will be reached where screens will be incapable of providing the level of melt cleanliness required to meet the casting quality demands. At that point, a ceramic foam honeycomb filter structure becomes the best option to achieve the required quality. The one characteristic sometimes missing in the discussion of aluminum filtration is the function of refractory material in the foam structure. Unlike iron, steel, and copper alloys, the majority of particulate impurities in aluminum alloys is aluminum oxide, the most prevalent impurity, which exhibits very little physical or chemical affinity for most common refractory materials. Unlike iron, for example, where sticky silicacious slags will physically stick to many suitable refractories, aluminum oxide (alumina) in aluminum is “dry” and does not exhibit a particular physical affinity to any refractory. For this reason, there has never been any reliable data presented that support a claim of one refractory being superior over another in nonmetallic reduction in molten aluminum alloys, all other attributes of the foam filter being equal. The refractory composition can be responsible for other filter behaviors, however, such as strength (both cold and hot), friability, thermal shock, wettability, thermal capacity (based on density of the material), thermal cycling (remelt behavior), and soon. For these aforementioned reasons, alumina, silicon carbide, mullite, and graphite have been used as media in foam filters for aluminum filtration. Although they offer different physical refractory performance for any given porosity and thickness, they all behave similarly in particulate-retention capability. Once the appropriate refractory is chosen, the most important physical characteristic of the foam filter is its porosity. At this time (2008), there is no true, accurate universal definition in the industry to describe the porosity of a foam filter. The nomenclature of pores per inch (ppi) is a carryover from the foaming process that is used to make all kinds of foam products, such as cushioning materials, packaging, air filtration substrates, carpet padding, and so on, as well as the reticulated polyurethane foam used as substrate for refractory foam filters.
In general, there are three basic classes of foam porosity most commonly used in the aluminum foundry industry: Coarse (10 ppi) Medium (20 ppi) Fine (30 ppi)
However, the porosity of the foam substrate is only a portion of the final story. No matter what foam is used, the refractory slurry applied to the foam can have different effects on the final foam filter flow characteristics, all else being equal. Therefore, comparing one manufacturer’s 10 ppi filter to another’s 10 ppi filter or even 15 ppi filter leads to confusion on the part of the foundry. Regardless of the ppi rating, what a foundry is really purchasing is the pressure drop that the filter provides the metal stream between the entrance face of the filter and the exit into the balance of the running system. For any given porosity rating and filter thickness, the pressure drop should be consistent, with minimum expected variation. As the pressure drop increases (as is the case with finer-grade filters), the filter has a greater ability to retain a higher and higher percentage of small and smaller particulates. Conversely, as the pressure drop increases, the expected flow rate in pounds per second will decrease. This often creates a basic dilemma. A common mistake made in filter printing is not allowing enough of an increase in the area of the filter versus the runner it is placed in in order to compensate for the filter pressure drop and correspondingly to sustain the necessary flow rate. Under normal circumstances, the rule of thumb is to allow for a 4-to-1 ratio of the usable filter area (excluding the print) versus the runner area in order to compensate for the filter pressure drop. When using coarser filters (10 ppi), this ratio can be somewhat reduced. When finer filtration is required (20 or 30 ppi) along with fast flow rates, the only alternative is to increase the usable surface area of the filter. Although most people focus on the particulate-removal capabilities of refractory foam filters as their primary function, it is, in fact, only a part of the contribution to overall casting quality improvement that they provide. In particular, unlike screens, properly printed foam filters have the capability to reduce the amount of turbulence and entrained air in the liquid metal stream, further minimizing the likelihood of the formation of regenerated oxides and dissolved hydrogen postfilter. In fact, an important additional function of an in-mold ceramic foam filter is to provide flow control. Unlike foam filters, screens actually split the metal stream into up to hundreds of small streams exiting the screen (Fig. 8, 9). This is particularly important in the first few seconds of mold filling prior to steady-state flow being achieved. The formation of the small streams dramatically increases the surface-area-tovolume ratio of the liquid metal, likewise increasing the surface area exposed to the air.
The more liquid metal surface area available to the atmosphere, the greater the potential for reoxidation to occur. Additionally, it can be shown that air bubbles can be easily aspirated into the mold through screens. Unlike screens, refractory foam filters inhibit the transmission of large air bubbles through the foam structure. An aluminum foundry should treat an air bubble as the encapsulated air it truly is. As damaging as an air bubble can be if it remains trapped in the solidified casting, it is equally damaging as a source of oxygen and hydrogen, which can
Fig. 8
Screen water model, split flow
Fig. 9
Ceramic foam filter, convergent flow
Molten-Metal Filtration / 199 react, forming aluminum oxides or dissolving in solution. One way to look at an air bubble in molten aluminum is as if it were a balloon on a string, the air bubble being the balloon and the string a trail of aluminum oxides following along wherever the balloon travels. Because refractory foam filters, via their unique pressure drop, can remove a significant portion of the energy from the metal stream as well as smooth the metal flow free of air bubbles, the metal exiting the refractory foam filter will be substantially less turbulent. Once this turbulence reduction is achieved, it will not remain so unaided by the foundry. The metal stream is the most nonturbulent when exiting the refractory foam filter immediately at the exit face. If anything is done to disrupt the metal stream as it travels away from the filter to the actual mold cavity, turbulence will increase. For this reason, anything that would increase the metal stream velocity postfilter should be avoided. This can include, but is not limited to, a choke postfilter, a further metal drop, a high-aspect-ratio runner (i.e., a runner that is substantially taller than it is wide), or any configuration that would tend to induce splash.
Types of Molten-Metal Filters Filter selection often depends on ease of use. The simplest filters are the strainer cores used in down sprues in conventional foundry sand casting. While serving to choke or control metal flow, they trap some particulates by surface attachment and a degree of cake filtration. For in-mold foundry applications, ceramic foam filters are common. Ceramic foams are lighter, easier to preheat, and provide easier flow capability than bonded-particle filters. In the primary and mill products casting industry, filter selection will depend on economics, space constraints, auxiliary metal treatment capabilities, desired filtration efficiency, and end-product application. Metal or Fiberglass Screens. Steel mesh screens and fiberglass cloth are also commonly used for aluminum. Molybdenum mesh is used for copper alloys. These materials act as planar filters. They provide moderate filtration, particularly of larger particles and inclusions exceeding 100 mm (0.004 in.) in size. The mesh screen, fiberglass cloth, and strainer cores are, of course, single-use filters. Bonded-particle filters consist of a refractory grain (either Al2O3 or SiC) bonded together to form a rigid structure. Bonded-particle filters provide a relatively tortuous path for metal flow through the filter (good depth filtration). These filters are often used vertically to separate melting or holding furnace hearth areas from the dip-out well. Silicon carbide vertical gate filters are quite durable and offer the potential of low thermal gradients between the hot face and cold face because of the relatively high thermal conductivity of silicon carbide compared to other materials, such as oxide ceramics. Temperature
gradients between inlet/outlet sides of the filter may start as low as 5 C but will build with use to gradients of 20 to 40 C (40 to 75 F) have been reported for these filters, versus 35 to 70 C (60 to 130 F) for oxide ceramic foams. Bonded- particle filters possess less porosity than for a ceramic foam (38 versus 75%, respectively) and greater tortuosity; therefore, they function more as depth filters. Cartridge Filters. One interesting application of the rigid filter concept is the cartridge filter. It is simply a shell-and-tube-type construction similar to a heat exchanger. This type of filter provides a great amount of surface area and usually has very fine pore size. Therefore, although a substantial priming head (up to 0.3 m, or 1 ft) of metal may be required before flowthrough begins, and although the flow rate is not very high, filtration efficiency is excellent. Cartridge filters are now receiving attention for such premium-quality applications as aircraft plate, computer disks, and foil. Efficiencies of 95% or better in removing particles less than 5 mm (200 min.) in size have been reported. Ceramic foam filters are the most common types of filters currently used in nonferrous casting. They are produced by slurry coating a reticulated polyurethane cellular foam, followed by drying and firing to burn out the precursor foam. This results in a ceramic replica of the original organic foam structure. Suitable ceramics for these filters are alumina, zirconia, and mullite. Some ceramic filters also employ chromic oxide. These filters are approximately 75% porous and are produced in a variety of pore sizes, as measured by pores per lineal inch (ppi). Ceramic foam filters are used in flat, platelike form. In the primary aluminum industry, plate sizes may range from 300 by 300 to 585 by 585 mm (12 by 12 to 23 by 23 in.), depending on flow rate (up to 680 kg/ min, or 1500 lb/min) and casting size (up to 34,000 kg, or 75,000 lb). Smaller filters are used in the die casting and foundry industries. Dual-structure or multilayered ceramic foam filters have been employed in the past but do not find present application. These use a finer pore size for greater inclusion-capture capability, coupled with a coarser pore size to permit greater throughput before the filter becomes saturated or clogged.
In-Mold Filtration Reduction of entrapped particulates in a casting is governed first by proper melting and handling practices and then by management of the casting process as the metal is poured, fed to the mold, and solidified. In-mold filtration is an important technology in aluminum foundry casting production. Filter application in the mold must follow sound principles to balance the pressure drop accompanying the resistance of a filter within the gating system with the necessary flow rate to minimize turbulence and therefore to achieve a sound, solidified cast structure. Proper selection of the specific filter
type and grade, sizing, and placement of the filter within the gating system will achieve significant particulate removal as well as metal flow control. Filter Selection and Application in Sand Gating. To achieve maximum filtration and flow modification benefits, the filter must be properly sized, placed, and coordinated with the other elements in the gating system. A filter that is too small can impede the flow of metal, increase fill times, and result in short pours if it blocks too soon. Also, remember that a filter is not an excuse to neglect good melting, treatment, and pouring practices. The selection of a filter size and porosity will depend on a variety of factors, including the pour weight of the casting, desired mold filling rate, alloy type, initial metal cleanliness, and pouring temperature. The most basic of guidelines for sizing based on predicted filter blockage for foam filter/flow modifiers is 0.2 to 0.4 g/cm2 (15 to 30 lb/in.2) of filter face before blockage. For flow rate, using the most common porosity foam filter/flow modifiers, the typical range for a 10 ppi filter is 50 to 100 g/cm2/s (0.7 to 1.4 lb/in.2/s) and for a 20 ppi filter, 30 to 80 g/cm2/s (0.4 to 1.1 lb/in.2/s). The filter (or filters in a multifilter system) should be large enough (total exit area and porosity) that the fill time is the same as it would be in an unfiltered casting with the same gating system dimensions. Also, the system should be unpressurized, and the filter should not become the choke in the system. If the filter becomes the choke, the metal will accelerate as it exits the filter and create undesired turbulence. If the ingate is the choke (a pressurized system), the flow modification effect of the filter will be lost as the metal sprays into the casting cavity. When blockage ultimately occurs from buildup of inclusions on and within the filter, it is not gradual but almost immediate. The selected filter must be large enough to maintain normal flow throughout the pour. As a starting point, the filter frontal area should be at least four times the cross-sectional area of the runner. Depending on foundry practice, it may be possible to reduce the filter size as casting experience and confidence are acquired. Filters may be placed horizontally, vertically, or at some intermediate angle within the runner system, depending on the molding method, pattern layout, or other molding restrictions. Filters in the vertical position are best placed just beyond the sprue well that will absorb the energy of the initial falling metal, as shown in Fig. 10. Filters in the horizontal position are best placed just off of the sprue well (Fig.11). Filters placed at the base of the sprue, where direct impingement of the metal stream will force the metal through the filter at high velocity until the sprue fills, may create turbulence in the initial metal entering the casting. Although the metal can be forced through the foam structure at these higher velocities, it can be defeating of the flow-modifying capability of the foam to do so.
200 / Molten Metal Processing and Handling
Fig. 11 Fig. 10
Horizontal installation of filter off of the sprue well
Vertical installation of filter in a mold feeding system
Within the basic guidelines for in-line filter application in sand molds, there is considerable latitude in positioning and placement (Fig. 12a–c). A single centrally located filter may be used, a filter may be placed in each runner, or a filter may be placed at each ingate for maximum effectiveness. Filters may be inserted manually or with core-setting equipment. Filter Selection and Application in Permanent Mold Casting. Most of the gating system design issues critical to sand casting of aluminum apply to permanent mold casting as well. There is even greater emphasis on the need for repeatable flow and tranquil filling characteristics and precisely controlled thermal properties to assure progressive solidification in the direction of the risers and feed heads. Proper gating design is vital to minimize turbulence as the mold fills and to avoid superheating the mold at any one point. The high density of the mold and total lack of permeability, compared to sand molds, put special emphasis on proper venting. Both horizontal and vertical gating and risering systems benefit from use of a tapered sprue to avoid air entrainment, an enlarged basin at the base of the sprue to absorb the kinetic energy of the falling metal, and an extended runner beneath the mold cavity to absorb entrained air and trap initial dross entering the runner system. The choke is normally at the base of the sprue, resulting in an unpressurized system. Quality issues with permanent mold castings may arise from entrained oxides and/or the reoxidation of treated metal, as is the case with sand castings. Filter screens are commonly used in an effort to intercept inclusions in permanent mold castings. However, screens have minimal filtration effect and damage the metal by splitting the stream and creating turbulence.
Development efforts are currently underway to apply foam filters to permanent mold castings to trap inclusions and minimize turbulence and reoxidation in vertical, horizontal, tilt, and low-pressure permanent mold applications. Filter selection is based on the same criteria (pour weight, desired fill rate, alloy type, initial metal cleanliness, pouring temperature) as sand casting. Direct Pouring. The primary purpose of runners and gating systems is to deliver metal to the casting cavity in a nonturbulent fashion. However, even though gating and running system components may be recycled, gating is not free. Significant costs of gating include: Cost of poured metal not turned into saleable
castings, including melt conversion cost, handling, and melt loss Total cost of remelting removed gating metal, including energy, melt loss, alloy loss, and other treatment costs Total cost of molding, including consumption of pattern and mold space, extra sand and binder, more metal, and runner maintenance due to runner erosion Pouring temperature cost, including superheat to compensate for loss of heat in runner system and oxidation aggravated by higher temperatures and turbulence Cleaning costs, including gating-removal labor, equipment and supplies, and handling and storage of removed gating Compromised thermal profile resulting from coldest metal ending up in the riser, and unnecessarily large risers needed for adequate thermal mass to accommodate shrinkage Oxidation and air entrainment resulting from turbulence in runners (rolling-back wave generation and air bubbles)
Fig. 12
Placement of filters (shaded). Filters may be placed in (a) each ingate, (b) one central location, or (c) each runner.
Molten-Metal Filtration / 201 By pouring directly through a sprue attached to the top of the casting cavity or mounted on a side entry close to the casting, many of the costs of a conventional gating system can be reduced or eliminated. Direct pouring will certainly reduce poured weight and increase yield. However, for automotive and other castings with stringent physical requirements, some means must be used to control filling turbulence and prevent mold erosion and/or core damage from direct impingement of the molten metal. A direct-pouring unit for sand or permanent mold casting (Fig. 13) has been developed and patented (FOSECO Metallurgical Inc.) that incorporates an insulating pouring sleeve (that may also act as a riser) with a foam filter to trap inclusions and modify the flow of the metal as it enters the casting cavity. Filter media other than foam have not been proven adaptable to the direct-pouring process. It is essential that turbulence reduction in the metal flow be possible from any device used to attempt direct pouring. In direct pouring, filtration takes a back seat to reduced turbulence in terms of primary operational characteristics.
Fig. 13
Patented direct-pouring filter. Courtesy of FOSECO Metallurgical Inc.
This direct-pouring system replaces all the traditional gating system elements and reduces feeding requirements. By eliminating the runner system and ingates, ingate hot spots requiring feeders are eliminated. Also, with metal entry through the top of the casting, the hottest metal ends up in the direct-pouring unit sleeve, reducing thermal mass requirements if the sleeve serves as a feeder as well. Example of In-Mold Filtration for Scrap Reduction in Sand Casting of Automotive Intake Manifold. Prior to the introduction of filtration/flow modification, a sand cast manifold experienced unacceptably high scrap rates resulting from oxides and other inclusions. Mold-filling turbulence was identified as a contributing factor. Minor modifications to the existing gating system enabled the addition of a single foam filter/flow modifier that led to a reduction of machining scrap and other oxiderelated scrap from 4.6 to 0.8%. With the casting valued at approximately $25 (United States) when the scrap was discovered, this scrap-reduction level represents a gross savings per casting of $0.95. Based on the typical production rate of 325,000 castings per year, this results in an annual gross savings of $308,750. Example of In-Mold Filtration for Scrap Reduction in Casting Oil Cooler Extension. Machining scrap and leakers were responsible for a combined scrap rate of 4.73%. Before the part could be leak tested, it required machining. At the point of machining, the part was valued at $16.75 (United States). After leak testing, the part was valued at $20. Placing a foam filter/flow modifier in the conventional gating system reduced the machining scrap rate from 1.81 to 0.4% and the leaker rate from 2.92 to 0.28%. The combined scrap rate reduction (machining and leakers) results in a gross savings to the foundry of $0.76 per casting or $45,600 based on the annual production rate of 60,000 pieces. Example of Direct-Pour Filtration in Casting Large Commercial Pump Body. Initially, this 30 kg (65 lb) casting included a conventional runner system and had a pour weight of 42 kg (93 lb) for a yield of 70%. By replacing the components of the runner system with a direct-pouring system, the pour weight was reduced by 9 kg (19 lb) and the yield increased to 88%. At $0.10/lb (United States), this accounted for a $1.90 savings per casting. The elimination of the conventional running and gating system also resulted in a reduction in cleaning room time from 6 to 3 min (removal of risers, flash, etc.), saving an additional $1.35 per casting. The filtration capability and flow-modifying effects of the foam filter in the direct-pouring unit reduced the (machining) scrap rate from 7 to 2%. With an approximate casting value of $75 at the time scrap was identified, this resulted in an additional savings of $3.75 per casting. With a total savings of $7.00 per casting and a production rate of 3000 castings per year, this resulted in a $21,000 gross
savings. At the selling price of $110, this represents a gross profit improvement of 6.4%.
In-Furnace Filtration of Molten Aluminum As described in the preceding section, filters are also placed in the feeding system for casting. However, it may not always be possible to employ direct filtering within the mold system. Consequently, in-furnace filtering is an alternative, such as with high-pressure die casting. Other processes, such as gravity permanent mold, low pressure, and sand casting, can also make use of filters directly into a melting or casting furnace. In-furnace filtration provides a very helpful means to the foundry in removing inclusions and thereby increasing the molten-metal quality to produce castings. In-furnace filters can be placed directly in either the melting or casting furnace, or sometimes both. In-furnace filtering products can be applied to most of the typical melting and holding/casting furnaces in use today (2008), with proper evaluation. The benefits of in-furnace filtration can be demonstrated using various analytical techniques on the molten metal itself prior to casting, or by monitoring and evaluation of machining and finishing operations and overall casting production, to determine reduction in the number of scrap castings that can be shown to be inclusion-related.
In-Furnace Filter Types There are several types of filter media that have been employed in a furnace. Historically, the earliest and simplest form of filter was to use a fiberglass sock at the tap-hole or outlet spout on the furnace to catch the dross or other debris. Additionally, use of a ceramic-media bed in a box just off the outlet allowed the metal to flow through and the ceramic media to collect not only the dross or other debris but also to filter some inclusions from the pouring metal itself. While these fiberglass filters at a tap-hole of a melting furnace are still employed and are indeed important, newer furnace filtering schemes have taken over in many instances. For in-furnace applications, either a ceramic foam filter or a bonded-particle filter is often used in modern practice. Many factors tend to favor the latter. The bonded-particle filter is an aggregate of silicon carbide granular particles together with a proprietary ceramic binder in a high-temperature firing process to give body to the filter. The forming process for the bonded-particle filter allows many different sizes, grades, and three-dimensional filter shapes to be produced for specific applications. This is especially useful in that there are many variants in size, refractory construction, and configuration in foundry melting and casting furnaces employed today (2008).
202 / Molten Metal Processing and Handling This type of filter is not suitable as a singleuse, in-mold filter but rather is useful and economical only in applications that allow for sustainable usage. A bonded-particulate filter has a close-packed structure, with just 40% nominal porosity. The open-pore volume at this level is insufficient to permit normal rules of solidification (Chvorinov’s rule) to be followed without an excessively-sized print in the gating to accommodate the flow rate required through the filter to achieve proper solidification in a casting. The bonded-particle filter provides outstanding longevity—often two to three months or more—in many aluminum casting operations because of its chemical stability afforded by the binder, which possesses excellent molten aluminum resistance. The inherent higher mass/density and strength of a bonded-particle filter versus a ceramic foam filter—which replicates a reticulated polyurethane foam structure—favors the application of the former in furnaces where long filter life is desirable. The bonded-particle filter is used either as a vertical gate filter in a melting furnace in front of a tap-hole or placed in front of a transfer pump to provide first-stage filtering. This firststage filtering, which lessens the burden on downstream filtering processes, minimizes entrainment of sludge and other debris associated with melting foundry returns and other materials with a high inclusion content. The molten metal is usually further delivered by ladle for subsequent treatment, such as degassing and/or flux injection, and then transferred to casting furnaces in the permanent mold or high-pressure die casting process or direct poured into sand or precision-sand molds. Figure 14 depicts a typical application of a vertical gate filter in a melting furnace in front of a tap-hole. In the dip-well of a casting furnace, further filtration is advantageous, and this can be accomplished either by another vertical gate filter (Fig. 15) or by an all-media box filter (Fig. 16), from which the casting can be poured by using either a manual ladle or an automatic ladle dipping from the interior. The box filter provides the additional advantage that the filtered metal is always available without regard to buildup and breakaway of inclusions/accretions on the sidewalls of the refractory. Both vertical gate and box filters can be produced in many sizes and geometric shapes to meet specific furnace and ladle dimensional requirements. The long-life potential of a bonded-particle filter allows for its long-term use as an in-furnace filter, complementing any further filtering that may be performed in-mold. In addition to long life, the silicon carbide composition provides greater thermal conductivity than an oxide-based filter, which is important in maintaining temperature in most casting processes. Additionally, the multifaceted particle structure, together with a fully interconnected pore structure, provides a tortuous metal-flow path through the filter. This tortuosity, coupled with
Fig. 14
Vertical gate filter in front of a tap-hole of a reverberatory melting furnace
Fig. 15
Vertical gate filter in a holding or casting furnace. (a) Filter separates heated chamber from dip-well. (b) Installation of a vertical gate filter in a new furnace dip-well
Fig. 16 dip-well
Box filters. (a) All-media bonded-particle box filter. (b) All-media box filter employed in a
Molten-Metal Filtration / 203 the chemical and physical affinity of the binder system, enhances the probability of inclusion capture and hence contributes to enhanced filtration efficiency.
In-Furnace Filter Installation Filter installation into reverberatory furnaces and use of in-furnace filters can be accomplished with careful effort. Measurements of the dip-well dimensions, or other areas in the furnace where the filter will be installed, should be taken carefully, with emphasis on the clean refractory, removing any buildup prior to taking the measurements. For a vertical gate filter, gaskets may be used on the filter edges to provide a seal and take up some degree of irregularity in the refractory sidewalls. It is necessary to fully scrape the sidewalls and bottom of the furnace vessel to provide a good fit. It is also very advantageous to create accepting slots in the furnace refractory for the filter, which can be done during a furnace relining. For either a vertical gate filter or a box filter, it is necessary to allow sufficient space for the ladling operation. Preheat of the filter is necessary to remove any moisture condensation and to prevent thermal shock, but with a thermally-conductive silicon carbide filter, this is straightforward and only requires that the filter be preheated to a temperature of 200 to 300 C (400 to 575 F). This may be accomplished by placing the filter on the furnace sill plate, covered with an insulating blanket, against a furnace hot wall or in a nearby oven. The filter may then be placed directly into the molten metal in the furnace as quickly as safety will allow, without the need to drain the furnace. When a melting or casting furnace is being originally manufactured, relined, or rebuilt, it is often desirable to create slots in the refractory sidewalls to accommodate a filter. In this case, the face-edges of the filter may have a compressible fiber gasket attached to further create a seal in the contact area between the filter medium and the sides or shoulder of the slotted construction. Crucible Furnace with Filter Baffle. Filter installation into crucible furnaces is somewhat similar in principle to installation into a reverberatory furnace, but there are specific differences. The filter can be viewed as a baffle separating the crucible into two sections, which may or may not be equal. Filter baffles have also been placed in the pouring ladle (Fig. 17). If a crucible furnace is used for both charging and degassing (and/or other melt treatment), space must be allowed for casting. In that instance, it is important to maintain both temperature and molten-metal level stability by matching intake with take-out volume, in order to avoid sludge formation and to maximize filtering surface area effectiveness. Further, the contours of the crucible place an extra imposition on the fabricated filter shape to become
sealed to the crucible wall to avoid metal bypass from unfiltered to filtered section, usually demanding extra gasketting. In addition, if due to space/volume constraints the filter must be placed offset to the main diameter, stops are necessary to avoid filter movement into the more open diameter. Figure 18 portrays a ceramic bonded-particle filter placement in a crucible. As an alternative to a baffle filter, there may be sufficient space available in casting crucible holding furnaces, especially larger ones, to accommodate a filtering vessel, such as an allmedia filter or one constructed of refractory materials with a smaller filter that can be replaced when spent. Low-pressure furnaces readily employ filters directly in or onto the stalk or riser tube. Both ceramic foam filters and bonded-particle filters are used in this application. In the lowpressure furnace, application of air forces metal up the stalk or riser tube and into the mold cavity with each cycle. Excess metal (the riser volume) drains back through the filter at the completion of each cycle. The filter medium must not create an undue restriction to metal flow, which would affect the pressurization cycle, nor the backflow of metal from the riser
during the drainage. The lower-porosity ceramic foam filters are often found to be more advantageous in this regard, with a disc shape being easily affixed into or onto the end of the riser tube. Therefore, bonded-particle filters with substantially lower porosity tend to be supplied in more three-dimensional shapes (Fig. 19) in this application to provide more filtering surface area and hence metal-flow capability. To accommodate a larger filter, there must be a sufficient gap or space between the end of the riser tube and the floor of the furnace, as shown in Fig. 19.
Factors That Determine Filter Life There are many factors that determine filter life, including overall incoming metal quality on the inlet side, the specific filter grade, metal flow rate and total throughput through the filter, and metal-level drawdown (in which the filter loses its prime). Other than simply gaining experience and knowledge about how long an in-furnace filter can be used, there are two observations that can be made. As the filter begins to clog, there will be a metal
