Brazing

© 2003 ASM International. All Rights Reserved. Brazing (#06955G)

Brazing Second Edition

Mel M. Schwartz

ASM International® Materials Park, Ohio 44073-0002 www.asminternational.org

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© 2003 ASM International. All Rights Reserved. Brazing (#06955G)

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Copyright © 2003 by ASM International® All rights reserved No part of this book may be reproduced, stored in a retrieval system, or transmitted, in any form or by any means, electronic, mechanical, photocopying, recording, or otherwise, without the written permission of the copyright owner. First printing, September 2003

Great care is taken in the compilation and production of this book, 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. Prepared under the direction of the ASM International Technical Book Committee (2002–2003), Charles A. Parker, Chair. ASM International staff who worked on this project include Scott Henry, Assistant Director of Reference Publications; Bonnie Sanders, Manager of Production; and Nancy Hrivnak and Jill Kinson, Production Editors. Library of Congress Cataloging-in-Publication Data Schwartz, Mel M. Brazing / Mel M. Schwartz.—2nd ed. p. cm. Includes bibliographical references and index. ISBN 0-87170-784-5 1. Brazing. I. Title. TT267.S39 2003 671.5′6—dc21

2003051963 ISBN: 0-87170-784-5 SAN: 204-7586 ASM International® Materials Park, OH 44073-0002 www.asminternational.org Printed in the United States of America

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Contents Preface . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . vii Chapter 1

Introduction to Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 Brazing versus Soldering . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 Historical Development of Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 Advantages and Limitations of Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 Mechanics of Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4 Brazing versus Other Welding Processes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5

Chapter 2

Brazing Fundamentals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 Adhesion, Wetting, Spreading, and Capillary Attraction . . . . . . . . . . . . . . . . . . . . . . . 7 Practical Experience, Work-Related Tips, and Problem-Solving . . . . . . . . . . . . . . . 12

Chapter 3

Elements of the Brazing Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Filler-Metal Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Base-Metal Characteristics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Filler-Metal Characteristics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 Surface Preparation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17 Joint Design and Clearance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 Temperature and Time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 Processes (Heat Sources) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21 Torch Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21 Furnace Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23 Induction Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 32 Controlled-Atmosphere Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 36 New Induction Brazing Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37 Resistance Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41 Carbon Resistance Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42 Direct Resistance Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42 Dip Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46 Infrared (Quartz) Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50 Exothermic Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51 Laser Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52 Braze Welding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53 Diffusion Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54 Microwave Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56

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The Future of Braze Processing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58 Practical Experience, Work-Related Tips, and Problem-Solving . . . . . . . . . . . . . . . 59 Chapter 4

Base Metals and Base-Metal Family Groups . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63 Metallurgical Reactions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63 Base-Metal Family Groups . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 70 Case Histories and Problem-Solving Examples . . . . . . . . . . . . . . . . . . . . . . . . . . . . 159

Chapter 5

Brazing Filler Metals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 177 Basic Characteristics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 177 Filler-Metal Selection Criteria . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 179 Filler-Metal Types . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 181 Specialized Brazing Filler Metals and Materials . . . . . . . . . . . . . . . . . . . . . . . . . . . 211 Filler-Metal Selection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 219 Filler-Metal Forms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 219 Case Histories and Problem-Solving Examples . . . . . . . . . . . . . . . . . . . . . . . . . . . . 231

Chapter 6

Fluxes and Atmospheres . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243 Atmospheres . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243 Fluxes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 266 Case Histories and Problem-Solving Examples . . . . . . . . . . . . . . . . . . . . . . . . . . . . 281

Chapter 7

Fixturing, Tooling, Stopoffs, Parting Agents, Surface Preparation, Surface Cleaning, and Repair . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 289 Fixturing and Tooling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 289 Stopoff Materials and Parting Agents . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 295 Surface Cleaning and Preparation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 297 Surface Preparation for Specific Base Metals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 302 Clean Rooms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 307 Vacuum Brazing Cleaning . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 307 Postbrazing Treatments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 308 Repair Techniques with Cleaning Agents . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 308 Case Histories and Problem-Solving Examples . . . . . . . . . . . . . . . . . . . . . . . . . . . . 308

Chapter 8

Joint Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 313 Types of Joints . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 313 Joint Clearance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 315 Design for Assembly . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 319 Effects of Brazing Variables on Clearance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 322 Strength . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 327 Joint Design and Ceramics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 329 Case Histories and Problem-Solving Examples . . . . . . . . . . . . . . . . . . . . . . . . . . . . 336

Chapter 9

Evaluation and Quality Control of Brazed Joints . . . . . . . . . . . . . . . . . . . . . . . . 339 Design and Quality System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 339 Quality Standards for Brazing and Brazing Processes . . . . . . . . . . . . . . . . . . . . . . . 340 Types of Common Defects . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 341 Brazing Process Planning and Control . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 341 Nondestructive Inspection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 342 Design Testing, Evaluation, and Feedback . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 344 Destructive Inspection and Testing Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 345 iv

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Chapter 10 Applications and Future Outlook . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347 Automation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347 Fluxless Brazing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 351 Novel and Emerging Brazing Processes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 352 Future Outlook . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 373 Case Histories and Problem-Solving Examples . . . . . . . . . . . . . . . . . . . . . . . . . . . . 376 Abbreviations and Symbols . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 383 Subject Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 385 Filler Metal Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 403

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Preface This updated and revised second edition of Brazing is intended to provide the reader with the information needed to braze materials that will be used in the 21st century. My goal has been to fuse experiences, basic understandings, theories, and practical information from the past with guidance about expected brazing applications and problem-solving considerations for the future and what it may hold for materials joining. It will be interesting to observe the continuing development and use of novel types of equipment and how engineers, scientists, and technology innovators will tackle the unknowns of this new millennium. In addition, this revised edition includes lessons learned on tooling, design, materials, atmospheres, processing, and equipment throughout the book. These examples should be helpful to the new brazing initiate as well as to more experienced brazing technologists and engineers. Several new and emerging topics are covered, including nanostructures and materials, microwave and laser brazing, more effective use of vacuum atmospheres, functionally gradient materials (FGM), and intermetallics, to name just a few. Coverage has been increased on beryllium alloys, aluminumlithium alloys, new titanium alloys, various composites (metal, ceramic, intermetallic), ceramic-tometal brazing, and ceramic-to-ceramic brazing. The enforcement of environmental regulations and the rising cost of metals such as copper and silver have forced manufacturers of brazing and soldering products to develop new fluxes and filler metals, while devising more efficient means of applying existing products. These trends have been reflected in the content of this new edition. Residual flux and its disposal have come under the scrutiny of regulatory agencies. To address this concern, a family of concentrated fluxes has come on the market that provides improved fluxing performance with smaller amounts being applied to the part. The quantities used can be reduced even more with automated dispensing, which places a small amount of flux exactly where it is required rather than the more expansive manual brushing of the flux. The increasing use of aluminum for parts has spawned the development of better aluminum filler metals and fluxes for brazing. The elimination of cadmium from braze filler metals is another area that has attracted attention in recent years. Cadmium has for years been used in certain brazing filler metals because of its characteristic as a temperature depressant, which facilitates melting and wetting of the filler metal. OSHA has established stringent regulations on allowable cadmium levels in the air. Some manufacturers have addressed this through the installation of sensitive filtration equipment and closing off the brazing area from other manufacturing operations. Another approach has been to utilize substitute cadmium-free braze filler metals now commercially available. These substitutes are similar in wetting characteristics to the cadmium types, but they do require more heat to produce the wetting. This factor should be taken into consideration when establishing the application of the heat. Other products that are considered are brazing powders and rods pre-coated with flux and equipment that automatically dispenses exact amounts of filler metals in paste form. vii

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Modeling can be used to increase reliability, repeatability, and efficiency of furnace operations. That is, the optimum cycle for a particular brazing process and workload is determined and a program is designed to allow the influencing parameters to be automatically compensated in real time, thereby achieving optimum operation. The basis for actual in-process control is modern computer methods and more sophisticated knowledge of the brazing process. The thinking and understanding of companies and their management must become global, and engineers must be able to reduce the technology development cycle for new materials and brazing systems from a norm of five to seven years to two to three years. One final area of concentrated development work is ceramic joining. Although the ability to fabricate ceramic materials that are reliably strong and tough is continuously advancing, the successful utilization of these materials will depend on the ability to assemble simple components into structures that will function effectively. While metal joining can be performed by starting at one end of the joint, ceramic joining requires the entire joint region to be brought into a reactive state simultaneously. Preliminary supplemental heating may be needed to prevent thermal shock during joining. A transfer of the experience gained in other material fields could prove helpful. Adaptation of physical and chemical deposition methods, surface engineering by laser beams, ion implantation, and chemical doping to improve surface reactivity and solid state adhesion are just a few of the techniques available for implementation. The development of generic joining technology for ceramic materials is vital and should be pursued in the context of metallurgical and ceramic sciences. Furthermore, it should be paralleled by an engineering effort to design and construct equipment specific to ceramic joining. Extending joining technology to ceramic materials will allow implementation of advanced technologies in circumstances where ceramics are the only materials having the requisite properties. Such a development provides an opportunity for the expertise of joining engineers plus venture capitalists to cash in on solving the underlying basic problems and thus augment the international competitiveness of U.S. products. If one tries to view the future and what it holds, woven through these fanciful visions are items that reflect trends observed in current reality. For instance, with the current pace of technological change, it can take less than six months for a state-of-the-art computer to become old technology. And software viruses can gobble up monthly production reports in seconds. But there is also good news. Within the next few years, micro-robots may crawl through intelligent joining systems, performing preventive maintenance in areas previously inaccessible. Solar power could become the cost-effective, environmentally safe way to fuel many manufacturing processes. Traditional methods of design and manufacture will give way to concurrent engineering (CE) strategies that enable improved communications between design and manufacturing. Automakers, aircraft manufacturers, and even job shops will design for manufacturability. These strategies will strongly influence development and deployment of advanced manufacturing technology well into the new millennium. However, with all the enthusiasm, enhancement, and advancements made by all the above industries, the basics in joining have not been altered. In spite of the move to automation, computer-controlled and monitored processes, and robots, the basic principles for all the various joining processes and fabrication techniques have not changed. You still need a heating media/source, material to be joined, a filler material to be added or preplaced in most cases, and a protective atmosphere/flux or vacuum in many situations. Over the past several years, many organizations have surveyed the changing face of science and engineering within the United States. In each case, the conclusions were the same: • Fewer engineering students will be available for industrial employment during the first decade of the 21st century.

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• As a result of the aging work force and benefits of early retirement programs, the demands for qualified joining (brazing) engineers will increase especially with the retirement in 2010–2012 of the baby boomers. • The changing face of industry will result in the need for more technically trained individuals. Department of Labor statistics predict an increased need in the engineering work force of 165% for the ten-year period from 1999 to 2009. As indicated in a recent National Science Foundation report, “the educational system in the United states has always been hierarchical. For this reason, the progression (and attrition) of students from primary education through graduate programs is conceptualized as a pipeline.” The feeder to engineering education resides in middle school (grades 6 to 8). It is here that students begin to take an interest in career opportunities and the educational requirements needed to enter these professional careers. This is where they begin planning their high school curriculum, which will either prepare them for a college education or vocational endeavors. The real solution will take the cooperative effort from each portion of industry, i.e., professional societies, industry and educational organizations. Professional societies and educational foundations can assist in image-building and career planning. Finally, the secondary schools and universities must begin to view themselves as equal partners with industry and societies in course requirement development and articulation agreements. A knowledge of metallurgy is basic to an understanding of the brazing process. This does not mean that the brazing engineer, brazer, and the brazing inspector must become a metallurgist. But it does mean that engineering schools of the world increasingly must face up to the fact that in the future engineers must play a key part in management’s strategic planning group. The university system recognizes that tomorrow’s engineers will have to have enhanced capabilities due to three significant factors: increased product sophistication and variation; a global manufacturing environment; and a multitude of social and economic changes. The 21st century engineer or technologist will require a radically different education from that of his predecessors. The workplace of the future will need systems integrators, not individuals classified as metallurgists, brazing technologists, manufacturing engineers, quality engineers, or industrial engineers; individuals will need knowledge and talent in each of these fields. The engineer of the future must be fully conversant with modern materials applications. Those who are involved in developing industrial computer networks must also be familiar with the manufacturing processes themselves. Improvements in joining science and technology must keep pace with advances in materials science and technology, or else the benefits of these new materials will not be achieved in the marketplace. Mel Schwartz

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Brazing Second Edition Mel M. Schwartz, p1-5 DOI: 10.1361/brse2003p001

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CHAPTER 1

Introduction to Brazing MANY PRODUCTS are assembled from two or more individual components that are often permanently joined to produce structurally sound assemblies. Joining methods include various fasteners, interference-type joints, adhesives, and, for the highest-integrity joints, the many techniques classified under welding. Brazing, although fundamentally different from fusion welding, is one such technique; soldering is another. The term brazing encompasses a group of welding processes that produce coalescence of materials by heating them to the brazing temperature in the presence of a filler metal having a liquidus above 450 °C (840 °F) and below the solidus of the base metal. (The liquidus, or melting point, is the lowest temperature at which a metal or an alloy is completely liquid, and the solidus is the highest temperature at which a metal or an alloy is completely solid.) The brazing filler metal is distributed between the closely fitted faying surfaces of the joint by capillary action. In the brazing process, the materials involved are heated to a temperature approximately 56 °C (100 °F) above the temperature of the filler material (usually a metal) that is being used. The filler metal turns liquid, covers all the mating surfaces, and creates an alloy bond with the faying surfaces. The filler metal can be preplaced, plated, or applied from an external source. If the filler metal comes from outside the mating surfaces, it requires capillary action to draw it between the faying surfaces (see Chapter 2, “Brazing Fundamentals”). When selecting a joining method, many factors, including requirements, must be considered, for example, service temperature and environment (corrosive, galvanic, etc.); service requirements for reliability, fatigue resistance,

impact resistance, and other cyclic conditions; and manufacturing considerations, such as cost, equipment required, and other assembly details. Brazing (usually conducted at 540 to 1620 °C, or 1000 to 2950 °F) is considered a hightemperature joining process compared to adhesive bonding, mechanical fastening, and soldering. Therefore, it cannot be used for plastics; however, it is used for metal-matrix composites and ceramic-matrix composites. Lap joints, which are the most common type of brazed joint, are usually as strong as or stronger than the base materials being joined as long as (a) the faying surfaces overlap for a distance equal to at least 3 times the thickness of the thinner of the two members being joined, and (b) the clearance between the two parts ( joint thickness) is kept to approximately 0.075 mm (0.003 in.) or less at braze temperature (Ref 1).

Brazing versus Soldering The basic distinction between brazing and soldering is that brazing is conducted at higher temperatures (soldering processes use filler metals having a liquidus not exceeding 450 °C, or 840 °F). The historical distinction between the processes has its origin in the earliest solders, which were based on tin, while brazes were based on copper-zinc alloys. Indeed, the word braze is a derivation of the Old English braes, meaning to cover with brass. On the other hand, the term solder is an adaptation of the Old French soudure, meaning to make solid. Brazing and soldering require the application of a number of scientific and engineering skills to produce joints of satisfactory quality and reliability. Brazing employs higher temperatures than soldering, but the fundamental concepts

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are similar, particularly with respect to metallurgy and surface chemistry (Table 1.1). However, joint design, materials to be joined, fillermetal and flux selection, heating methods, and joint preparation can vary widely between the two processes. Economic considerations involving filler-metal and process technology are also varied, particularly in relation to automated techniques and inspection and testing. Brazing and soldering are performed in many industries, from exotic applications in the electronics and aerospace field to everyday plumbing applications. The type of metallurgical reaction between a filler and parent metal is sometimes used to differentiate soldering from brazing. Solders usually react to form intermetallic phases, that is, compounds of the constituent elements that have different atomic arrangements than the elements in solid form. By contrast, most brazes form solid solutions, which are mixtures of the constituents on an atomic scale. However, this distinction does not have universal validity. For example, silver-copper-phosphorus brazes react with steels to form the interfacial phase of Fe3P in a similar manner to the reaction of tin-base solders with iron or steels to form FeSn2. On the other hand, solid solutions form between silverlead solders and copper just as they do between the common silver-base brazes and copper. Soldering and brazing involve the same bonding mechanism, that is, reaction with the parent material, usually alloying, so as to form metallic bonds at the interface. In both situations, good wetting promotes the formation of fillets that serve to enhance the strength of the joints. Similar processing conditions are re-

quired, and the physical properties are comparable, provided that the same homologous temperature is used for the comparison.

Historical Development of Brazing Early metalworkers, stimulated by a desire to produce structures that were difficult or impossible to build using methods then in existence, realized that it was possible to fill the joint between two metal pieces with molten metal and allow it to solidify. These artisans soon learned by experience that, in order to achieve adherence, the metals to be joined and the filler metal had to be kept free of oxides and the filler metal had to have a lower melting point, and, furthermore, that a given filler metal would not necessarily adhere to all metals. From these basic requirements, brazing and soldering grew into crafts whose practitioners were well versed in what to do and what not to do in order to produce sound joints. Just as these joining techniques developed empirically, so did the lower-melting-point filler metals. Workers first used lead and tin solders as well as silver and copper-arsenic ores, which were readily available and had low melting points. Later, the alloy brass was developed and found to be more desirable for joining copper, silver, and steel structures, because it provided higher-strength joints and could withstand higher temperatures. Early silversmiths, probably wanting to produce white solder joints for aesthetic reasons, melted brass and silver together and found it to have an even lower melting point than brass,

Table 1.1 Comparison of soldering, brazing, and welding Process Parameter

Joint formed Filler-metal melt temperature, °C (°F) Base metal Fluxes used to protect and to assist in wetting of base-metal surfaces Typical heat sources Tendency to warp or burn Residual stresses

Soldering

Brazing

Welding

Mechanical <450 (<840)

Metallurgical >450 (>840)(a)

Metallurgical >450 (>840)(b)

Does not melt Required

Does not melt Optional

Optional

Soldering iron; ultrasonics; resistance; oven Atypical

Furnace; chemical reaction; induction; torch; infrared Atypical

...

(a) Less than melting point of base metal. (b) Less than or equal to melting point of base metal

...

...

Plasma; electron beam; tungsten and submerged arc; resistance; laser Potential distortion and warpage of base metal likely Likely around weld area

Chapter 1: Introduction to Brazing / 3

good adherence, and better corrosion resistance. Although innumerable combinations of silver, copper, and zinc subsequently evolved, primarily to meet melting-point requirements, these silver-brass, brass, and lead-tin alloys were essentially the only brazing and soldering filler materials available for generations.

Advantages and Limitations of Brazing Advantages. Brazing has many distinct advantages, including the following: • Economical fabrication of complex and multicomponent assemblies • Simple method to obtain extensive joint area or joint length • Joint temperature capability approaching that of base metal • Excellent stress distribution and heat-transfer properties • Ability to preserve protective metal coating or cladding • Ability to join cast materials to wrought metals • Ability to join nonmetals to metals • Ability to join metal thicknesses that vary widely in size • Ability to join dissimilar metals • Ability to join porous metal components • Ability to fabricate large assemblies in a stress-free condition • Ability to preserve special metallurgical characteristics of metals • Ability to join fiber- and dispersion-strengthened composites • Capability for precision production tolerance • Reproducible and reliable quality-control techniques Strong, uniform, leakproof joints can be made rapidly, inexpensively, and even simultaneously. Joints that are inaccessible and parts that may not be joinable at all by other methods often can be joined by brazing. Complicated assemblies comprising thick and thin sections, odd shapes, and differing wrought and cast alloys can be turned into integral components by a single trip through a brazing furnace or a dip pot. Metal as thin as 0.01 mm (0.0004 in.) and as thick as 150 mm (6 in.) can be brazed. Brazing allows for the substitution of complex castings (Fig. 1.1) (Ref 2) for plate, bar, or tubular shapes. The cost and times for manufac-

turing enter the picture here; one must consider the relative cost of material, availability of material, schedule to manufacture the assembly, and the production equipment. The brazing equipment could be torch, induction, resistance, or furnace. Figure 1.2 illustrates an assembly with several internal brazed joints that can be simultaneously joined in one production load in, for example, a furnace using fluxed joints and/or an inert or vacuum atmosphere where no flux is used. Figure 1.3 illustrates the joining of two copper plates whereby large surface areas are covered by the liquid filler metal. The criteria of strength and corrosion resistance were the key requirements that convinced the joining engineers to select brazing over several other joining techniques. Brazed joint strength is high. The nature of the interatomic (metallic) bond is such that even a simple joint, when properly designed and

Fig. 1.1

Use of brazing to enable replacement of complex castings with assemblies of basic components. The casting shown (top) needs to be faced, drilled, and tapped in three places. It is much easier to braze three threaded couplings/ tubes into a machined block (bottom). Source: Ref 2

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made, has strength equal to or greater than that of the as-brazed parent metal. The natural shapes of brazing fillets are excellent. The meniscus surface formed by the fillet metal as it curves across corners and adjoining sections is ideally shaped to resist fatigue. (It should be noted that in brazed joints using eutectic-type filler metal, fillets often contain an excessive amount of brittle intermetallic compounds. In fact, fillets are 5 to 10 times thicker than the joint and thus have a much higher volume of the liquid phase from which these brittle phases crystallize. Therefore, crack nucleation often originates in fillets.) Complex shapes with

Fig. 1.2

Assembly in which several internal brazed joints are accomplished at the same time in one furnace brazing operation by the use of internal brazing performs. Source: Ref 2

greatly varied sections can be brazed with little distortion, and precise joining is comparatively simple. Unlike welding, in which the application of intense heat to small areas acts to move the parts out of alignment and introduces residual stresses, brazing involves fairly even heating, and thus part alignment is easier. Limitations. A brazed joint is not a homologous body but rather is heterogeneous, composed of different phases with differing physical and chemical properties. In the simplest case, it consists of the base-metal parts to be joined and the added filler metal. However, partial dissolution of the base metal, combined with diffusion processes, can change the composition and therefore the chemical and physical properties of the boundary zone formed at the interface between base metal and filler metal and often of the entire joint. Thus, in addition to the two different materials present in the simplest example given previously, a complicated transitional or even completely different zone must be considered. In determining the strength of such heterogeneous joints, the simplified concepts of elasticity and plasticity theory—valid for a homogeneous metallic body where imposed stresses are uniformly transmitted from one surface or space element to the adjacent ones—no longer apply. In a brazed joint formed of several materials with different characteristics of deformation resistance and deformation speed, the stresses caused by externally applied loads are nonuniformly distributed.

Mechanics of Brazing Brazing involves a limited dissolution or plastic deformation of the base metal (Ref 3). Brazing proceeds through four distinct steps:

Fig. 1.3

Brazing used to join large surface areas. Channels were grooved into these two copper plates prior to brazing to serve as internal cooling channels. Source: Ref 2

1. The assembly or the region of the parts to be joined is heated to a temperature of at least 450 °C (840 °F). 2. The assembled parts and brazing filler metal reach a temperature high enough to melt the filler metal (foil, wire, paste, platings, etc.) but not the parts. 3. The molten filler metal, held in the joint by surface tension, spreads into the joint and wets the base-metal surfaces. 4. The parts are cooled to solidify, or “freeze,” the filler metal, which is held in the joint by capillary attraction and anchors the parts

Chapter 1: Introduction to Brazing / 5

together by metallurgical reaction and atomic bonding.

Brazing versus Other Welding Processes The mere fact that brazing does not involve any substantial melting of the base metals offers several advantages over other welding processes. It is generally possible to maintain closer assembly tolerances and to produce a cosmetically neater joint without costly secondary operations. Even more important, however, is that brazing makes it possible to join dissimilar metals (or metals to ceramics) that, because of metallurgical incompatibilities, cannot be joined by traditional fusion welding processes. If the base metals do not have to be melted to be joined, it does not matter that they have widely different melting points. Therefore, steel can be brazed to copper as easily as to another steel. Brazing also generally produces less thermally induced distortion, or warping, than fusion welding. An entire part can be brought up to the same brazing temperature, thereby preventing the kind of localized heating that causes distortion in welding. Finally, and perhaps most important to the manufacturing engineer, brazing readily lends itself to mass-production techniques. It is relatively easy to automate, because the application of heat does not have to be localized, as in fusion welding, and the application of filler metal is less critical. In fact, given the proper clearance conditions and heat, a brazed joint tends to make itself and is not dependent on operator skill, as are most fusion welding processes. Automation is also simplified by the fact that there are many means of applying heat to the joint, including torches, furnaces, induction coils, electrical resistance, and dipping. Several joints in one assembly often can be produced in one multiple-braze operation during one heating cycle, further enhancing production automation.

As noted in Table 1.1, essentially no melting of the base metal occurs in brazing; however, the temperatures involved can affect the properties of the metals being joined. For example, base metals whose mechanical properties were obtained by cold working may soften or undergo grain growth if the brazing temperature is above their recrystallization temperatures. Mechanical properties obtained by heat treatment may be altered by the heat of brazing. On the other hand, materials in the annealed condition are usually not altered by brazing. As with other welding processes, brazing produces a heat-affected zone (HAZ) with a strongly altered microstructure due to intensive mutual mass transfer between base metal and filler metal. The width of this zone varies with the heating process used. In torch and induction brazing, for example, only a localized zone is heated; in furnace and dip brazing, the entire part is subjected to the brazing temperature. As a rule, the HAZ produced during brazing is wider and less sharply defined than those resulting from other fusion-related processes. ACKNOWLEDGMENT

Portions of this article are adapted from M.M. Schwartz, Introduction to Brazing and Soldering, Welding, Brazing, and Soldering, Volume 6, ASM Handbook, ASM International, 1993, pages 109–113.

REFERENCES

1. Brazing Handbook, 4th ed., American Welding Society, 1991 2. W.D. Kay, Ten Reasons to Choose Brazing, Weld. J., Sept 2000, p 33–35 3. E. Lieberman, Modern Soldering and Brazing Techniques, Business News Publications, 1988

Brazing Second Edition Mel M. Schwartz, p7-13 DOI: 10.1361/brse2003p007

Copyright © 2003 ASM International® All rights reserved. www.asminternational.org

CHAPTER 2

Brazing Fundamentals BRAZING does not involve any melting or plastic state of the base metal. Brazing comprises a group of joining processes in which coalescence is produced by heating to suitable temperatures above 450 °C (840 °F) and by using a ferrous and/or nonferrous filler metal that must have a liquidus temperature above 450 °C and below the solidus temperature(s) of the base metal(s). The filler metal is distributed between the closely fitted surfaces of the joint by capillary attraction. Brazing is distinguished from soldering in that soldering employs a filler metal having a liquidus below 450 °C. Brazing has four distinct characteristics: • The coalescence, joining, or uniting of an assembly of two or more parts into one structure is achieved by heating the assembly or the region of the parts to be joined to a temperature of 450 °C or above. • Assembled parts and filler metal are heated to a temperature high enough to melt the filler metal but not the parts. • The molten filler metal spreads into the joint and must wet the base-metal surfaces. • The parts are cooled to freeze the filler metal, which is held in the joint by capillary attraction and anchors the part together.

Adhesion, Wetting, Spreading, and Capillary Attraction Metals More than 195 years ago, Thomas Young (Ref 1) proposed treating the contact angle (θ) of a liquid as the result of the mechanical equilibrium of a drop resting on a plain, solid surface under the action of three surface tensions (Fig. 2.1). The surface tensions are γlv at the interface

of the liquid in equilibrium with its saturated vapor, γsl at the interface between the solid and the liquid, and γsv at the interface of the solid in equilibrium with the saturated vapor of the liquid. Hence: γsv = γlv cos θ + γsl

(Eq 1)

It is important to keep in mind that phases are supposed to be mutually in equilibrium. The designation γsv is a reminder that the solid surface near the liquid should have an equilibrium film of vapor due to the film pressure. Young’s equation has been used extensively in literature, which reflects its general acceptance. However, Eq 1 has never been verified experimentally. The problem is that surface tensions of solids are not easy to measure due to the inevitable presence of the interfacial tension between a solid and its liquid. More importantly, there is the difficulty that any tensile stresses existing in the surface of the solid would prevent the system from being in equilibrium. The surface tension at the solid-vapor interface (γsv) has a relationship with surface tension of a solid in vacuum (γs) as follows: γsv = γs – πe

(Eq 2)

where πe refers to the spreading pressure. Consequently, Young’s equation may be rewritten as: γs = γlv cos θ + γsl + πe

(Eq 3)

Because most of the solids have a negligible πe, particularly when the contact angle (θ) is greater than 10°, Young’s equation becomes: γs = γlv cos θ + γsl

(Eq 4)

A decrease of the contact angle causes an increase of the liquid drop surface area and thus increases the total liquid surface free energy.

Chapter 2: Brazing Fundamentals / 9

expansion. It can be shown thermodynamically that in the absence of a reaction, the driving force for wetting does not exceed γlv, resulting in a steady-state contact angle (Ref 3). The driving force with the contribution of the free energy of reaction in most cases exceeds the resisting force represented by γlv, because θ is 0° during spreading. A condition of an expanding drop during a reaction is defined as spreading. It can be seen that the free energy of a reaction in which the substrate is a passive participant does not contribute to the driving force for wetting; thus, spreading does not occur. The contact angle, however, adjusts to conform with the surface-energy changes of the liquid caused by composition changes due to the reaction. Example: Copper-Silver System. The equilibrium phase diagram for the copper-silver binary system (Fig. 2.2) can be used to illustrate examples of wetting and spreading (Ref 3). The system has a eutectic at 780 °C (1430 °F), with 72 wt% Ag. At 900 °C (1650 °F), the solid-solution limit is 5 wt% Ag in copper and 8 wt% Cu in silver. Several compositions are identified in the phase diagram by the letters A to D. When a drop of liquid C is placed on solid B at 900 °C (1650 °F), wetting occurs, with a contact angle of 11° and no chemical reaction, because the phases are in chemical equilibrium. This behav-

Fig. 2.2

ior corresponds to γsv > γlv, because, in a given system, the surface free energy of a liquid is less than that of a solid, due to its lack of long-range order. The liquid thus has the opportunity to rearrange its surface structure to a lower freeenergy state. However, when liquid C is placed on solid A, spreading occurs, because substrate A (as an active participant in the reaction) changes its surface composition toward B. The third equation in Fig. 2.1 applies in this case. Another example is that of liquid D on solid B. Liquid D is not in equilibrium with B and dissolves some of the substrate to change its composition to C. Even though a reaction occurs, there is no spreading, because B is a passive participant with no change in composition, even though it is being dissolved. However, with liquid D on solid A, spreading occurs, because both are active participants as they change to equilibrium compositions C and B, respectively. In both of the latter examples, liquid D is an active participant, because it dissolves some of the substrate to reach equilibrium compositions. It does not, however, contribute to spreading, which is controlled by the active participation of the substrate.

Ceramics Joining dissimilar materials invariably results in high interfacial energy; that is, the work of

Stable phase equilibrium diagram for the copper-silver system

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adhesion is not sufficient to maintain the joint integrity. Most structural ceramic-metal interfaces are no exceptions; the liquid metal does not readily wet the ceramic surface. Therefore, the first challenge in metal-to-ceramic joining is to alter the interfacial thermodynamics to render the ceramic surface wettable. There are two approaches available today (Ref 4) to accomplish this task: metallization of the ceramic surface and reaction wetting. A thin layer of metal alloy is deposited on the ceramic by vapor deposition or sputtering. Brazing is then carried out by appropriate filler material or by simply melting the deposited layer (Ref 5). This two-step approach is not as enthusiastically embraced by industry as is direct reaction brazing. In reaction brazing, the filler metal is carefully chosen so as to facilitate compound formation at the interface. A small percentage of reactive metals, such as aluminum and titanium, are added to the otherwise inert base alloys (Ref 6, 7). The compounds that form are commonly spinels for the oxide ceramics and complex nitrides for the ceramic nitrides (Ref 8, 9). It is important to realize that wetting in such systems is time dependent. Successful bond formation relies on rapid transport of the reactive metal to the interface and a rapid rate of compound formation. Reaction wetting may not be the solution to all metal-ceramic joining

problems, because interfacial compound formation could create fragile layers (Ref 9), further complicating the development of good joint efficiency. The weakest among the adhesive strength at the ceramic-compound boundary, the cohesive strength of the compound, or the adhesive strength between compound and metal would determine the final joint strength. Table 2.1 (Ref 10) summarizes the various reaction products that have been identified in literature for the common metal-ceramic systems. Many researchers have concentrated their study on alumina surfaces, because it is one of the very few ceramics for which essential thermodynamic data are available. One of the first commercial applications of metal-to-ceramic brazed components is the turbocharger rotor (Ref 8), where a silicon nitride turbine blade is brazed to a stainless steel shaft. Other ceramics of interest to brazers are silicon carbide and zirconia. Ceramics exhibit very different thermal expansion behavior compared to metals; hence, considerable residual stress can build up during cooling. This thermal expansion mismatch more or less dictates the use of a ductile filler material. Most commercial brazing systems are therefore silver and copper base. The soft interlayer might not be sufficient to compensate for large differences in thermal expansion coeffi-

Table 2.1 Ceramic-metal interface formation and reaction products Bonding conditions System

Temperature, K

Load, MPa (ksi)

Al2O3(a)-Nb(a) Al2O3(b)-Nb(c) Al2O3(d)-Nb(a)

1925 1973 1973

20 (3) 10 (1.5) 6.4 (1.0)

0.1 2 1

Argon Vacuum 10–4 Pa Vacuum 10–3 Pa

(Al)Nb (O)Nb NbOx, (Al,O)Nb

Al2O3(b)-Nb(a) Al2O3(e)-Ti(a) SiC(b)-Ti(a) SiC(b)-Zr(a) SiC(a)-Al/Ti/Al(a) SiC(a)-Al(a) Si3N4(a)-Fe(a) SiO2-Al(a)

1973 1250 1773 1773 1273 1373 1683 875 ... 1270 1773

6.4 (1.0) ... 0.34 (0.05) 0.56 (0.08) 0.56 (0.08) 0 (0) 3000 (435) 10 (1.5) ... 20 (3) 20 (3)

1 60 1 1 1 1 1 10–4 24 0.25 0.25

Vacuum 10–3 Pa Vacuum 10–4 Pa Vacuum 10–3 Pa Vacuum 10–3 Pa Air Vacuum 10–2 Pa Argon 75N2/25H2 (H2/H2O = 102) H2/H2O = 2 × 105 H2/H2O = 105

1773 1473 1273 ...

20 (3) 1 (0.15) 0 (0) ...

0.25 0.25 0.6 8

H2/H2O = 5 × 105 H2/H2O = 2 × 103 H2 H2

NbOx, (Al,O)Nb Ti3Al, TiO, (AlO)Ti Ti3SiC2, Ti5Si(C), TiSi2 ZrSi, ZrC + (Si)Zr TiC, TiAl3Si Al4C, (Si)Al Fe3Si, (Si)Fe Not observed α-Al2O3, (Si)Al None None. After 6 h, (Al)Pt; after 1000 h, Pt3Al Pt3Al, (Al)Pt Pt3Si (Zr)Pt Pt3Zr, (Zr)Pt

Al2O3(b)-Cu(a) Al2O3(b)-Pt(a) Al2O3(b)-Pt(a) SiO2-Pt(a) ZrO2(a)-Pt

Time, h

Plane of interaction: (a) Polycrystal. (b) (0001). (c) (110). (d) (1010). (e) (1100). Source: Ref 10

Atmosphere

Reaction products

Chapter 2: Brazing Fundamentals / 11

cients (e.g., Si3N4 as compared to stainless steel). In such situations, laminated interlayers that provide a continuous gradient thermal expansion coefficient are used (Ref 6). Thermodynamic phenomena that occur at the interface can be studied in terms of the contact angle, θ, and the work of adhesion, W. These terms can be related to various surface or interfacial energies. The general case for a liquid metal in contact with a solid ceramic is shown in Eq 1, where a balance of surface tension forces results in the familiar Young’s equation. The Dupree equation is easily derived from Eq 1: W = γlv (1 + cos θ)

(Eq 7)

Attempts at understanding the nature of the force of adhesion across the interface have not been very successful. In 1965, researchers (Ref 11) rationalized, on the basis of the work of adhesion data for an alumina-metal interface, that the observed work of adhesion was the sum of two independent contributions arising from the van der Waals forces and a primary chemical bond. Predicting adhesion data in joining an alumina-metal interface is of great importance in many applications. The objective of proposed research to predict wettability and bond strength from measurable parameters and bridge the gap between a theoretical understanding and technology of observed work was undertaken (Ref 12). Researchers (Ref 13, 14) attempted to explain the entire work of adhesion across the metalceramic boundary in terms of physical forces using the dielectric principle. Such models are not of much use to the brazing industry, because most commercial metal-to-ceramic bonds are based on chemical bond formations.

Effects of Capillary Attraction and Wetting on Brazing Capillary attraction makes leak-tight joints a simple proposition for brazing. In a properly designed joint, the molten filler metal is normally drawn completely through the joint area without any voids or gaps, and brazed joints remain liquid- and gas-tight under heavy pressures, even when the joint is subjected to shock or vibrational types of loading. Capillary action results in the phenomenon where surface tension causes molten braze filler

metal to be drawn into the area that covers the parallel surfaces that are to be brazed. Capillarity is a result of surface tension between base metals(s), filler metal, flux, or atmosphere and the contact angle between base metal and filler metal. In actual practice, fillermetal flow characteristics are also influenced by dynamic considerations involving viscosity, vapor pressure, gravity, and metallurgical reactions between filler metal and base metal. As a matter of fact, present-day brazing practices have evolved as the result of an empirical approach to the phenomena of wetting and spreading, which are of prime importance in the formation of brazed joints. Classical, physical, and chemical principles led to equations governing the shape of liquid surfaces and the rate of filling a capillary gap in systems that do not interreact. However, the extension of theory to practical systems necessitates the consideration of a number of complicating factors, which often arise in everyday practice. A few of these factors include the condition of the solid surface as to the presence of oxide films and their effects on wetting and spreading, surface roughness, alloying between the filler metal and base metal and the extent to which this affects the thermodynamic properties of the liquid and solid surfaces, and the condition and properties of the brazing atmosphere. The factors that control the rate at which wetting, spreading, and capillary flow occur are of great practical, as well as theoretical, interest. Studies have indicated profound influences of various kinds of surface activation that cannot be explained in terms of surface energies or alterations in equilibrium contact angle (Ref 15, 16). Some of the most spectacular of these effects have been observed in systems in which a finite contact angle is thermodynamically unstable, because the solid-vapor surface energy exceeds the sum of the liquid-solid surface energies—that is, a system in which thermodynamics would predict complete spreading. In actual fact, spreading may or may not occur in this type of system, and the rate of spreading can be markedly dependent on surface chemistry, although the fundamental mechanisms of this dependence are not all clear. Wetting is, perhaps, best understood by example. If a solid is immersed in a liquid bath and wetting occurs, a thin, continuous layer of liquid adheres to the solid when it is removed from the liquid. Technically speaking, in the

12 / Brazing, Second Edition

wetting process, the force of adhesion between the solid and the liquid is greater than the cohesive force of the liquid. In practical terms, with respect to brazing, wetting implies that the liquid filler metal spreads on the solid base metal instead of balling up on its surface (Fig. 2.3). It has been demonstrated that wetting actually depends on a slight surface alloying of the base metal with the filler metal. A comprehensive theory of the wetting or spreading of liquids on solid surfaces is presented in Ref 17 and 18. It can be concluded that wetting is the ability of the molten filler metal to adhere to the surface of a metal in the solid state and, when cooled below its solidus temperature, to make a strong bond with that metal. Wetting is a function not only of the filler metal but also of the nature of the metal or metals to be joined. There is considerable evidence that in order to wet well, a molten metal must be capable of dissolving, or alloying with, some of the metal on which it flows. Wetting is only one important facet of the brazing process. A very important factor affecting wetting is the cleanliness of the surface to be wetted. Oxide layers inhibit wetting and spreading, as do grease, dirt, and other contaminants that prevent good contact between the filler metal and the base metal. One of the functions of a flux is to remove the oxide layer on the joint area and to expose clean base metal.

Fig. 2.3

Wetting and dewetting

Good wetting and spreading of the liquid filler metal on the base metal are necessary in brazing, because the mechanics of the process demand that the filler metal be brought smoothly, rapidly, and continuously to the joint opening. If the conditions within the capillary space of the joint do not promote good wetting, the filler metal is not drawn into the space by capillary attraction. It all boils down to the fact that, for successful joining of components by brazing, the filler metal selected must have a melting point above 450 °C (840 °F) and must also wet the base metal without melting it. Then, the joint must be designed so that the mating surfaces of the components are parallel and close enough together to cause capillary attraction.

Practical Experience, Work-Related Tips, and Problem Solving In order to braze tungsten carbide (WC) granules or diamonds to 1010 carbon steel wheels with BNi-2 filler metal, avoid the problem of filler metal sagging. Because the wheels must rest flat in the furnace, the diamonds or carbides are then on the vertical surface of the wheel diameter. Typical brazing takes place at 1040 °C (1900 °F) in a pure dry hydrogen atmosphere. There are a large number of variables that must be taken into consideration. It is much easier and more practical to use temperature as the controlling variable. Key variables that affect braze quality include: • The chemistry which is controlled by a specification and as a result the melting and flow characteristics of the filler metal are controlled by the chemistry. • Partial pressure of nitrogen will affect the melting characteristics of the filler metal. • A variation in the partial pressure of oxygen will affect the melting and flow characteristics of the filler. • The length of time in the oxidation range of 540 to 925 °C (1000 to 1700 °F) can also alter the melting and flow characteristics of the filler metal. • The heating rate, particularly at the high temperature where diffusion takes place, can alter the melting and flow characteristics.

Chapter 2: Brazing Fundamentals / 13

• The maximum brazing temperature is the best variable to control, because any one of the previously mentioned variables can change, requiring a change in the brazing temperature. A better braze filler metal for this type of application would be a very widemelting-range material and a filler metal of Cr-Ni-B-Si-Fe. This filler metal has a melting range of 970 to 1160 °C (1780 to 2120 °F). • Therefore, the large number of variables presents a problem, but considering and taking into account the various variables, control of the flow of the filler metal is feasible. REFERENCES

1. T. Young, Philos. Trans. R. Soc. (London) A, Vol 95, 1805, p 65 2. A.W. Adamson, Physical Chemistry of Surfaces, 4th ed., John Wiley & Sons, 1982, p 339 3. P.R. Sharps, A.P. Tomsia, and J.A. Pask, Wetting and Spreading in the Cu-Ag System, Acta Metall., Vol 29 (No. 7), 1981, p 855–865 4. M. Erg and A.W. Hennicke, Ceramics in Advanced Energy Technologies, A. Krockel et al., Ed., Dreidel Publishing, 1982, p 138 5. M.E. Twentyman and P. Hancock, in Surfaces and Interfaces in Ceramic and Ceramic-Metal Systems, Vol 14, Materials Science Research, J.A. Pask and A.G. Evans, Ed., Plenum Press, 1981, p 535 6. H. Mizuhara, Vacuum Brazing Ceramics to Metals, Adv. Mater. Process., Vol 131 (No. 2), Feb 1987, p 53–55 7. A.J. Moorhead and A. Keating, Direct Brazing of Ceramic for Advanced HeavyDuty Diesels, Weld. J., Oct 1986, p 117

8. R.E. Loehman, Interfacial Reactions in Ceramic-Metal Systems, Ceram. Bull., Vol 68 (No. 4), 1989, p 891 9. M.G. Nicholas and R.J. Lee, Joining Dissimilar Materials, Met. Mater., Vol 5 (No. 6), 1989, p 348 10. J.T. Klomp, in Ceramic Microstructures 86: Role of Interfaces, J.A. Pask and A.G. Evans, Ed., Plenum Press, 1988, p 307 11. J.E. McDonald and J.G. Eberhart, Adhesion in Aluminum Oxide-Metal Systems, Trans. AIME, Vol 233, 1965, p 512 12. G.R. Edwards and J.J. Moore, “Investigation of Brazing Alloys for Ceramic Substrates,” Research Proposal CSM 3264, Colorado School of Mines, Feb 1990, p 54–70 13. R.G. Barrera and C.B. Duke, Dielectric Continuum Theory of the Electronic Structure of Interfaces, Phys. Rev. B., Vol 13 (No. 10), 1976, p 4477 14. A.M. Stoneham and P.W. Tasker, in Ceramic Microstructures 86, Vol 21, Materials Science Research, J.A. Pask and A.G. Evans, Ed., Plenum Press, 1988, p 155 15. C.M. Adams, Jr., “Dynamics of Wetting in Brazing and Soldering,” Technical Report WAL TR 650/1, Army Materials Research Agency, Watertown Arsenal, Watertown, MA, July 1962 16. S. Weiss and C.M. Adams, Jr., The Promotion of Wetting, Weld. J., Vol 46 (No. 2), Feb 1967, p 49s–57s 17. W.D. Hawkins, Physical Chemistry of Surface Films, Reinhold, 1952, p 1–413 18. M.M. Schwartz, Fundamentals of Brazing, Welding, Brazing, and Soldering, Vol 6, ASM Handbook, ASM International, 1993, p 114–125

Brazing Second Edition Mel M. Schwartz, p15-62 DOI: 10.1361/brse2003p015

Copyright © 2003 ASM International® All rights reserved. www.asminternational.org

CHAPTER 3

Elements of the Brazing Process IN ORDER TO PRODUCE satisfactory brazed joints, a careful and intelligent appraisal of the following elements is required: • • • • • • •

Filler-metal flow Base-metal characteristics Filler-metal characteristics Surface preparation Joint design and clearance Temperature and time Rate and source of heating

This chapter reviews the important considerations related to each of these elements.

Filler-Metal Flow Wetting is only one important facet of the brazing process. If the molten filler metal does not flow into the joint, the effectiveness of the filler metal is greatly restricted. Flow is facilitated by capillary attraction, which in turn results from surface-energy effects (Ref 1, 2). It is therefore apparent that a high liquid surface tension, a low contact angle, and low viscosity are desirable for promoting filler-metal flow. Thus, a low contact angle, which implies wetting, is a necessary but not a sufficient condition for flow. Viscosity is also important. Filler metals with narrow freezing ranges that are close to the eutectic composition generally have lower viscosities than those with wide freezing ranges. Flow is the property of a filler metal that determines the distance it will travel away from its original position because of the action of capillary forces. In practice, reactions between the filler metal and the base metal are usually minimized by (a) selecting the proper filler metal, (b) keeping the brazing temperature as low as possible but high

enough to produce flow, (c) keeping the time at temperature short, and (d) cooling the brazed joint as quickly as possible without causing cracking or excessive distortion.

Base-Metal Characteristics The base metal has a prime effect on joint strength. A high-strength base metal produces joints of greater strength than those made with softer base metals (other factors being equal). When hardenable metals are brazed, joint strength becomes less predictable. This is because there are more complex metallurgical reactions involved between hardenable base metals and the filler metals. These reactions can cause changes in the base-metal hardenability and can create residual stresses (Ref 3). Also vitally important are the coefficients of thermal expansion (CTEs) where different materials make up the assembly and gaps can open or close as heating proceeds to the joining temperature. Also, during cooling after the filler metal has solidified, differences between the CTEs generate residual stresses that may produce distortion. The relative ease of brazing some base metals and other materials is shown in Table 3.1. There are several metallurgical phenomena that influence the behavior of brazed joints and, in some instances, necessitate special procedures. Included among these base-metal effects are (a) alloying; (b) carbide precipitation; (c) stress cracking; (d) hydrogen, sulfur, and phosphorus embrittlement; and (e) oxide stability. In addition to the base-metal effects mentioned previously and the normal mechanical requirements of the base metal in the brazement, the effect of the brazing cycle on the base metal

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and the final joint strength must be considered. Base-metal alloys that are strengthened by cold working will be annealed and the joint strength reduced when the brazing-process temperature and time are in the annealing range of the base metal being processed. Hot-cold-worked heat resistant base metals also can be brazed; however, only the annealed physical properties will be available in the brazement. The brazing cycle by its very nature will usually anneal the coldworked base metal, unless the brazing temperature is very low and the time at elevated temperature is very short. It usually is not practical to cold work the base metal after the brazing operation. When it is essential to design a brazement having strength above the annealed strength of the base metal after the brazing operation, specifying a heat treatable base metal is necessary. The base metals can be of the oilquench type, the air-quench type that can be brazed and hardened in the same or separate operations, or the precipitation-hardening type in which the brazing cycle and solution treatment cycle may be combined. Parts can be hardened and then brazed with a low-temperature filler metal employing short times at elevated temperature in order to maintain the mechanical properties (prevent softening by annealing). The strength of the base metal has a profound effect on the strength of the brazed joint; thus, this property must be clearly kept in mind when designing the joint for specific properties. Also, some base metals are easier to braze than others. This is the case particularly for certain base metals joined by specific brazing processes. For example, reactive metals (titanium, beryllium, etc.) brazed in vacuum or in an inert atmosphere in a furnace are easier to braze than with a torch process.

Table 3.1 Relative ease of brazing various base materials Degree of difficulty

Impossible Difficult Fair

Easy

Materials

None Ti, Zr, Be, and their alloys; ceramics; graphite; glass; TiC Al, W, Mo, Ta, alloys with more than 5% metals forming refractory oxides, cast iron, WC Cu, Ni, Co, and their alloys; steels; precious metals

Filler-Metal Characteristics The second material involved in joint structures is the filler metal. The term brazing filler metal is essentially synonymous with the commonly employed term brazing alloy. Its selection is important but not for the reasons many engineers think. A specific filler metal cannot be chosen to produce a specific joint strength, which is unfortunate but true. Actually, brazing can provide strong joints with almost any good commercial filler metal if brazing methods and joint design are selected and applied correctly. Several characteristics that filler metals must possess or are desirable are: • Proper fluidity at the brazing temperature to ensure flow by capillary action and to provide full alloy distribution • Stability to avoid premature release of lowmelting-point elements in the filler metal • Ability to wet the base-metal joint surfaces • Low-volatilization characteristics of the alloying elements of the filler metal at the brazing temperature • Ability to alloy or combine with the parent metal to form an alloy with a higher melting temperature • Controllability of the washing or erosion between the filler metal and the parent metal within the limits required for the brazing operation • Depending on the service requirements, the ability to produce or avoid base-metal/fillermetal interactions One of the most broadly misunderstood facts relating to filler metals is that brazed-joint strength is completely unrelated directly to the melting method used. This fact is hard to accept, because it seems to contradict a long-established metallurgical truth with regard to the manufacture of steels and other constructional metals. The effect of melting practice on filler metals, however, is not the same as that of melting practice on metals during primary manufacturing. If constructional metals are produced by vacuum melting, for example, there is a definite relationship between the vacuum-melting practice and the final strength of the ingot, bar, or rolled sheet. That is not true with a filler metal, because the joint strength depends on such factors as joint design, state of stress, brazing temperature, amount of filler metal applied, loca-

Chapter 3: Elements of the Brazing Process / 17

tion and method of application, heating rate, holding time at the peak temperature, and many other considerations that make up what is termed brazing technique. The process by which filler metals penetrate and alloy with base metals during brazing is referred to as diffusion. In applications requiring strong joints for high-temperature, highstress service conditions (such as turbine rotor assemblies and jet engine components), it is generally good practice to specify a filler metal that diffuses readily and alloys with the base metal. When the assembly is constructed of extremely thin base metals (as in honeycomb structures and some heat exchangers), good practice generally calls for a filler metal with a low-diffusion characteristic relative to the base metal being used. Diffusion is an essential and normal part of the metallurgical process that contributes to good brazed joints. In choosing a filler metal, the first criterion is the working temperature. Very few filler metals possess distinct melting points. Filler metals in which the solidus and liquidus are close together do not usually exhibit a strong tendency to separate, and they are relatively fluid. They flow readily and should be used with small joint clearances. Other filler-metal selection criteria include corrosion resistance, such as oxidation and galvanic corrosion with other parts of the assembly and the service environment; color match to the base metal; electrical and thermal conductivity; joint-filling capacity; hardness and machinability; ductility and fracture toughness; and ability to form fillets. Additionally, the designer must consider the extent of alloying with the base metal. The improvement in mechanical properties of the joint and the increase in remelt temperature obtained by alloying could be beneficial. Structural changes in the interface layers of the base metals; aggressive, extensive alloying between the base metal and filler metals; formation of brittle intermetallic compounds; and erosion of components can be deleterious, and, at times, catastrophic. All such effects, beneficial or otherwise, vary greatly with the joint gap, temperature and time of brazing, and compositions of the base metals and the filler metal. Although the mechanical properties of filler metals in massive form can provide a guide to their suitability for use in different capillaryjoining applications, in general, designers can-

not work with the mechanical properties of assemblies brazed using different joint configurations or cycles of time and temperature. Finally, the placement of the filler metal is an important design consideration, not only because the joint must be accessible to the heating or filler-metal placement method chosen, but also because, in automatic heating setups, the filler metal must be retained in its location until molten. Filler metals are available in different forms, and selection may depend on which form is suitable for a particular joint design. The most common filler-metal forms are wire, powder, foil, flux paste, strip, shim, and nonfluxing paste. Preforms made from wire, strip, and foil can be used. Several general rules apply in filler-metal placement. Wherever possible, place the filler metal on the most slowly heated part of the assembly in order to ensure complete melting of the filler metal. Gravity may be used to assist filler-metal flow, particularly for those filler metals having wide ranges between their solidus and liquidus temperatures. Filler metals can be chosen to fill wide gaps or to flow through joint configurations where the gap can vary, for example, around a corner. Unless movement between the components being joined is unimportant or can be corrected manually, through self-jigging or by fixtures after the filler metal is molten, the filler metal should be placed outside the joint and allowed to flow through it and should not be placed between the joint members. To avoid forming a void, the filler alloy should not flow from two sides to a center point. If erosion of thin members is possible, the filler metal should be placed on the heavier sections, which heat up more slowly, so that flow proceeds toward the thin sections. Apart from suiting the placement method selected for the joint, the form chosen for the filler metal may be needed to gage accurately the amount applied—not just for economy and reproducibility but also to regulate and maintain joint properties and configuration.

Surface Preparation A clean and nearly oxide-free surface is imperative to ensure uniform quality and sound brazed joints. A sound joint may be obtained more readily if all grease, oil, wax, dirt, and nearly all oxides have been carefully removed

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from the base and filler metals before brazing, because only then can uniform capillary attraction be obtained. Oils and greases should be removed before cleaning to remove oxides. Even after proper cleaning, all metal surfaces exposed to air will have a thin oxide film, which does not impede filler-metal flow during brazing. Brazing as soon as possible after the material has been cleaned is recommended. The length of time the cleaning remains effective depends on the metals involved, atmospheric conditions, storage and handling practices, and other factors. Cleaning is commonly divided into two major categories: chemical and mechanical. Chemical cleaning is the most effective means of removing all traces of oil or grease. Trichlorethylene and trisodium phosphate are the usual cleaning agents employed. Various types of oxides and scale that cannot be eliminated by these cleaners are removed by other chemical means. The selection of the chemical cleaning agent depends on the nature of the contaminant, the base metal, the surface condition, and the joint design. Regardless of the cleaning agent or the method used, it is important that all residue or surface film be removed from the cleaned parts by adequate rinsing to prevent the formation of other equally undesirable films on the faying surfaces. When faying surfaces of parts to be brazed are prepared by blasting techniques, several factors should be understood and considered. The purposes of blasting parts to be brazed are to remove any oxide film and to roughen the mating surfaces so that capillary attraction of the filler metal is increased. The blasting material must be clean and must be of a type that does not leave on the surfaces to be joined any deposit that restricts filler metal or impairs brazing. The particles of the blasting material should be fragmented rather than spherical, so that the blasted parts are lightly roughened rather than peened. The operation should be done so that delicate parts are not distorted or otherwise harmed. Vapor blasting and similar wet blasting methods require care because of possible surface contamination. Mechanical cleaning may be adequate, in which case the design must permit this during manufacture. In some cases where chemical cleaning is required, it may be followed by protective electroplating, necessitating access to the faying surface by the liquids involved.

Another technique in surface cleaning and protection is the use of solid and liquid brazing fluxes. At temperatures up to approximately 1000 °C (1830 °F), fluxes often provide the easiest method of producing or maintaining surface cleanliness, and, in such cases, the design must not only permit easy ingress of the flux but also allow the filler metal to wash it through the joint. Above 1000 °C, the flux residues can be difficult to remove, and surface cleaning by, for example, a furnace atmosphere is desirable, but the design must permit the gas to penetrate the joint. Apart from cleanliness and freedom from significant oxides, surface roughness is important in determining ease and evenness of flow of the filler metal. This varies with different manufacturing methods and may influence the engineer’s choice, or it may require access to a surface treatment/roughening process. Generally, a liquid that wets a smooth surface will wet a rough one even more. A rough surface will modify filler-metal flow from laminar to turbulent, prolong flow time, and increase the possibility of alloying and other interactions. Surfaces often are not truly planar, and, in some instances, surface roughening will improve the uniformity of the joint clearance. Conversely, the designer and engineer may require that filler metal not flow onto some surfaces. Stopoff materials often avoid this, but the design must permit easy application of the stopoff without danger of contaminating the surfaces to be joined. Self-fluxing filler metals in a suitably protective environment, such as vacuum, may provide the essential surface wetting. Examples include copper-phosphorus on copper and silver-copper-lithium on stainless steels; additional examples are given in Chapter 4, “Base Metals and Base-Metal Family Groups.”

Joint Design and Clearance A brazed joint is not a homogeneous body. Rather, it is a heterogeneous assembly that is composed of different materials with different mechanical, physical, and chemical properties. It is the simplest case; it consists of the basemetal parts to be joined and the added filler metal. Diffusion processes, however, change the composition and therefore the chemical and physical properties of the boundary zone formed at the interface between base metal and

Chapter 3: Elements of the Brazing Process / 19

filler metal. Thus, in addition to the two different materials present in the simplest example given previously, further dissimilar materials must be considered. Why should small clearances be used? The smaller the clearance, in situations where there is not extensive alloying and erosion, the easier it is for capillarity to distribute the filler metal throughout the joint area, and there is less likelihood that voids or shrinkage cavities will form as the filler metal solidifies. Small clearances and correspondingly thin filler-metal films make strong joints. The soundest joints are those in which 100% of the joint area is wetted and filled by the filler metal. They are at least as high in tensile strength as the filler metal itself, and often higher. If brazing clearances ranging from 0.03 to 0.08 mm (0.0012 to 0.003 in.) are designed, they are designed for the best capillary action and the greatest joint strength. Before the detailed design for a part to be brazed is made, the first decision is how and where the components are to be joined. Because brazing relies on capillary attraction, the design of a joint must provide an unobstructed and unbroken capillary path to enable flux, if used, to escape from it as well as to allow the filler metal to get into the joint. Where filler metal is added to a joint by hand, such as by feeding in a rod or wire, the joint entry must be visible and accessible. If preplaced rings or shims are used, the joint must be designed so that the preform can be placed in position easily and remains in place until molten. Some of the more important factors influencing joint design are the required strength and corrosion resistance, the necessary electrical and thermal conductivity, the materials to be joined, the mode of application of the filler metal, and postjoining inspection needs. The actual design itself should follow from a consideration of all the previously mentioned factors. Consideration also should be given to the ductility of the base metal, the stress conditions of the joint, and the relative movements of the two surfaces during joining, which may introduce problems and inaccuracies requiring careful consideration by the designer. The lack of concentricity of circular components may cause gaps to vary from excessive to over-tight. This can be avoided by designs and techniques such as knurling, splining, use of shims, and machining. Viscosity, surface tension, and specific gravity of the filler metal are not the only factors that

determine the gap-filling capability of a given filler metal. Many other considerations are involved, such as the tendency of the filler metal and parent materials to alloy with one another. Joint strength increases as joint gap decreases, down to a minimum. For stressed applications, optimal joint clearance may have to be designed inside the gap-filling range of the filler metal. Table 3.2 shows allowable joint clearances for various filler-metal systems. Other factors influencing optimal joint gap with a specific filler metal are joint length, brazing temperature, and base-metal reactions. It is important to remember that an assembly expands during heating and that the joint gap may either widen or close by the time the filler metal starts to melt and move. It is desirable to design the joint so that the solidifying filler metal is exposed to compressive rather than tensile stresses. This is much more important in brazing than in soldering, because brazing temperatures are higher, increasing the total expansion. With cylindrical joints, the component with the larger coefficient of expansion should, whenever possible, be on the outside. It is equally important to make sure that at the joining temperature the gap does not become impossibly wide. If the components in the assembly have to be reversed, design modifications can reduce stresses, but a sufficient gap must be provided to ensure that narrowing of the joint gap at the brazing temperature does not provide insufficient clearance for filler-metal flow. Finally, it is important to ensure that there is sufficient filler metal to absorb room-temperature tensile stresses in order to compensate for any reduction in joint gap.

Table 3.2 Preferred gaps for different brazing filler metals Brazing filler-metal system

Al-Si alloys(a) Mg alloys Cu Cu-P Cu-Zn Ag alloys Au alloys Ni-P alloys Ni-Cr alloys(b) Pd alloys

Joint clearance, mm (in.)

0.15–0.61 (0.006–0.024) 0.10–0.25 (0.004–0.010) 0.00–0.05 (0.000–0.002) 0.03–0.13 (0.001–0.005) 0.05–0.13 (0.002–0.005) 0.05–0.13 (0.002–0.005) 0.03–0.13 (0.001–0.005) 0.00–0.03 (0.000–0.001) 0.03–0.61 (0.001–0.024) 0.03–0.10 (0.001–0.004)

(a) If joint length is less than 6 mm (0.240 in.), gap is 0.12 to 0.75 mm (0.005 to 0.030 in.). If joint length exceeds 6 mm (0.240 in.), gap is 0.25 to 0.60 mm (0.010 to 0.024 in.). (b) Many different nickel brazing filler metals are available, and joint-gap requirements may vary greatly from one filler metal to another.

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Joint clearance is probably one of the most significant factors in all types of brazing operations. Naturally, joint clearance receives special consideration when joints are designed at room temperature. Actually, joint clearance is not the same at all phases of brazing. It has one value before brazing, another value at the brazing temperature, and still another value after brazing, especially if there has been diffusion of the filler metal into the base metal. To avoid confusion, it has become general practice to specify joint clearance as the value at room temperature before brazing. The recommended joint clearances given in Table 3.2 are based on joints having members of similar metals and equal mass. When dissimilar metals and/or metals of widely differing masses are joined by brazing, special problems arise that necessitate more specialized selection among the various filler metals, and the joint clearance suitable for the job at hand must be carefully determined. Although there are many kinds of brazed joints, selection of joint type is not as complicated as it may seem, because butt and lap joints are the two fundamental types. All others, such as the scarf joint, are modifications of these two. The scarf joint is identical with the butt joint at one extreme of the scarf angle and approaches the lap joint at the other extreme of the scarf angle. Selection of joint type is influenced by the configuration of the parts as well as by stress requirements and other service requirements,

Fig. 3.1

such as electrical conductivity, pressure-tightness, and appearance. Also influential in selection of joint type are fabrication techniques, production quantities, method of feeding the filler metal, and other factors. Lap joints are generally preferred for brazing operations, particularly when it is important that the joints be at least as strong as the weaker member. For maximum strength, the lap-joint length should equal 3 times the thickness of the thinner member. Figure 3.1 shows good and bad butt-joint and lap-joint designs. Butt joints are usually used where strength requirements may not be critical or where the use of a lap joint would be objectionable. The lap joint shown in Fig. 3.1 is the easiest type of brazement that can be made. The reasonable amount of overlap for lap joints is 3 to 6 times the thickness (T) of the thinner of the two members being joined.

Temperature and Time The temperature of the filler metal naturally has an important effect on the wetting action, because the wetting and alloying action improves as the temperature increases. Of course, the temperature must be above the melting point of the filler metal and below the melting point of the parent metal. Within this range, a temperature generally is selected that gives the best filler-metal wetting and flow. Usually, the lowest satisfactory brazing temperatures are preferred to (a) economize on heat

Comparison of some good and bad ways to assemble lap and butt joints. BFM, brazing filler metal. T, thickness

Chapter 3: Elements of the Brazing Process / 21

energy required, (b) minimize the heat effect on the base metal (annealing, grain growth, or warpage, for example), (c) minimize basemetal/filler-metal interactions, and (d) increase the life of fixtures, jigs, or other tools. Higher brazing temperatures may be desirable to (a) use a higher-melting but more economical or otherwise superior filler metal; (b) combine annealing, stress relief, or heat treatment of the base metal with brazing; (c) permit subsequent processing at elevated temperatures; (d) promote base-metal interactions in order to modify the filler metal (this technique is usually used to increase the remelt temperature of the joint); (e) more effectively remove surface contaminants and oxides with vacuum brazing; and (f ) avoid stress cracking. The time at brazing temperature also affects the wetting action, particularly with respect to the distance the filler metal can creep. If the filler metal has a tendency to creep, the distance generally increases with time. The alloying action between the filler metal and parent metal is, of course, a function of both temperature and time. In general, for production work, both temperature and time are kept at a minimum consistent with good quality. In conclusion, the filler metal and brazing process must be selected with a true understanding of both the physical metallurgy of the base material and the interactions of the base material with the filler metal.

Processes (Heat Sources)

joining requires efficient transfer of heat from the heat source into the joint. The heat capacity and thermal conductivity of the assembly must be considered. A 0.025 mm (0.001 in.) diameter wire cannot be brazed to a large mass of copper with a small torch. The size of individual assemblies, the numbers required, and the rate of production necessary influence selection of the heating method. Many other factors must be considered before the choice is made. The rate of heating, thermal gradients, and cooling rates, both external and internal, vary tremendously with different methods of heating, and the effects of these on dimensional stability, distortion, and metallurgical structure must be considered. The joint can be heated in many ways, which are commonly categorized by the actual method of heating. There are six commonly used methods: • • • • • •

Torch brazing Furnace brazing Induction brazing Dip brazing Resistance brazing Infrared brazing Lesser known heating methods include:

• • • •

Laser brazing Exothermic brazing Weld brazing Microwave brazing

Torch Brazing

The heating methods available often place a constraint on the designer and engineer in selecting the best type of capillary joint. In principle, there are many methods of heating available for brazing (Table 3.3). Effective capillary

Manual torch brazing is the method most frequently used for repairs, one-of-a-kind brazing jobs, and short production runs as an alternative to fusion welding. Heating the assembly with a

Table 3.3 Characteristics of various heating methods Characteristics(a) Method

Torch (flame) Electrical resistance Induction Furnace (atmosphere) Furnace (vacuum) Dip (flux bath) Infrared (a) H, high; M, medium; L, low

Capital cost

Running cost

Basic output

Flux required

Versatility

Operator skill required

L/M M M/H M/H H L/M M

M/H M M M/H L M/H L

L M/H M/H H H L/M M

Yes Yes Y/N Y/N No Yes Y/N

H L M M M L L

Yes No No No No Yes No

22 / Brazing, Second Edition

gas (oxyfuel) flame, either a hand-held torch or machine-mounted burners, is the most common method for brazing. Automatic setups comprising several burners can produce several hundred assemblies per hour. The torch brazing technique is relatively simple and can be mastered by the mechanically adept in a short time. Those already experienced in torch welding and the brazing of other metals generally encounter no difficulty learning torch brazing. Flux is normally required with the process. An exception is the joining of pure copper base metal with phosphorus-bearing filler metal. The phosphorus joins with the oxides on the surface to promote wetting of the filler metal. Torch brazing is widely used because of its relatively low cost and portability. The flame is generated by the combustion of a combination of oxygen and a fuel gas. The same equipment can be used as with oxyfuel gas welding. The process lends itself to the use of low-melting filler metals, which have excellent flow characteristics. Torches. Torch brazing employs the same type of torch, controls, and gases used for gas cutting and welding. To braze, the operator uses a brazing torch tip. In the oxyfuel torch, fuel gas and oxygen flow through separate tubes and combine in a mixing chamber, then travel through the torch tip before igniting. The outer barrel of the handle is often knurled or embossed to give a good grip. The typical torch is 305 to 600 mm long (12 to 24 in.) and 1.38 to 2.76 kg (3 to 6 lb) in weight. Handles are of brass or aluminum and stainless steel; tips are of copper alloys. Adjusting the gas flow alters the blend of gases to give the desired flame. Equal-pressure mixers, used with fuel gas above 13.7 kPa (2 psi), receive fuel gas and oxygen at nearly equal pressures to produce a neutral flame. The lowpressure or injector-type torch, for low fuel-gas pressures, receives oxygen at 69 to 207 kPa (10 to 30 psi); fuel gas aspirates into the mixing chamber. Injector torches come in versions for specific gases and pressure ranges. Torch Tips. Brazers fit torches with tips larger in diameter than those used for welding. Tip style varies with fuel gas—brazing with acetylene requires a tip with a squared-off orifice face; orifices of tips for natural gas and propane are recessed (Ref 4).

Fuel Gases. Depending on the temperature and heat required, all commercial gas mixtures can be used to fuel the torch: oxyacetylene; oxyhydrogen; oxynatural gas; acetylene and air; hydrogen and air; propane, methane, and natural gas and air; and certain proprietary mixes. Flame temperature and cost are major factors in selecting a fuel gas. Oxyacetylene and oxynatural gases are the mixtures most often used commercially and are preferred in that order. The adjustment of the flame is very important. Generally, a slightly reducing flame is desirable. The oxyacetylene combination produces the highest temperature. The other gases are cooler, and their flames are less concentrated. Thus, they are easier to use and advantageous on lightgage material. Flame Characteristics. As the ratio of oxygen to fuel gas changes, the nature of the flame changes. The most common flame conditions

Fig. 3.2

Conditions for an oxyacetylene flame

Chapter 3: Elements of the Brazing Process / 23

used are reducing and neutral. Oxidizing flames are not recommended, because they oxidize the filler and base metals, and such oxidation impedes wetting and flow (Fig. 3.2). Carburizing Flame. When the ratio of oxygen to fuel gas is small, the flame is carburizing, and soot is produced. As oxygen is added to the fuel gas, the flame becomes luminous, and sooting disappears. As the oxygen content is increased, the luminous part of the flame becomes smaller and is centered near the torch tip. A blue zone, consisting of an excess of fuel gas, forms around the outside edge of the flame. This flame can be used for brazing. Reducing Flame. As the oxygen content is increased further, the luminous area becomes smaller and consists of an inner cone with a feathery trail extended out toward the flame end. This condition indicates a slight excess of fuel gas and is an excellent flame for brazing. Neutral Flame. When the oxygen addition reaches the ratio necessary for the fuel gas to be completely combusted, the feather that extended out from the bright inner cone disappears. This flame is used in brazing when an excess of carbon in the reducing flame is detrimental to the base metals or when maximum flame temperature is required. Oxidizing Flame. When the oxygen-to-fuelgas ratio exceeds that needed for complete combustion, the flame becomes oxidizing. The flame produces a hissing sound. The inner cone in the flame appears to be constricted. As stated previously, oxidizing flames are not recommended for brazing. Fuel Flame. When hydrocarbon fuel gases are burned without oxygen or are added through the torch, they typically produce a yellowish flame. Soot particles are usually present, because the oxygen in the air is not sufficient to support complete combustion. This flame is not useful in brazing. Apparatus. The torch-brazing operator also needs gas cylinders and the appropriate accessories, including regulators, valves, hoses, check valves, and flashback arrestors. For manual brazing, some operators prefer oxygen combustion for its high heating rate. Compressed air can be used, although flame temperature and heating rate are lower. In manual torch brazing, the filler metal is usually face fed in the form of wire or rod, or preplaced. In the latter case, care must be exercised in the placement of the filler metal and the

guidance of the torch to preclude premature melting of the filler metal. One way to prevent overheating is to use flux with a melting temperature not too far below that of the filler metal. The proper brazing temperature is indicated when the flux becomes liquid. To ensure uniform heating throughout the joint, which is very important, it may be advisable to use a multipletip torch or more than one torch. Mechanization. In automating the torch brazing process, fabricators use devices that move the workpiece or torches. Typically, the workpiece indexes—by conveyor or turntable—through one or more torch stations. At any station, brazing may be manual or automatic. Stations can be fitted with automated feeders of flux and filler metal or can position filler-metal rings at the joint. Single or multiple burners can be mounted at each station. Other stations can house cooling devices, preheating torches, and loading and unloading devices. Figure 3.3 shows a specially designed tenstation rotary brazing system that has enabled a major manufacturer of air-conditioning components to automate production of copper manifold assemblies. One operator loads the copper assemblies; all other functions are automatic. The new system completes 250 parts/h and features adjustable fixtures that allow the machine to accommodate 22 different part cycles. The key to the versatility of the system is the use of a paste filler as the brazing medium. Finely atomized filler metal, proper flux, and a creamy neutral binder are blended into one homogeneous mixture. After paste application, the assemblies move through three natural-gas-fired heating stations, where the copper-phosphorus alloy liquefies at 795 °C (1465 °F). The brazed parts are then gradually cooled via an air quench and a water quench before being manually unloaded by the operator.

Furnace Brazing The popularity of furnace brazing derives from the comparatively low cost of equipment, the adaptability of the furnace, and the minimal jigging required. With many brazing assemblies, the weight of the parts alone is sufficient to hold them together. With other configurations, one or two rectangular blocks of metal are all the fixturing needed.

24 / Brazing, Second Edition

Furnace brazing is a medium- to high-volume production process for self-fixtured assemblies with preplaced filler metal. Operators lay filler metal as a thin solid sheet or as a clad layer into the joint before inserting the assembly into the furnace. The furnace is then purged with a gaseous atmosphere or evacuated of air and heated to a temperature above the liquidus of the filler metal but less than the melting point of the base metals. The brazements are then cooled or quenched by appropriate methods to minimize distortion and produce the required properties in the filler and base materials. This cycle is designed to produce the required melting and solidification of the filler metal to join the components without melting or impairing the properties or shape of the base metals. The process joins parts that would otherwise distort from the localized heating of torch brazing. Compared to other brazing processes, capital equipment cost can be high, but it may be easily recovered by the high throughput possi-

Fig. 3.3

ble with batch processing in furnaces. Jointdesign skill is critical for furnace brazing, because brazing occurs without observation or intervention by an operator. There are two basic types of furnaces used for brazing: batch furnaces that process mediumproduction work lots and continuous conveyorfed furnaces that turn out the highest-volume production. Heating is typically by resistance, but furnaces that combine radiation and naturalgas convection heating can decrease heating time by 50%, resulting in a more uniform heating of the part, and can reduce energy costs. The parts should be self-jigging or fixtured and assembled, with filler metal preplaced near or in the joint. The preplaced filler metal may be in the form of wire, foil, powder, paste, slugs, or preformed shapes. Fluxing is used except when a reducing atmosphere, such as hydrogen, and either exothermic or endothermic combusted gas can be introduced into the furnace. In some instances, both flux and a reducing atmosphere

Automated brazing system using a special paste alloy to braze up to 250 parts/h. Source: Ref 5

Chapter 3: Elements of the Brazing Process / 25

may be necessary. Pure, dry, inert gases, such as helium and argon, are used to obtain special atmospheric properties. For brazing in a batch furnace, the operator loads all of the assemblies at one time into the furnace. Loading is from the top, side, or bottom of the furnace. In loading from the bottom, the furnace is raised above a pit-type work area; after loading, the furnace is lowered over the work area. With cycle times as short as 15 min, batch furnaces can produce several hundred brazements a day. The simplest type of batch furnace is the muffle type. These supply heat by gas combustion. For better quality, brazers employ electricresistance or radiant-tube furnaces or furnaces that combine resistance and combustion heating. Box-type furnaces load and unload from the side (Fig. 3.4). These connect the heating chamber and a water-encased cooling station, each with its own source of controlled atmosphere. A refractory-lined baffle door separates the chambers. Work moves through the furnace and cooling station in trays or baskets. Most batch furnaces are lined with refractory material. With the development of nickel brazing came the need for a furnace that could withstand high heat—the retort-bell-type combustion furnace. For brazing in a purified-hydrogen atmosphere, these comprise an inner container of a heat-resistant alloy sealed from outside air and the products of combustion. They are twopiece containers—an operator loads the workpiece on a base, and a cover lowers over the base. Sand or an O-ring seals the base-cover joint. Dry hydrogen enters the top of the cover; the container is purged and then lowered into a pit-type

Fig. 3.4

Electrically heated batch box-type brazing furnace

furnace. Hydrogen flows through the container during preheating, brazing, and cooling. When continuous-type furnaces are used, several different temperature zones may be used to provide the proper preheating, brazing, and cooling temperatures. The speed through a conveyor-type furnace must be controlled to provide the appropriate time at the brazing temperature. It is also necessary for the assembly to be properly supported so that it does not move while traveling on the belt. This may require special fixtures, but most often, brazements are designed to be self-supporting. Compared to batch-furnace brazing, part quality may be higher and manual labor greatly reduced. Typical production volume is 100 to 1000 kg (220 to 2200 lb) of brazed assemblies per hour. There are various types of continuous furnaces. For example, a newly designed and built furnace is a hydrogen-atmosphere hump meshbelt electric furnace rated for temperatures up to 1177 ± 11 °C (2150 ± 20 °F). The furnace has a nitrogen curtain located on the furnace entrance ramp that is acceptable by the Occupational Safety and Health Administration (OSHA) code to isolate explosive gases. The furnace has two heating zones, the first with 18 electric heating elements and the second with 12 heating elements. The hot zone measures 2.86 m (9.4 ft) in length, with a 0.9 m (3.0 ft) slow-cooling zone before the exit. The furnace walls consist of an internal wall of ceramic blanket, an intermediate wall of ceramic-fiber board, and an outer steel shell. The innovative design of the hydrogenatmosphere hump mesh-belt furnace system has produced parts such as those shown in Fig. 3.5.

26 / Brazing, Second Edition

Innovative designs such as this continue to provide insight into the advanced technology being made available to the thermal processing industry and display the potential for development of advanced materials processing systems (Ref 6). A second type of semicontinuous controlledatmosphere brazing (CAB) furnace is shown in Fig. 3.6. This type of furnace was designed for brazing aluminum heat exchangers. Additionally, this type of furnace and control system was designed to accommodate the production demands in the small-to-intermediate-sized brazing shops. By monitoring the temperature differential through the load and adjusting the fan speed (i.e., heating rate) and cycle time accordingly, the semicontinuous furnace system has the ability to successfully braze parts of various dimensions or those loaded in different configurations under a single set of parameters. Part-to-part temperatures can be maintained at the recom-

Fig. 3.5

mended temperature range of 600 to 620 °C (1110 to 1150 °F) for 3 to 5 min, with a temperature of ±3 °C (±5 °F). Although CAB brazing of aluminum heat exchangers is a mature process, significant improvements in both product yield and quality can be achieved through process modifications that include both furnace design and proper materials selection. A prototype CAB braze furnace design is shown in Fig. 3.7 and is currently envisioned for use in a typical aluminum heat-exchanger manufacturing plant. The “intelligence” of the furnace centers around a fuzzy logic controller and a CAB thermal model that are based on a personal computer or Unix software architecture. To meet the future braze quality requirements of aluminum heat-exchanger manufacturing, it will be necessary for furnace designers and suppliers to consider making CAB braze furnaces that are smarter, more adaptive, and highly costeffective. This can be done by using logic con-

Examples of aerospace brazed manifolds and tube assemblies. The part on the right is approximately 61 cm (24 in.) in length.

Gas fired burners Fan assembly

Atmosphere curtains

Fan assembly

Airflow baffle

Muffle

Atmosphere curtains

Air blast curtains

Muffle tension cylinder

Load area/dryoff oven

Fig. 3.6

Entrance purge chamber

Convection chamber

Fan assemblies

Water jacketed cooling chamber

Schematic of the semicontinuous controlled-atmosphere brazing furnace system. Source: Ref 7

Air blast chamber

Wire brush assembly

Chapter 3: Elements of the Brazing Process / 27

trollers that can self-adjust furnace conditions for optimal efficiency, using image systems to increase potential problem awareness, using new materials that can provide durability and are cost-effective, and by coming up with new designs that reduce floor space requirements without reducing the quality of the process. It should be noted that several furnace manufacturers have installed twin silica (woven tape) curtains at the entrance and exit ends of brazing furnaces and have reduced natural-gas consumption by approximately one-third (Fig. 3.8) (Ref 8). The curtains serve both as a barrier to and combustion chamber for escaping gases. An additional benefit arises from the fact that the product line speed can be increased during continuous furnace operation. The tape curtains used perform at temperatures varying from 760 to 1100 °C (1400 to 2000 °F). In this particular application, they lasted for an average of one week—very good service, considering that quite a number of other materials evaluated over the years could not withstand the high temperatures involved. The natural-gas savings are very significant. Although the silica-based tape was found to be a little more expensive than some of the other filament tapes, its ability to withstand the high

temperature and the added service life, coupled with the reduced energy costs, made it very attractive. The effective cooling of components in continuous nitrogen-hydrogen atmosphere furnaces depends on the heat-transfer efficiency of the furnace atmosphere. The cooling of parts in a continuous furnace is controlled by the convective heat transfer, which is greatly influenced by the composition of the nitrogen-hydrogen atmosphere. The heating of parts is dominated by radiant heat transfer from the furnace walls and is not greatly influenced by the composition of the atmosphere. The cooling capacity or the productivity of a continuous furnace can be manipulated simply by changing the concentration of hydrogen in the atmosphere. This provides heat treaters added flexibility to optimize their atmosphere requirements, depending on the production demands. A model has been developed based on heattransfer principles to determine the cooling capability of a continuous furnace. The model indicates that the cooling capability can be increased by increasing hydrogen concentration in a nitrogen-hydrogen atmosphere or by using a high-convection cooling system. The selection of a higher hydrogen concentration or a

4. Infrared imaging cameras

5. Innovative muffle & element designs 6. Innovative belt & fixture designs

3. Pre-braze convection 2. Optical part sensor

1. Intelligent fuzzy logic

Fig. 3.7

Proposed next-generation furnace for controlled-atmospheric brazing. Source: Ref 8

28 / Brazing, Second Edition

high-convection cooling system would, however, depend on the desired increases in cooling capability and on the overall economics (Ref 9). A large volume of furnace brazing is performed in vacuum, which prevents oxidation and often eliminates the need for flux. Vacuum brazing has found wide application in the aerospace and nuclear fields, where reactive metals are joined or where entrapped fluxes would be intolerable. If the vacuum is maintained by continuous pumping, it removes volatile constituents liberated during brazing. There are

several base metals and filler metals that can be harmed by brazing in a vacuum, because their low-boiling-point or high-vapor-pressure constituents cause part of the metal to be lost. Vacuum is a relatively economical method of providing an accurately controlled brazing atmosphere and is an effective means of screening the work to be brazed from oxidizing gases and other impurities. The vacuum pressures used for brazing generally range from 0.13 to 1.3 Pa (2 to 20 × 10–4 psi). This range corresponds to a gas that is several hundred times purer than the purest gas used for atmosphere brazing. Vacuum brazing does not allow as wide a choice of filler metal as does atmosphere brazing. Vacuum furnaces are invariably heated by electricity in any one of a number of forms. Vacuum brazing takes place in either a singleor double-pumped retort or in a batch-type furnace. A single-pumped retort is loaded with the brazement, evacuated, and heated externally by resistance or by gas or oil combustion. The maximum furnace temperature is 1150 °C (2100 °F), although most operate at 870 °C (1600 °F) or lower. For higher-temperature brazing, manufacturers can employ a doublewall retort in which the workload sits in an inner high-vacuum (1.33 Pa, or 20 × 10–4 psi, or lower) container that sits inside a rough-vacuum

Fig. 3.8

Brazing furnace using silica-based woven tape curtains to reduce natural-gas consumption

Fig. 3.9

Diagram of cold-wall vacuum furnace showing the various components

Chapter 3: Elements of the Brazing Process / 29

(1.33 to 13.33 × 102 Pa, or 0.19 to 1.9 psi) chamber. Thermal insulation and the electrical heating elements fit between the walls of the outer and inner retorts. Batch-type vacuum furnaces can be a hot- or cold-wall type; the cold-wall furnaces are more popular. The furnace diagram in Fig. 3.9 illustrates the major components of a simple batch-type fluxless vacuum furnace for aluminum brazing. Figure 3.10 illustrates the microstructure of a fluxless brazed aluminum joint, while Fig. 3.11 shows a graphic sequence of events for a batchtype furnace. The cold-wall-type furnace in Fig. 3.9 consists of a horizontally charged, water-cooled

Fig. 3.10

1200

+3

650 Door opens Braze

760 mm atmospheric pressure

10 +2

Mg evaporation

Soak

1000

10 –2

Vacuum cool Convection cool

800

Temperature ˚F

Chamber vacuum torr

Rough vacuum

10 –0

10 –1

540

Mg sublimation

10 +1

10 –3

425

Temperatures: 600

Hot start cycle repeat

Program profile 100 micron staging pressure

315

Furnace actual Work surface Work internals

400

205 Vacuum break

10 –4 200

95

10 –5

High vacuum

Cold start 10 –6

0

0 5

10

15

Time, minutes

Fig. 3.11

Sequence of events during vacuum furnace brazing cycle

20

25

Temperature ˚C

10

Microstructure of furnace-brazed joint after fluxless aluminum brazing

vacuum furnace chamber with an internal resistance-heated hot zone, reflective metallic heat shields, an adjustable power supply, a single set of vacuum pumps, a water recirculating system, and a control panel. The operation can be initiated by automatic or manual sequencing. Depending on the geometry of the workpiece and the desired production rate, the hot-zone size can vary from less than 0.28 m3 (10 ft3) up to 17 m3 (600 ft3) or more. The hot zone can be rectangular, cylindrical, or flat in shape. There may be singular or multiple temperature-control zones within the hot zone. The designs to be considered depend on the workpiece shape, size, desired production rate, and user preference. Batch furnace cycle times can vary from 8 min to longer than a day. Small, lightweight parts that would fit into a 0.028 m3 (1.0 ft3) hot zone can be heated and brazed in approximately 8 to 20 min. Cycle times for most aerospace, automotive, and electronic components vary from 10 to 45 min. Large components, such as cryogenic heat-exchanger cores, have cycle times that vary from several hours up to 24 to 36 h. In the latter case, heat-exchanger cores measuring 1.2 × 1.2 × 6.1 m (3.9 × 3.9 × 20.0 ft) and weighing in excess of 9075 kg (20,000 lb) are routinely brazed in vacuum. Cold-wall-furnace configurations include horizontal designs with front loading, vertical designs with top or bottom loading, or clamshell

30 / Brazing, Second Edition

designs with front loading. Heated by radiation, temperatures peak at 2200 °C (4000 °C), and vacuums reach 0.00013 Pa (20 × 10–6 psi). In addition to batch-type furnaces, there are semicontinuous furnaces with entrance and exit vacuum locks in series with single or multiple heating stations. In all of these variations, the basic brazing process is essentially the same. The parts to be brazed are exposed to high vacuum, and heat is applied in a uniform and controlled manner (Ref 10). Semicontinuous furnaces have found application in high-volume production situations, such as the automotive industry, where this type of operation is customary. Figure 3.12 illustrates a system comprised of three chambers in series. Each chamber section is roughly equivalent in size to a single batchtype furnace chamber. The entrance and exit lock chambers are isolated from the heating chamber and from the atmosphere by large, automated gate-valve-type doors that are water cooled. Vacuum-sealed shafts are used for driving and positioning these doors each time a load of parts is charged. A workload may be present in all three of the chamber sections. A workload in the entrance lock could be undergoing evacuation and preheating while the workload in the furnace section is heated and brazed in a high-vacuum environment. Simultaneously, vacuum is broken in the exit lock, the workload is discharged, and the exit lock is closed, sealed, and then re-evacuated. Although the furnace chamber is exposed to intermediate vacuum levels and gaseous contamination from incoming work when the entrance-lock furnace door is opened and the work is indexed, the workload is not exposed to intermediate vacuum during the critical brazing portion of the cycle.

Plan view

Mechanical pumps (3) Booster pumps (2/3) Diffusion pumps (1/3)

Fan cool Vacuum tight doors (4) Exit vacuum lock (optional) Heat zone

Fig. 3.12

Power supply Entrance vacuum lock with heating capability Control panel

Semicontinuous vacuum furnace with threechamber design

The appeal of this vacuum furnace arrangement stems, in part, from the fact that the furnace chamber is not exposed to atmospheric conditions, thereby eliminating moisture and gaseous adsorption The quality of the brazed work is comparable to that obtained in a conventional batch furnace that is properly designed for aluminum brazing under vacuum. In the brazing of automotive aluminum heat exchangers, the vacuum furnace is equipped with a sophisticated residual gas analyzer for monitoring key brazing factors, such as magnesium evolution and oxidizing gas gettering. There are a number of other reasons, based on process mechanisms, why brazing aluminum under vacuum is advantageous. Oxides form on the surface of aluminum when it is exposed to air, which restricts the flow of molten filler metal. When aluminum is heated, the differential expansion of nascent aluminum and its thin oxide causes cracks in the oxide layer. If oxygen is available, the aluminum oxide envelopes readily reform. In vacuum, however, magnesium (a getter) in the filler metal vaporizes as it reaches the liquidus temperature, and the magnesium vapors react with any traces of residual oxygen and that absorbed in components within the chamber, in accord with these most prevalent reactions: Mg + H2O 3 MgO + H2 2Mg + O2 3 2 MgO Mg + CO2 3 MgO + CO

Good operating practice allows the maintaining of broken oxide film, with resultant wetting by the fluid aluminum filler, to form fillets. The remaining platelets, on breaking of the oxide film, become minute oxide inclusions without any significant deleterious effect. Heating Elements and Radiation Shields. Elements must withstand low vapor pressures and high temperatures and must have a large heating surface for good heat transfer. The most commonly used materials for elements, and their maximum operating temperatures, are molybdenum, 1900 °C (3450 °F), or 1590 °C (2900 °F) in high vacuum; tungsten, 2480 °C (4500 °F); and tantalum, 2205 °C (4000 °F). Radiation shields (Fig. 3.13) must perform the dual function of containing the heat and protecting the rubber vacuum seals that are incorporated in the door of the furnace and are fitted around all power leads or other controls that pass

Chapter 3: Elements of the Brazing Process / 31

through the furnace wall. A less obvious consequence of effective insulation is that the cooling rate is severely retarded. Because of this, the overall time of a vacuum brazing operation is usually much longer than that of other batch brazing techniques; this results in expensive equipment being tied up for long periods in processing comparatively small workloads. It should be remembered, however, that most of the metals on which vacuum brazing excels are also costly, and that any process that permits material economies through fabrication and ensures the necessary joint properties is justified. Equipment developments that have reduced processing times include transfer mechanisms, multizone furnaces, and inert-gas quenching. Controls. Brazing furnaces must be fitted with automatic temperature-controlling-and-recording devices. Control cabinets house vacuum gages and temperature instruments and controllers. Microprocessor-based units monitor and control temperature and holding times for the complete brazing cycle through heating, brazing, cooling, and quenching. They control the atmosphere, signaling gas-flow timing, and pressures. The previously mentioned devices and systems are able to control these parameters: temperature, temperature ramp rate, vacuum leak check, vacuum level, cooling, cycle events, and soak times, along with the capability of logging

Fig. 3.13

Vacuum furnace hot zone showing radiation shields

the various heating cycles performed by the vacuum furnaces and logging of process alarms and cycle modifications that are new additions to brazing furnace operations. Every production vacuum furnace operation should have its own fully automatic processing programmer, complete with data-logging instrumentation; run at vacuum levels of 0.00013 Pa (20 × 10–6 psi) to temperatures of at least 1450 °C (2650 °F); quench up to 100 kPa (2 bars) gas pressure; and run in partial pressures of nitrogen, argon, hydrogen, and helium. In addition, all horizontal vacuum furnaces have rail-guided work loaders to maintain part integrity, complete with a full overhead crane system. Cooling Systems. Some furnace brazements, depending on the specific alloys involved, must cool rapidly to ensure good mechanical properties and corrosion resistance. To quench the workpiece after brazing, backfill inert gases flow into the furnace. The backfill gases used are argon, nitrogen (except if base materials contain titanium, beryllium, zirconium, or aluminum), helium, and hydrogen. There are two preferred methods that are used for back filling: • With an internal gas-quenching setup, inert gases flow into a plenum behind the insulated hot furnace zone. Baffles and nozzles direct

32 / Brazing, Second Edition

the gas toward the brazement, and hot gases recirculate through a heat exchanger that cools the gas and redirects it to the work. • In vacuum furnaces with turbocharged gas quenching, backfill gases flow into the furnace hot zone in an orbital pattern. An internal turbofan recirculates this gas around the work, and a heat exchanger keeps cool gas flowing. Atmospheres. Furnace brazing takes place in vacuum or in a controlled atmosphere of high-purity inert or reducing gas. Both reduce the amount of flux needed or, in some cases, eliminate the need for fluxes altogether, to prevent oxidation during heating. Accurate, uniform temperature control and consistent atmospheric protection during brazing and cooldown ensure sound joints. Vacuum brazing often is preferred for joining alloys that contain chromium, aluminum, or titanium and for dissimilar-metal combinations of titanium, zirconium, niobium, molybdenum, or tantalum. To avoid contamination, these alloys also can be brazed in inert-gas atmospheres, if the gases are of high purity. Selecting the best controlled atmosphere for furnace brazing depends on base and filler materials and the capability of the furnace to maintain atmosphere quality through the cycle. Controlled atmospheres, which prevent formation of oxides to ensure good filler-metal flow and wetting, can be active or reducing, inert or relatively inert. Compared to vacuum brazing, atmosphere brazing enables use of a wider selection of filler metals, because some fillers contain low-boiling-point or high-vapor-pressure metals that can evaporate during vacuum brazing. In controlled-atmosphere brazing, gases pumped into the furnace are generated from natural gas or gases based on dissociated ammonia, nitrogen, hydrogen, argon, or helium. The escalating cost of natural-gas derivatives has led to increased use of nitrogen or hydrogen-nitrogen blends. Gas selection depends on base-material properties. For example, some grades of copper that contain copper oxide should not be brazed in reducing hydrogen atmospheres, because steam can form as the brazing process reduces the dissolved copper oxide. The compositions of American Welding Society (AWS)-designated controlled atmospheres recommended for brazing cover a wide

range (Table 3.4). These data are not intended as a comprehensive tabulation of atmospheremetal combinations but rather as a general outline of some of the more widely used combinations. Hydrogen in atmospheres reduces most metal oxides at elevated temperatures but can cause hydrogen embrittlement in some materials. Carbon monoxide reduces oxides of iron, nickel, cobalt, and copper. It is toxic, however, and must be properly vented or burned off. Carbon dioxide, properly mixed with carbon monoxide, inhibits decarburization and maintains a stable atmosphere when brazing steels. When using CO2-CO atmospheres, furnaces must be airtight; otherwise, the CO2 content can rise. Nitrogen in the atmosphere displaces air from the furnace and acts as a carrier gas for other atmosphere gases. High concentrations of nitrogen in the atmosphere are avoided when filler metals susceptible to nitriding—chromium, molybdenum, titanium, and zirconium—are present.

Induction Brazing Resistance to the flow of electricity induced by coils placed around the workpiece provides the heat for the induction brazing process. Resistance heating melts preplaced filler metal. Power densities range from 0.5 to 1.5 kW/cm2. Some advantages of the process include selective heating of the workpiece for brazing only where needed, quick attainment of brazing temperature, and the use of a variety of fluxes. This process lends itself to high-volume production applications that can be controlled remotely. High-strength components can be induction brazed with little loss of strength because of the precise heating capabilities of the process. Heating is localized to the part surface or just below, which is an advantage when joining components where metallurgical changes cannot be tolerated and on parts that allow minimal or no distortion. Most induction brazing occurs in air. Inductor design limits the complexity of assemblies, often making furnace brazing a more viable alternative for complex assemblies of several joints at one time. The process requires close, accurate part fit-up. Initial equipment cost can be high. The process relies on high-frequency alternating currents flowing through the induction

Combusted fuel gas (decarburizing)

Combusted fuel gas, dried

Combusted fuel gas, dried (decarburizing) Dissociated ammonia

Cryogenic or purified N2 + H2

Cryogenic or purified N2 + H2 + CO Cryogenic or purified N2 Deoxygenated and dried hydrogen Heated volatile materials (inorganic vapors—zinc, cadmium, lithium, volatile fluorides Purified inert gas (e.g., helium, argon)

Purified inert gas + H2 (e.g., helium, argon) Vacuum above 266.6 Pa (2 torr)

Vacuum from 66.65 to 266.6 Pa (0.50 to 2 torr) Vacuum from 0.13 to 66.65 Pa (0.001 to 0.50 torr)

Vacuum of 0.13 Pa (0.001 torr) and lower

2

3

4

6A

6B

9A

10A

10C

...

...

...

...

...

...

...

–68 (–90) –59 (–74)

–29 (–20)

–68 (–90)

–54 (–65)

–40 (–40)

–40 (–40)

RT

RT

Maximum dew-point of incoming gas, °C(°F)

...

...

...

...

1–10

...

...

... 100

2–20

1–30

75

38–40

15–16

14–15

5–1

H2

...

...

...

...

...

...

...

100 ...

70–97

70–99

25

41–45

73–75

70–71

87

N2

...

...

...

...

...

...

...

... ...

1–10

...

...

17–19

10–11

9–10

5–1

CO

...

...

...

...

...

...

...

... ...

...

...

...

...

...

5–6

11–12

CO2

Composition of atmosphere, %

BNi, BAu, BAlSi, titanium alloys

BCu, BAg

BCu, BAg

BCuP, BAg

...

Same as 5

BAg

Same as 5 Same as 5

Same as 5

BAg(b), BCuP, RBCuZn(b), BCu, BNi Same as 5

Same as 2

Same as 2

BCu, BAg(b), RBCuZn, BCuP

BAg(b), BCuP, RBCuZn(b)

Filler metal

Heat- and corrosion-resisting steels, aluminum, titanium, zirconium, refactory metals

Carbon and low-alloy steels, copper

Low-carbon steels, copper

Copper

...

Same as 5 plus titanium, zirconium, hafnium

Same as 3 Same as 5 plus cobalt, chromium, tungsten alloys, and carbides(e) Brasses

Same as 4

Same as 3

Copper(c), brass(b), low-carbon steel, nickel, Monel, mediumcarbon steel(d) Same as 2 plus medium- and highcarbon steels, Monel, nickel alloys Same as 2 plus medium- and highcarbon steels Same as 1–4 plus alloys containing chromium(e)

Copper, brass

Base metals

...

...

...

...

Special purpose. Parts must be very clean and atmosphere must be pure ...

Special purpose. May be used in conjunction with 1–5 to avoid use of flux

... ...

...

...

...

Referred to commonly as exothermic-generated atmospheres Decarburizes. Referred to commonly as endothermicgenerated atmospheres Referred to commonly as endothermic-generated atmospheres Carburizes

Remarks

RT, room temperature (a) types 6, 7, and 9 include reduced pressures down to 266.6 Pa (2 torr). (b) Flux required in addition to atmosphere when alloys containing volatile components are used. (c) Copper ahould be fully deoxidized or oxygen-free. (d) Heating time should be minimized to avoid objectionable decarburization. (e) Flux must be used in addition to the atmosphere if appreciable quantities of aluminum, titanium, silicon, or beryllium are present.

10B

10

9

8

6C 7

5

Combusted fuel gas (low hydrogen)

Source(a)

1

Brazingatmosphere number

Table 3.4 Atmospheres for furnace brazing

Chapter 3: Elements of the Brazing Process / 33

34 / Brazing, Second Edition

coils to create an electromagnetic field around the workpiece. The opposing currents that are induced into the workpiece generate the heat for brazing. The filler metal can be preplaced and can be in the form of wire, strip, or powder. The preform should be in good contact with the joint. Joint clearances of 0.05 to 0.13 mm (0.002 to 0.005 in.) have proven to be well suited for production work, but clearances up to 0.20 mm (0.008 in.) have been acceptable. Joint design should take thermal expansion into consideration. Large shear areas in the brazed joint provide maximum strength. Types of Joints. The most common joints— as with other heating methods for brazing—are butt and lap. A lap joint would be used to join two tubes, particularly if the joint requires strength and pressure-tightness (such as in the case of pumping a liquid through the pipes). In this example, the lap joint is easier to apply and provides more surface area to bond the assembly. Irregular-shaped parts do not lend themselves to induction brazing very readily, due to the complexity of coil configuration and heating

Fig. 3.14

Induction brazing of a steel base to a cast iron nose

uniformity. Any new user would, therefore, do well to seek the advice of the equipment manufacturer. Regardless of the type of joint, a good repeatable fit is desired. In assembly, one must consider that assembly is done at room temperature while brazing is at a higher temperature and that the parts expand. In the brazing of a steel base to a cast iron nose, for example, the coefficients of expansion of the two materials (cast iron, 18.9 × 10–6/°C, or 10.5 × 10–6/°F; steel, 14.9 × 10–6/°C, or 8.3 × 10–6/°F) must be considered in assembly design and heating for brazing (Fig. 3.14, 3.15). At 760 °C (1400 °F), the change of the joint gap is 0.06 mm (0.002 in.), and the gap is reduced. Coils. Induction brazing requires a coil, induction generator, and fixturing. Coils are made of copper tubing 4.76 to 9.52 mm (0.19 to 0.37 in.) in diameter, round or flattened. Depending on joint configuration, coils may be of square- or rectangular-cross-section tubing and are typically custom designed for each application. Size, shape, contour, number of turns, and

Chapter 3: Elements of the Brazing Process / 35

turn spacing all affect the strength of the electromagnetic field and the heat pattern. Examples of various coil designs are shown in Fig. 3.16. Proper coil design is important to the effectiveness of the brazing process. The geometry of the joints and the conductivity of the base metal must be taken into consideration. The heating pattern should be such that the joint area is heated uniformly. The joint area should reach the brazing temperature first to allow proper flow of the filler metal.

Induction Generator. Generators are rated by output power and frequency. For example, units can supply 25 to 200 kW power at 10 Hz and 15 to 50 kW at 25 kHz. The power rating of the induction generator and its ability to control current in the inductor directly influence heating rate. A low-power generator decreases heating rate, which allows time for temperature in the brazing joint to equalize. Generators may be one of three types: motor generator, tube oscillator, or solid state. Solidstate types have all but replaced motor-generator machines. Each of these generators operates at different frequency ranges—solid state from a few to 100 kHz and tube oscillators at 150 to 450 kHz. While frequency is often not of great importance in brazing, because it affects only the rate of surface heating as opposed to depthof-joint heating, brazing thin sections with a high-frequency (450 kHz) tube generator will likely produce better results than with a low-frequency solid-state unit. Conversely, heating with a solid-state generator with frequencies in the 8 to 10 kHz range works best on heavy sections. For sizing the generator needed for a particular application, fabricators take into account required production rate, base materials and their ability to absorb power (Fig. 3.17), and radiation from the work. A large number of induction-brazing defects stem from improper selection of power level, resulting in excessive heating rate. Calculate power absorbed by the workpiece from this equation: P = WTC/0.95t

Fig. 3.15

Diagram of assembly of cast iron and carbon steel components for induction brazing

Tube

where P is absorbed power in kilowatts, W is kilograms of material heated by the induction coil and by thermal conduction away from the

Induction coil

Tubing

Split Induction coil

Flange Alloy ring

Alloy ring Tube Induction coil

Fig. 3.16

Plug

Induction coil designs for specific applications

Tubing

36 / Brazing, Second Edition

joint, T is temperature rise in °C; C is mean specific heat of the base material, and t is heating time in seconds required to meet production requirements. Because determining W can be difficult due to the temperature gradient that develops in the work, use the accompanying graph to size an induction generator. The graph plots absorbed kilowatt per pound (kilogram) of workpiece versus temperature. Knowing the brazing temperature and base material, find the corresponding kilowatt per hour, and multiply by desired production rate, in pounds per hour (kilogram per hour), to determine absorbed power. Fixturing for induction brazing can be simple, especially if the assembly is self-jigging. If not, tack welding often suffices. If fixtures are necessary, construct any portion of the fixture within approximately 50 mm (2 in.) of the inductor coil of heat-resistant nonmetallic materials, such as ceramics or quartz. Any metallic materials used should be nonmagnetic, such as austenitic stainless steel or aluminum. Good fixturing contributes to repeatability of brazed parts. Flux and Atmospheres. Flux is used to clean the part while heating and to reduce the formation of oxides during heating. It is important to first clean the parts and flux them immediately prior to brazing. If the parts are fluxed and the flux is allowed to dry, it might

Brazing temperature, °C 540

760

1205

980

0.11

0.25

0.09

0.20

er

0.15

an

St

pp

Co

mi

0.06

Tit

nu m

ee l

0.07

ium

Ni

0.08

ck el

0.10

0.05

0.10

0.04 0.03

0.05

0.02

Absorbed energy, KW/kg of base material

315

Alu

Absorbed energy, kW/lb of base material

0 95

flake off. Critical to the success of the process is how the flux is applied and the amount that is applied (neither too little nor too much). Sometimes, it is possible to eliminate flux by using a controlled atmosphere (hydrogen-nitrogen-argon), which prevents oxidation. In brazing copper (and copper alloys), use can be made of silver-copper-phosphorus alloys. These alloys do not require flux; the phosphorus acts as a flux.

Controlled-Atmosphere Brazing Controlled-atmosphere induction brazing combines the unique benefits of induction heating with the advantages of brazing in a protected atmosphere. Induction heating provides: • Highly localized heating. Heating of the entire part is not required, as in furnace brazing. • Rapid heating • Precise control for repetitive results • Brazing of sequential joints in a given assembly as a result of localized heating With the addition of a controlled atmosphere, the following benefits are provided: • Avoid oxidation and/or volatilization of hightemperature metals and refractory materials • Joints produced are dependable, corrosionand oxide-free, and strong • Production of complex electronic devices requiring sequential brazing • Useful in critical structural assemblies with inaccessible areas • Leaves smooth fillets • Highly suitable for applications where brazed joints are hard to clean • Eliminates use of flux where residual flux can be particularly corrosive, as in complex vacuum tubes and joints of aircraft components. Also eliminates cost of fluxing and subsequent cleaning from flux

0.01 0

0 200

600

1000

1400

1800 2200

0

Brazing temperature, °F

Fig. 3.17

Graph that can be used to size induction generators

Reference 11 contains numerous applications where, on a production basis, the benefits of induction heating have been combined with the advantage of a controlled atmosphere. The examples cited include many copper alloys as

Chapter 3: Elements of the Brazing Process / 37

well as steel and stainless steel assemblies. The applications include a megawatt magnetron used in airborne radar systems, klystron tube components, and shielded grid triodes in specialized vacuum tubes.

New Induction Brazing Systems Major advances have been made in recent years in generator and temperature-control design. Solid-state generators have replaced motor generators and vacuum tube generators operating with intermediate power levels at frequencies from 50 to 350 kHz. Improvements have been made in generator control design, allowing precise and instantaneous responses to changes from sensors. The proper-sized generator required for a given operation is best determined by laboratory trial. However, a preliminary estimate can be made by considering the power absorbed by the pieces to be joined, the amount of radiation lost from the heated piece, and the required production rate. The estimate accuracy depends on good coil design and the electrical match between the generator and the workload. The most common frequencies for future use are 10, 25 to 50, and 50 to 350 kHz with solidstate generators, and 250 to 450 kHz with tube equipment. The lower frequencies heat more deeply with more uniform heating of matching surfaces if one of the components has a large mass. Higher operating frequencies, such as 450 kHz, may be suitable for brazing nonferrous metals or for brazing steel to a nonferrous metal or a nonmagnetic (austenitic) steel. Under certain circumstances, higher frequencies provide more efficient heating, and tube equipment operating at 2.5 to 8 MHz may be necessary for very thin components. The availability of a “sandwich” braze using a clad brazing strip (i.e., copper clad with silver filler metal on both sides) can overcome a problem of thermal expansion and the stress effects that have occurred in the past with induction brazing applications. Stress concentration and high residual stresses in the joint become critical factors when the joint members are stronger than the filler metal or when differential contraction takes place with dissimilar materials. The difference in thermal expansion for tungsten carbide may be troublesome when the carbide is placed in tension, resulting in cracked carbide. In this

case, stresses are minimized due to plastic deformation in the low-yield-strength copper layer. The flexibility of induction brazing has been enhanced by a hand-held induction heating system. Because it is not permanently mounted, the portable, lightweight system enables the operator to use this technique on nonuniform structures. Furthermore, the capability of taking the coil to the work, rather than the work to the coil, extends the area of applicability of this heating method. Not only are more varied heating uses possible, but also the new degree of flexibility lends itself to computerized, automated systems, including cellular operations. The unit has a small operating head and control equipment. The head is 102 × 152 × 330 mm (4 × 6 × 13 in.) and weighs only 36 kg (79 lb). The system delivers up to 25 kW into a workpiece. Full power is obtainable up to 129 m (425 ft) from the solid-state power supply using dry, flexible cable. The system operator is completely safe, because arcing to the workpiece does not occur if the coil should come in contact with the work. Controls for the system are mounted in the lightweight housing and can be operated through a computer interface. Some typical data provide a means for statistical process control, recording time, frequency, power, volts, amps, status, and date. The system has been used successfully in both refrigeration and compressor manufacturing. Tube-in-Place Induction Brazing. The development of manned spacecraft, maneuverable deep-space probes, and supersonic and hypersonic aircraft has resulted in the need for highly reliable, leak-free, and lightweight fluidtransmission systems. The obvious answer for this need is a system without joints, constructed from thin-wall, high-strength tubing. The acceptance of permanent joining methods for aerospace fluid systems was slow. The first uses of brazed joints in tubing for aerospace applications were in the B-70 (Ref 12) and A-4 vehicles. The first totally brazed system put into service was the hypergolic fuel and oxidizer system in the Gemini spacecraft (Ref 13). Since 1964, there has been a general acceptance of brazed systems by the aerospace industry. Brazed systems have been employed in a series of spacecraft programs, from Apollo to Space Shuttle to Space Lab to International Space Station (Ref 14–16). The advantages of brazed systems over welding and threaded connectors for all the aforementioned vehicles include the following:

38 / Brazing, Second Edition

• Freedom from leaks • Compatibility with system fluids, operating environments, and production techniques • Repairability and replaceability of system components in the field (or on the launch pad, if necessary) • In-place assembly and joining to provide a basically stress-free system

Fig. 3.18

• Potential for designing small tools, using thinner-wall tubing and less drastic heataffected zones Joint Design. The joint design illustrated in Fig. 3.18 provides for capillary flow of the filler metal in two directions in the joint to produce a maximum surface area of sound bonding. This enhances strength and provides a reliable seal. The filler metal, which, in the majority of applications, has been 82Au-18Ni, melting at 950 °C (1740 °F), flows between the outside wall of the tubing and the inside wall of the fitting being brazed. The maximum temperature attained does not exceed this temperature by more than 110 °C (200 °F), remaining below 1060 °C (1940 °F). Brazing is performed by an automated process in an atmosphere of dry argon, and the completed braze is of high quality with a high degree of reliability. Materials. Table 3.5, which is based on many studies (Ref 13), shows which tubing materials

Joint design for tube-in-place induction brazing. Source: Ref 15, 17

Table 3.5 Recommended filler metals for brazing of fluid systems

347

321

AM355

17-4 PH

PH 15-4

AM350

21-6-9

CP titanium

Ti-6Al-4V

Ti-0.15Pd

CP niobium

6061 aluminum

304L corrosionresistant alloy 304 corrosionresistant alloy 347 corrosionresistant alloy 321 corrosionresistant alloy AM350 corrosionresistant alloy 17-7 PH corrosionresistant alloy PH 15-7 Mo corrosionresistant alloy 21-6-9 corrosionresistant alloy CP titanium Ti-6Al-4V Ti-3Al-2.5V Ti-0.15Pd CP niobium LB2-coated niobium CP molybdenum Silicide-coated Mo Beryllium

304

Tube material

304L

Fitting material

1–6

1,6

1,6

1,6

...

...

...

...

...

...

...

...

...

...

1,6

1,4–10

...

...

...

...

...

...

...

...

...

...

...

...

1,6

1,6

1,6

1,6

...

...

...

...

...

...

...

...

...

...

1,6

1,6

1,6

1,4–10

...

...

...

...

...

...

...

...

...

...

...

6

...

6

1,6

...

...

6

...

...

...

...

...

...

...

...

...

...

...

1,6

...

...

...

...

...

...

...

...

...

...

...

...

...

...

1,6

...

...

...

...

...

...

...

1

...

...

...

...

...

...

...

1

...

...

...

...

...

... ... ... ... ... ... ... ... ...

... ... ... ... ... ... ... ... ...

... ... ... ... ... ... ... ... ...

... ... ... ... 1,6 1,6 6 6 ...

... ... ... ... ... ... ... ... ...

... ... ... ... ... ... ... ... ...

... ... ... ... ... ... ... ... ...

... ... ... ... ... ... ... ... ...

... ... ... ... ... ... ... ... ...

11,12 ... ... ... ... ... ... ... ...

... 11–14 11 ... ... ... ... ... ...

... ... ... 11,12 ... ... ... ... ...

... ... ... ... 15 ... ... ... ...

... ... ... ... ... ... ... ... 16

Note: 1, 82Au-18Ni; 2, 72Au-22Ni-6Cr; 3, 60Au-20Ag-20Cu; 4, 99.5Ag-0.5Li; 5, 92Ag-7.8Cu-0.2Li; 6, 72Ag-27.8Cu-0.2Li; 7, 92.6Ni-7.0Cr-2.9B-0.6C (max); 8, 85.1 Ni-7.0Cr-5.0Si-2.9B; 9, 79Ni-11.5Cr-3.5Si-3.0B-3.5Fe-0.15C (max); 10, 72.2Ni-20Co-4.5Si-3.3B; 11, 95Ag-5Al; 12, 81Pd-14.4 Ag-4.6Si; 13, Ti-15Cu-15Ni; 14, Ti49Zr-4Be; 15, Beta titanium; 16, Al-12Si

Chapter 3: Elements of the Brazing Process / 39

are the most compatible with the various corrosive hypergolic propellants. The materials 304L and AM350 have been used for most applications (Ref 12, 14, 15). The majority of the various types of fittings, including unions, elbows, tee fittings, reducer fittings, crosses, and valves, have been made of 304L corrosion-resistant steel. Additional studies (Ref 13, 16), also represented in Table 3.5, evaluated filler metals and their compatibility with fuels and base metals. Tooling. In general, the tube-in-place system is a total tube-brazing system. It is restricted to induction heating methods, because these methods permit fluxless brazing and the capability of completing joints in areas of very limited access. There are two distinct types of tooling employed for fluxless brazing of tubular components: • Open-coil tooling, in which the inductor is an unshielded multiturn solenoid coil and the

Fig. 3.19

Open-coil setup and typical joints

atmosphere-control chamber is constructed separately from the coil (Fig. 3.19) • In-place tooling, in which the inductor comprises either a split solenoid coil or opposing pancake coils, which are integral with the atmosphere-control chamber Tooling for in-place brazing combines the elements of the open-coil method into an integral tool. This type of tooling is available in either a clamshell-type (Fig. 3.20) or a plierstype (Fig. 3.21) configuration. The coil itself is the most critical part of the tooling. Split-solenoid coils require power transfer through pin connections that present arcing problems and limit the tool to a method of cooling external to the coil. Shaped pancake coil inductors provide the capability for using the water-cooling system of the induction generator to cool the induction coils and permit the use of pliers-type tools. In general, pliers-type tools provide the smallest

40 / Brazing, Second Edition

envelopes and the least cumbersome cooling, power, and gas connections. Brazing Procedure. Brazing is accomplished in a very straightforward manner. The tubing

Fig. 3.20

Clamshell-type tooling for in-place brazing

Fig. 3.21

Pliers-type tooling and typical joints

details and fittings are positioned on the structure and clamped where necessary. Once the brazing sequence is established, the internal tubing details and fittings are positioned on the structure and clamped where necessary. Once the brazing sequence is established, the internal argon-gas protection is provided, and the fitting is carefully positioned. The pliers-type tool is then positioned around the joint, and the brazing cycle is initiated. The automatically controlled brazing cycle consists of an argon-preflow period of at least 10 s, a heating period of from 10 to 30 s, and a postheating argon-flow period of from 60 to 90 s. The brazed joint is visually inspected for proper filler-metal flow and external filleting. If it is acceptable, the next joint is brazed. If the joint is unacceptable, the operator can try to promote the required filler-metal flow by reheating the joint with an identical brazing cycle. Only two reheats are permitted, because

Chapter 3: Elements of the Brazing Process / 41

there is a 10% tolerance in the brazing cycle settings, and because, if filler-metal flow is not promoted by proper use of the tooling, the chances of promoting flow are negligible. Excessive reheating when reactive filler metals are used also increases the chance for perforation of the tube wall through base-metal erosion. Nondestructive Testing. Each brazed joint is inspected visually to determine whether it is within dimensional tolerances and whether the filler metal is visible in the fillet around the entire circumference of the tube at the fitting end. Any melting of the base metal is cause for rejection. One-hundred percent radiographic inspection is performed, and all joints are required to meet rigid specification acceptance standards. This requirement is imposed on all procedure-qualification test specimens and also on all production brazes made for installation, repair, or modification. A new-generation induction brazing machine that features automatic temperature control has been developed. This brazing tool head contains a fiber-optics sensing device that functions in combination with a photoelectric cell, which feeds a signal through an amplifier to a controller recorder for automatic temperature control of the induction brazing cycle. Such developments, as a result of the use of solid-state electronics in the electrical circuitry, will continue to improve, and eventually lead to automation of, these brazing techniques.

Resistance Brazing Resistance brazing is most applicable to relatively simple joints in metals that have high electrical conductivity. For resistance brazing, the workpiece, with filler metal preplaced, is part of an electric circuit. Brazing heat comes from either placing carbon electrodes in contact with the brazement to conduct heat into the work or by relying on resistance of the brazement to generate heat. Electrode resistance brazing is best suited to joining of high-conductivity materials, such as copper-base alloys. Applied to steels, which are poor conductors, electrodes can cause local overheating. Using this process, often only one part of an assembly is heated by the electrode, the other by conduction. This procedure is most effective when brazing dissimilar materials. During direct resistance heating of the work, current flows for such a short time and heating

is so localized that the temperature of the workpieces does not change except in the local joint area. Advantages of resistance brazing are that it localizes heat and is flameless, non-contaminating, fast, and easily controlled. Disadvantages are that at least one workpiece, preferably all, must conduct electricity; large work, anything weighing over 2.3 kg (5 lb) or cross sections larger than 33 cm2 (5 in.2) may require so much current as to make the process too slow to be practical; and the process does not easily heat joints of nonuniform cross section. In the usual application of resistance brazing, the heating current, which is normally alternating current, is passed through the joint itself. Equipment is the same as that used for resistance welding, and the pressure needed for establishing electrical contact across the joint is ordinarily applied through the electrodes. The electrode pressure also is the usual means for providing the tight fit needed for capillary behavior in the joint. The component parts are generally held between copper or carbon-graphite electrodes. The heat for resistance brazing can be generated mainly in the workpieces themselves, in the electrodes, or in both, depending on the electrical resistivity and dimensions. The flux used must be conductive. Normally, fluxes are insulators when cool and dry, but they may become conductive from the heat of brazing. The process is generally used for low-volume production in joining electrical contacts and related electrical elements. Resistance brazing requires a transformer, electrodes, and fixtures. Standard resistancewelding transformers suffice for brazing, with the required capacity depending on joint size, heating capacity of the base material, and desired brazing time. In high-volume production applications, spot-welding equipment may suffice to resistance braze, using a transformer, upper and lower electrodes, and process controls. For manual brazing, operators wield handheld tongs that clamp the workpiece and hold the electrodes. Transformer power is typically rated at 10 to 25 kVA. Electrodes are of standard resistance-welding-grade copper alloys or of carbon. Most common are Resistance Welding Manufacturers Association (RWMA) class 2, chromium copper, and class 14, refractory (molybdenum) metal electrodes. With higher electrical resistivity, they generate more heat than do RWMA electrodes and

42 / Brazing, Second Edition

perform well for brazing highly conductive materials. Electrodes of high electrical conductivity braze low-conductivity materials, such as steels. Filler metals used for resistance brazing need to melt at relatively low temperatures to avoid oxidation. Most common are BAg-1, 1A, 2, 7, 8, and 18 and BCuP-1, 2, and 5. Preplace a strip or shim on large flat joints, a ring on cylindrical joints, paste or powder on irregular surfaces, and a rod for feeding additional filler to supplement preplaced filler in large joint gaps (Table 3.6). In resistance brazing, more so than in other brazing processes, attention is given to selecting from the metallurgically compatible filler metals the one having the lowest brazing temperature, because in resistance brazing, it is necessary to keep the maximum local temperature reached by the work as low as possible while providing uniform heating of the abutting joint surfaces and the filler metal. Fluidity of the filler metal is not critical in most resistance brazing, because the filler metal is usually preplaced, and the bond area is relatively large.

Carbon Resistance Brazing This method is generally used for joining of heavier sections. When this technique is applied to metals of low conductivity, such as steels, local overheating is likely to occur at the points of contact. The extent of overheating can be minimized by shaping the electrodes to the contour of the work, by heating at a lower rate, or by using only lower-melting-point filler metals. Carbon resistance heating of copper-base alloys is much easier. The choice of filler metals is no longer restricted to silver-base filler metals, and filler metals that contain phosphorus and no flux can also be employed. When a flux is used, care must be taken to prevent the carbon from

Table 3.6 General types of filler metals usually selected for resistance brazing of various classes of work metals Work metal

Steels, stainless steels, heat-resisting alloys, copper, copper alloys, nickel alloys Aluminum alloys Copper and copper alloys

Brazing filler metal

Silver filler metals (BAg type) Al-Si filler metals Cu-P filler metals

Note: These types of filler metals all have relatively low brazing temperatures.

becoming impregnated and to avoid the formation of an insulating film between the electrodes and the work. It is not always necessary for the current to pass through both the component parts. Frequently, only one part is resistance heated, and the second part is heated by conduction. This approach is particularly effective when either dissimilar metals or components of different masses are being joined. The equipment used for carbon resistance brazing may take a variety of forms, depending on the application involved. Hand-operated portable tongs are favored for many applications, particularly for joining heavy assemblies or when a number of joints, at different angles and locations, are required on one main unit. Heavy electrical-control-and-generating equipment provides many such applications. The fluxless method is used extensively for making connections of terminals or assemblies to stranded or braided copper electrical conductors. Carbon electrodes used in carbon resistance brazing are of two general types: carbon graphite and electrographite (artificial graphite). These electrode materials are made by simultaneously heating and blending the finely divided raw materials with coal tar pitch, which serves as a binder. Carbon or graphite blocks or rods are most frequently used, because filler metals do not wet them as they do metals. Various grades of these materials are available with ranges of resistance and hardness that permit them to be adapted to various jobs. The carbon or graphite is easily worked to the desired shape. These parts, however, must be considered expendable, because they waste away quite rapidly. Carbon resistance brazing is a far more attractive process when applied to copper than when applied to steel.

Direct Resistance Brazing Direct resistance heating is based on the same principles as resistance spot and projection welding, and the same machines can be used for both processes. The current must always pass through both of the parts that are being joined, and the highest temperature occurs at the interface between the parts. In most instances, the current flow is of such a short duration, and

Chapter 3: Elements of the Brazing Process / 43

heating is so localized, that no detectable increase in temperature takes place either in the electrodes or in the mass of the component. It is not unusual for the operator to hold the parts while the joint is being formed. When two metals of different resistivities are being joined, one part heats more rapidly than the other, unless precautions are taken. A simple cure is to introduce the current in short bursts and to allow time between the pulses for the temperature to balance out. Alternatively, using electrodes of different compositions can counteract unbalanced heating. By having the low-resistance electrode against the high-resistance member, and vice versa, the heating potential on both sides of the component interface is balanced. Direct resistance heating differs from carbon resistance heating, and from all other methods of brazing, in certain important respects. As already mentioned, severe temperature gradients are promoted within the workpiece, and only small localized areas at the joint interface attain the required brazing temperature. The area of joining is, therefore, small and has little mechanical strength. Attempting to increase the area is unlikely to be successful, even if sufficient power is available. Varying current densities over the larger area produces inconsistent results and severe local overheating. Also, because the highest temperatures occur at the interface, the filler metal must be preplaced between the mating faces of the component parts. Filler metals in the form of shims usually are the logical first choice for use with direct resistance heating. Electrodes for direct resistance brazing are made of high-resistance electrical conductors, such as chromium copper, tungsten, or molybdenum rods or inserts, or even steel in some instances. Direct resistance brazing is extensively used in the manufacture of electrical and electronic devices, many of which are of an extremely delicate construction by comparison with the heavier electrical equipment for which carbon resistance heating is preferred. The termination of fine wire conductors is a typical example of the use of direct heating. Applications. The use of flux-free (inertgas) brazing has been introduced in new uses of direct eutectic resistance brazing. Examples include brazing steel ring-piece pipe joints using resistance or induction heating, brazing chromium-nickel alloy steel fuel lines using resistance heating, and brazing steel servosteering parts in automated production pro-

cesses. Advantages of flux-free brazing over traditional flame and protective-gas furnace processes include: • Lower energy costs • Brazing without using flux • Easy integration of heating equipment into automated production processes • Meeting tight tolerances in the brazing of bent pipes Figure 3.22 shows ring pieces made of Deutsche Industrie-Normen (DIN) 7642 are used as connection elements for pipes/tubes (e.g., oil and fuel lines) in automobiles and mechanical equipment. The ring piece typically is joined using high-temperature brazing in a protective-gas, continuous-type furnace. This requires fixing the joint position via spot welding and fluxing before brazing. Disadvantages of the process include total or partial loss of brazing paste and the inability to be sure whether enough brazing paste is available to completely fill the brazing gap. Thus, when brazed parts must meet higher safety standards, expensive filler-metal rings often are substituted for brazing paste. Protective-gas, local-heating methods are more effective when brazing fewer pieces and

Fig. 3.22

Brazing ring-piece pipe joints using direct electric-resistance heating

44 / Brazing, Second Edition

thin-walled parts. Thus, brazing of ring-piece pipe joints using resistance heating is more economical. The brazing device (Fig. 3.23) essentially consists of a protective-gas chamber, two part electrodes (one stationary and one mobile), and one in the height-adjustable pipe electrode. A height-adjustable device between the part electrodes picks up the ring piece. The ring piece is placed in the fixture together with a pipe previously fitted with an alloy ring. To start the automatic brazing process, a part electrode is driven against the ring piece, which closes the secondary circuit. At the proper temperature, the filler metal liquefies and flows through capillary action into the braze gap. One braze joint is produced every 41 s, and 28 s are required to bend the pipe and check the seal. After a 10 s hold under the protective gas, the chamber is opened, and the brazed piece is removed. A ring piece joined to a steel pipe using a copper-tin alloy ring and resistance heating is shown in Fig. 3.24. Figure 3.22 shows resistance brazing apparatus used to join ring-piece pipe joints for an 8 to 12 mm (0.3 to 0.5 in.) diameter pipe. Servosteering parts traditionally have been brazed in protective-gas, continuous-type furnaces with a low failure rate of brazed parts.

Fig. 3.23

Direct electric-resistance-heating brazing device

Failures that occur are generally due to a partial loss of filler metal while in the oven. Because the required annual quantity totals over 1 million pieces, the brazing production process must be automated. Mounting the filler metal automatically and monitoring at the brazing temperature can reduce the failure rate. Both local heating processes are suitable for this application. In automatic resistance-heating brazing, the pipes and the ring pieces are isolated, automatically mounted, and brazed under a technical nitrogen gas. The accuracy of applying brazing wire is automatically examined by video inspection. Short and long pieces are alternately brazed with ring pieces after they have been automatically upset and collected opposite the brazing station. An automatic wire-feed device feeds the filler metal. The wire feed is warmed and follows the natural action of the filler metal even before reaching the working temperature of the brazed joint. The alloy is fed out after reaching this temperature. As a result, the heated filler metal melts immediately on the second dispensing. This differs from traditional methods of applying cold wire, where the wire often is pushed past the brazing joint. Following brazing, assembled components are transported by conveyor belt to the testing

Fig. 3.24

Ring piece joined to a steel pipe using a coppertin alloy ring and resistance heating

Chapter 3: Elements of the Brazing Process / 45

station, where a camera system checks the concave fillet and heated pipe area for flaws. Heating of the pipe area and the quality of the concave fillet are automatically tested. Parts detected having defects in the concave fillet exceeding 0.12 mm (0.005 in.) are removed from the process using a special procedure. The brazed assembly is turned four times. Flaws are displayed as black picture points, and the sum of these is used as a threshold value. When the camera registers fewer than 60 black points (0.015 mm2, or 0.0002 in.2) in a region, the program approves the detail picture. If the program counts more than 50 black points, a failure report is given. A part that passes inspection continues to the bending station. Brazing in three brazing stations enables a joining time of 15 s. During this time, the brazed assembly is upset, brazed, tested, and bent. The brazing process in reconditioning shafts of mechanisms and machines is based on producing an additional repair component in the form of a strip; deposition on its surface of a pasty filler metal; joining the strip to the reconditioned surface of the component by step-seam heating with a high-intensity pulse current and a low voltage, with simultaneous application of pressure; and cooling the brazing zone (Ref 18–20). As a consequence of the resultant transition resistance (Fig. 3.25) of the strip/fillermetal/component type, during a current pulse, the main component and filler strip are locally heated, and the filler metal is melted. Under the effect of the applied compressive force, the molten filler metal wets the surfaces being joined and is compacted, and after completing heating, it solidifies into a weld and forms a permanent joint. At the moment of formation of bonds between solid and liquid metals, the atoms of the parent and filler metals in the filler metal (Fig. 3.26) are excited. Consequently, the reaction rate, in a general form, can be expressed by the equation (Ref 19, 20):



dx Q  = (N0 – x) ν exp –  dt kT

cal bonds, k is the Boltzmann constant, and T is absolute temperature. One of the elements of the technological process is the temperature conditions of brazing, which take into account the temperature

Fig. 3.25

Diagram of electric resistances. Rt Rt1, and Rt2 are a transition resistance between the , electrodes on the surface of the strip, strip filler metal, and the filler metal of the main component, respectively. RM1, RM2, and RM3 are the resistance of the metal of the strip, filler metal layer, and the parent metal, respectively. (1) Roller electrodes. (2) Steel strips. (3) Filler metal. (4) Reconditioned shaft. Source: Ref 20



where x is the number of atoms entering into a chemical bond, N0 is the number of contacting atoms on the surface of the parent and additional materials of the component/strips, ν is the frequency of inherent vibrations of the atoms, Q is the energy of activation of formation of chemi-

Fig. 3.26

Diagram of the structure of the brazed joint

46 / Brazing, Second Edition

conditions of operation of the components, the melting point of the main metals joint, and the brazing temperature of the joint. Analyzing these temperatures with requirements made for the strength-reserve factor, the selected brazing temperature range (and consequently, the filler metal) was selected at 850 to 1150 °C (1560 to 2100 °F). The nickel, copper-zinc, silver, and copper-nickel filler metals are used most extensively in this range. All these filler metals satisfy the main requirements on brazing steels; namely, they wet efficiently the brazed surfaces, have sufficient fluidity and fill brazing gaps, ensure high strength of the brazed joint, and have minimum deformation of the distance at working temperatures. The results show that the size of the heataffected zone (HAZ) does not exceed 0.5 mm (0.02 in.), and in depositing the second and subsequent layers, the HAZ spreads only to the thickness of the first layer and does not reach the parent metal. It is thus possible to increase greatly the endurance limit of the brazed joint in comparison with arc surfacing methods. The cooling rate greatly affects the structure and, consequently, the strength of the brazed joint. This is reflected mainly in the formation of the primary structure and the distribution of the components of the filler metal in the brazed joint. Therefore, the proposed technology of reconditioning dynamically loaded worn components by electric resistance brazing a multilayer coating is characterized by the hyperactivity and low energy requirement of the process, by formation of joints with a small HAZ, and by retaining the initial properties of the metal of the component with a high strength of the brazed joint (Ref 20).

Dip Brazing As the name implies, during dip brazing, the assembly is immersed in a heated bath of either molten metal or a flux bath of molten salt. There are two methods that are considered to be dip brazing methods. One employs a molten metal bath, the other, a chemical (flux) bath. The molten metal bath dip brazing technique, which is not widely used, could be described as the high-temperature equivalent of dip soldering. It is used for the manufacturing of electronics and similar very small components.

In both of these processes, the parts being joined are held together and immersed in a bath of molten bonding metal that flows into the joints when the parts reach a temperature approaching that of the bath. In the chemical bath process, molten flux is used rather than molten metal. A different salt (flux) is often required if either the parent metal or the filler metal is changed, which may not be the case with the molten metal bath. However, one filler metal, such as a silver-base alloy, may be satisfactory for joining many brasses and steels. Molten Metal Bath Dip Brazing. The molten metal bath process is simple, cheap, and effective but is limited to brazing of small assemblies, such as wire connections or metal strips. A crucible, usually made of graphite or ceramics, is heated externally to the required temperature to maintain the filler metal in fluid form. A cover of flux may be maintained over the molten filler metal. The size of the molten bath (crucible) and the heating method must be such that immersion of parts to be brazed does not lower the temperature of the bath below that necessary for brazing. Parts to be brazed must be clean and are often protected with flux prior to their introduction into the bath. One potential future use for this process is in providing an alternative to soldering for electrical connections on motors and generators. The demand for increased output and/or reduced weight can now be met only by running the soldering machines hotter, and the softening temperature of the soldered joints is often the limiting factor. Another likely use of molten metal bath dip brazing is for pretinning of small inserts of difficult-to-braze metals that are to be subsequently brazed to heavier parts. Molten Chemical (Flux) Bath Dip Brazing. Dip brazing in molten salt (flux) involves the immersion of an assembly in a bath of molten salt, which provides the heat and may also supply the fluxing action for brazing. The bath temperature is maintained above the liquidus of the filler metal but below the melting range of the base metal. The advantages (Ref 21) of molten chemical bath (salt bath) dip brazing are: • Time for heating is approximately one-fourth of that required in a controlled-atmosphere furnace. • A protruding joint can be selectively brazed by partly immersing the assembly.

Chapter 3: Elements of the Brazing Process / 47

• A cocoon of frozen salt forms instantly around the assembly when it is immersed in the molten salt, which usually prevents premature melting of the filler metal by providing temporary insulation. • Brazing can usually be combined with carburizing or hardening, without the necessity for a separate reheating operation. • More than one assembly or joint in an assembly can be brazed at the same time, because production is limited only by the size and heating capacity of the furnace. • The workpiece is protected from scaling or decarburization by a thin film of salt that adheres to the surface of the assembly when it is removed from the salt bath. • Removal of the salt film is accomplished by dissolving during quenching or washing operations. • Because the density of the molten salt supports a considerable portion of the weight of the workpiece, the assembly weighs less when immersed, which can reduce the likelihood of distortion during heating. Limitations (Ref 21) of the process are: • The process is not generally used for intermittent operation, being better suited for work that requires daily production. • Joints that do not protrude from the assembly cannot be selectively brazed by partial immersion; most or all of the assembly must be heated to the brazing temperature in order for such joints to be brazed.

Fig. 3.27

Externally heated furnaces. Source: Ref 17

• The workpieces must be completely dry, because the molten salt reacts violently with moisture, splattering and possibly even exploding. If moisture is present, all work requires preheating. • The shape of the part must be designed to avoid trapping of air or salt and to drain completely after removal from the salt bath. • The assemblies should not require large, complicated fixtures. • Part cleaning may be difficult. • Salt residues, especially chloride, can corrode assemblies. • Proper maintenance of a salt bath furnace is difficult and has special problems, such as solidification during power outages. • Salt vapors may present health hazards unless properly ventilated. Equipment. Some salt bath furnaces are externally heated by gas, oil, or electrical resistance; this type of furnace lends itself more readily to intermittent operation and is not widely used for high-volume production. On the other hand, furnaces that are internally heated by immersed or submerged electrodes are not well suited to intermittent operation; therefore, they are used for high-volume production. Figures 3.27 and 3.28 illustrate the principal types of furnaces used for salt bath dip brazing. The externally heated furnaces shown in Fig. 3.27 are usually gas fired or oil fired and, less frequently, are heated by means of electricalresistance elements. When electrical-resistance heating is used, pot failure may result in total destruction of the heating elements. Figure 3.28

48 / Brazing, Second Edition

shows two types of internally heated furnaces that are energized with alternating current supplied by a transformer. The molten salt is an electrical conductor, and heat is generated within the salt between the electrodes by resistance to the passage of current. By closely spacing the electrodes, an electromagnetic stirring action of the salt is obtained that assists in maintaining temperature uniformity and controlling the temperature to within ±3 °C (±5 °F) (Ref 22). Theory and Start-Up. Figure 3.29 illustrates the theory of electrodynamic circulation. Maxwell’s well-known law of electromotive force states that any conductor that carries current, when placed in a magnetic field, tends to move at right angles to the direction of the magnetic field and at right angles to the direction of

Fig. 3.28

Internally heated furnaces. Source: Ref 17

the current. Applying this rule to the salt bath, two electrodes (A and B) are immersed vertically and close together in molten salt (S). A current of high magnitude (3000 A, for example) is passed between them. Each electrode is surrounded by approximately circular lines of force (D). High-density current flows between the electrodes at all points below the surface of the bath. At any given point between the immersed electrodes, such as point P, the salt element (represented by the heavy line) acts as a conductor of current between A and B. Because the electromagnetic field between the two electrodes extends outward from the plane of the drawing at point P, the liquid salt, which is acting as a conductor of current, is forced downward between the electrodes, as indicated by the

Chapter 3: Elements of the Brazing Process / 49

black arrows. If the direction of current is reversed, the direction of the magnetic field also reverses, and the direction of movement is still downward (Ref 17, 22). Once started, a molten flux bath is usually idled at a temperature high enough to maintain the salts in the liquid state. Electrodes. The electrodes in general use are made of carbon, wrought nickel, or Inconel 600. Such electrodes are less prone to attack than copper and copper-bearing electrodes and cause far less bath contamination. Carbon electrodes are the least costly. A carbon electrode lasts three to six months. Carbon has half the electrical conductivity of nickel, and therefore, for a given current density, carbon electrodes must have twice the cross section of nickel electrodes for equal capacity. Nickel electrodes are more costly than carbon electrodes. They may or may not be water cooled, depending on current density. Nickel electrodes wear most rapidly at their air-flux interface, and over-the-top electrodes generally need to be replaced in three to six months. They have the advantage of being easily removed. They are simply unbolted and lifted out of place with a forklift or other powerlift equipment. Over-the-top electrodes develop considerably more magnetohydrodynamic (MHD) force and therefore produce considerably more MHD flux circulation than sidewall- and bottom-positioned electrodes. This is due to the greater length of these electrodes and their attendant magnetic field in the flux.

Fig. 3.29

Schematic illustration of electrodynamic circulation. A, B, electrodes; S, molten salt; D, lines of force; P, point. Source: Ref 17

Submerged electrodes last much longer. Their average life is one to two years, with some sets lasting three years. Bottom-placed electrodes are more effective in producing convection currents in the molten flux and are less likely to be accidentally struck by dipped parts. Fluxes. The choice of a flux for dip brazing is not as much a matter of joint quality as of other important, although secondary, considerations. The aluminum dip brazing bath normally consists of a chloride mixture with small amounts of aluminum or sodium fluoride. The fluoride acts as a reducing agent, reacting with any oxides in the bath. This necessitates an exhaust system to remove volatile, corrosive fumes. Lithium lowers the melting point of flux and is an important additive to the salt bath—in the form of lithium chloride (LiCl)—for brazing of aluminum alloys. The fluxes containing LiCl are stable and require little chemical adjustment during operation. Salt bath dip brazing with silver-base filler metals is performed in a bath of neutral salts (Ref 17). Usually, a small amount of brazing flux is applied to the joint with the preplaced filler metal. After the moisture has been dried from the salt in a preheating cycle, the work is immersed into the molten salt for heating and brazing. On the other hand, aluminum is brazed in a molten salt bath consisting of active salts that serve as brazing flux and make it unnecessary to apply additional fluxes to the work before immersion. Process Operations. The salt bath process (Ref 17) is essentially the same regardless of whether aluminum, magnesium, titanium, steel, or nickel is being brazed. The process consists of six distinct steps: deburring and filing of machined or ground edges, cleaning, assembly, preheating, immersion, and flux removal. Deburring and Filing of Machined or Ground Edges. Because the brazed joint is made by virtue of filler metal flowing into the joint as a result of capillary action or gravity, burrs are removed, because they could restrict fillermetal flow and cause an imperfect union. Cleaning. The joint surface must be free of grease, oil, paint, oxide, and scale that would prevent the filler metal from wetting the workpiece surfaces. Assembly. During the assembly of parts, the filler metal, depending on its form, is applied to

50 / Brazing, Second Edition

the joints. The details can be held in position by tack welding, spring-loaded fixtures, spring clamps, staking, self-locking joints, or spot welding. (For more information on tooling, see Chapter 7, “Fixturing, Tooling, Stopoffs, Parting Agents, Surface Preparation, Surface Cleaning, and Repair.”) Preheating of an assembly before brazing serves several purposes. If prefluxing is used, preheating dries the flux and vaporizes all moisture from the assembly and the fixture. (Even a slight amount of moisture can cause spattering in contact with molten salt.) Preheating decreases the temperature drop of the salt bath, thus reducing brazing time, and also minimizes the premature melting of externally placed filler metal. For assemblies consisting of both heavy and light sections, preheating reduces thermal gradients and subsequent distortion and improves the wetting action on the heavier parts as well. To be effective, the preheating temperature must be at least 55 °C (100 °F) lower than the melting temperature of the filler metal. If oven preheating is used, oxidation must be avoided by using temperatures below 480 °C (900 °F); otherwise, an inert atmosphere is desirable. Finally, the preheating operation drives off all moisture from the assembly and fixture, thereby removing any possibility of any explosion, which might occur if water were introduced into the molten flux. Immersion. General ranges of brazing temperatures are used with various salts, and specific brazing ranges are employed for the individual filler metals. The time in the molten salt bath differs from one job to another. When only thin-section parts are to be brazed, the holding time may be as short as 1 min (Ref 17, 22). If assemblies are permitted to remain in the flux longer than 3 min, strength is substantially lowered, due to gradual transformation of the parent material from the as-wrought to the as-cast condition of the joint at the brazed interface (Ref 17). After the workpieces have been in the bath for the required time, they are carefully lifted from the salt bath. A uniform motion is necessary during removal from the bath; jerky movements can cause the liquid filler metal to be displaced from the joint. Flux Removal. A certain amount of flux adheres to the assembly after brazing, and this must be drained off while the parts are hot. The assembly is allowed to cool and then is

immersed in hot running water to remove the frozen salt cocoon. The assembly is normally removed from the fixture and subsequently dipped in a series of acid-to-water rinses to complete the flux-removal process.

Infrared (Quartz) Brazing Infrared brazing may be considered to be a form of furnace brazing in which heat is supplied by light radiation below the visible red rays in the spectrum. The principal heating is done by the invisible radiation from high-intensity quartz incandescent lamps. Heat sources (lamps) capable of delivering up to 5000 W of radiant energy are commercially available. The lamps do not necessarily need to follow the contour of the part in order to heat it, but the heat input varies inversely as the square of the distance from the source. Suitable reflectors, and sometimes parabolic focusing features, can present a unique source of energy for brazing. Because this process resembles furnace brazing, the brazing techniques are similar. The major difference is the source of heat for brazing. The quartz lamps supply brazing heat, in conjunction with electronic controls (including amplified feedback), to provide continuous, precise control of temperature and time to fractions of degrees and seconds. Infrared brazing setups are generally not as fast as induction brazing, but the equipment is less expensive. Honeycomb panels have been successfully brazed using opposing banks of lamps, and spot brazing of smaller parts has been accomplished using parabolic mirrors. An example of joining of Ti-6Al-4V by a rapid infrared processing technique has been investigated at temperatures between 1010 and 1250 °C (1850 and 2280 °F) and times of up to 120 s. With this infrared technique, joining is typically completed in seconds in an ambient atmospheric pressure of argon. Researchers (Ref 23) studied the effect of joining temperature and time on the microstructure and strength of the joints. The filler metal used in this study was a Ti-15Cu-15Ni wt% alloy. Results showed that the joint shear strength reached 554 MPa (80 ksi) when processed at approximately 1100 °C (2010 °F) for 60 s. Microstructural examinations of the joint with both an optical microscope and a scanning electron microscope indi-

Chapter 3: Elements of the Brazing Process / 51

2010

2190

500

75

400

60

300

45

200 900

(a)

1000

1100

1200

1300

Joint shear strength, ksi

Joint shear strength, MPa

1830 600

Joining temperature, °C Joining temperature, °F 2010

2190

500

75

400

60

300

45

200 900

(b)

1000

1100

1200

1300

Joint shear strength, ksi

1830 600

Joining temperature, °C Joining temperature, °F 1830

2010

2190

600

500

75

400

60

300

45

200 900

(c)

Fig. 3.30

1000

1100

1200

1300

Joint shear strength, ksi

The use of exothermic reactions for industrial brazing applications has been limited, because little information has been made available concerning exothermic brazing materials and the economics of their use. Exothermic brazing usually requires preplaced filler metal, as compared with other processing methods currently in practice that use controllable exothermic reactions. Exothermic brazing is a process in which the heat required for melting and flow of a commercial filler metal is generally produced by a solidstate exothermic chemical reaction. An exothermic chemical reaction is defined as any reaction between two or more reactants in which heat is given off due to the free energy of the system. Nature has provided us with countless numbers of these reactions, but only the solid-state or nearly solid-state metal/metal oxide reactions are suitable in exothermic brazing units (Ref 17). Exothermic brazing employs simplified tooling and equipment. The process uses the reaction heat in bringing an adjoining or a nearby metal interface to a temperature at which preplaced filler metal melts and wets the metal interface surfaces. The filler metal can be any commercially available material having suitable flow temperatures. The only limitations may be the thickness of the metal that must be heated through and the effects of this heat, or of any previous heat treatment, on the metal properties. The first step is to select a heat source and a filler metal. Then, the parts are cleaned chemically

Joining temperature, °F

Joint shear strength, MPa

Exothermic Brazing

(and mechanically, if necessary), the exo-reactant mixture suspended in alcohol is painted on the outside of the parts to be brazed, and the completed assembly with preplaced filler metal

Joint shear strength, MPa

cate that perfect wetting exists between the filler metal and Ti-6Al-4V for most joints. The brazeaffected zone increased with increasing joining temperature and time but did not show a direct influence on the joint strength (Fig. 3.30). Meanwhile, the Ti-6Al-4V base alloy exhibited no noticeable microstructural changes due to the rapid processing cycle of the infrared heating process. Based on this study, it is concluded that, compared with traditional joining methods, the rapid infrared joining technique has the following advantages: fast heating, little energy consumption, easy operation, no need for vacuum, little metallurgical modification to the base metal, and low cost.

Joining temperature, °C

Joint shear strength of infrared-bonded Ti-6Al-4V specimens as a function of temperature for: (a) 30 s, (b) 60 s, and (c) 120 s

52 / Brazing, Second Edition

is wrapped in a Fiberfrax (Unifrax Corporation) blanket and heat-lamp dried. The assembly is inserted into a furnace containing an inert atmosphere, and the furnace temperature is increased to the point at which ignition of the heat source can occur. For example, if the boron-vanadium pentoxide system were selected, the ignition temperature would be approximately 480 to 510 °C (900 to 950 °F). Finally, the external reaction products are washed away with hot water, and the assembly is complete and ready for use. Exothermic joining has proved applicable in numerous potential areas. It has been used to braze honeycomb sandwich panels, to braze inserts into panels locally, to attach small parts to larger structures, to make tube connections for aircraft hydraulic lines, and is considered a field-repair mechanism for aircraft structural honeycomb panels. Exothermic joining appears to be capable of fulfilling difficult requirements in producing reliable and reproducible joints, and its techniques are extremely economical and simple when compared with those of heat-source systems that depend on solar-energy conversion. These systems, except when the rays of the sun are focused directly on the work area, require the conversion of solar energy to electricity and the associated storage and conversion of the electrical energy to heat. The effectiveness of

Fig. 3.31

exothermic joining in vacuum has also been demonstrated (Ref 17).

Laser Brazing Laser brazing has been used in a limited number of applications. For example, a laser has been used where a very small, localized area of heat was required, such as for brazing of small carbide tips on printer heads for electronic printers. A finite-element model (FEM) was developed for the thermal analysis of a stud-to-plate laser brazing joint (Fig. 3.31, 3.32), and the transient temperature fields were analyzed by using a three-dimensional model (Ref 24, 25). Temperature-dependent thermal properties, effect of latent heat, and the convection and radiative heat losses were considered. The brazing parts used were American Iron and Steel Institute (AISI) 304 stainless steel stud and aluminum Al 5052 plate, and the filler metal 88Al12Si was used (Fig. 3.32). The brazing was done in open air with the aid of a powder flux, and neither a prebrazing nor a postbrazing heat cycle was attempted to control the thermal history of the joint. The studs to be joined function as guideposts or shafts for rollers, and the major quality factors are the perpendicularity of the

Experimental setup of stud-to-plate laser brazing system. VCR, videocamera recorder; IR, infrared; CNC, computer numerical control

Chapter 3: Elements of the Brazing Process / 53

stud with the plate surface, and the joint strength. The CO2 laser beam strikes the workpieces to heat locally in an area more than twice the diameter of the stud. The thermal diffusivity of aluminum is over 10 times higher than that of stainless steel, thus requiring more energy input to the aluminum plate. The beam impingement area was approximately 7.5 mm (0.3 in.) in diameter centered on the stud. This enabled both sides to be heated simultaneously for the filler metal to spread and penetrate into the braze gap. The braze flux melts within a few tenths of a second and covers the beam impingement area with a liquid film, and the filler metal being covered with the flux melts at the preplaced location. The filler metal, once melted at the periphery of the stud, instantly forms a small drop and moves toward the center of the joint geometry, that is, the highest-temperature region. A pseudo-TM01 mode (the so-called doughnut mode) of the continuous-wave CO2 laser beam was used as a heat source, for which TM00 mode generated by beam oscillator was opti-

cally modulated using an axicon lens. Relocation of the filler metal during the brazing process, including its wetting and spreading, was examined by using a high-speed motion analyzer, and the results were incorporated in the FEM for defining the solution domain and boundary conditions. The numerical results were obtained for typical process parameters and were compared with experimental ones determined by using the infrared and thermocouple measurements. The joining of a stainless steel stud to an aluminum plate by laser brazing is a new technology that is still under development. For a laser brazing performed in a few seconds by localized surface heating, control of thermal conditions in the workpiece is critical, even more so for a dissimilar-metal braze joint. In future work, the FEM used here will be refined to more accurately simulate the filler-metal behavior and further developed to fulfill a coupled solution of the heat conduction and filler-metal flow problem.

Braze Welding

Fig. 3.32

Schematic diagram of stud-to-plate laser braze (top) and a cross section of workpieces (bottom). O.D., outside diameter; SS, stainless steel

Braze welding offers considerably greater possibilities for joining composite materials, which are used on an increasing scale in various areas of technology (from sports equipment to space systems). The use of composite materials enables strength, stiffness, and impact loading resistance to be increased and the weight of structures to be reduced. Metal-matrix composites are used extensively, and these materials have a number of advantages in comparison with other materials: higher working temperature, shear and transverse strength under off-center loading and compression, higher impact toughness, and so on. Of these, composite materials with an aluminum matrix, reinforced with fibers of steel, boron, and carbon, as well as materials hardened with SiC and Al2O3 particles are used extensively. The method combines heating with a passing current with compression of the bond zone. Braze welding is carried out in projection welding machines using aluminum filler metals. The best results have been obtained using filler metals of the Zn-Al-Cu system. Filler metals based on tin have shown satisfactory spreading over the surface of the parent metal, but the strength was too low. Filler metals based on aluminum have high strength, but their melting point is

54 / Brazing, Second Edition

close to the melting point of the matrix alloy. At these brazing temperatures, debonding of the composite has been detected in the tinning stage, and brittle phases also grew. When using a filler metal of the Zn-Al-Cu system, the required strength of the joint was obtained by selecting the size of the lap. The results show that for a composite 1.2 mm (0.05 in.) thick, the thickness of the lap was 12 to 15 thicknesses of the material. However, compression during brazing reduced this value. Technology was developed for braze welding sheet composite materials (Ref 26, 27) and pipes and rods made of fiber-reinforced composites. Specialized equipment was developed for braze welding pipes and rods that enabled the process to be carried out with volume compression of the bonding zone (Ref 28). Braze welding technology has also been used to join composites to aluminum alloys, such as pipes of composites with endpieces of AMg6 (aluminum-magnesium) alloy. This technology was applied in the fabrication of frame structures, for example, a bicycle frame made of aluminum-boron composite. The frame weighs 1.2 kg (2.6 lb) and the fork 0.4 kg (0.9 lb). In the frame, the endpieces and pipes made of the composite were joined by braze welding. The technology of braze welding dispersionhardened composites does not differ from that used for braze welding fiber-reinforced composites. However, with an increase of the content of the hardening phase, the electrical resistance of the composite increases, thus reducing the time required to reach the brazing temperature. At braze welding temperatures, the gas saturation of the composite has no effect on pore formation (Ref 28). Automotive manufacturers have recently planned to introduce a new model and to use the gas metal arc braze welding process to join the roof panel to the front and rear quarter panels of the auto body. At the prototype stage, it was soon realized that this welding process was not meeting quality requirements. Some of the problems encountered were excessive weld reinforcement, panel distortion from heat buildup, ejection of braze metal in the form of spatter, and pinhole porosity in the braze-metal surface. The plasma arc braze welding process was suggested, although very little background was available on the process. A specialized torchand-wire feed system has been developed that would have the ruggedness needed for the assembly line.

The process has proven to have many advantages. Wire feed speed can be set separately of the current flow, thereby allowing the panels to be joined with current levels less than 50 A, which eliminates distortion. The process is capable of depositing 1.5 kg/h (3.3 lb/h) of braze metal, and brazing times are short: 10 and 20 s, respectively, for each joint. Because the plasma arc does not need to heat an electrode wire to melting, its force is not as violent as a gas metal arc. Therefore, spatter is not a problem, and porosity has been reduced to very low levels. Overall, the plasma arc braze welding process has proven capable of producing the quality required on this application (Ref 29). Structure tests (local buckling) on different stiffened structures (panel) made by braze welding were carried out at the temperature level from 360 to 750 °C (680 to 1380 °F). The materials of the panels were the titanium alloy Ti 1100, the intermetallic alloy Ti-48Al-2Cr, and the powder-metallurgy-produced DS-aluminum 620. The aim of the project was to consolidate the knowledge about the theory of the superposition of thermal loading and pressure loading used on structures such as hypersonic planes (Fig. 3.33) (Ref 30).

Diffusion Brazing Diffusion brazing is a hybrid joining process that combines the features of liquid phase joining and diffusion bonding and has the beneficial features of both techniques (Ref 31). Diffusion brazing and its lower-temperature analog, diffusion soldering, use a molten filler metal to initially fill the joint clearance, but during the heating stage, the filler diffuses into the material of the components to form solid phases, raising the remelt temperature of the joint. The steps involved in making a diffusion-brazed (or diffusion-soldered) joint are shown in Fig. 3.34. This process provides the ready means to fill joints that are not perfectly smooth or flat (a feature of liquid phase joining) while offering greater flexibility with regard to service temperature. The process also provides the following consequential advantages: • Facilitating the achievement of exceptionally good joint filling in large area joints • Allowing edge spillage from the joints to be tightly controlled and kept to a minimum

Chapter 3: Elements of the Brazing Process / 55

• Attaining high thermal conductivity with copper, silver, and gold systems, because the joint produced is composed of primary metal An alloy system suitable for diffusion soldering or brazing should have the following characteristics: • Preferably be a binary alloy, to keep the joint design and joining process as simple as possible • Have a phase constitution that includes a relatively low-melting-point eutectic reaction to initiate the melting process • Have as few brittle intermetallic compounds as possible, which should all melt at temperatures below or comparable to the joining temperature. This reduces the establishment of

Fig. 3.33

Vacuum weld brazed joint

diffusion barriers that can impede the process and lead to the formation of brittle interlayers. • The terminal primary metal phase should possess a wide range of solid solubility of the other constituents. This minimizes the risk of intermetallic phases precipitating during cooling of the assembly from the processing temperature and provides a greater process tolerance to the amount of filler metal introduced into the joint. Examples of alloy systems that satisfy these conditions and lend themselves to viable diffusion-soldering and -brazing processes are silver-tin, gold-tin, and nickel-boron (Ref 31). A study (Ref 32) of copper-tin diffusion brazing was made to identify the crucial process parameters needed to optimize this joining

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process. Key parameters that were identified in practical trials were the thickness of the tin layer and the loading applied to the joint during the brazing cycle. It was established that the tin layer thickness must be controlled to 2 µm, within a tolerance of –0.5 µm, to obtain strong joints. A compressive load of 4 MPa (0.6 ksi) is adequate, while one of 3 MPa (0.4 ksi) is too low. The precise joining temperature is less critical, provided that it is 680 °C (1260 °F) or higher, sufficient to destabilize the brittle Cu3Sn intermetallic compound. There is a risk of reprecipitating this intermetallic phase on cooling if the heating operation does not adequately disperse the tin into the copper layers, due to the diminishing solubility of tin in copper as the joined assembly is cooled down to room temperature. This fact helps to explain why the thickness of the tin layer is highly critical, in

contrast with the silver-tin diffusion-soldering process, where the solubility of tin in silver is essentially maintained constant as the temperature is reduced. With regard to the application of this diffusion-brazing process to plasma-facing components for nuclear reactors, the initial concern about the relatively long reaction time at elevated temperatures (required for diffusion brazing and the resulting promotion of interfacial copper-beryllium intermetallic phases) was not borne out in practice, as indicated by the high strengths of the joints obtained in berylliumstrengthened copper assemblies (up to –230 MPa, or –35 ksi), both when produced under a pressure of 150 MPa (22 ksi) in a hot isostatic press and 4 MPa (0.6 ksi) in a uniaxial press.

Microwave Brazing

Fig. 3.34

Schematic of steps to make a diffusion-soldered (or diffusion-brazed) joint

The process of microwave heating is fundamentally different from any other form of heating. An understanding of the theoretical and practical aspects of the technology has been achieved (Ref 33). The joining of ceramic materials using microwave radiation has received increasing attention. There are potential economic benefits in terms of reduced energy costs as well as physical benefits due to localized heating. However, there is still a considerable amount of work to be done. Several groups have reported successful attempts to bond ceramics using microwaves (Ref 34–37). All bonds have been produced in the solid state and can be compared with diffusion bonding. The general mechanism of bonding is also the same as diffusion bonding, that is, the generation of heat across an interface, leading to atomic migration. Even when the parent ceramic does not couple well to microwave energy, for example, Al2O3 and aluminosilicate-base ceramics, bonds can still be produced by seeding the interface. Materials such as SiC, which heats very quickly in a microwave field, have been placed at the interface to induce high localized temperatures and aid bonding. In Japan, a research program (Ref 37), similar to that described in Ref 35 and 36, used a 6 GHz source with a maximum available power of 3 kW and described microwave heating and joining of Al2O3 and Si3N4 ceramics. Using ceramic rods 3 mm (0.12 in.) in diameter and 100 mm (4 in.) in length, butt joints were pro-

Chapter 3: Elements of the Brazing Process / 57

duced with and without intermediate layers. The joining operation was conducted in air for Al2O3 and in nitrogen for Si3N4. The samples were heated at a rate of 40 °C/s (70 °F/s) to the joining temperature (typically 1400 to 1850 °C, or 2550 to 3360 °F), held for 3 min with a pressure of 0.6 MPa (0.09 ksi), and then cooled at 15 °C/s (25 °F/s). Joints were evaluated by fourpoint bend testing. It was found that the bend strength increased with the rise in joining temperature. Three grades of Al2O3 ceramic—92, 96, and 99% (the balance being intergranular glass)— were evaluated. An average joint strength of 420 MPa (61 ksi) was obtained for 92% Al2O3. This value is approximately equal to the strength of the parent material. In general, joint formation became easier as the glass content of the material increased. This led the researchers to suggest a model for the microwave joining of Al2O3, assuming the microstructure of Al2O3 ceramics consists of Al2O3 particles surrounded by an intergranular glassy phase. The model relies on the intergranular glass heating preferentially to the Al2O3 particles and melting. Such melting (or softening) occurs over a temperature range of 1400 to 1800 °C (2550 to 3270 °F). The Al2O3 is not melted at these temperatures; consequently, the bond is formed via the glassy phase, and, on cooling, the bond line is not observed. The model is supported by observations that 99% Al2O3, which normally does not produce a bond to a similar material, does so if an intermediate layer containing a relatively high fraction of glass is used. The general conclusions that may be drawn from the work worldwide are that ceramics may be joined using microwaves. In particular, the presence of a glass phase, or interlayer, appears important. This is not detrimental, bearing in mind that most engineering ceramics do contain secondary glassy phases. The results also indicate that the use of microwaves allows ceramics to be joined in very short times. Typical bonding times are 5 to 15 min, compared to conventional bonding durations of 30 min and greater. If the only effect of microwave radiation was to heat the material, then it would be expected that joining times would be independent of the method used. In more recent work (Ref 38), a microwaveheating technique was developed for making a braze joint between a tungsten carbide (WC) support and a surface layer of polycrystalline

diamond or, alternatively, between a WC support and a relatively thin WC backing layer with polycrystalline diamond on its working surface. The technique would be used to fabricate diamond-covered cutting tool bits. Such bits could be used, for example, to drill geothermal wells and would be improved versions of some of the diamond-covered bits now used to drill oil and gas wells. Whereas the braze joints of the oiland gas-well versions become weakened at temperatures of 700 °C (1290 °F), the braze joints of the improved drill bits would be designed to withstand hard-rock-drilling temperatures up to 900 °C (1650 °F). The major problem in fabricating the improved drill bits is to use higher-melting-temperature brazing materials and to heat the braze joints accordingly to effect brazing, without overheating the diamond. Overheating in this context means heating to a temperature of approximately 1200 °C (2190 °F), causing the diamond to become graphitized and thereby to lose resistance to wear. The basic idea of this technique is to use the selective heating characteristics of microwaves to develop the required brazing temperature without overheating the diamond. Selective heating would be possible, because the commercially fabricated diamond is a very good absorber of microwaves, while the proposed brazing materials would be moderateto-good absorbers. Experience teaches that the best capillary action and shear strengths in braze joints on diamond/WC tool bits are achieved with fillets of 0.08 to 0.8 mm (0.003 to 0.03 in.), and that braze interlayers should be thick enough (at least 0.02 mm, or 0.0008 in.) to relieve stresses caused by differential thermal expansion between diamond and WC. The brazing material must be able, at the brazing temperature, to wet or diffuse into both the diamond surface layer and the WC substrate or into the WC backing layer and WC substrate, as the case may be. In preparation for a typical fabrication process according to this technique (Fig. 3.35), a diamond disk 2 to 3 mm (0.08 to 0.12 in.) thick is placed on top of a braze interlayer 0.08 to 0.8 mm thick on top of a WC substrate. This assembly of components is mounted in a region of strong electric field in a microwave-processing chamber. A pyrometer is focused on the diamond surface layer; during the subsequent microwave heating, the output of the pyrometer is used to monitor the temperature of the diamond and is used as a feedback signal to control

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the microwave power to achieve the desired brazing temperature. The dimensions of the braze interlayer, the components to be brazed, and the process tooling are chosen, along with the temperature-versus-time heating curve, to obtain the strongest possible braze joint with minimal residual stress from differential thermal expansion. The braze interlayer could consist of a foil of a filler metal. Alternatively, the braze interlayer could be made of a combustion-synthesis compound, in which case microwave heating would be used to ignite a combustion wavefront with temperatures of thousands of degrees. When this wavefront reached the interfaces with the adjacent backing and substrate layers, it would provide sufficient local heating to form the desired braze joint. In conclusion, the feasibility of joining ceramic materials using microwave heating is proven, using interfacial agents or seeding interlayers, such as glasses, and advantages are expected in cost-effectiveness and improved microstructure and properties.

The Future of Braze Processing Automation of the brazing process is the key to maximizing production and quality and minimizing costs. Typical machines that are currently available are capable of automatic brazing by torch, furnace, induction, infrared, and resistance processes. Rotary and shuttle machines automatically apply paste and filler metal, heat the parts, torch braze the assembly,

Fig. 3.35

cool the brazement, and remove some or most of the flux. For small parts, automatic unloading is common. Many engineers are reluctant to try to automate brazing operations for fear of upsetting a well-established process and causing quality problems because of a method change. Their attitude, in short, is, “If it’s working and producing parts, don’t change it; don’t upset it.” However, brazing processes, when properly understood, are just as adaptable to increased productivity through automation as any other process; on the other hand, of course, they are also subject to similar limitations. Regardless of how attractive the idea of automating a brazing operation may be, the first step usually is to check out the economics. If automation is economically justified, the next step is to investigate the mechanics of how to automate the operation. In summary, the following steps are necessary for successful automation: • Make an economic analysis to justify automation. • Break down the total operation into its several parts and determine which can be eliminated, which can be combined, and which can be carried on in conjunction with other operations. • Determine the required type of automated equipment (rotary indexing machine, constant rotary machine, in-line conveyor, inline indexing conveyor, shuttle machine, or racetrack-type conveyor). The types of parts to be brazed and the required production rate

Microwave heating with temperature feedback control makes it possible to braze polycrystalline diamond to an underlying layer of tungsten carbide without overheating the diamond.

Chapter 3: Elements of the Brazing Process / 59

normally determine what type of mechanized equipment is most feasible. • Apply specialized skills and knowledge to the various design and engineering problems involved in realizing the objectives of automation: mechanics (in the selection of a mechanical movement and also in the design of the part-holding fixture), selection and application of flux and filler metals, type and position of heat application, cooling, additional mechanics of automatic loading and unloading, and electrical control systems.

Practical Experience, Work-Related Tips, and Problem Solving Example 1: Torch Brazing of a 304L Stainless Steel Tube to a Fitting Joint. An attempt was made to braze a tube to a fitting joint that does not have suitable strength, and when it pulled out of the joint, the filler metal did not appear to flow the full length of the joint. The fillets were good, and it appears to have been brazed well before testing. The base metal was 304L stainless steel, and the filler metal was BAg-7 (a silver filler metal). A white flux was used to cover the parts, which were then assembled and torch brazed. The filler metal was fed from the outside of the joint. It appears that the brazement was not heated uniformly, and while the outer end of the fitting was heated up adequately, the back portion of the fitting did not reach a temperature sufficient to allow the filler metal to flow. The front end of the joint was considerably thinner and lighter and thus heated up quite rapidly. However, the back section was considerably heavier and more difficult to heat. The molten filler metal froze before it reached the back end of the joint because the metal was cooler, and the filler metal solidified in the joint. To solve this problem, one must control the torch so that heating occurs closer to the back end of the fitting to ensure that this portion of the joint is up to temperature before the filler metal is applied. A second technique is to supply a ring at the back of the prefluxed joint, then assemble the tube and heat the assembly. Using 304L base metal is significant since the L grade is low carbon and would not form carbide precipitation and, potentially, avoid corrosion problems. The selection of the BAg-7 filler metal, which has a brazing range of 652 to 760 °C

(1206 to 1400 °F), was significant. This filler metal is one of the lower melting materials that does not contain cadmium. To ensure good strength of the joint, the suggested overlap distance should be 3 to 4T, where T is the thickness of the thinner member (which in this case is the tube). With this overlap distance, the strength of the joint should be adequate for this type of fitting. Example 2: Furnace Brazing of a Leaded Steel Fitting to a Copper Heat Exchanger Shell. Brazing a leaded steel with a silver filler metal could create a problem. If the parts were freshly machined, the BAg-4 filler metal with standard white paste flux for silver filler metals should be adequate for brazing the heat exchanger components. A second filler metal that could be suitable is BAg-24, which also contains silver-copperzinc-nickel. This filler metal has a narrower melting range and thus is good for brazing tighter joints. One point to consider when silver brazing a steel fitting to a copper shell of a heat exchanger is the possibility of galvanic corrosion. Depending on the electrolyte, a steel part may show a higher tendency to corrode. Example 3: Brazing Parts Exposed to Synthetic Machining Oils. The newer synthetic oils that are used in machining bear watching. Even though some brazing problems have been traced to oils and lubricants, a definitive answer on this subject has not been formulated to date. Example 4: Application of Filler Metal for Continuous Brazing. In another situation, if one is brazing 304 stainless steel with copper paste filler metal using a continuous brazing furnace with an Inconel muffle through the hot zone, one must be alert to any holes developing in the bottom of the retort in the hot zone. When there is excess copper, it can ball up at the bottom of the part or drip off, and it can drip through and land on the retort. Metallurgically, nickel and copper are mutually soluble in all proportions. Since the Inconel retort is a very high nickel alloy, the copper that pools in the retort actually creates a coppernickel alloy. If the concentration of copper gets high enough to cause melting, the hole in the retort occurs. In order to prevent the above problem, the best way is to tighten all the surface fits so that the clearance on all joints is zero. The best joint strength is obtained with the press-fit or metalto-metal surface where the copper in the joint is

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less than 0.03 mm (0.0012 in.) thick when metallographically inspected. Example 5: Brazing of 409 Stainless Steel. In brazing 409 stainless steel for automotive tubing and fittings, on some occasions the brazed parts come out bright and shiny and sometimes with different purple hues, which interferes with the flow of the filler metal. The brazing work is performed in a continuous furnace with a nitrogen/hydrogen atmosphere with a dewpoint of approximately –46 °C (–50 °F). Copper filler metal is used on some assemblies and nickel-base filler metals on others. To solve the problem and the cause of the variation in color on these brazed parts and why it interferes with the flow of the filler metal is to utilize a nickel plate with approximately 0.01 mm (0.0004 in.) of electrolytic nickel. To provide for easier brazing, one could change to a 409Cb alloy. The niobium (columbium) addition to the 409 also ties up carbon, and the niobium does not produce an oxide on the surface like titanium that interferes with brazing. Example 6: Brazing of Leaded Brass. Lead is more readily vaporized than zinc, and thus any open globules of lead from a machine cut will outgas the lead. There is then a high lead content in the dust taken out of the furnace. This dust becomes a hazardous material to work with and dispose of, and must be handled properly. Example 7: Atmosphere Control. A vacuum furnace operates at 0.0013 Pa (2 × 10–7 psi) at a brazing temperature of 1070 °C (1950 °F). How does the operator determine that the atmosphere in the furnace is acceptable? One easy way to test atmosphere quality as well as base metals of titanium, zirconium, niobium and tantalum, is to use a 0.13 mm (0.005 in.) titanium foil, which is put in the vacuum furnace with or without the load. To check the furnace conditions, use the titanium foil without additional parts. Be certain that the titanium foil rests on a dense piece of alumina or is supported by clean ceramic paper. With a very good atmosphere, the titanium foil will be the same color after the heating cycle as it was when it went in the furnace. Next, the test is to take the foil and bend it back on itself. If the atmosphere was not good, the foil will fracture. If the atmosphere is very good, the foil can be opened without fracturing. The residual gas analyzer (RGA) is an excellent addition to the vacuum furnace. It can pro-

vide a great deal of information as to the residual gases in the vacuum furnace. As is typical of the conditions that can exist in furnaces, one is able to braze titanium very readily at 0.13 Pa (2 × 10–5 psi) in a very clean furnace. Both titanium and zirconium foil can be used for the same application. However, the titanium foil is more readily available. Example 8: Magnesium Buildup in a Vacuum Furnace Used to Braze Aluminum. Several people and companies are brazing aluminum in a vacuum furnace and are having problems with magnesium buildup on the walls and doors between the charge and exit vestibules. The magnesium tenaciously sticks to the metal surfaces and is difficult to remove. There is a material stopoff that has been used in the brazing industry for more than 50 years to limit the flow of filler metal. It has also been found to work well in the salt bath furnaces used in the brazing of aluminum. There are many types of stopoffs available for various applications. Stopoff materials are very interesting and necessary materials because not much material is needed, but they do prevent filler metals from flowing into unwanted areas. (Additional information on stopoff materials is provided in Chapter 7.) Stopoffs are available in solvent-based and water-based versions. The water-based type is recommended for use inside vacuum aluminum brazing furnaces because solvent vapors are undesirable in this confined area. The stopoff is painted on the metal surfaces where the magnesium usually condenses. Stopoff materials can be obtained in a painttype consistency that can be applied by brush or roller to the inside of the furnace in areas where the magnesium is condensing. Apply the stopoff right over what is already there to allow future condensed magnesium to be readily removed. Example 9: High-Frequency Induction Brazing of Stainless Steel. This example involves high-frequency induction brazing of stainless steel parts with BNi-7 filler metal. The parts are inside a quartz tube and use a vacuum atmosphere of 0.013 Pa (2 × 10–6 psi). The braze has been successfully performed; however, some areas of the part reach only a low-red heat. These areas discolor, and organic solvents will not remove the discoloration. At the joint, where the brazing temperature was 1050 °C (1920 °F), the part is bright, shiny silver. The cement being used to hold the filler metal in

Chapter 3: Elements of the Brazing Process / 61

place is not causing this discoloration. Chromium-containing base metals such as 304L will start to oxidize at approximately 540 °C (1000 °F) and will continue to increase the oxide thickness up to 760 °C (1400 °F) before getting progressively lighter. Good brazing quality will increase. When a base metal such as 304L is hot rolled or forged and pickled, the chromium is depleted from the surface, leaving an iron-nickel alloy at the surface. This makes for a much easier braze and requires a less critical atmosphere than the freshly machined chromium-containing stainless steels. To completely eliminate discoloration where there is chromium at the surface, the entire part must be heated in an adequate protective atmosphere close to the brazing temperature, and then a specific type of stopoff is applied to eliminate the discoloration.

REFERENCES

1. R.G. Gilliland, Wetting of Beryllium by Various Pure Metals and Alloys, Weld. J., June 1964, p 248–258 2. D.R. Milner, A Survey of the Scientific Principles Related to Wetting and Spreading, Br. Weld. J., Vol 5, 1958, p 90–105 3. Welding, Brazing, and Soldering, Vol 6, Metals Handbook, 9th ed., American Society for Metals, 1983, p 956 4. Weld. Des. Fabr., June 1993, p 16 5. Weld. J., Oct 1989, p 55 6. P. Kosir and D.K. Patrick, Data Collection and Modern Controls for Hump Mesh Belt Electric Brazing Furnace, Ind. Heat., Oct 1999, p 105–109 7. J. Boswell, C. Field, et al., Production Versatility Increased in Semi-Continuous Controlled Atmosphere Brazing Furnace, Ind. Heat., May 1999, p 73–77 8. T.V. Evans, R.P. Johnson, et al., Next Generation Controlled Atmospheric Braze Processing of Aluminum Heat Exchangers, Ind. Heat., Jan 1997, p 28–30 9. D. Garg, C.E. Baukal, et al., Effective Cooling of Components, Ind. Heat., Feb 1997, p 43–47 10. L.L. Ashburn, Fluxless Vacuum Furnace Brazing of Aluminum Particularly Advantageous for More Critical Applications, Ind. Heat., Mar 1994, p 47–51

11. P.F. Gerbosi and J.F. Libsch, ControlledAtmosphere Brazing with Induction Heating, Weld. J., Oct 1989, p 32–37 12. Brazed Joints, Thin-Wall Tubing Streamline B-70 Plumbing, Space/Aeronaut., May 1960, p 75–90 13. “Brazed Fluid Transmitting Systems,” Gemini Project, Contract NAS9-170, Progress Report 11, McDonnell Douglas, 12 Feb 1963 14. E.C. Nezbeda, “Lunar Excursion Module Manufacturing,” National Aeronautic and Space Engineering and Manufacturing Meeting (Los Angeles, CA), 4–8 Oct 1965 15. W.E. Clautice, Induction Brazing at the Kennedy Space Center, Weld. J., Vol 53(No. 10), Oct 1974, p 612–622 16. Compatibility of Au-Cu-Ni Braze Alloy with NH3, NASA Tech. Briefs, Vol 3(No. 2), 1978, p 226–227 17. M.M. Schwartz, Brazing, ASM International, 1987 18. L.B. Roginskii et al., Reconditioning by Brazing Cast Iron Crankshafts, Svar. Proizvod., Vol 6, 1993, p 14–15 19. S.V. Lashko and N.V. Lashko, Brazing of Metals, Publ. Mashinostroenie, Moscow, 1988 20. S.S. Nurkhanov and L.B. Roginskii, Reconditioning and Hardening Rolls by Electric Resistance Brazing, Weld. Int., Vol 11(No. 1), 1997, p 75–77 21. Q.D. Mehrkam, Never Underestimate the Power of a Salt Bath Furnace, Weld. Des. Fabr., AJAX Reprint 184, Mar 1968 22. Aluminum Brazing Handbook, 1st ed., The Aluminum Association, Jan 1971 23. C.A. Blue and R.Y. Lin, Rapid Infrared Joining of Ti-6A1-4V, Proc. Adv. Mater., Vol 4, 1994, p 21–28 24. J.S. Park and J.M. Kim, Finite Element Modelling for Thermal Analysis of Studto-Plate Laser Brazing for a Dissimilar Metal Joint, NTIS Alert, 1 Aug 1997, p 3 25. J.-S. Park and S.-J. Na, Heat Transfer in a Stud-to-Plate Laser Braze Considering Filler Metal Movement, Weld. J., Apr 1998, p 155–163 26. V.F. Khorunov and V.S. Kutchuk-Yatsenko, “Brazing of Sheet Composite Materials with an Aluminum Matrix,” International Symposium (Montreal, Canada), 1990.

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27. I.S. Duckno and V.S. Kutchuk-Yatsenko, A Method for Making Permanent Joints of Aluminium-Based Metallic Composite Materials to Aluminium Alloys and Ce ramics, Second Japan. International SAMPE Symposium, Society for the Advancement of Material and Process Engineering, 1991, p 758–765 28. I.V. Zvolinskii, V.S. Kutchuk-Yatsenko, et al., Special Features of Brazewelding Metal Matrix Composites, Weld. Int., Vol 9(No. 1), 1995, p 41–43 29. R.V. Hughes et al., Weld. Met. Fabr., Vol 63(No. 3), 1995, p 110–111 30. T. Fleischer, Thermal Loaded Metallic Structures for Hypersonic Planes, NTIS Alert, 15 June 1997, p 27 31. G. Humpston and D.M. Jacobson, Principles of Soldering and Brazing, ASM International, 1994, p 129 32. S.P.S. Sangha, D.M. Jacobson, and A.T. Peacock, Development of the Copper-Tin Diffusion-Brazing Process, Weld. J., Oct 1998, p 432–438

33. J.A. Fernie, “Introduction to Microwave Heating and Its Use for the Joining of Ceramics,” 457/1992, TWI, Aug 1992 34. T.T. Meek and R.D. Blake, CeramicCeramic Seals by Microwave Heating, J. Mater. Sci. Lett., Vol 5, 1986, p 270–274 35. D. Palaith, R. Silberglitt, and E.L. Libelo, “Microwave Joining of Ceramic Materials,” Second International Conf. on Ceramic Materials and Components for Engines, 14–17 Apr 1986 (LubeckTravemunde, W. Germany) 36. D. Palaith, R. Silberglitt, C.C.M. Wu, et al., Microwave Joining of Ceramics, Mat. Res. Soc. Symposium, Vol 124, 1988, p 255–266 37. H. Fukishima, T. Yamana, and M. Matsui, Microwave Heating of Ceramics and Its Application to Joining, Mat. Res. Soc. Symposium, Vol 124, 1988, p 267–272 38. M. Barmatz, H.W. Jackson, and R.P. Radtke, Microwave Brazing of Polycrystalline Diamond onto Drill Bits, NASA Tech. Briefs, Dec 1998

Brazing Second Edition Mel M. Schwartz, p63-176 DOI: 10.1361/brse2003p063

Copyright © 2003 ASM International® All rights reserved. www.asminternational.org

CHAPTER 4

Base Metals and Base-Metal Family Groups THIS CHAPTER BEGINS by describing the general metallurgical considerations related to the selection of different base metals. The remainder of the chapter is devoted to describing specific considerations related to the groups of alloys that are most commonly joined by brazing.

Metallurgical Reactions Some metals and alloys exhibit metallurgical phenomena that influence the behavior of brazed joints and base-metal properties and, in some cases, necessitate special procedures. These phenomena may be classified as: • Base-metal effects, including carbide precipitation • Hydrogen embrittlement • Heat-affected zone and oxide stability • Sulfur embrittlement • Filler-metal effects, such as vapor pressure • Base-metal/filler-metal interactions, including alloying • Phosphorus embrittlement • Stress cracking Other factors that cause interactions between base metals and filler metals include postbrazing thermal treatments, corrosion resistance, and dissimilar-metal combinations. The extent of interaction varies greatly, depending on compositions (base metal and filler metal) and the duration and extent of the thermal cycles in the processing. There is always some interaction, except when mutual insolubility permits practically no metallurgical interaction. In addition to the base-metal effects mentioned previously and the normal mechanical

requirements of the base metal in the brazement, the effect of the brazing cycle on the base metal and the final joint strength must be considered. Cold-worked-strengthened base metals are annealed and the joint strength reduced when the brazing process temperature and time are in the annealing range of the base metal being processed. Hot-cold-worked heat-resistant base metals can also be brazed; however, only the annealed physical properties are available in the final brazement. The brazing cycle, by its very nature, usually anneals the cold-worked base metal, unless the brazing temperature is very low and the time at heat is very short. It is not practical to cold work the base metal after the brazing operation. When a brazement must have strength after brazing that is above the annealed properties of the base metal, a heat treatable base metal should be selected. The base metal can be an oilquench type, an air-quench type that can be brazed and hardened in the same or a separate operation, or a precipitation-hardening type that can be brazed and solution treated in a combined cycle. Parts already hardened may be brazed with a low-temperature filler metal, using short times at temperature to maintain the mechanical properties. The strength of the base metal has a profound effect on the strength of the brazed joint; thus, this property must be clearly kept in mind when designing the joint for specific properties. Some base metals also are easier to braze than others, particularly by specific brazing processes. Carbide Precipitation. If stainless steels are heated to temperatures from 425 to 815 °C (800 to 1500 °F), the carbon in the base metal combines preferentially with chromium to form

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chromium carbide, usually at the grain boundaries. This chromium depletion reduces the corrosion resistance of the stainless steel. This condition has been defined as sensitization by some investigators. In certain corrosive environments, the mechanical properties may be impaired, with little or no apparent surface attack. A short brazing cycle keeps the chromium carbide precipitation to a negligible level with normal types of stainless steels. When this is not possible, one of the special grades of stainless steel may have to be used if its corrosion resistance is to be preserved after brazing. Precipitated carbides in stainless steels may be redissolved by heat treating at 1010 to 1120 °C (1850 to 2050 °F), followed by rapid cooling. Another stabilizing treatment that disperses the unprecipitated chromium uniformly throughout the structure consists of heating to 870 °C (1600 °F) for 2 h, followed by furnace cooling to 540 °C (1000 °F) and subsequent air cooling. If the cooling from the brazing temperature is rapid, no appreciable amount of carbides is precipitated. Where this cannot be done due to mass and it is necessary to braze stainless steels for corrosive service, one of the stabilized compositions, such as type 347 or 321, or an extra-lowcarbon grade, such as 304L, should be used. There are several ways to prevent or minimize the deleterious effects of carbide precipitation. First, because the reaction is time dependent, keeping the brazing thermal cycle as short as possible can minimize carbide precipitation. With short cycle times, such as would result from torch or induction brazing of small parts, even the unstabilized grades can be brazed without severe losses in corrosion resistance. The susceptibility to carbide precipitation also depends on carbon content. Thus, type 304 is less susceptible than type 302, and the extralow-carbon grades, such as type 304L, are relatively insensitive to carbide precipitation. For critical applications, type 347, the niobium-stabilized grade, is recommended. It has good high-temperature strength and can be brazed without danger of impaired corrosion resistance. Type 321 is also a stabilized grade, but it has slightly lower general corrosion resistance than type 347 and is more difficult to braze, because titanium is used as the carbidestabilizing element. When high-melting-point filler metals are used, precipitated carbides can be redissolved by heat treatment after brazing.

Alternatively, corrosion resistance can be restored by diffusing chromium back into the depleted area around the carbide precipitates. Two hours at 870 °C is the recommended homogenizing heat treatment. Hydrogen Embrittlement. Hydrogen can also be a source of trouble. Because of its small atomic size, it is able to diffuse quite rapidly through many metals, and the rate of diffusion increases with temperature. When hydrogen diffuses into a metal that has not been completely deoxidized, it may reduce the oxide of the metal, if the temperature is high enough. Metallic sponge and water vapor are the end products of this reaction. Once hydrogen has diffused into the metal, several things can happen. If oxygen is present, the hydrogen may combine with it to produce water vapor. The water-vapor molecule, unlike the hydrogen molecule, is too large to diffuse out of the metal, and the high vapor pressures that develop can literally tear the metal apart by starting many fissures and blisters, mainly at the grain boundaries. The ultimate result is hydrogen embrittlement. It commonly occurs in copper and copper-base alloys that have not been deoxidized. Pressures developed for tough pitch copper have been calculated to be as high as 620 MPa (90 ksi). Electrolytic tough pitch copper, silver, and palladium, when they contain oxygen, are subject to hydrogen embrittlement if heated in the presence of hydrogen. If tough pitch copper is to be brazed without embrittlement, hydrogen must not be present in the heating atmosphere. A better practice is to use deoxidized copper or oxygen-free copper where brazing is to be performed. Oxygen-free copper, if improperly heated, may also be oxidized and become subject to hydrogen embrittlement. It is impractical to salvage hydrogen-embrittled copper. A recently completed study (Ref 1) examined several commercial filler metals containing zinc, cadmium, or phosphorus that were found to cause embrittlement by migration of (copper) oxide to the grain boundaries, causing void formation and rupture of grain boundaries. (Oxides do not migrate as such but rather dissolve in the grains. The oxygen diffuses to the grain boundaries, where it recombines, forming lessstressed particles.) The brazing was performed both by the conventional fluxed and fluxless methods without the presence of any source of hydrogen; however, this still resulted in the

Chapter 4: Base Metals and Base-Metal Family Groups / 65

same embrittlement. Therefore, it was concluded that the influence of flux is insignificant, because embrittlement also persisted in joints brazed without flux under argon. To narrow the possibilities of embrittlement, a 72Ag-28Cu eutectic filler metal was used to fill several joints, and there was no such embrittlement, even in the most drastic brazing conditions employed. In comparing this filler metal with the others, the only difference was in the composition. While BAg-8 contains silver as the only addition to copper, the other filler metals are a ternary or quaternary formulation, containing additions of zinc, cadmium, and phosphorus. The difference between silver and these other alloying additions is in their ability to reduce copper oxide; silver cannot act as a reducing agent. This indicates that embrittlement occurs as a result of the interaction of the other alloying elements with the copper base metal and not directly because of other factors, such as flux, atmosphere, and the time-temperature cycle. These parameters were identical for all the filler metals, and their influences were as expected; that is, the embrittlement was enhanced when more drastic conditions were employed. Furthermore, embrittlement occurred only in tough pitch copper (containing oxygen as Cu2O precipitates) but not in phosphorus-deoxidized copper, which is completely free of oxides. As a consequence, the coexistence of the additional alloying elements of zinc, cadmium, and phosphorus, together with oxygen in the substrate, is the necessary prerequisite for embrittlement. The results indicate that the responsible elements are those that are capable of reducing copper oxide by a mechanism analogous to the hydrogen embrittlement of tough pitch copper. It seems quite certain that the thermodynamic activity of zinc, cadmium, and phosphorus is sufficiently high to cause the following reaction to take place: CuO + (X) 3 Cu + XO

where X stands for one of these elements in the filler metal. The mechanism suggested by this study is embrittlement induced by the coexistence of copper oxides in the base material together with certain alloying additions in the filler metal, such as cadmium, zinc, or phosphorus, capable of reducing the copper oxides. Steel is especially prone to another mecha-

nism for hydrogen embrittlement. In this type, hydrogen diffuses into the steel as atomic hydrogen in the same manner as it diffuses into copper, but it tends to accumulate in small voids, such as those around nonmetallic inclusions and at grain boundaries. Water vapor is not formed, as in copper, but the hydrogen atoms combine to form hydrogen molecules, which are less mobile and remain trapped at the discontinuities and, as a result, increase the concentration of molecular hydrogen, increase the vapor pressure, and lower the ductility of steel when stressed. However, steel and other ferrous alloys may be salvaged by allowing the hydrogen to diffuse out by baking at slightly elevated temperatures (95 to 205 °C, or 200 to 400 °F) or by permitting the steel to stand for long periods of time until the ductility is regained. A third type of embrittlement can occur when hydrogen combines with the metal to form hydride. The hydride lowers the notch toughness and affects the strain rate of the metal. Titanium, zirconium, niobium, tantalum, and their alloys are subject to this form of hydrogen embrittlement. Ductility can be restored if proper thermal treatments are followed after brazing; however, an inert or vacuum atmosphere should be used for brazing to avoid any embrittlement. Most other metals and alloys whose oxides may be reduced by hydrogen contain an excess of deoxidizing elements and are not subject to hydrogen embrittlement. Heat-Affected Zone. The heating of base metals may cause changes in their properties, particularly if the metals are heated above their annealing temperatures. Base metals whose mechanical properties were obtained by cold working (hard tempers) may soften or undergo an increase in grain size if the brazing temperature is above the recrystallization temperature. Where mechanical properties are obtained by thermal treatment, they may be altered by the brazing operation. Materials in the annealed condition generally experience no appreciable change due to brazing. The width of the zone through which these changes may occur varies with the process used. If the heating is localized, as in torch or induction brazing, the effects are confined to a narrow zone. If the whole assembly is heated, as in furnace brazing, the entire assembly is affected. In general, the heat-affected zone produced during brazing is wider and less sharply defined than those resulting from other welding processes.

66 / Brazing, Second Edition

Oxide Stability and Formation. When clean metals are heated to brazing temperature, their surfaces may form metal oxides if the atmosphere around the part contains oxygen. Oxidized metal surfaces are usually difficult to wet with most filler metals. Fluxes and special atmospheres are designed to prevent oxide formation or to reduce at elevated temperature any oxidation that occurs during initial heating (Ref 2). Chromium, aluminum, titanium, silicon, magnesium, manganese, and beryllium all have oxides that are difficult to remove, and, therefore, these metals usually require special preparation (Ref 2). Fluoride-bearing fluxes can reduce some oxides; hydrogen gas of sufficient purity can reduce them above certain temperatures, and techniques such as vacuum brazing may have to be used. Ideally, oxide formation should be prevented by brazing in low-dewpoint or vacuum atmospheres. Sulfur Embrittlement. Nickel and certain alloys containing appreciable amounts of nickel, if heated in the presence of sulfur or compounds containing sulfur, may become embrittled. This occurs when a low-melting nickel sulfide is formed preferentially at the grain boundaries; this sulfide, being brittle and weak, cracks if subsequently stressed. Material so embrittled is usually scrapped, because the damage that has occurred cannot be salvaged. Nickel and nickel-copper alloys are most susceptible to this attack, whereas alloys containing chromium are less susceptible. It is important that alloys in which nickel is the major component be clean and free of sulfur-containing materials (such as oil, grease, paint, and drawing lubricants) prior to heating and that heating be done in relatively sulfur-free atmospheres. Vapor Pressure. Every metal is in equilibrium with its vapor pressure; some amount of the metal is present in the gaseous state. For most metals, at normal temperatures, this vapor pressure is so small as to be considered nonexistent. For vacuum-tube applications, some metals, such as zinc and cadmium, have relatively high vapor pressures, give off undesirable gases at normal brazing temperatures, and therefore cannot be permitted as constituents of the filler metal. Accordingly, special vacuum-tube-grade filler metals have become commercially available (see Chapter 5, “Brazing Filler Metals”), and special fluxes are used in some situations. Base-Metal/Filler-Metal Interactions. There are always some interactions between the filler metal and the base metal. Although some

of this interaction aids in wetting the base metal, other detrimental effects may occur. Such effects include: • Formation of brittle intermetallic compounds that lower joint strength • Diffusion of the filler metal into the base metal to produce color changes • Creation of a new alloy—with a higher melting point than that of the original filler metal—that chokes off the flow of the filler metal Researchers (Ref 3–6) have reported on their work whereby modified brazing processes for nickel-base materials were used to reduce the formation of brittle phases in the braze joint and also to speed the joining operation. The three modified brazing processes developed by the researchers included: • Brazing under defined load • High-speed brazing • Application of mechanical-excitation brazing Additives (silicon, boride) used to reduce the melting point of nickel-base materials brazed with nickel filler metals cause brittle phases, which exert a negative influence on the mechanical properties of the brazed joints. Diffusion annealing and subsequent aging merge the brittle phases in the braze joint, whereas the aging causes hardening of the base material. After this, the mechanical properties of the joint are comparable to those of the base material. However, applying this heat treatment may cause the formation of coarse grains. Brazing under defined load pressures differs from conventional brazing in that a previously defined load is set up quickly after the brazing temperature is reached. Brazing temperatures of 1150 and 1180 °C (2100 and 2150 °F) with temperature retention times of 1 and 10 min, respectively, have proved to be the best parameter combination. This type of pressure-brazed joint with a more homogeneous microstructure produces good strength properties. Brazing time is of little significance; in contrast, the brazing temperature is of great importance to the strength properties. However, strength values that can be attained without diffusion annealing are comparable with those of conventionally brazed specimens using cost-intensive and time-consuming heat treatments.

Chapter 4: Base Metals and Base-Metal Family Groups / 67

The formation of brittle phases can be influenced or even avoided by combining a considerable reduction of the brazing time with a simultaneous increase in the brazing pressure. The only useful method to achieve the required high gradient of temperature is a conductive heat treatment technique. Another method to prevent the formation of brittle phases is by the mechanical excitation of the components during the brazing process. While brazing, a transducer directly connected to one specimen transfers high-frequency energy into the brazing couple. Using ultrasonic vibrations of approximately 30 kHz (amplitude of approximately 2 µm in the longitudinal direction of the specimens), the accumulation of brittle phases is prevented. The wetting of the base metal with filler metal is improved, because the ultrasonic vibrations destroy any existing surface oxides. The superposition of mechanical excitation produces seams of a quality comparable with those joints heat treated (1100 °C, or 2010 °F, for 20 h) after conventional brazing. Alloying is one of the significant basemetal/filler-metal interactions that can determine the behavior of brazed joints. The extent of interaction varies greatly, depending on the compositions of the base metal and the filler metal and on thermal cycles. There is always some interaction, except where mutual insolubility permits practically none. The term alloying is a general term covering practically every aspect of interaction. Some of these aspects are as follows. First, the molten filler metal can dissolve the base metal. Second, constituents of the filler metal can diffuse into the base metal, either through the bulk of the grains or along the grain boundaries, or can penetrate the grain boundaries as a liquid. The results of such base-metal dissolution or filler-metal diffusion may be to raise or lower the liquidus or solidus temperature of the filler-metal layer, depending on composition and thermal cycle. Examples include nickel, cupronickel, or Monel joined with pure copper filler metal; enough dissolution and diffusion occur so that the solidus of the copper filler metal is increased and flow is terminated. This also means that the remelt temperature of the filler-metal layer is higher than its original solidus temperature. In brazing of ferrous-base high-temperature alloys with filler metals containing boron, grainboundary penetration of the base metal by a low-

melting complex can cause joint degradation. This effect is particularly damaging if the base metal is thin, as in the case of brazed honeycomb sandwich panels. Producers and experienced users of high-temperature filler metals should be consulted during the design of parts for which the use of these filler metals is anticipated. Formation of intermetallic compounds as a result of interactions between constituents of the base and filler metals can occur, and these compounds are usually brittle. Whether or not such compounds form depends on base-metal and filler-metal compositions, time, and temperature, and just because intermetallic compounds do form, it does not necessarily follow that the joint is so embrittled as to lose engineering utility. This depends on the nature of the specific compound, its quantity, and its distribution. Phosphorus Embrittlement. Phosphorus combines with many metals to form brittle compounds known as phosphides. For this reason, copper-phosphorus filler metals are not usually used with iron- or nickel-base alloys; however, two nickel-base filler metals—10 to 12 P, 0.10 maximum C, balance Ni; and 13 to 15 Cr, 9.7 to 10.5 P, 0.08 maximum C, balance Ni—have been used in some applications for brazing heat-resisting alloys. The first filler metal is extremely free flowing, exhibits a minimum amount of erosion with most nickel- and iron-base alloys, and is good for use in exothermic atmospheres. The second is used for brazing of honeycomb structures, thin-wall tube assemblies, and other structures that are used at high temperatures. Erosion can be controlled because of low solubility with iron- and nickelbase alloys, and this filler metal produces strong, leakproof joints with heat-resistant base metals at relatively low brazing temperatures. Furthermore, it is recommended for nuclear applications where boron cannot be used. Stress Cracking. There are many highstrength materials, such as stainless steels, nickel alloys, and copper-nickel alloys, that have a tendency to crack during brazing when in a highly stressed condition and in contact with molten filler metal. Materials with high annealing temperatures—and particularly those that are age hardenable—are susceptible to this phenomenon. Such cracking occurs almost instantaneously during the brazing operation and is usually readily visible, because the molten filler metal follows the crack and completely fills it. This process has been described as stress-corrosion cracking (SCC), where the molten filler

68 / Brazing, Second Edition

metal is considered to be the corrosion medium. Cracking of stressed steel in a caustic solution or stressed brass in an ammonia solution are widely known examples of SCC. Sufficient stress to cause stress cracking can be produced

by cold work prior to brazing or by an externally applied stress from mechanical or thermal sources during the brazing operation. When stress is encountered, its cause can usually be determined from a critical analysis of the

Thermal expansion curves for some common materials. These nomographs assume a case wherein α1 exceeds α2, so that scale value for (α1 – α2) is negative. Resultant values of ∆CD are therefore also negative, signifying that the joint gap is reduced on heating. Where (α2 – α1) is positive, values of ∆CD are read as positive, signifying enlargement of the joint gap on heating. Clearance to promote filler-metal flow must be provided at brazing temperature. D, nominal diameter of joint, mm (in.); ∆CD, change in clearance, mm (in.); ∆T, brazing temperature minus room temperature, °C (°F); α1, mean coefficient of thermal expansion, male member, mm/mm · °C (in./in. · °F); α2, mean coefficient of thermal expansion, female member, mm/mm · °C (in./in. · °F). Source: Ref 7, 8

Fig. 4.1

Chapter 4: Base Metals and Base-Metal Family Groups / 69

brazing procedure. The usual remedy is to remove the source of stress. Stress cracking has been eliminated by: • Using annealed-temper rather than hard-temper material • Annealing cold-worked parts prior to brazing • Removing the source of externally applied stress, such as improper fit of parts or jigs that exert stress on the parts

Fig. 4.1 (continued)

• Redesigning parts or revising joint design • Heating at a lower rate • Heating the fluxed and assembled parts in a torch brazing application to a temperature high enough to effect stress relief, cooling to the brazing temperature, and then hand feeding the filler metal The age-hardenable high-nickel alloys are very susceptible to SCC. These alloys should be

70 / Brazing, Second Edition

brazed in the annealed or solution-treated condition with a relatively high-melting filler metal (preferably above 870 °C, or 1600 °F) that has sufficient strength to withstand handling during the age-hardening treatment. Postbrazing Thermal Treatments. A postbrazing thermal treatment to improve mechanical properties in brazed assemblies is frequently desired. In ferrous alloys, this treatment entails quenching from an elevated temperature, followed by tempering at some lower temperature. In other alloys, such as beryllium copper, 17-7 PH, Inconel X, and some Monels, the treatment consists of heating to some intermediate temperature, followed by a controlled rate of cooling. When a thermal treatment is performed subsequent to brazing, it is important that the filler metal selected have sufficient strength at the thermal treatment temperatures to withstand the necessary handling. It is also important for the base metal, filler metal, and postbrazing thermal treatment to be compatible relative to temperatures in heating and cooling. Postbrazing thermal treatments may generate residual stresses in brazed joints and may result in lowered joint strength. Dissimilar-Metal Combinations. There are many dissimilar-metal combinations that may be brazed. In fact, brazing can often be used where metallurgical incompatibility precludes the use of other joining processes. One of the most important factors to consider in brazing dissimilar metals is rate of thermal expansion. If a metal having high thermal expansion surrounds a low-expansion metal, clearances that are satisfactory for promotion of capillary flow at room temperature are excessive at brazing temperature. Conversely, if the low-expansion metal surrounds the high-expansion metal, no clearance may exist at brazing temperature. For example, in brazing of a molybdenum plug in a copper block, the parts must be press fitted at room temperature. However, if a copper plug is to be brazed in a molybdenum block, a properly centered loose fit at room temperature is required (Ref 7–9). Figure 4.1 shows thermal expansion of copper, molybdenum, and other common materials. Nomographs have been developed that are useful for learning the actual changes in clearance in ring-and-plug joints between dissimilar metals (Fig. 4.2). The equation may be used in cases where more accuracy is important or where one of the variables is off the nomograph scale. In more complex joint configurations, it is

usually best to prepare preproduction samples to establish ideal clearances. A technique often used in brazing of material with different coefficients of expansion is sandwich brazing. A common application of this technique is the manufacture of carbide-tipped metal cutting tools. A relatively ductile metal is coated on each side with the filler metal, and the composite is used in the joint. This places a third material in the joint that creeps during cooling and reduces the stresses caused by differential contraction. In some variations of the technique, wire mesh is used in place of the foil. There are other factors that also must be considered for successful brazing of dissimilar metals. The filler metal must be compatible with both the base metals. Wide differences in basemetal melting points must be considered when choosing the filler metal (Ref 7, 8). Where corrosion or oxidation resistance is needed, the filler metal should have properties at least equal to the poorest of the two metals being brazed. In addition, under conditions of the application, galvanic couplers that may promote crevice corrosion should be avoided. Filler metals that form low-melting phases with the base metals are not recommended unless, as in some special cases, techniques are used to form the final filler metal in situ. The metallurgical reactions that occur during brazing or subsequent thermal treatments between the filler metal and the base metal are important. One example is brazing of aluminum to copper. The copper reacts with the aluminum to form a low-melting brittle compound. Such problems can be overcome by coating one of the base metals with a compatible metal. In the case of aluminum to copper, the copper can be coated with silver or a high-silver alloy and then the joints made with a standard aluminum filler metal.

Base-Metal Family Groups In the ensuing list of base-metal groups, it is apparent that prerequisites for ductile joint behavior include freedom from related hard phase bands in the brazed joint and a matching of elastic-plastic behavior in the brazed joint and base material. It thus follows that not only is the optimal structural development required for a joint capable of deformation, but also the elastic-plastic behavior of the whole system of filler metal and base metal must be considered

Chapter 4: Base Metals and Base-Metal Family Groups / 71

(Ref 10–12). This results in a ductile behavior in the brazed joint and an elastic and plastic behavior in the base metal. When the elastic and plastic properties of the brazed joint and base metal are well matched, this leads to the demand for filler metals whose mechanical properties in the joint are well matched to those of the base metals (Ref 13).

Aluminum and Aluminum Alloys The non-heat-treatable wrought aluminum alloys that are brazed most successfully are the ASTM 1xxx and 3xxx series (e.g., 1350, 1100, 3003, 3004, 3005) and low-magnesium alloys of the ASTM 5xxx series (e.g., 5005, 5050, 5052). Available filler metals melt below the solidus temperatures of all commercial wrought non-heat-treatable alloys. The non-heat-treatable wrought aluminum

Nomograph for equation ∆ C = D∆ T (α – α )

mm D (in.) 150

6

125

5

alloys have manganese or magnesium as major alloying elements. Some alloys contain only one alloying element; others contain several. The alloys with higher magnesium contents are more difficult to braze by the usual flux methods because of poor wetting by the filler metal and excessive penetration into the base metal. Filler metals are available that melt below the solidus temperatures of most commercial non-heattreatable wrought alloys. The heat-treatable wrought alloys most commonly brazed are the ASTM 6xxx series (6151, 6951, 6053, 6061, 6063). The ASTM 2xxx and 7xxx series of aluminum alloys are low melting and, therefore, not normally brazeable, with the exception of the 7072 and 7075 alloys. The rest of the wrought aluminum alloys can be strengthened by heat treatment and, to a much smaller degree, by cold working. They are

D

2

1

(α – α ) 2

1

6

6

10 mm/mm/°C (10 in./in./°F) 100

75

4

mm

3

∆ T °C

(°F) 1800

950 50

900

2

–0.050

–1.00

–0.040

–0.80 –0.60 –0.50 –0.40

1600

–0.30

750

ex li

ne

–11

–9 –0.015

1300

–8 –7 –6 –5

–8 –7 Solution to sample problem

1400

1

–9

–10 –0.020

–0.20

–6 –3

Index lin

e2

–0.10

–4

–5

–0.08

1 650 1200

20

–12 –0.030

–0.005 700

25

–16 –15 –14 –13

–0.010

1500 800 Ind

D (in.)

–1.20

1700

850

∆C

–2 –0.04

600 0.75

–4

–0.06

1100

–3 –0.03 –0.001

550 15

1000 0.50

500

–2

900

–1 Sample problem

10

Given: D = 50 mm ∆ T = 680°C –6 (α – α ) = –5X = –5 x 10 mm/mm/°C 2

7

Solution:

1

∆

C = –0.17 mm D

0.25

Fig. 4.2

Nomograph for determining changes in diametral clearance caused by heating of dissimilar-metal joints. Source: Ref 7, 8

72 / Brazing, Second Edition

called heat treatable alloys. The heat treatable alloys generally contain magnesium and silicon. These alloys, because of their higher total alloy content, begin to melt at lower temperatures than higher-purity aluminum alloys and consequently are generally brazed at lower temperatures than the non-heat-treatable alloys. The heat treatable alloys, which include the commonly used 6061 alloy, are thermally treated during fabrication or after brazing to attain their high mechanical properties. Both groups of alloys lose strength at increasingly rapid rates as they are heated. At their respective annealing temperatures, they exhibit distinctly different behavior. The non-heattreatable alloys soften and return immediately to the O temper. The heat treatable alloys must be held at annealing temperature for a minimum of 20 min before an appreciable portion of their temper disappears. The heat treatable alloys can be returned to temper by heat treatment and aging. The non-heat-treatable alloys must be cold worked to be tempered. Aluminum is brazed at temperatures between 555 and 645 °C (1030 and 1195 °F). Obviously, a certain amount of annealing cannot be avoided during brazing. Aluminum alloys that have solidus temperatures above 590 °C (1090 °F) are easily brazed with commercial binary aluminum-silicon filler metals. Stronger, lower-melting alloys can be brazed with proper attention to filler-metal selection and temperature control, but the brazing cycle must be short to minimize penetration by the molten filler metal. High-quality castings are no more difficult to braze than equally massive wrought alloys. Aluminum sand and permanent mold casting alloys with high solidus temperatures are brazeable; brazeable aluminum casting alloys include 443.0, 356.0, 406, 710, 711, and 850.0. Alloys 443.0, 356.0, and 406 are used for both sand and permanent mold casting. Alloy A710.0 is primarily a sand casting metal. Alloys 711 and 850.0 are used for permanent mold casting. Alloys 443.0, 356.0, and 710.0 are the casting alloys most frequently brazed. Die casting alloys are difficult to braze; the castings are not easily wetted by the molten filler metal and tend to blister when brought to brazing temperature because of their high gas content and entrapped lubricants. Not all aluminum alloys can be brazed. The high-strength wrought aluminum alloys and certain casting alloys contain high amounts of alloying ingredients. These alloys have lower

melting temperatures than those of commercially available filler metals. Some experimental lower-melting filler metals have been produced that can be used for successful brazing of the not-readily-brazeable alloys 2024, 2219, and 7075. Low-melting filler metals contain zinc, copper, or other metal with or without silicon alloyed with aluminum. Although some of these combinations have been used successfully, none has been developed with adequate corrosion resistance to be recommended for commercial use. Therefore, the 2xxx and 7xxx series of aluminum alloys are not normally brazeable, with the exception of 7072 (used as a cladding material only) and 7005. All commercial filler metals for brazing of aluminum alloys are aluminum-base alloys themselves. These filler metals are available as wire or shim stock. A convenient method of preplacing filler metal is to use brazing sheet, which is an aluminum alloy base metal that is coated with a filler metal. Brazing sheet is coated on one or both sides. Core alloys 3003 and 6951 (a heat treatable alloy) are generally used. A third method of applying filler metal is to use a paste mixture of flux and filler-metal powder. Common aluminum filler metals contain silicon as the melting-point depressant with or without additions of zinc, copper, and magnesium. Commercial filler metals for brazing of aluminum are aluminum-silicon alloys containing 7 to 12% Si. Lower melting points are attained, with some sacrifice in resistance to corrosion, by adding copper and zinc. Filler metals for vacuum brazing of aluminum usually contain magnesium. More information on filler metals is presented in Chapter 5, “Brazing Filler Metals.” The optimal brazing temperature range for an aluminum-base filler metal is not only determined by the melting range of the filler metal and the amount of molten filler metal needed to fill the joint but is also limited by the mutual solubility of the filler metal and the base metal being brazed. The brazing temperature ranges of some filler metals are related to those of some base metals. Aluminum can be brazed by most of the standard practices. Most aluminum brazing is done by the torch, dip, or furnace process. Furnace brazing may be done in air or controlled atmosphere, including vacuum. Other methods, including induction, infrared, and resistance brazing, are used for specific applications, usually of the readily brazeable aluminum alloys. Regard-

Chapter 4: Base Metals and Base-Metal Family Groups / 73

less of the process, the temperature must be closely controlled for successful brazing. With dip or furnace brazing, automatic proportioning temperature-control devices are available that can maintain the flux bath within ±3 °C (±5 °F) and the furnace atmosphere within ±6 °C (±11 °F) of the desired brazing temperature. In manual torch, induction, or resistance brazing, operator skill and judgment are used to maintain the required temperature range for brazing based on flux color and on melting and flow of the filler metal. Automated torch and induction brazing utilize paste filler metals and preforms or wire with dispensable fluxes. On the increase is the use of vacuum furnace brazing for aluminum fabrication. Because it is done without flux, the joints are free from the corrosion problems commonly associated with residual or entrapped flux. Moreover, brazed assemblies containing inaccessible recesses can be fabricated efficiently. Furnaces operating in the 0.0013 Pa (10–5 torr) range are used. The success of the operation depends on the use of magnesium vapor as a getter of oxygen on the aluminum surface and magnesium alloyed in the filler-metal (aluminum-silicon) coating. Most furnace brazing of aluminum is done in air, but some benefits can be gained by brazing in a controlled atmosphere; by controlling moisture in the furnace at low levels, less flux is needed to achieve brazing (Nocolok process; Nocolok is a trademark of Solvay Fluor GmbH). New types of brazing sheet designed for fluxless brazing require nitrogen or other inert atmospheres. Furnaces operating at 0.0013 Pa have been used, and the success has stemmed from the use of magnesium as vapor, which acts as a getter of oxide on the aluminum brazing surface. Work has shown that magnesium in the parent aluminum alloy and in the brazing sheet can produce satisfactory brazed joints. Aluminum alloys are sometimes brazed to ferrous materials; in such cases, a nonoxidizing atmosphere is used to protect the ferrous material. Even though dip brazing of aluminum is still a popular and successful process, the search for improved methods of brazing complex aluminum structures, such as heat exchangers, that eliminate or minimize flux removal and corrosive residue problems has continued for many years. Aluminum alloys of the 1xxx, 3xxx, 5xxx, 6xxx, and 7xxx series can be vacuum brazed using No. 7, 8, 13, or 14 brazing sheet that is clad with 4004 filler metal. When additional

filler metal is required, 4004 in wire or sheet form can be introduced. The joint designs used for brazing with flux can also be used for fluxless vacuum brazing. The diffusion brazing process (Ref 14), which uses diffusion between the base metal and the filler metal, has been tried for joining aluminum-silicon alloy castings. The diffusion brazing process with a copper or brass preform, described in Ref 14, can apply for all hypoeutectic, eutectic, and hypereutectic alloys of aluminum-silicon system castings; the minimum temperature where the braze interface showed a liquid phase structure was 530 °C (990 °F) for the copper preform and 510 °C (950 °F) for the brass preform. The shear strength of the diffusion-brazed joint was dependent on the chemical compositions of the base metal, the type of material for the preform, and the brazing temperature and time. The maximum strength of the diffused-brazed joint under optimal conditions was 130 to 150 MPa (19 to 22 ksi) for the base metal of both Al-7Si and Al-12Si alloy castings and 100 to 130 MPa (15 to 19 ksi) for the base metal of Al-20Si alloy casting. The effect of brazing time on the strength of the brazed joint varied depending on the type of preform. The strength of the joint with the copper preform was determined by the characteristics of the braze interface structure. With the brass preform, the strength was controlled by the growth of silicon grains in the braze interface structure. Figure 4.3 shows the effect of brazing temperature and time on the strength of the diffusion-brazed joints of various aluminum-silicon system alloy castings. The 8019 (Al-8.3Fe-4.0Ce) aluminum alloy, a high-performance alloy, was weld brazed by researchers (Ref 15), and the work indicated that no thermal degradation was observed in alloy 8019 after 450 °C (840 °F) for up to 4 h; however, clear degradation could be found at 500 °C (930 °F) after only 100 s. The joining temperature for alloy 8019 was kept below 450 °C. Weld brazing is an innovative new approach that combines brazing with welding to provide synergistic benefits. The results of preliminary studies on 6061-T6 alloys showed that the technique could improve tensile-shear, tensile-fracture, and fatigue strengths. The new generation of aluminum alloys containing lithium exhibit improved specific stiffness and strength compared with conventional aluminum-copper (2000-series) and Al-Zn-Mg

74 / Brazing, Second Edition

(7000-series) alloys. Such alloys are therefore of particular interest to the aerospace industry, where weight reduction is a major requirement. The compositions of some important alloys are given in Table 4.1. High-strength diffusion-bonded joints have been obtained in aluminum-lithium alloy sheet using both solid-state and liquid phase tech-

30

200 Insert metal: Cu

100

500 °C 530 °C 540 °C

15

0 200 Insert metal: Cu - 30% Zn

500 °C 530 °C 540 °C

15

100

0

30

7 % Si 12 % Si 20 % Si

Shear strength, ksi

Shear strength, MPa

7 % Si 12 % Si 20 % Si

0

5

20

10 15 Brazing time, min

Fig. 4.3

Effects of brazing time and temperature on the shear strength of diffusion-brazed joints in relation to the types of preforms and aluminum-silicon alloy castings used

niques (Ref 16). The strengths are much greater than values for adhesive-bonded joints, and in the solid state, the bond microstructure and corrosion resistance are expected to be similar to those of the base metal. The highest shear and peel strengths and minimum variability have been obtained for joints fabricated under very clean conditions in vacuum, and diffusion bonding followed by superplastic forming has been demonstrated for aluminum-lithium 8090 alloy (Ref 16). In another study (Ref 17), diffusion bonds were produced between sheets of an Al-Li-CuMg-Zr alloy using Al-4%Cu vapor-deposited metallic interlayers. The joining was performed at a bonding temperature of 530 °C (990 °F), a pressure of 4 to 5.5 MPa (0.6 to 0.8 ksi), times in the range of 10 to 30 min, and a bonding atmosphere with partial oxygen pressure lower than 2 × 10–3 Pa (3 × 10–7 psi). Microstructural changes were analyzed that occurred both in the parent alloy and in the bond interface after diffusion-bonding cycles and postbonding heat treatments. For previous bonding conditions, elimination of the continuous interfacial oxide layer is possible, and only discrete oxide particles (probably lithium-rich spinels) were detected in the bond interface. This oxide elimination is accelerated if the bonding surfaces are chemically cleaned with a commercial deoxidant. In spite of the stability and high resistance to migration of the bonding interface, it is possible to obtain zones of local recrystallization, especially in superplastic joined sheets. It is probably due to mechanisms of local dynamic recrystallization associated with highly deformed zones.

Table 4.1 Compositions and densities of some important aluminum-lithium alloys Composition, wt% Production route

Ingot

Melt spun Powder atomized Mechanically alloyed Source: Ref 16

Alloy

Li

Cu

Mg

Si

Fe

8090 (U.K.) 8091 (U.K.) 2090 (U.S.) 2091 (France) Weldalite 049 (U.S.) 1420 (Russia) 678 (U.S.) . . . (U.S.) Cospray (U.K.) IN 905 XL (U.S.)

2.1–2.7 2.4–2.8 1.9–2.6 1.7–2.3 1.3 1.5–2.6 3.22 4.1 4.33 1.5

1.0–1.6 1.8–2.2 2.4–3.0 1.8–2.5 4.5–6.3 ... 0.54 ... ... ...

0.6–1.3 0.5–1.2 0.25 1.1–1.9 0.4 4.0–7.0 3.07 1.0 ... 4.0

0.2 0.3 0.1 0.2 ... ... ... ... ... ...

0.3 0.5 0.12 0.3 ... ... ... ... ... ...

Zr

0.04–0.16 0.08–0.16 0.08–0.15 0.04–0.10 0.14 0.05–0.3 0.83 0.2 0.11 ...

Density, 103 kg/m3

2.53 2.54 2.60 2.58 2.73 2.50 ... ... 2.41 2.58

Chapter 4: Base Metals and Base-Metal Family Groups / 75

The thermal bonding cycles applied to the alloys do not cause serious surface oxidation problems, although surface lithium depletion occurs by sublimation. In addition, the grain size of the superplastic alloy is kept in the range necessary for superplastic behavior. Diffusion-bonding trials were carried out using the same alloy (AA8090), both in the nonsuperplastic (T6) and superplastic conditions. The liquid phase bonding in air of unreinforced and SiC-fiber-reinforced aluminum using interlayers of copper-silver alloy was investigated (Ref 18). The bond strengths were measured using a simple shear jig, and the associated microstructures were characterized by electron microscopy and electron probe microanalysis. A shear strength of 65 MPa (9.4 ksi) was achieved when bonding unreinforced aluminum at 510 °C (950 °F) using a 50 µm thick alloy interlayer and a pressure of 10 MPa (1.5 ksi) for 30 min; the bonded region covered ~85% of the area of the joint. With reinforced aluminum, a bonding pressure of 20 MPa (3 ksi) was required to achieve sufficient contact and a similar area of bond; after application of pressure for 30 min at 510 °C, shear strength of 54 MPa (8 ksi) was developed at the joint. A mechanism for bonding that reflects diffusion behavior and interphase reactions is shown in Fig. 4.4. A novel brazing technique (a simplified and cost-effective method) using an alloy powder mixture instead of a clad surface has been developed for brazing aluminum, copper, and brass (Ref 19, 20). In these applications, at least one of the aluminum components is clad with filler metal consisting of an aluminum-silicon alloy of neareutectic composition, such as AA4045, AA4047, or AA4343 (Table 4.2) (Ref 21). These alloys contain 9 to 13 wt% Si and are characterized by a melting temperature (in a narrow range near 577 °C, or 1071 °F) (Ref 22) considerably lower than that of the core alloy

(~660 °C, or 1220 °F). Joining is carried out in the presence of a noncorrosive flux, such as a fluoroaluminate salt (Ref 23, 24), to remove native surface oxide films from the contacting aluminum surfaces. Oxide removal enhances wetting by the molten aluminum-silicon eutectic alloy at the brazing temperature and eases liquid metal penetration of the joint. The present brazing technique uses the eutectic bonding approach described in earlier publications (Ref 25, 26) but avoids the need to coat the base-metal surface with an intimately adhering layer of the eutectic-forming metal by electroplating or vacuum deposition. In the present technique, at least one of the aluminum surfaces

Al

Al FeAlx

Ag - Al Cu - Al Cu - Ag

Cu - Ag

Cu - Al Ag - Al Al 50 µm

(a)

Al

(b)

Al

Al

Al - Cu - Ag liquid

Al - Cu - Ag liquid

Ag2Al

Ag2Al

Al

(c)

Al

(d)

Al

Al

Al - Cu - Ag liquid

Al - Cu - Ag liquid

Ag2Al

Ag2Al

Al

(e)

Al

(f)

Fig. 4.4

Proposed mechanism for aluminum bonded with copper-silver alloy interlayer. (a) Start of process, 490 °C (915 °F). (b) Solid-state interdiffusion, 0 to 5 min. (c) Formation of Al-Cu-Ag liquid, 5 min, 502 °C (935 °F). (d) Expulsion of liquid, 5 to 10 min, 510 °C (950 °F). (e) Further reaction, 5 to 15 min, 510 °C (950 °F). (f ) Isothermal solidification, >15 min, 510 °C (950 °F)

Table 4.2 Composition limits for selected aluminum alloys Content, wt% Aluminum alloy

1050 1100 3003 3102 4045 6061 X-800

Si

Fe

Cu

Mn

Mg

Zn

Cr

Ti

Others

0.25

0.4

0.05 0.05–0.2 0.05–0.2 0.1 0.3 0.15–0.4 0.31

0.05 0.05 .1.0–1.5 0.05–0.4 0.05 0.15 1.11

0.05 ... ... ... 0.05 0.8–1.2 0.27

0.05 0.1 0.1 0.3 0.1 0.25 ...

... ... ... ... ... 0.04–0.35 ...

0.03 ... ... 0.1 0.2 0.15 0.008

... 0.15 0.15 0.15 0.15 0.15 0.15

0.95 Si+Fe 0.6 0.4 9–11 0.4–0.8 0.05

0.7 0.7 0.8 0.7 0.19

76 / Brazing, Second Edition

is coated with a thin layer of a powder mix consisting of an element capable of forming a lowtemperature eutectic with aluminum (e.g., silicon, copper, germanium, zinc) and a flux capable of dissolving surface oxide films (Ref 27), as illustrated in Fig. 4.5(a). A commonly available noncorrosive flux (Ref 24) was used in the present work. This flux consists of a mixture of KAlF4 and K2AlF5 · H2O powders in a molar ratio of the respective salts of approximately 13 to 1 (Ref 28), with a particle dimension of the order of 1 µm. Brazing is carried out by heating the joint at approximately 600 °C (1110 °F) in nitrogen gas at near-atmospheric pressure for a few minutes. During temperature ramp-up, the flux melts at ~562 °C (1044 °F) and dissolves the surface oxide layers on the aluminum, as illustrated in Fig. 4.5(b). Sufficient flux must always be present to remove these oxides. Oxide dissolution must occur more rapidly than reoxidation of the aluminum surface, allowing the silicon particles to come into intimate contact with the bare metal. At this juncture, the large, elemental concentration gradients at the aluminum-silicon

Fig. 4.5

interface cause the aluminum and silicon to interdiffuse (Fig. 4.5c). At temperatures exceeding 577 °C (1071 °F), it is found that the silicon particles diffuse rapidly into the aluminum surface and generate in situ a layer of aluminum-silicon liquid alloy of near-eutectic composition (Fig. 4.5d). The filler metal penetrates the joint of interest by capillary action and forms a fillet, thus producing a metallurgical bond on cooling. Any unused filler metal remains on the aluminum surface to form a layer of aluminum-silicon alloy of near-eutectic composition (Fig. 4.5e). In this process, only one of the joined surfaces need be covered with the silicon-flux mix, because the molten flux spreads rapidly across the joint to remove surface oxide films from the mating surfaces. The quantity of filler material formed from one coated surface is generally sufficient to yield a good metallurgical bond. Because fillets are formed through capillary flow of the filler metal, brazing requires only minimal contact force at the joint interface (Ref 29). In recent years, the use of rapid-solidification powder metallurgy has made it possible to

Successive steps in the novel brazing process. (a) Deposition of a silicon-flux powder mix on the aluminum surface. The silicon particle dimensions range from ~1 to 100 µm; the flux particle dimensions do not exceed 1 µm. (b) Melting of the flux at 562 °C (1044 °F) and dissolution of surface oxide films. (c) At 562 °C 577 °C (1070 °F), rapid dissolution of silicon to form localized pools of filler metal of neareutectic composition, followed by coalescence of liquid metal pools. (e) End of filler-metal generation and solidification

Chapter 4: Base Metals and Base-Metal Family Groups / 77

develop a new family of aluminum alloys exhibiting unique properties. One of these materials, dispersion-strengthened aluminum, derives its high strength from nanoscale AlN particles embedded in an aluminum matrix. Unlike precipitate particles in conventional aluminum alloys, these ceramic particles resist dissolution and/or coarsening during thermal treatment; therefore, the material is capable of maintaining much of its strength at high temperatures (Ref 30–32). Due to these properties, dispersionstrengthened aluminum is expected to be well suited as construction material for high-temperature applications where weight reductions are of particular concern. Two different brazing filler metals, based on the aluminum-silicon system, were investigated with regard to the brazeability of dispersionstrengthened aluminum (Ref 33). One of the filler metals was a commercial type with eutectic aluminum-silicon composition, whereas the other one was especially made for the present investigation (Table 4.3). The alloy was made with an addition of 1% Mg, which has been reported in numerous papers to have a beneficial effect on the wetting behavior of the filler metals in vacuum (Ref 34–38). During brazing, magnesium evaporates and has a gettering effect on the residual gas contaminants in the furnace atmosphere. This reduces the oxidation of aluminum and simultaneously breaks up the surface oxide layer through precipitation of indigenous MgO crystallites. The results on the wetting behavior of dispersion-strengthened aluminum under conditions applicable to brazing showed that a eutectic aluminum-silicon filler metal completely wets the base metal both under high-vacuum conditions and in controlled argon atmospheres, provided that the partial pressure of oxygen is sufficiently low. The main problem appears to be the stability of the matrix grain structure. In general, the

process of grain erosion and coarsening can be controlled by restricting the supply of the filler metal so that only a small metal volume is exposed to erosion. In addition, there is a great potential for reducing the thermodynamic driving force of the erosion reaction by proper adjustments of the filler-metal composition and/ or the brazing temperature (Ref 33). Still, grainboundary liquidation may be a problem, which, in turn, may require additions of surface-active elements to the filler metal to control the wetting behavior (Ref 39).

Beryllium and Beryllium Alloys Joint design, filler-metal selection, and a choice of brazing cycle are complicated by the low ductility of beryllium as well as its reactivity with many of the filler metals used for brazing. The low ductility of beryllium is aggravated by the presence of structural discontinuities, such as surface scratches, notches, and asymmetrical stress patterns produced by single-lapped joints. Beryllium parts must be handled carefully, and joint configurations such as butt, scarf, step, and double-lapped joints should be considered when beryllium structures are designed. Because beryllium reacts with the constituents of most filler metals, brazing should be done under conditions that minimize the formation of intermetallic compounds: rapid heating and cooling cycles, low brazing temperatures, minimum time at brazing temperature, and minimum amounts of filler metal. However, these procedures are not always practical, because many filler metals do not wet or flow well on beryllium surfaces. As a result, filler metal frequently must be preplaced between the joint members in amounts sufficient to produce the brazed joint. Longer brazing times than desired may be needed to ensure wetting and flow of the

Table 4.3 Compositions of filler metals investigated for brazing of dispersion-strengthened aluminum Content, wt% Filler metal

No. 1 No. 2 Source: Ref 33

Si

Mg

Ti

Br

Mn

Zn

Cu

Fe

Al

11–13 11

0.03 1

... 0.2

... 0.04

0.07–0.2 ...

0.05 ...

0.05 ...

0.4 ...

bal bal

78 / Brazing, Second Edition

filler metal; under such conditions, the diffusion that occurs between the base metal and the filler metal may tend to adversely affect the joint properties. Brazing is the preferred method for metallurgically joining beryllium. From a stiffness/ weight standpoint, beryllium is unexcelled by any bulk metal or alloy in existence. However, due to its chemical and metallurgical reactivity, braze-joining techniques must be highly specialized, and suitable filler-metal systems and brazing temperature ranges include: • Zinc (427 to 454 °C, or 801 to 849 °F) • Aluminum-silicon (566 to 605 °C, or 1051 to 1120 °F) • Aluminum (645 to 655 °C, or 1195 to 1210 °F) • Silver-copper (640 to 904 °C, or 1185 to 1659 °F) • Silver (882 to 954 °C, or 1620 to 1749 °F) Strictly speaking, zinc, with a melting point of 420 °C (788 °F), does not quite meet the accepted definition for liquidus temperature for a filler metal (425 °C, or 800 °F). Nevertheless, it is generally accepted as the lowest-melting filler metal for brazing of beryllium. Using rather specialized techniques, zinc brazing of beryllium has found important applications. Its successful use was demonstrated in production joining of over 2000 fuselage longeron components for the beryllium spacer for an intercontinental ballistic missile. Brazing was selected as the joining process for the longeron assembly because the brazed design had the lowest cost, the least weight, and it minimized clearance problems in the assembly (Ref 40). Zinc was selected as the filler metal because it caused no detrimental reaction with the beryllium base material. The dip brazing of beryllium with zinc filler metal offered the following advantages: • Zinc has no undesirable reaction with beryllium. • The brazing temperature was low enough that no recrystallization effects took place in beryllium sheet. • Shear strength at room temperature was consistent and high (128 MPa, or 19 ksi). • Shear strength at room temperature decreased linearly to a value of approximately 17 MPa

(2.5 ksi) at 425 °C (800 °F), indicating moderate temperature applications. • Sophisticated equipment was not required. Aluminum and aluminum-silicon filler metals that contain 7 to 12% Si can provide highstrength brazed joints for service at temperatures up to 150 °C (300 °F). These filler metals have been used widely in high-strength wrought beryllium assemblies, because joining is performed well below the base-metal recrystallization temperature. Fluxless brazing requires stringent processing but has been used effectively. A significant advantage of aluminumbase versus silver-base filler metals is that metallurgical interaction is minimal. This is of prime concern when thin beryllium sections or foils are to be joined. When beryllium assemblies are brazed using the aluminum-silicon-type filler metal, by virtue of the low furnace brazing temperature (605 °C, or 1120 °F), the wrought sheet beryllium base properties do not deteriorate but maintain the desirable isotropic mechanical properties of beryllium. Silver and silver-base filler metals, such as Si-7Cu-0.2Li, find use particularly in structures exposed to elevated temperatures. An added advantage with these systems is that atmosphere brazing is used and may be performed in inertgas atmospheres or vacuum. The lithium is added to improve wettability. Silver interacts favorably with beryllium, in that hard or brittle intermetallics do not remain on cooling. Brazing temperatures vary, depending on the mass of the parts to be joined, joint precision, and pressure applied to the components. Both the silver-base and aluminum-silicon filler metals, however, exhibit poor capillary flow, and preplacement is recommended. The use of the 7% Cu-0.2%Li sterling silver filler metal markedly improves the ability of the filler metal to make well-defined fillets and to bridge gaps of several mils. Copper beryllides, which generally form at the faying surfaces, inhibit rapid penetration of silver into the beryllium components, and, hence, a more stable liquid phase is present at the brazing temperature. Properly executed beryllium brazements may be made to exhibit useful strengths at temperatures up to approximately 870 °C (1600 °F) (Ref 40, 41). Research on brazing beryllium to a CuCrZr vabotron element (Ref 42) found that beryllium forms brittle intermetallic compounds with most

Chapter 4: Base Metals and Base-Metal Family Groups / 79

other metals; however, beryllium and copper form a eutectic at ~850 °C (1560 °F), and beryllium recrystallizes above 730 °C (1350 °F). The alloy CuCrZr is precipitation hardenable, requiring annealing (>900 °C, or 1650 °F) and hardening at ~500 °C (930 °F) for optimal strength, while beryllium forms an oxide layer on its surface. The use of BAg-18 (Ag-30Cu-10Sn) as a filler metal was successful in joining the beryllium component at 650 to 680 °C (1200 to 1255 °F) for 10 to 45 min under a pressure of ~0.3 MPa (0.04 ksi). Also used was the Incusil active brazing alloy (ABA) filler metal (Wesgo Metals) with vacuum induction brazing at 720 °C (1330 °F) for 30 s to 3 min at a pressure of 0.4 MPa (0.06 ksi).

Copper and Copper Alloys The copper alloy base metals include copperzinc alloys (brass), copper-silicon alloys (silicon bronze), copper-aluminum alloys (aluminum bronze), copper-tin alloys (phosphor bronze), copper-nickel alloys, and several others. The previously mentioned alloys are commonly brazed with copper- and silver-base filler metals (Table 4.4). The brazeability of copper and its alloys is generally rated from good to excellent. With

some alloys, however, difficulties may be encountered. For example, some lead-containing alloys can form dross that interferes with wetting, and tin-containing alloys, if not stress relieved before brazing, may crack when subjected to rapid localized heating. Softening of the base metal frequently occurs during brazing, because many copper-base alloys derive their properties as a result of heat treatment at relatively low temperatures, cold work, or both. The degree of softening increases with higher temperatures and longer times of exposure to elevated temperatures. Softening of areas in proximity to the braze can be minimized by cooling the assembly, except for the area to be brazed, by immersion in water, packing with wet rags or wet asbestos, or otherwise providing a heat sink to keep the overall temperature of the part as low as possible. In all cases, brazing with a low-melting filler metal for a minimum time reduces softening. The importance of uniform and controlled heating cannot be overemphasized, especially when dealing with brasses, cold-worked phosphor bronzes, and cold-worked silicon bronzes. They are especially susceptible to cracking. Tough pitch coppers are subject to embrittlement when heated in reducing atmospheres containing hydrogen. Consequently, although tough pitch coppers are generally rated as hav-

Table 4.4 Filler metals and other parameters for brazing of copper alloys Base material

Coppers C10100 to 14200 High coppers C15000 to 19400 Red brasses C20500 to 24000 Yellow brasses C25000 to 29800 Leaded brasses C31000 to 38500 Tin brasses C40500 to 45500 Phosphor bronzes C50200 to 52900 Aluminum bronzes C60600 to 64200 Silicon bronzes C64700 to 66100 Copper nickel C70100 to 82000 Nickel silver C73200 to 79800

Atmosphere

Flux

BCuP-2, BCuP-5 BCuZn; BAg-1, -1a, -2, -5, -6, -18 BAg-8, -1

Filler metal

None 1, 2, 5

None FB3A, C, D, E, I, J FB3A

BAg-1, -1a, -2, -5, -6; BCuP-5, -3; BCuZn BCuP-3, -4, -5; BAg-1, -1a, -5, -6 BAg-1, -1a, -2, -7, -18; BCuP-5 BAg-1, -1a, -2, -5, -6; BCuP-5, -3 BAg-1, -1a, -2 BCuP-3, -5; BAg-5, -6 BAg-1, -1a, -2, -3

1, 2, 5

BAg-1, -1a, -2 BAg-1, -1a, -2, -5, -18; BCuP-3, -5 BAg-1, -1a, -2, -5, -6; BCuP-3, -5

...

Notes

Do not braze oxygen-containing alloys in hydrogen atmospheres.

3, 4, 5

FB3 A, C, D, E, I, J FB3A, C, E

3, 4, 5

FB3A, C, E

3, 4, 5

FB3A, C, E

None 1, 2, 5 4, 5

None FB3A, C, E FB4A

Stress relieve prior to brazing

4, 5

FB3A, C, E

1, 2, 5

FB3A, C, E

Clean surface well. Stress relieve prior to brazing Stress relieve prior to brazing

3, 4, 5

FB3A, C, E,

Stress relieve prior to brazing

Braze cycle should be short to avoid softening Short braze cycle. Stress relieve prior to brazing ...

Apply flux with furnace brazing

Note: For information about brazing copper alloys under Atmosphere, see Table 3.4; under fluxes, see Chapter 6, “Fluxes and Atmospheres.”

80 / Brazing, Second Edition

ing good to excellent brazeability, they should not be brazed in a furnace that contains hydrogen with a reducing potential, such as dissociated ammonia, or in an exothermic-based or endothermic-based atmosphere. Heating by open flame or by torch also may result in hydrogen diffusion and embrittlement. Phosphorus-deoxidized and oxygen-free coppers can be brazed without flux in hydrogencontaining atmospheres without risk of embrittlement, provided that self-fluxing filler metals (copper-phosphorus) are used. The use of flux is required, however, when silver-base filler metals that contain additives such as zinc, cadmium, or lithium are used to braze these coppers to each other or to copper alloys or other metals. The coppers, including those that contain small additions of silver, lead, tellurium, selenium, or sulfur (generally no more than 1%), are readily brazed with the self-fluxing copperphosphorus filler metals, but wetting action is improved when a flux is used and when a sliding motion between components is provided while the filler metal is molten. Another copper group includes those with enhanced mechanical properties due to the addition of small amounts of alloying elements. These precipitation-hardenable copper alloys contain beryllium, chromium, or zirconium and form oxide films that impede the flow of filler metal. To ensure proper wetting action of the joint surface by the filler metal, beryllium-copper alloys should be freshly machined or mechanically abraded before being brazed. Removal of beryllium oxide from joint surfaces requires the use of a high-fluoride-content flux. Brazing precipitation-hardenable coppers in the aged condition reduces their mechanical properties. The high-strength beryllium-copper alloys (2% Be) can be furnace brazed and simultaneously solution treated at 790 °C (1455 °F). The temperature is lowered to 760 °C (1400 °F) to solidify the filler metal, and then the assembly is rapidly quenched in water and aged at 315 to 345 °C (600 to 655 °F), which develops adequate hardness in the base material. The 72Ag28Cu eutectic filler metal is generally used with flux. When the sections to be brazed are thin and can be cooled very rapidly, solution heat treated beryllium-copper may be brazed in the temperature range of 620 to 650 °C (1150 to 1200 °F). Other beryllium-copper alloys (0.5% Be) can be silver brazed rapidly with filler metals containing 45 to 50% Ag, 15% Cu, 16% Zn, and 18 to 24% Cd.

The chromium-coppers are brazed with silver-base filler metal and fluoride flux and simultaneously solution treated at 900 to 1010 °C (1650 to 1850 °F) and cold worked; aging can take place as a subsequent operation to develop improved mechanical properties. This is not the case with zirconium coppers. They do not precipitation harden without the benefit of prior cold working, a sequence that is incompatible with brazing. In the absence of cold working, the strength of zirconium coppers is not improved by aging treatments. Brasses. The copper-zinc alloys (brasses) are produced with varying ratios of the two elements to provide desired properties and casting characteristics. Other elements occasionally are added to enhance particular mechanical or corrosion properties; these special brasses are identified by the added element—for instance, lead, which gives leaded brass. Additions of manganese, tin, iron, silicon, lead, and aluminum, either singly or collectively, rarely exceed 4%. Red brasses (low) contain up to 20% Zn and are readily brazed with a variety of filler metals. Flux is normally required for best results, especially when the zinc content is above 15%. All the brasses can be brazed with the silver and copper-phosphorus filler metals, and the highermelting-point (low-zinc) brasses can also be brazed with the copper-zinc filler metal. It is recommended that protective atmospheres be used in furnace brazing of brasses; however, even in a protective atmosphere, flux should be used to promote good wetting by the filler metal and to reduce zinc fuming. When heated above 400 °C (750 °F), brass tends to lose zinc by vaporization. This loss can be reduced by fluxing the parts during furnace brazing or by using an oxidizing flame during torch brazing. Brasses are also subject to cracking and should therefore be heated carefully and uniformly. Yellow brasses (high) contain 25 to 40% Zn and are readily brazed, but low-melting filler metals should be used to avoid dezincification of the base metal. Leaded brasses are formed when lead is added to red or yellow brass in amounts up to 5%. Alloys containing more than 5% Pb are usually not brazed. The lead forms dross on heating that can seriously impede wetting and the flow of the filler metal. Consequently, in brazing of leaded brasses, the use of a flux is mandatory to prevent dross formation in the joint area. Naval brass, leaded naval brass, and admiralty brass contain up to 1% Sn and may contain other

Chapter 4: Base Metals and Base-Metal Family Groups / 81

alloying elements such as lead, manganese, arsenic, nickel, and aluminum. Except for the aluminum-containing alloys, these brasses are readily brazed; they have greater resistance to thermal shock and are less susceptible to hot cracking than the high-lead brasses. For proper wetting, brasses that contain aluminum require a special flux. Bronzes. Alloys of copper and tin are properly termed tin bronzes. However, these coppertin alloys contain small amounts of phosphorus (up to 0.25%) added as a deoxidizer. The tin bronzes have come to be known commercially as phosphor bronzes. Although susceptible to hot cracking in the cold-worked condition, phosphor bronzes have good brazeability and are adaptable to brazing with any of the common filler metals that have melting temperatures lower than that of the base metal. To avoid cracking, phosphor bronzes should be stress relieved at 290 to 345 °C (555 to 655 °F) before brazing. After this stress relief or anneal, the parts should be supported in a stress-free condition during brazing, and slow heating cycles should be used to avoid thermal shock. When the tin content is high or when there are appreciable lead additives, adequate flux protection during brazing is necessary. All the phosphor bronzes can be brazed with the Ag-Cu-Zn-Cd and copper-phosphorus filler metals. The copper-zinc filler metals are appropriate for brazing the low-tin varieties. Aluminum bronzes are copper-aluminum alloys with high copper contents and 3 to 13% Al, with or without varying amounts of iron, nickel, manganese, and silicon. Aluminum bronzes are generally considered difficult to braze because of their aluminum content, which results in the formation of refractory aluminum oxide at brazing temperature in alloys containing more than 8% Al. However, alloys containing 8% Al or less are brazeable, provided that appropriate fluxes are used to dissolve the aluminum oxide. The oxide, which inhibits the flow of filler metal, cannot be reduced in dry hydrogen. Electroplating of copper on the surface to be brazed is one technique used to enhance wetting of the low-melting silver filler metal, while use of flux and a protective atmosphere in a furnace is a technique employed if unplated surfaces are used. Silicon bronzes that contain up to approximately 3.25% Si and are in a highly stressed condition are susceptible to hot shortness and

stress cracking by molten filler metal and should be well fixtured during heating and brazing to prevent excessive stressing. To avoid cracking, these alloys should be stress relieved at approximately 290 to 345 °C (555 to 655 °F) before brazing. For best brazing results, joint surfaces of copper-silicon alloys should be freshly machined, cleaned, and flux coated or copper plated before brazing to prevent the formation of refractory silicon oxide. Mechanical cleaning is recommended, or, for light oxide, the material can be pickled. Silver-bearing filler metals and flux generally are used. Other Copper Alloys. Commercially available copper nickels may contain from 5 to 40% Ni and are susceptible to both hot cracking and stress cracking by molten filler metal. To prevent cracking, copper-nickel alloys should be stress relieved before brazing. Stresses should not be introduced during brazing. The Ag-CuZn-Cd filler metals and flux are preferred for brazing the copper nickels. Nickel silvers, which are brasses (Cu-Zn-Ni) that contain up to approximately 20% Ni but do not contain silver, are highly susceptible to hot cracking and should be stress relieved at approximately 290 °C (555 °F) before being brazed. The nickel silvers can be brazed readily with the same procedures used for brazing brass. When copper-zinc filler metals are used, however, care is required because of the relatively high brazing temperatures. Brazing Processes for Copper Alloys. Typical brazing processes used for copper and its alloys include torch, furnace, and induction. The advantages of furnace brazing that are applicable to the joining of copper and copper alloys relate to the furnace as a source of heat and to the cooling chamber that is provided on conveyor-belt furnaces as a means for cooling assemblies from the brazing temperature to 150 °C (300 °F) or below. To a lesser degree, the protective atmospheres used most conveniently in furnace brazing—notably, the exothermic-based and endothermic-based atmospheres—constitute another advantage of brazing deoxidized coppers and copper alloys in a furnace. For copper alloys susceptible to hot cracking, control of cooling rate in the cooling chambers of brazing furnaces is as important as control of the rate of heating to the brazing temperature. Torch brazing of copper and its alloys is selected largely on the basis of feasibility and cost. Torch brazing is often used when workpiece limitations preclude the use of alternative

82 / Brazing, Second Edition

brazing processes. Low equipment cost is a major advantage in manual torch brazing, which is particularly useful for assemblies involving unequal masses. A brazer with moderate skill can adjust and apply the heating flame so that unequal masses are brought uniformly to brazing temperature. The brazer can also apply the heat selectively to the joints of assemblies involving both large and small areas. The size of brazing flames may range from those of extremely small torches the size of a hypodermic needle used for electronic leads up to those of large torches used in brazing assemblies weighing hundreds of kilograms. Precautions should be taken in torch brazing certain coppers and copper alloys. Where it is necessary to braze oxide-containing coppers, a reducing atmosphere in the flame must be avoided, because it can promote hydrogen embrittlement. For these coppers, a neutral or slightly oxidizing flame and a short brazing cycle are necessary. Brasses are subject to volatilization of zinc when overheated or when held too long at brazing temperature. Application of flux suppresses zinc volatilization. Alloys containing elements that readily form refractory oxides (aluminum, beryllium, chromium, and silicon) must be protected by flux and should not be exposed to an oxidizing flame. The efficiency of heating by induction varies directly with the electrical resistivity of the alloy. Brass, because it has higher electrical resistivity, can be heated more efficiently than copper; steel, which has even higher resistivity, can be heated more efficiently than brass. In terms of the high-frequency power input and the time required to heat 0.45 kg (1 lb) of metal in a joint assembly at a brazing temperature of 705 °C (1300 °F) and a power input (450 kHz) of 15 kW, steel required 16 s, whereas brass required 30 s and copper approximately 55 s. The main advantages of induction brazing of copper and its alloys over torch brazing are minimized warpage, reduced postbraze cleaning, and less required operator skill. Induction heating is suited for mass producing brazed assemblies, primarily because inductors can be designed that heat a line of assemblies as they are carried through the induction field by conveyor belt or turntable. One of the general limitations of induction brazing is the cost of induction heating equipment, which far exceeds the cost of torch brazing equipment and usually exceeds the cost of equipment for resistance brazing or

dip brazing in molten salt. Because the efficiency of heating copper and copper alloys by induction is generally low, costs to achieve given production rates are high. The low longterm operating and maintenance costs of induction units, however, may outweigh the higher initial costs. Other general limitations of induction brazing may relate to the size and shape of assemblies that can be brazed, the design of inductors (cooling coils used to convey heat to the assembly), and the requirements for matching impedances that apply to brazing of copper and copper alloys. In a recently completed study (Ref 43) on the use of advanced particulate-reinforced composites and dispersion-hardened materials as candidate materials to replace pure annealed copper for the electromagnetic windings of the compact ignition tokamak nuclear fusion reactor, it was found that direct brazing of such materials may lead to brittle joints and that indirect brazing may need to be employed for joining such materials. As a result, 0.25 vol% alumina dispersionhardened copper (ADHC) was brazed satisfactorily using induction brazing in an argon atmosphere using 72Ag-28Cu eutectic and pure silver electroplate as interlayer materials. Both interlayers completely wet ADHC. For both interlayers, the application of additional stress (20 MPa, or 3 ksi) on the joint improved the joint quality. By plating a 17.5 µm layer of copper on the ADHC prior to brazing, other researchers have obtained shear strengths as high as 223 MPa (32 ksi). Resistance brazing is often used in joining copper conductors, terminals, and other parts in lap joints for electrical connections where heating must be localized and closely controlled during brazing and where the brazed joint must have low electrical resistance. Generation of heat in the filler metal and nearly complete filling of the joint with a thin layer of the filler metal help meet all these objectives. Normally, the filler metal most frequently used is copperphosphorus, which is used without a flux— especially in brazing of copper. The success of resistance brazing for joining small copper electrical conductors to massive copper assemblies has led to the use of the process for attaching armature leads to commutator bars on large electric motors and generators. Joints made by this method provide a large conducting cross section that prevents signifi-

Chapter 4: Base Metals and Base-Metal Family Groups / 83

cant resistive heating at the connections in service. The joints are made at a much lower temperature than would be possible by resistance welding. The use of conventional resistance welding equipment and high-resistivity electrodes with specially contoured tips makes it possible to concentrate the heating at the joint, to keep the heating time to a minimum, and to obtain efficient handling and comparatively high production rates. Another use of portable resistance welding machines for resistance brazing is in attaching copper bus-bar terminals or similar strip connectors to large electrical equipment that cannot be brought to, or positioned for brazing in, a conventional fixed-position resistance welding machine. Electrical connections to such equipment can often be made more economically by resistance brazing than by mechanical means and are made more readily by resistance brazing than by arc welding. Resistance brazing done with portable resistance welding machines is a convenient and economical way of interconnecting large copper electrical bus bars or of attaching either large or small copper bus bars to motor generators, transformers, and other electrical equipment. The low melting temperature of the filler metal helps to avoid overheating and excessive annealing of the work, and the usual selection of self-fluxing copper-phosphorus filler metals avoids corrosion problems and the need for flux removal. In some applications, filler metal for resistance brazing can be provided in the form of a coating already present on one or both members to be joined, eliminating not only the use of flux but also the operation of placing filler metal at the joint. This is done when copper wire is joined by high-speed resistance brazing to copper terminals clad with copper-phosphorus filler metal. Salt bath furnaces have been used for brazing of copper and copper alloys with silver filler metals. The same neutral salts, operating temperatures, and brazing procedures used for steel are used for dip brazing of most copper alloys. Applications of dip brazing of copper alloys in molten salt include waveguides and waveguide hardware, flowmeter hardware, and capillary tube and bellows assemblies. Copper, Other Metals, and Composites. Copper-steel and steel sections of engine systems produced by high-temperature brazing are

large, two-layer structures consisting of a ribbed wall and a smooth external jacket. They have the form of rotating bodies of complicated geometrical form with cooling circuits operating at high pressures. Therefore, special requirements on strength, corrosion resistance, and efficiency at high and low temperatures are imposed on brazed joints. The design of sections requires the production of brazed joints in combinations of copper and creep-resisting alloys with steels. In producing brazed sections, the filler metals are represented mainly by specially developed copper-manganese alloys of the PM17 type and nickel-manganese alloys of the G70NKh, G40NKh, and so on types, produced in the form of strips and foils 0.1 to 0.3 mm (0.004 to 0.012 in.) thick. Their use ensures strength characteristics and corrosion resistance of brazed joints close to those of the parent material (Ref 44). High-temperature brazing of sections is carried out in vacuum compression systems (VCSs) using induction heating, with automatic control of the temperature-time parameters of brazing and rotation of the brazed section around the horizontal axis. Figure 4.6 shows the diagram of a VCS for brazing sections of engine systems in the horizontal position, with rotation and compression of the surfaces of components during heating for brazing. The system consists of a brazing chamber with an induction heater, a system for loading and unloading sections, a vacuum system, and a system for forced pressurizing and

Fig. 4.6 position

Diagram of a vacuum compression system for brazing sections of engine systems in the horizontal

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cooling. Brazing of engine system sections is carried out at 1000 to 1200 °C (1830 to 2190 °F) (depending on the grade of the alloy used) and with the pressure of the shielding gas up to 490 kPa (71 psi). However, when producing brazed sections with a small cross section of the channels, the use of filler metals in the form of strips, even 0.005 mm (0.0002 in.) thick, often leads to clogging of the channels and unstable hydraulic resistance. In addition, the technology of depositing filler metals in the form of strips on the brazed surfaces by resistance welding is very complicated, laborious, and not always efficient. Therefore, it is more efficient to deposit metallic coatings, which are components of the filler metals, on the surfaces to be brazed. In brazing, as a result of contact interaction of the metal of the coating with the parent material, a liquid phase with a lower melting point than in the brazed material forms and plays the role of the filler metal. This brazing method, the so-called resistance-reactive brazing, is used widely for producing sections where, in brazing, it is necessary to ensure accurate amounts of the filler metal to obtain brazed joints of the required quality. Metallic coatings can be deposited by electroplating or thermovacuum spraying on the surfaces of the brazed components. When brazing copper-steel sections of the engine systems, the main component of the filler metal is manganese, which can also be deposited on the ribbed surface of the copper wall. The optimal conditions of brazing coppersteel sections through a manganese coating include: thickness of the deposited coating layer, 12 to 15 µm; brazing temperature, 1010 ± 10 °C (1850 ± 18 °F); holding time, 10 min; and residual pressure, 6.65 Pa (10–3 psi). The tensile strength of the brazed joints is 200 to 225 MPa (29 to 33 ksi). X-ray spectrum microanalysis of the distribution of elements in the brazed joint shows that the composition of the brazed joint is identical with that of brazed joints obtained when brazing copper-steel sections with copper-manganese filler metals in compact form (strips, foils). Steel sections of engine systems are brazed with alloys where the main components are nickel and manganese, which are deposited in layers on one of the components to be brazed. In heating for brazing, manganese and nickel and melted by resistance-reactive heating, and a liquid phase forms.

The optimal brazing conditions of steel sections using a nickel-manganese coating as the filler metal include: thickness of the layer of the nickel coating deposited on each surface to be brazed, 10 to 15 µm; thickness of the layer of the manganese coating deposited on one of the brazed surfaces, 25 to 30 µm; brazing temperature, 1200 ± 10 °C (2190 ± 18 °F); holding time, 10 min; and residual pressure, 6.65 Pa (10–3 psi). The tensile strength of the brazed joints is 580 to 700 MPa (84 to 102 ksi). Locomotives used in the channel tunnel connecting Britain and France have a rotor made of parallel copper bars held together by an end ring (Fig. 4.7). This ring, which is made of Cr-Cu-Zr material, is brazed at each end to the copper bars. This braze must withstand the stress of rotor speeds up to 6000 rpm. It was essential to develop a brazement that retains a high strength during the brazing process. Flame brazing was the method chosen, because it provided a uniform heat over and around irregular shapes. It was also thought to be better for odd joint configurations. The gas chosen was an air-acetylene mixture. This mixture gave the desired flame characteristics for wraparound, depth of heating, and rate of heating. The number of heat devices varied depending on the size of the rotor being brazed; one setup had 42 and another had 64. Uniformity of heating was attained by rotating the rotor assembly. The part of the rotor that fit into the end ring was designed with a series of rectangular cavities adjacent to the joint area where the filler

Fig. 4.7

Rotors made of copper bars

Chapter 4: Base Metals and Base-Metal Family Groups / 85

metal and flux could be placed. The filler metal was a silver alloy that melted at 620 °C (1150 °F). The foil form was cut into small segments and placed at the joints. The flux was hand brushed onto the braze area and filler metal. Four pilot burners were used to ignite the flames in four quadrants around the ring. Brazing time, on average, was 21/2 to 41/2 min, with the unit rotating at 1 rpm. The rotors were quenched in cold water after brazing. Hardness before brazing was 148 to 154 HV. After brazing, it was 135–143, which was within acceptable limits (Ref 45). The first commercially available brazed copper-brass car and truck radiators have made their debut in the global automotive market. The design innovations and technological advances of copper-brass (CuproBraze) radiators have shown that these radiators are 35 to 40% lower in weight compared to traditional (nonoptimized) copper-brass radiators, because they are designed and manufactured with far less materials in their fins and tubes. Also, the heavy lead-base solder traditionally used in copper-brass radiators is replaced with a lightweight filler metal. Copper-brass radiators can be designed and manufactured to suit the diverse cooling requirements of the automakers of the world. They can be produced using the same furnaces as employed to braze aluminum radiators. Other advantages of copper-brass radiators include the inherent superiority of copper over aluminum in thermal conductivity, internal corrosion resistance, durability, ease of fabrication, longer manufacturing tool life, and simpler joining techniques. In laboratory performance tests, copper-brass radiator prototypes have lasted 8000 h without failure. This equals more than 800,000 km (500,000 miles) of service (Ref 46). Chief among the technologies that provide advantages in the manufacture of copper-brass radiators are no-flux brazing and electrophoretic coating. Newly designed materials and fluxless brazing of radiator components develop superior strength in fin, tube, and header joints and overall strength of copper-brass radiators. They comprise thinner fin and tube material; brazed copper fins are 0.05 mm (0.002 in.) thick or less, while brazed brass tubes are 0.127 mm (0.005 in.) thick. A filler metal of a nontoxic, low-temperature alloy (based on the Cu-Ni-Sn-P system) has been successfully employed. These alloys have a melting temperature as

low as 600 °C (1110 °F) and are self-fluxing in an inert atmosphere, so no rinsing is required after the brazing operation. Reference 47 reports on an investigation into a composite carbon-fiber/copper matrix. This type of composite manifests a number of advantages originating from the properties of the component materials, such as thermal and electrical conductivity, which make it suitable for, for example, resistance welding electrodes and electrical contacts or dilatation bases in semiconductor power elements. In order to make full use of the advantages of this material, it is necessary to join it to ceramics and to metals and alloys used in engineering. The research work examined the joining of carbon-fiber/copper composite to molybdenum and to FeNi42 alloy. These two materials were selected in view of the potential future applications of the resulting joints in the electronics industry, namely, for power diodes and thyristors (joints involving molybdenum) and for electrical contact pieces (joints involving FeNi42). The filler metal used was 72.5Ag-19.5Cu5In-3Ti (wt%). The joints were made using the diffusion-bonding method in a diffusion-bonding press. In the bonding of the volumetric composite carbon fiber/copper to molybdenum, the following process parameters were used: temperature, 950 °C (1740 °F); time, 30 min; pressure, 50 MPa (7 ksi); and vacuum, 2.66 × 10–4 Pa (3 × 10–8 psi). The laminar composites, consisting of carbon fiber in Ag-Cu-In-Ti alloy, were bonded to FeNi42 alloy and to Al2O3 ceramics under the following conditions: temperature, 850 °C (1560 °F); pressure, 0.6 MPa (0.09 ksi); time, 5 min; and vacuum, 2.66 × 10–4 Pa (3 × 10–8 psi). The experiments showed that it was possible to produce joints displaying good strength properties of carbon-fiber/copper composites to molybdenum and of carbon-fiber/AgCu alloy composites to FeNi42 alloy. In the examination of the microstructure, of the linear distribution of elements and of x-ray diffraction patterns, the bond of the composite carbon-fiber/AgCu alloy to molybdenum is likely to be of the diffusion type, which results from diffusion of silver and titanium out of the composite as well as from diffusion of iron and nickel in the opposite direction. In the case of the composite carbonfiber/copper bonding to molybdenum, the joint was formed mainly as an effect of nickel diffusion into molybdenum.

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tion is performed in a flux bath. Furnace brazing offers the greatest potential for magnesium combustion, because the entire structure is heated to the brazing temperature. However, because the M1A alloy, with its high solidus temperature (648 °C, or 1198 °F), is the only magnesium alloy suitable for furnace brazing, the risk of combustion is minimal, provided that the furnace temperature is controlled within the prescribed limits and that brazing time is the minimum necessary to achieve filler-metal flow. Electric or gas heating furnace equipment with automatic temperature controls capable of holding the furnace temperature within ±6 °C (±10 °F) of the brazing temperature for magnesium base metal should be used to minimize incipient melting of the base metal and to reduce the potential of a magnesium fire. The best results in furnace brazing are obtained if dry powdered flux is sprinkled along the joint. Brazing times depend somewhat on the thickness of the materials used and the amount of fixturing necessary to position the parts. Torch brazing is accomplished using a neutral oxyfuel or air-fuel gas flame. Because of the close proximity of the initial melting point (solidus temperature) of the magnesium alloy base metal and the flow point (liquidus temperature) of the filler metal, manual torch brazing is generally preferred, using the AZ125A magnesium filler metal. Natural gas is also well suited for torch brazing, because its relatively low flame temperature reduces the danger of overheating. Dip brazing of magnesium alloys is accomplished by immersing the assembly in a molten brazing flux and holding the parts and fixture at

Magnesium and Magnesium Alloys Brazing techniques similar to those used for aluminum are used for magnesium alloys. Furnace, torch, and dip brazing can be employed, although dip brazing is the most widely used. Magnesium alloys that are considered brazeable, and recommended filler metals, are listed in Table 4.5. Furnace and torch brazing experience has been limited to the M1A alloy. Dip brazing has been used successfully for AZ10A, AZ31B, K1A, M1A, and ZK21A alloys. Other magnesium alloys have not been brazed with present filler metals and techniques developed, due to their low solidus temperatures. The two filler metals AZ92A or BMg-1 (9% Al, 2% Zn, 0.1% Mn, 0.00005% Be, balance Mg) and AZ125A or BMg-2 (12% Al, 0.5% Zn, 0.00005% Be, balance Mg) are both suitable for torch, dip, or furnace brazing. Because the AZ125A filler metal has a lower melting range (Table 4.5), it is usually preferred in most brazing applications. The ease of ignition of magnesium depends to a large extent on the size and shape of the material as well as on the size or intensity of the source of ignition. Fabricated assemblies, ingots, and castings are difficult to ignite, because heat is conducted rapidly away from the source of ignition. The probability of combustion occurring during torch brazing of magnesium is extremely remote, because only the joint area, and not the entire structure, is heated, and because the fluxes used prevent burning in the heated area. The dip brazing process presents no combustion problems with magnesium, because the entire opera-

Table 4.5 Brazeable magnesium alloys and recommended brazing filler metals Solidus ASTM alloy designation

Liquidus

Brazing range

Suitable brazing filler metal

°C

°F

°C

°F

°C

°F

BMg-1 or AZ92A

BMg-2 or AZ125A

632 566 649 648 626

1170 1050 1200 1198 1159

643 627 650 650 642

1190 1160 1202 1202 1187

582–616 582–593 582–616 582–616 582–616

1080–1140 1080–1100 1080–1140 1080–1140 1080–1140

X ... X X X

X X X X X

830 770

599 566

1110 1050

604–616 582–610

1120–1140 1080–1130

... ...

... ...

Base metals AZ10A AZ31B K1A M1A ZK21A

Brazing filler metals AZ92A(a) AZ125A(b)

443 410

(a) Available as wire or rod. (b) Available as wire, rod, strip, or powder

Chapter 4: Base Metals and Base-Metal Family Groups / 87

the desired brazing temperature. The flux serves the dual function of heating and fluxing. The immersion time in the flux bath can vary from 30 to 45 s up to 1 to 3 min. Large assemblies with more metal mass and/or fixturing require the longer times. Because of the large volume of flux and uniform heating, more consistent results are achieved with dip brazing than with other brazing procedures. The corrosion resistance of brazed joints in magnesium alloys depends primarily on the thoroughness of flux removal and the adequacy of joint design to prevent flux entrapment. Because the filler metal is a magnesium-base alloy, the problem of galvanic corrosion is minimized. Protection of magnesium alloys from corrosion is provided by applying appropriate chromate and other coatings. Torch brazing has been used to join hydraulic lift floats, whereas dip brazing has been used to produce battery containers and microwave antennas.

Nickel, Cobalt, and Heat-Resistant Alloys Nickel and the high-nickel alloys may be divided into the following classes: • • • • •

Commercially pure nickel Nickel-copper alloys Nickel-chromium-iron alloys Nickel-chromium-molybdenum alloys Thorium-dispersed nickel alloys

Commercially pure nickel and nickel-copper alloys are used primarily for applications where corrosion resistance is important or where product purity must be maintained. The other categories of alloys also have good corrosion resistance in many media and, in addition, have high strength and oxidation resistance at elevated temperatures. To use these properties, it is important to determine the effects of service conditions on the brazed joint. Nickel and the high-nickel alloys are embrittled by sulfur and low-melting metals, such as zinc, lead, bismuth, and antimony. Nickel and nickel alloy parts should be thoroughly cleaned prior to brazing to ensure the absence of substances that may contain any of these elements. Sulfur and sulfur compounds must also be excluded from the brazing atmosphere. Nickel and its alloys are subject to stress cracking in the presence of molten filler metals.

Parts should, therefore, be annealed prior to brazing to remove residual stresses or be carefully stress relieved during the brazing cycle. In selection of a brazing process and a filler metal for a nickel-base alloy, the characteristics of the alloy must be carefully considered. Nickel alloys differ significantly not only in physical metallurgy (precipitation strengthened versus solid-solution strengthened) but also in process history (cast versus wrought). These characteristics can have a profound effect on brazeability. The filler metals normally used for ferrous metals are suitable for joining nickel and highnickel alloys. It also is important to consider any heat treatments that may be required for the base metal, because the brazed joint must withstand the temperatures involved. In corrosive environments, high-silver filler metals are preferred, while cadmium-free filler metals are chosen to avoid stress-corrosion cracking (SCC). Most high-nickel alloys are capable of being brazed with pure copper filler metal, which is similar to the brazing of carbon and low-alloy steel, except that the copper filler metal characteristically alloys to a greater extent with nickel than with iron. Alloying during brazing makes capillary flow difficult. The copper does not flow far before it has picked up enough nickel to raise its liquidus and reduce its fluidity. To eliminate this problem, the filler metal should be placed as close to the joint as possible, and there should be a sufficient reservoir to fill the joint. Secondly, the assembly should be heated as rapidly as practicable to the brazing temperature. Phosphorus combines with many metals to form brittle compounds known as phosphides. For this reason, the copper-phosphorus filler metals usually are not used with any iron- or nickel-base alloy; however, the BNi-6 and BNi7 filler metals are used for brazing nickel-base alloys. Nickel-base filler metals offer the greatest corrosion and oxidation resistance and elevated-temperature strength. Nickel-base filler metals, containing palladium or platinum, and gold-base, palladiumbase, and platinum-base filler metals have also been used successfully to braze high-nickel alloys. These filler metals generally have good wetting and flow characteristics. They have a low interaction rate with most nickel-base metals and are used to advantage in many special applications where joints are required to have good ductility, high strength, and good oxidation resistance. Filler metals consisting of base

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material plus additions of silicon and boron have been used successfully for brazing the precipitation-strengthened alloys of nickel. When appreciable amounts of aluminum and titanium appear (greater than 1%) in the precipitation-strengthened nickel-base alloys, the oxides of aluminum and titanium are almost impossible to reduce in a controlled atmosphere (vacuum or hydrogen). Therefore, nickel plating or the use of a flux is necessary to obtain a surface that allows wetting by the filler metal. Most commercial brazing processes may be used on nickel and high-nickel alloys. The most common of these are torch, furnace, induction, and resistance brazing; salt bath dip brazing has limited application. Whereas the silver-base filler metals may be used in torch brazing, the copper and nickel filler metals are usually used in controlled-atmosphere brazing. Attempting to braze over the refractory oxides of titanium and aluminum that may be present on precipitation-hardenable nickel-base alloys must be avoided. Procedures to prevent or inhibit the formation of these oxides before and/or during brazing include special treatments of the surface to be joined, and brazing in a highly controlled atmosphere. Surface treatments include electrolytic nickel plating and reduction of the oxides to metallic form. As stated earlier, a typical practice is to nickel plate the joint surfaces of any alloy that contains aluminum and/or titanium. For vacuum brazing, when aluminum and titanium are present in trace amounts, plating 0.01 mm (0.0004 in.) thick is considered optional. Alloys with up to 4% Al and/or Ti require plating 0.015 mm (0.0006 in.) thick, whereas alloys with aluminum and/or titanium contents greater than 4% require plating 0.02 to 0.03 mm (0.0008 to 0.0012 in.) thick. When brazing is done in a pure dry hydrogen atmosphere, thicker plating (0.03 to 0.04 mm, or 0.0012 to 0.0016 in.) is desirable for alloys with greater than 4% Al and/or Ti contents. Brazing of base materials containing more than a few percent of aluminum, titanium, zirconium, or other elements that form very stable oxides requires vacuums of 0.13 Pa (2 × 10–5 psi) or lower. Vacuum furnaces for this type of brazing usually employ a diffusion pump backed by a mechanical pump. Consideration must be given to the effect of the brazing thermal cycle on the base metal. Filler metals that are suitable for brazing nickelbase alloys may require relatively high thermal cycles. This is particularly true for the filler-

metal alloy systems most frequently used in brazing nickel-base alloys—the Ni-Cr-Si and Ni-Cr-B systems. It is significant to point out the effects of thermal cycles on solid-solution-strengthened and precipitation-strengthened nickel-base alloys. An example of the former is Inconel 600, a high-temperature nickel-base alloy containing 15.5% Cr, 76% Ni-Co, and 8% Fe that resists oxidation up to 1175 °C (2150 °F) and is used in furnace parts and fixtures, heat exchangers, chemical-handling equipment, and turbine and reactor parts. This alloy may not be adversely affected by nickel-base filler-metal brazing temperatures of 1010 to 1230 °C (1850 to 2250 °F). An example of the latter is Inconel 718, a nickel-base superalloy containing 50 to 55% Ni, 17 to 21% Cr, 4.75 to 5.5% Nb and Ta, 1% Co, approximately 0.20% Fe, and greater than 2% Al and Ti. This alloy is oxidation resistant up to 980 °C (1800 °F) and is used in aircraft turbine parts, pumps, and rocket motors. This alloy may, however, display adverse property effects when exposed to brazing cycles higher than its normal solution heat treatment temperature. Inconel 718, for example, is solution heat treated at 955 °C (1750 °C) for optimal stressrupture life and ductility. Brazing temperatures of 1010 °C (1850 °F) or above result in grain growth and an attendant decrease in stress-rupture properties, which cannot be recovered by subsequent heat treatment. The applicability of the transient liquid phase (TLP) metallic bonding method for joining finegrained Inconel 718SPF superalloy sheets by inserting a nickel-phosphorus or a Ni-Cr-P amorphous interlayer has shown that a joint with uniform chemical composition could be obtained for the Inconel 718SPF superalloy with a nickel-phosphorus interlayer at 1100 °C (2010 °F) for 8 h. When a Ni-Cr-P interlayer was used under the same metallic bonding conditions, the concentrations of nickel, iron, and niobium in the bond region and in the base metal had a difference of more than 2 wt%. This means that longer bonding time was required to homogenize the chemical compositions of bonds with a Ni-Cr-P interlayer. The shear strength of the TLP bonds with a nickel-phosphorus interlayer indicated higher bond strengths than did that with the Ni-Cr-P interlayer under the same bonding conditions. The second phases, which existed at the grain boundaries, retarded grain growth in the bonding region. Homogenization during the TLP metallic bonding process was thus limited (Ref 48).

Chapter 4: Base Metals and Base-Metal Family Groups / 89

Consideration of base-metal property requirements for service enables selection of an appropriate filler metal. For lower melting temperatures (below 1040 °C, or 1900 °F), filler metals are available within the nickel-base filler-metal family and within other filler-metal systems (Tables 4.6, 4.7) (Ref 49–51). Brazing is one of the preferred methods for joining dispersion-strengthened nickel alloys that must function at elevated temperatures. High-strength brazements have been made with special nickel-base filler metals tested up to 1315 °C (2400 °F). Procedures have been developed for brazing of thoria-dispersed nickel and thoria-dispersed NiCr foils for high-temperature service using the candidate filler metals shown in Table 4.8 (Ref 49, 51). Heat-resistant alloys are suitable for use under moderate to high loading in the temperature range from 540 to 1100 °C (1000 to 2010 °F). These metals are complex austenitic alloys based on nickel, cobalt, or both. They have often been termed superalloys. Their greatest use is in the construction of gas turbine engines and hot airframe components. Heat-resistant alloys are generally brazed in hydrogen-atmosphere or high-vacuum furnaces with nickel-base or special filler metals. Because the brazing temperatures are high, the effect of the brazing thermal cycle on the base metals should be taken into account. The nonheat-treatable alloys suffer moderate strength losses due to grain growth during brazing. Coldworked alloys should not be brazed unless the severe loss in strength from annealing during brazing is considered in the design. Brazing of cobalt-base alloys is readily accomplished by the same techniques used for nickel-base alloys. Because most of the popular cobalt-base alloys do not contain appreciable amounts of aluminum or titanium, brazingatmosphere requirements are less stringent. These materials can be brazed in either a hydrogen atmosphere or a vacuum. Filler metals are usually nickel- or cobalt-base alloys or goldpalladium compositions. Silver or copper filler metals may not have sufficient strength and oxidation resistance in many high-temperature applications. Although cobalt-base alloys do not contain appreciable amounts of aluminum or titanium, an electroplate or flash of nickel is often used to promote better wetting of the filler metal. Nickel-base filler metals such as BNi-3 (Table 4.7) have been used successfully on cobalt-base alloys (Ref 51, 52) for honeycomb

structures. After brazing, a diffusion cycle was used to raise the brazed-joint remelt temperature to 1260 to 1315 °C (2300 to 2400 °F). Additionally, it may be desirable to heat treat the brazement after brazing to attain optimal base-metal properties. The BCo-1 filler metal (Table 4.7) appears to offer a good combination of strength, oxidation resistance, and remelt temperature for use on cobalt-base alloy foil (Haynes 25) (Ref 53). Cobalt alloys, much like nickel alloys, can be subject to liquid metal embrittlement or SCC when brazed under residual or dynamic stresses. This frequently is observed when silver or silver-copper filler metals are used. Liquid metal embrittlement of cobalt-base alloys by copper filler metals occurs with or without the applications of stress; therefore, copper filler metals should be avoided in brazing of cobalt-base alloys. Superalloys can be subdivided into two categories: conventional cast and wrought alloys, and powder metallurgy (P/M) products. Powder metallurgy products may be produced in conventional alloy compositions and oxide-dispersion-strengthened (ODS) alloys. Oxide-dispersion-strengthened alloys are P/M alloys that contain stable oxide evenly distributed throughout the matrix. The oxide does not go into the solution in the alloy even at the liquidus temperature of the matrix. However, the oxide is usually rejected from the matrix on melting of the matrix, which occurs during fusion welding, and cannot be redistributed in the matrix on solidification; therefore, these alloys are usually joined by brazing. There are two commercial alloy classes of ODS alloys: the dispersion-strengthened nickel-chromium and dispersion-strengthened nickel, discussed previously, and the mechanically alloyed (MA) alloys. An example is Inconel MA 754, a turbine vane material made from powder. The material has extremely good creep and oxidation resistance at temperatures up to 1050 °C (1920 °F) and is used in gas turbines (vanes, nozzle, burners, combustors), chemical plants (liners, heat exchangers, combustion chambers), and so on. They owe their oxidation resistance to a stable and tenacious oxide film, and their creep resistance to the combination of a large and directional grain structure plus a dispersion of submicron-sized particles consisting chiefly of yttrium oxide. There is no joining method that can preserve the coarse and elongated grain structure of the recrystallized parent metal across a joint line

90 / Brazing, Second Edition

Table 4.6 Commercially available nickel-base filler metals Solidus (melting point)

Chemical composition, % Ni

99.26 98.88 97.25 96.00 95.50 95.20 94.60 93.72 93.60 92.80 92.60 92.54 92.40 92.30 91.70 91.60 91.30 91.22 90.80 90.66 88.95 88.80 88.40 87.10 85.90 84.39 84.08 83.40 82.97 82.50 82.10 81.50 81.50 81.10 80.85 79.25 79.20 78.87 78.65 78.40 77.00 76.80 76.30 75.60 74.85 74.60 74.40 73.90 73.62 73.40 72.44 72.20 72.00 71.90 70.00 69.20 68.59 68.10 67.00 66.85 65.50 65.00 63.00 61.95 61.50 61.40 60.35 59.15 57.10

Liquidus (flow point)

Cr

C

P

Fe

B

Si

Co

Mn

Cu

Other

°C

°F

°C

°F

0.04 ... ... ... ... ... ... ... 2.20 2.30 2.70 ... 2.90 2.25 2.30 2.50 2.70 ... 3.00 0.20 5.00 3.80 5.60 5.00 5.60 9.00 5.50 8.00 6.50 11.50 7.00 11.40 15.00 11.50 10.00 11.00 10.00 13.30 13.00 13.00 15.00 15.20 7.00 13.50 ... ... 15.00 14.80 17.10 17.00 15.00 ... ... 15.00 16.00 17.50 18.00 10.00 20.00 19.00 ... ... ... 12.00 19.00 19.50 11.50 19.50 3.50

0.04 0.015 ... ... ... ... 0.03 0.03 ... ... ... 0.06 ... ... ... ... ... 0.03 ... 0.14 0.25 ... ... ... ... 0.35 0.12 ... 0.03 ... 0.03 ... ... ... ... 0.50 0.35 ... ... 0.70 ... ... ... 0.15 0.15 ... 0.10 0.70 ... ... 0.06 ... 0.03 0.60 ... 0.80 ... 0.40 ... 0.15 0.01 0.07 ... 0.55 ... ... 0.55 ... ...

0.30 1.10 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 4.00 ... ... 1.50 6.00 ... ... ... ... ... ... 7.50 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ...

... ... ... ... ... ... ... 1.25 1.00 1.00 1.10 ... 1.30 1.00 1.20 1.10 1.20 1.25 1.30 1.50 1.00 1.70 2.40 0.60 2.40 2.10 2.50 0.40 3.00 ... 3.00 ... ... ... 0.25 3.50 4.00 ... ... 3.50 ... ... 3.00 3.50 ... ... 3.00 3.80 ... ... 4.00 ... ... 4.00 2.00 4.50 4.50 3.50 3.00 3.00 ... ... 5.00 3.50 4.50 ... 3.40 2.25 1.00

... ... 0.75 1.50 1.50 1.80 1.90 1.50 1.60 1.50 1.90 2.90 1.90 1.70 1.80 1.80 1.80 3.00 1.90 3.00 1.80 2.20 2.80 1.20 2.50 1.66 2.50 0.80 3.00 ... 2.90 0.30 3.50 0.40 0.50 2.25 2.10 0.23 0.35 2.90 ... ... 3.20 3.25 ... 3.15 3.00 3.00 0.08 0.10 3.50 3.30 3.60 3.50 3.50 3.50 ... 2.50 ... ... ... ... 2.00 2.50 ... ... 2.50 ... 0.90

... ... 2.00 2.50 3.00 3.00 3.50 3.50 2.60 2.40 2.80 4.50 2.80 2.75 3.00 3.00 3.00 4.50 3.00 4.50 3.00 3.50 3.20 2.10 3.60 2.50 3.80 1.40 4.50 6.00 5.00 6.80 ... 7.00 0.90 3.50 4.35 7.60 8.00 1.50 8.00 8.00 4.50 4.00 8.00 4.25 4.50 3.80 9.20 9.50 5.00 4.50 4.50 5.00 3.50 4.50 9.00 3.50 10.00 10.00 7.00 7.00 ... 3.50 9.80 9.60 3.20 9.50 2.50

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 18.00 ... ... ... ... ... 20.00 20.00 ... ... ... ... ... ... 0.50 ... ... ... ... 5.20 ... 2.50 ... ...

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 17.00 ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.50 23.00 23.00 ... ... ... 9.50 ... 9.50 35.00

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 2.50 ... ... ... ... ... 4.50 5.00 ... ... ... ... ... ... ...

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 6.00 W ... ... ... ... ... ... ... ... ... ... ... 2.50 Mo ... ... 12.00 W ... ... ... ... 30.00 Mo 16.00 W ... ... 16.00 W ... ...

1082 1038 1021 1010 1010 1004 982 982 954 954 954 982 954 954 954 954 954 954 954 982 971 954 954 954 954 1010 982 982 971 1082 1038 1082 1054 1082 954 971 971 1082 1082 1082 1082 1082 977 976 1024 982 971 977 1082 1082 982 982 982 977 1093 949 1082 971 1079 1082 982 982 1066 971 1082 1082 1010 1082 987

1980 1900 1870 1850 1850 1840 1800 1800 1750 1750 1750 1800 1750 1750 1750 1750 1750 1750 1750 1800 1780 1750 1750 1750 1750 1850 1800 1800 1780 1980 1900 1980 1930 1980 1750 1780 1780 1980 1980 1980 1980 1980 1790 1789 1875 1800 1780 1790 1980 1980 1800 1800 1800 1790 2000 1740 1980 1780 1975 1980 1800 1800 1950 1780 1980 1980 1850 1980 1800

1121 1204 1066 1066 1066 1066 1066 1066 1038 1038 1038 1024 1038 1038 1038 1038 1038 1024 1038 1038 1182 1038 1038 1038 1038 1066 1038 1049 999 1138 1065 1138 1054 1138 1038 1160 1149 1138 1138 1138 1138 1138 1038 1052 1065 1024 1077 1038 1138 1138 1024 1024 1024 1052 1149 1065 1138 1093 1135 1138 1010 1010 1138 1104 1138 1107 1121 1121 1066

2050 2200 1950 1950 1950 1950 1950 1950 1900 1900 1900 1875 1900 1900 1900 1900 1900 1875 1900 1900 2160 1900 1900 1900 1900 1950 1900 1920 1830 2080 1950 2080 1930 2080 1900 2120 2100 2080 2080 2080 2080 2080 1900 1925 1950 1875 1970 1900 2080 2080 1875 1875 1875 1925 2100 1950 2080 2000 2075 2080 1850 1850 2080 2020 2080 2025 2050 2050 1950

(continued)

Chapter 4: Base Metals and Base-Metal Family Groups / 91

Table 4.6 (continued) Solidus (melting point)

Chemical composition, %

Liquidus (flow point)

Ni

Cr

C

P

Fe

B

Si

Co

Mn

Cu

Other

°C

°F

°C

°F

50.20 50.10 36.00 17.00 17.00 16.00 12.00 9.00 9.00 9.00 5.00 2.00

4.00 ... ... 21.00 19.00 ... 18.00 ... ... ... ... 16.00

... ... ... 0.80 0.40 ... ... 0.03 ... 0.03 ... 0.20

... 3.50 ... ... ... ... 0.10 ... ... ... ... ...

... 30.00 ... ... ... ... 66.90 ... ... ... ... 81.80

0.80 ... ... 3.25 0.80 1.00 ... ... ... ... ... ...

... 11.00 ... 3.00 8.00 ... 1.00 ... ... ... ... ...

... ... ... 44.95 50.80 16.00 ... ... ... ... 10.00 ...

45.00 ... ... ... ... 67.00 2.00 23.50 38.50 37.00 22.00 ...

... ... 59.00 ... ... ... ... 67.50 52.50 52.00 63.00 ...

... 5.40 Mo 5.00 In 10.00 W 4.00 W ... ... ... ... ... ... ...

996 954 1038 1038 1107 1004 1121 910 880 871 943 1149

1825 1750 1900 1900 2025 1840 2050 1670 1615 1600 1730 2100

1079 1001 1204 1121 1149 1021 1163 932 927 927 957 1204

1975 1835 2200 2050 2100 1870 2125 1710 1700 1700 1755 2200

(Ref 54, 55). However, some sacrifice of the maximum properties may be allowable if such a joint were to be sited outside the area of maximum stress; alternatively, a joint parallel to the major stresses might be acceptable in certain cases. Although brazing introduces a change in composition across the joint line, and although it does melt some of the parent metal and thus causes agglomeration of the dispersoid, nevertheless, these effects can be minimized by a suitable choice of composition and by close control of brazing and postbrazing conditions. Furthermore, brazing does not introduce mechanical distortion, and it can be carried out at temperatures well below the recrystallization range. Inconel MA 754 is one of the easiest alloys to braze in the family of ODS alloys. Vacuum, hydrogen, or inert atmospheres can be used for brazing. Prebraze cleaning consists of grinding or machining the faying surfaces and washing with a solvent that evaporates without leaving a residue. Generally, brazing temperatures should not exceed 1315 °C (2400 °F) unless demanded by a specific application that has been well examined and tested. The filler metals for use with these ODS alloys usually are not classified (Table 4.8); in most cases, they have brazing temperatures in excess of 1230 °C (2250 °F). These include proprietary filler metals containing nickel, cobalt, gold, or palladium. A series of filler-metal evaluation tests (Ref 55) found that brazing fine-grained MA 754 at 1200 °C (2190 °F) by induction heating for less than 30 s with a thin β-free coating equivalent to BNi-5 gave joints of excellent appearance and without recrystallization.

A second nickel-base ODS alloy is Inconel MA 6000. Like Inconel MA 754, it is γ strengthened. Inconel MA 6000 has a solidus temperature of 1300 °C (2370 °F); therefore, the brazing temperature should be no higher than 1250 °C (2280 °F). Nickel and cobalt filler metals have been used to join this alloy. Other mechanically alloyed ODS alloys include iron-base creepresistant MA 956, which was also recently examined (Ref 56). Although it was possible to produce a reasonable brazed joint in MA 956 by means of sputter-coated brazes based on the nickel filler metal BNi-la (72.5Ni-16.5Cr-4Fe4Si-3B), an agglomeration of the dispersed oxide at the braze interface was unavoidable. Furthermore, as in the brazing of nickel-base ODS alloys, the presence of boron in the braze causes local recrystallization of the parent material. This has led to future work whereby diffusion bonding was to be examined as the only joining process that did not require either a foreign material or a melted or disrupted region in the joint. Diffusion bonding therefore offered the ability to form a bond in an iron-ODS alloy in which the massive grain structure continued unbroken across the bond. To decrease the completion time during TLP bonding, an iron-base filler metal, comprising MA 956 base metal containing 7% Si and 1% B (wt%), was developed to join MA 956 base material. Transient-liquid-phase-bonded joints free of microvoids and bond line intermetallic phases were obtained using a bonding temperature of 1563 K, a holding time of 2 to 16 ks, and an applied pressure of 7.0 MPa (1 ksi). The bond line region in the TLP-bonded MA 956 base metal had a bamboolike microstructure,

Cr

B

Cr

37.0–38.0 79.5–80.5 34.5–35.5 81.5–82.5 29.5–30.5

Fe

C

P

Ni

18.0–20.0 16.0–18.0

7.5–8.5

Si

bal bal bal ... ...

Cu

3.5–4.5

W

Pd

1.0

Fe

... ... ... ... 33.5–34.5

Composition, wt %

... ... 2.5–3.5 bal 35.5–36.5

Ni

0.02 0.02 0.02 0.02 0.02 0.02 0.02 0.02 0.02 0.02

S

C

P

0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05

Al

0.7–0.9 0.35–0.45 0.02

B

Composition, wt %

(a) If determined, cobalt is 0.1% maximum unless otherwise specified. Source: AWS A5.8

BCo-1

Si

2.75–3.50 4.0–5.0 4.0–5.0 0.6–0.9 0.02 2.75–3.50 4.0–5.0 4.0–5.0 0.06 0.02 2.75–3.50 4.0–5.0 2.5–3.5 0.06 0.02 2.75–3.50 4.0–5.0 0.5 0.06 0.02 1.5–2.2 3.0–4.0 1.5 0.06 0.02 0.03 9.75–10.50 . . . 0.10 0.02 ... ... ... 0.10 10.0–12.0 0.01 0.10 0.2 0.08 9.7–10.5 ... 6.0–8.0 ... 0.10 0.02 3.25–4.0 ... 1.5 0.06 0.02

Cobalt-base alloy filler metals

AWS classification

BAu-1 BAu-2 BAu-3 BAu-4 BAu-5

Au

13.0–15.0 13.0–15.0 6.0–8.0 ... ... 18.5–19.5 ... 13.0–15.0 ... 13.5–16.5

Precious metals

AWS classification

BNi-1 BNi-1a BNi-2 BNi-3 BNi-4 BNi-5 BNi-6 BNi-7 BNi-8 BNi-9

Nickel-base alloy filler metals(a)

AWS classification

Composition, wt %

Mn

Cu

0.02

S

0.15 0.15 0.15 0.15 0.15

Other elements total

0.05

Al

990 890 975 950 1135

°C

0.05

Ti

Zr

0.05

Zr

°F

bal bal bal bal bal bal bal bal bal bal

Ni

bal

Co

1815 1635 1785 1740 2075

0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05

Solidus

0.05 ... ... 0.05 ... ... 0.05 ... ... 0.05 ... ... 0.05 ... ... 0.05 ... ... 0.05 ... ... 0.05 0.04 ... 0.05 21.5–24.5 4.0–5.0 0.05 ... ...

Ti

Table 4.7 American Welding Society (AWS) brazing alloys for elevated-temperature service

0.50

°C

1120

°C

2050

°F

Solidus

Liquidus

1790 1790 1780 1800 1800 1975 1610 1630 1800 1930

°F

Solidus °C

975 975 970 980 980 1080 875 890 980 1055

1015 890 1030 950 1165

Other elements total

0.50 0.50 0.50 0.50 0.50 0.50 0.50 0.50 0.50 0.50

Other elements total

1790 1970 1830 1900 1950 2075 1610 1630 1850 1930

1150

2100

°F

Liquidus °C

1860 1635 1885 1740 2130

°F

975 1075 1000 1040 1065 1135 875 890 1010 1055

°F

Liquidus °C

1860–2000 1635–1850 1885–1995 1740–1840 2130–2250

°F

1150–1230

°C

2100–2250

°F

Brazing range

1015–1095 890–1010 1030–1090 950–1005 1165–1230

°C

Brazing range

1950–2200 1970–2200 1850–2150 1850–2150 1850–2150 2100–2200 1700–2000 1700–2000 1850–2000 1950–2200

°F

Brazing range

1065–1205 1075–1205 1010–1175 1010–1175 1010–1175 1150–1205 925–1095 925–1095 1010–1095 1065–1205

°C

92 / Brazing, Second Edition

Chapter 4: Base Metals and Base-Metal Family Groups / 93

which was considered to be peculiar to bonded ODS alloys. For this reason, the growth mechanism of the solid phase during isothermal solidification was evaluated. During tensile testing at 923 K, the joints TLP bonded at 1563 K for 2 to 16 ks fractured in the base-metal zone, and consequently, the mechanical properties of the joint region and the base metal were similar. The creep-rupture properties of the joint regions were close to the base-metal properties in the transverse direction. In order to produce complex and high-temperature-resistant components, diffusion bonding/brazing has been investigated. The joining of a new iron-base ODS alloy, PM2000, was developed (Ref 57). Oxide-dispersion-strengthened superalloys are very promising metallic materials for different high-temperature-exposed components. The ferritic ODS material is especially interesting and suitable for high-temperature-exposed stationary components in turbine and diesel engine constructions. The combination of excellent oxidation and corrosion resistance and the high melting point (1480 °C, or 2700 °F) of the FeCr-Al alloy with a Y2O3-dispersion-strengthened matrix permits higher operating temperatures for the iron-base ODS alloy than ever attainable with the nickel-base superalloys. In iron-base Fe-Cr-Al ODS alloys, the Y2O3 dispersoid does not coarsen and coagulate during the joining. At room temperature, a bending strength of approximately 70% of base-material strength is attainable. A secondary recrystallization sometimes occurs in the bonding zone. Generally, the bond interface is formed as a straight grain boundary, which can limit the strength of the bond. It is very important to develop high-quality electric power systems that have excellent

energy efficiency. With this in mind, electric power systems combining gas and steam turbines must be improved to increase the energy efficiency. A key requirement for land gas turbines is that higher inlet gas temperatures can be applied. Single-crystal base material is one of the most suitable candidate materials for highquality land gas turbines. Nickel-base singlecrystal superalloys possess the excellent mechanical properties at elevated temperatures (because the melting temperature is increased when grain-boundary strengthening elements are not present). The elevated-temperature properties of single-crystal base material deteriorate when grain boundaries are formed in the joint region. Therefore, it is vital that singlecrystal base material is bonded without creating such a problem. Bonding and the crystallization behavior of a nickel-base single-crystal superalloy base material, CMSX-2, was investigated during TLP bonding using MBF-80 insert metal (Ref 58). Joint strength was evaluated using tensile and creep-rupture testing at elevated temperature. The TLP bonding of CMSX-2 was carried out at 1373 to 1548 K for 0 to 19.6 ks in vacuum (Table 4.9). Some of the conclusions include: • The dissolution width of the base metal increased when the bonding temperature and holding time increased. Base-metal dissolution was consistent with a derivation based on the Nernst-Brunner theory. • Material grew epitaxially into the liquid phase from substrates during the isothermal solidification, and single-crystal properties were maintained in completed joints. • The elevated-temperature tensile properties of completed joints were almost identical to those of base metal. The creep-rupture

Table 4.8 Filler metals for brazing thoria-dispersed (TD) nickel and TD-NiCr alloys Brazing temperature Base metal

TD-Ni

TD-NiCr

Source: Ref 51

Brazing filler metal

°C

°F

Nominal composition

TD-20 TD-6 J8600 Ni-Pd Td-6 CM50 NX77 NSB

1302 1302 1177 1246 1302 1066–1121 1177–1191 1288

2375 2375 2150 2275 2375 1950–2050 2150–2175 2350

Ni-16Cr-25Mo-4Si-5W Ni-22Cr-17Mo-4Si-5W Ni-33Cr-25Pd-4Si Ni-60Pd Ni-22Cr-17Mo-4Si-5W Ni-3.5Si-1.9B Ni-5Cr-7Si-1B-1W-4Co Ni-2Si-0.8B

94 / Brazing, Second Edition

strength and rupture lives of completed joints were also almost identical to those of the base metal. Joint fracture always occurred in the base material (Fig. 4.8).

Molybdenum, Niobium, Tantalum, Tungsten, and Their Alloys The refractory metals have a history of use in applications where structural integrity at high temperatures is required. For example, the filaments for incandescent lamps are made from tungsten. This metal is used extensively in vacuum tubes for cathode structures. Similarly, molybdenum is used for grids and supporting members in vacuum tubes and for lead wires in incandescent lamps. In current applications of the refractory metals in the aircraft, space, nuclear, and electronics industries, the joints must be of significant dimensions, with properties that are adequate to meet exacting requirements of strength, ductility, corrosion resistance, and so on. Because of its limited effect on base-metal properties, brazing has been accepted widely as a method of joining assemblies fabricated from refractory metals. The characteristics that affect the brazeability of refractory metals include ductile-to-brittle transition behavior, recrystallization temperature, and reactivity with oxygen, nitrogen, hydrogen, and carbon. The refractory metals include those metals that have melting points in excess of 2205 °C (4000 °F), that is, niobium, iridium, molybdenum, tantalum, ruthenium, tungsten, osmium, and rhenium. Sometimes vanadium, hafnium, rhodium, and chromium are included in this group of metals. The most important refractory metals, from a structural standpoint, are molybdenum, tantalum, niobium, and tungsten. Niobium, molybdenum, tantalum, and tungsten have much in common; however, there are differences that have a bearing on the manner in which these metals are used and joined. All of the refractory metals have a body-centered cubic crystal structure, with very high melting

temperatures, high to very high densities, low specific heats, and low coefficients of thermal expansion. Mechanical strength and structural integrity at high temperatures are excellent. The mechanical properties of refractory metals are affected markedly by ductile-to-brittle transition behavior, recrystallization temperature, and reactions with carbon and various gases. These characteristics must be considered when brazing procedures are established. Refractory metals, because of their body-centered cubic structure, have a well-defined transition from ductile to brittle behavior. With these metals, a large decrease in energy absorption occurs over a narrow temperature range, and an associated change in the type of fracture from tough to brittle is also evident. This temperature range is not a fixed property of the metal, because it varies with strain rate, alloying addition, impurities, heat treatment, and method of fabrication. The transition-temperature ranges for the pure refractory metals are as follows: • • • •

Niobium, –200 to –75 °C (–330 to –105 °F) Molybdenum, 150 to 260 °C (300 to 500 °F) Tantalum, <–195 °C (–320 °F) Tungsten, 260 to 370 °C (500 to 700 °F)

Therefore, molybdenum and tungsten are brittle at room temperature and must be handled carefully to avoid damage. Also, these metals must be brazed in a stress-free condition. The strength and ductility of refractory metals are adversely affected by microstructural changes that occur when the recrystallization temperatures of these metals are exceeded. Recrystallization-temperature range also varies with alloying additions, interstitial content, fabrication method (including degree of cold working), and time at temperature. Recrystallization-temperature ranges for unalloyed refractory metals are: • Niobium, 985 to 1150 °C (1805 to 2100 °F) • Molybdenum, 1150 to 1200 °C (2100 to 2190 °F)

Table 4.9 Composition of materials used in brazing a nickel-base single-crystal superalloy Composition, mass%

Base metal: CMSX-2 Insert metal: MBF-80 Source: Ref 58

Ni

B

Cr

Co

Mo

W

Ti

Al

Ta

bal bal

... 3.7

8.0 15.5

4.6 ...

0.6 ...

8.0 ...

1.0 ...

5.6 ...

6.0 ...

Chapter 4: Base Metals and Base-Metal Family Groups / 95

• Tantalum, 1100 to 1400 °C (2010 to 2550 °F) • Tungsten, 1200 to 1650 °C (2190 to 3000 °F) Some applications permit brazing with filler metals that melt below the recrystallizationtemperature range. Other applications require the use of filler metals that melt above this temperature range. The joint also must be designed to accommodate the loss in mechanical properties associated with recrystallization. Research has improved the high-temperature mechanical properties of the refractory metals and has led to an increase in recrystallization temperatures through alloying. For example, titanium, zirconium, and hafnium can be used to strengthen molybdenum and increase its recrystallization temperature. The recrystallization temperature of unalloyed molybdenum is approximately 1150 to 1200 °C (2100 to 2190 °F). In contrast, the recrystallization temperature of Mo-0.5Ti-0.07Zr molybdenum alloy (TZM) is approximately 1480 °C (2700 °F), and the stressrupture strength of this alloy at 980 to 1095 °C (1800 to 2000 °F) is several times that of unalloyed molybdenum. The third consideration is reactions with gases and carbon. The environment in which refractory metals are brazed is determined by the reactivity of these metals with oxygen, hydrogen, and nitrogen and the effects of these elements on mechanical properties. Tantalum and niobium are embrittled by the presence of hydrogen at relatively low temperatures. In contrast, molybdenum and tungsten

Stress (σ), MPa (ksi)

1000 (145) 900 (131) 800 (116) 700 (102) 600 (87)

Base metal CMSX-2 Insert metal MBF-80

Joint 1523 x 1.8ks (2.3MPa)

500 (73) 400 (58) 300 (44)

200 (30) Base metal

100 (15) 23

24

25

26

27

28

29

30

31

Larson–Miller parameter, T(20+Log t)x10–3

Fig. 4.8

Creep-rupture strength of superalloy CMSX-2 joints. Source: Ref 58

can be brazed in a hydrogen atmosphere. Molybdenum, tantalum, and niobium are embrittled by nitrogen at high temperatures. All of the refractory metals form carbides in the presence of minute quantities of carbon and its compounds. Based on the various characteristics of refractory metals, the following principles are applicable to brazing of refractory metals: • If maximum joint strength is required, refractory metals must be brazed at temperatures below those at which recrystallization occurs. • However, brazing at much higher temperatures may be necessitated by service requirements, and some decrease in joint properties must be anticipated. • All of the refractory metals can be brazed in a vacuum or in an argon or helium atmosphere with a very low dewpoint. • Graphite fixturing should not be used to position refractory metal parts during brazing, because these metals readily form carbides. However, graphite tooling may be acceptable if coated with a refractory material. • Ceramics can also be used for fixturing, but care in their selection must be exercised. Some ceramics cannot be used in a vacuum because of outgassing, whereas others react with refractory metals at high temperatures. • Refractory metals with higher melting temperatures than the one being brazed also can be used for fixturing. The filler metals used for brazing of refractory metals must be selected on the basis of the service condition, the specific application, and their compatibility with the base metal and coating (if a coating is used) (Table 4.10). Molybdenum and Its Alloys. Copper- and silver-base filler metals can be used to braze molybdenum for low-temperature service. For high-temperature applications, molybdenum can be brazed with gold, palladium, and platinum filler metals, nickel-base filler metals, reactive metals, and refractory metals that melt at lower temperatures than molybdenum. Nickel-base filler metals have limited applicability for high-temperature service, because nickel and molybdenum form a low-melting eutectic at approximately 1315 °C (2400 °F). There are three basic limits on brazing of molybdenum and its alloys for use above 980 °C (1800 °F):

96 / Brazing, Second Edition

• Recrystallization of molybdenum • Formation of intermetallics between refractory metals and filler metals • Relative weakness of filler metals at elevated temperatures The formation of intermetallics between molybdenum and filler metals is detrimental to joint soundness, because the intermetallics become brittle and may fracture at relatively low loads when the joint is stressed. The relative weakness of filler metals at elevated temperatures limits the use of brazed molybdenum assemblies. Most of the nickelbase elevated-temperature filler metals melt between 980 and 1150 °C (1800 and 2100 °F), where the superior elevated-temperature strength of molybdenum begins to manifest itself. Two binary filler metals, V-35Nb and Ti30V, have been evaluated for use with the Mo0.5Ti molybdenum base metal. Tee-joints were brazed in a vacuum for 5 min at temperatures of 1650 °C (3000 °F) for the Ti-30V filler metal and 1870 °C (3400 °F) for the V-35Nb filler metal. The filler metals had excellent metallurgical compatibility with the molybdenum base metal, and minimum erosion of the base metal occurred during brazing. Brazements made on Mo-0.5Ti alloy with the diffusion-sink filler metal Ti-8.5Si, which melts at approximately 1330 °C (2425 °F), exhibited excellent filleting and wetting, joint ductility, and freedom from cracks. Specimens with molybdenum powder added to the filler-metal powder were also evaluated and brazed at 1400 °C (2550 °F) (Ref 59). Because of its excellent high-temperature properties and compatibility with certain environments, molybdenum is a prime candidate for use in isotopic power systems for components of nuclear reactors and chemical-processing

systems, and the best filler-metal base on overall performance has been Fe-15Mo-5Ge-4C-1B (Ref 60, 61). The molybdenum alloy TZM (0.5Ti-0.08ZrMo) has also been brazed successfully at 1400 °C (2550 °F), with molybdenum powder added to the Ti-8.5Si filler metal (Ref 59, 62); other filler metals also can be used (Ref 63). A Ti25Cr-13Ni filler metal with a brazing temperature of 1260 °C (2300 °F) has produced the highest remelt temperature on TZM. Remelt investigations conducted to develop vacuum brazements of molybdenum and tungsten that can be used in seal-joint applications up to 1600 °C (2910 °F) were completed with the following filler metals: • • • • • •

Ti-65 wire Vanadium wire MoB-50MoC powder mixture V-50Mo powder mixture Mo-15MoB2 powder mixture Mo-49V-15MoB2 powder mixture

Brazing temperatures ranged from 1625 to 2255 °C (2960 to 4095 °F). Molybdenum joints made with Ti-65V, pure vanadium wire, V50Mo, and MoB-50MoC were as strong, or stronger, than the base metal at elevated temperature. The most resistant joints tested were brazed with Ti-65V, V-50Mo, and MoB50MoC. Without further testing, it probably would be best to limit the maximum service temperature to 1395 °C (2545 °F). A more recently completed wettability study (Ref 64) found that: • Filler metals generally wet molybdenum significantly better than they wet TZM. This is seen in Tables 4.11 and 4.12. The wettability index shown for each filler metal indicates that indexes >0.05 are indicative of good per-

Table 4.10 Filler metals for brazing of graphite, ceramics, and refractory and reactive metals Chemical composition, %

Solidus (melting point)

Liquidus (flow point)

Ti

Zr

Be

Ni

Al

Cu

°C

°F

°C

°F

43 45 48 43 70 48

43 45 47 47 ... 48

2 2 5 5 ... 4

12 8 ... ... 15 ...

... ... ... 5 ... ...

... ... ... ... 15 ...

799 899 893 927 950 1000

1470 1650 1640 1700 1742 1832

816 899 904 927 1000 1050

1500 1650 1660 1700 1832 1922

Chapter 4: Base Metals and Base-Metal Family Groups / 97

formance during brazing, and wettability indexes >0.10 are indicative of excellent performance during brazing. • The decrease in wettability of the filler metals on TZM versus molybdenum is believed to be due to the presence of titanium in the oxide on the surface of TZM. • Some nickel-containing filler metals have been shown to embrittle both molybdenum and TZM. Liquid metal embrittlement is believed to be the mechanism by which this embrittlement occurs. Torches, controlled-atmosphere furnaces, vacuum furnaces, and induction and resistance heating equipment can be used to braze molybdenum with the variety of filler metals described previously. In a report (Ref 66) on the vacuum brazing of P/M molybdenum TZM alloy plate and drawn molybdenum tubing assembly for spacecraft radiator panels, the full-scale test panel successfully passed a battery of tests, including heat transport, thermal shock, and launchlike vibration testing. Brazing was successfully employed despite the recrystallization of both the tubing and the isogrid substrate panel. Platinum filler metal demonstrated its ability to obtain high-strength reliable joints.

Niobium and Its Alloys. Niobium is used mostly for nuclear and aerospace applications. One aerospace application required the development of manufacturing procedures for fabricating flat and curved niobium alloy sandwich heatshield panels (Ref 59). Titanium and the titanium-base alloy Ti-11Cr-13V-3Al (B120VCA) produced joints with excellent ductility and less base-metal reaction (Ref 67, 68). The flow characteristics of the B120VCA filler metal were more sluggish than those of pure titanium; as a result, the filler metal could be used to bridge wide joint clearances. The previously mentioned honeycomb sandwich panels were successfully brazed with pure titanium at 1815 °C (3300 °F). Heat-shield panels have been brazed with B120VCA filler metal at 1650 °C (3000 °F) in the same vacuum brazing environment. Simulated re-entry environment tests have shown that the heatshield panel could be used at temperatures up to 1315 °C (2400 °F). Room-temperature and elevated-temperature tests of the structural honeycomb panel indicated that it possessed useful properties up to 1260 °C (2300 °F) (Ref 59). Other developments investigated and evaluated the high-remelt-temperature brazing techniques based on diffusion-sink and reactive brazing concepts. The diffusion-sink technique involves either permitting the filler metal to

Table 4.11 Wettability index values for commercial filler metals used in brazing of molybdenum alloys Temperature, °C (°F) Filler metal

750 (1380)

800 (1470)

825 (1520)

850 (1560)

875 (1610)

900 (1650)

950 (1740)

975 (1790)

1000 (1830)

1025 (1880)

1050 (1920)

1075 (1970)

Palcusil 5 Palcusil 10 Palcusil 15 Palcusil 20 Palcusil 25 Nicusil 3 Nicusil 8 T50 T51 T52 Cusiltin 5 Cusiltin 10 Braze 580 Braze 630 Braze 655 Nioro Nicoro 80 Palniro 7 Gapasil 9

... ... ... ... ... ... ... ... ... ... ... ... 0.030 ... ... ... ... ... ...

... ... ... ... ... 0 ... ... 0 ... ... ... 0.037 ... ... ... ... ... ...

... ... ... ... ... ... ... ... ... ... ... 0.021 ... 0 0.061 ... ... ... ...

0.012 0.011 ... 0.015 ... 0.024 0.020 0.023 0.023 0.024 0.008 0.037 0.045 0.066 0.075 ... ... ... ...

... ... ... ... ... ... ... ... ... ... 0.012 0.031 ... ... 0.069 ... ... ... ...

0.034 0.030 0.044 0.037 ... 0.053 0.052 0.052 0.049 0.057 0.057 ... ... 0.080 ... ... ... ... 0.017

0.045 0.045 0.045 ... 0.102 ... 0.096 0.069 ... 0.075 ... ... ... ... ... ... 0.024 ... 0.061

... ... 0.152 ... 0.119 ... ... ... ... ... ... ... ... ... ... 0.158 ... ... ...

... ... ... ... 0.418 ... ... ... ... ... ... ... ... ... ... ... ... ... ...

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.071 0.023 ...

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.119 ...

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.102 ...

Source: Ref 65

Source: Ref 65

580 Nicusil 3 T51 630 Cusiltin 655 Palcusil 5 Palcusil 10 Nicusil 8 T50 T52 Cusiltin 5 Palcusil 20 Ticusil Cusil Palcusil 15 Gapasil 9 Silcoro 60 Palcusil 25 Nioro Nicoro 80 Ticuni Silver Incuro 60 Silcoro 75 Palnicusil Altizirbe Palniro 7 Gold Palniro 1 Palniro 4 Palco

Filler metal

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ...

750 (1380)

0.007 0.008 0.007 0.015 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ...

800 (1470)

... ... ... ... 0 0.009 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ...

825 (1520)

0.009 0.009 0.009 0.018 0 0.009 0 0.007 0.008 0.014 0.016 0 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ...

850 (1560)

0.018 ... ... ... 0 0.025 ... ... ... ... ... 0 0.009 0.024 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ...

875 (1610)

... 0.015 0.010 0.016 0 ... 0.011 0.016 0.018 0.020 0.017 0 0.017 0.101 0 0.023 0 0 ... ... ... ... ... ... ... ... ... ... ... ... ... ...

900 (1650)

... ... 0.020 ... ... ... ... ... ... ... ... ... 0.034 ... 0.009 ... ... 0.020 ... ... ... ... ... ... ... ... ... ... ... ... ... ...

925 (1700)

... 0.022 0.018 ... ... ... 0.027 0.028 0.024 0.025 0.026 ... ... ... 0.009 0.039 0.025 0.023 0.096 0 0.020 0.093 ... ... ... ... ... ... ... ... ... ...

950 (1740)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.105 0.070 ... ... ... 0.008 0 0.089 0.034 ... ... ... ... ...

975 (1790)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.50 0.041 ... 0.204 0.225 0.029 ... ... 0.012 0 ... ... ... ... ... ... ...

1000 (1830)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.013 0 0.318 ... 0 ... ... ... ...

1025 (1880)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.049 ... 0 ... ... ... ... 0.107 ... ... ... ...

1050 (1920)

Temperature, °C (°F)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.013 ... ... ... ... 0.133 0.088 ... ... ...

1075 (1970)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.013 ... ... ... ... ... 0.107 ... ... ...

1100 (2010)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.037 ... ...

1125 (2060)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.053 ... ...

1150 (2100)

Table 4.12 Wettability index values for commercial filler metals used in brazing of molybdenum alloy TZM

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.070 0.037 ...

1175 (2150)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.044 ...

1200 (2190)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.047 ...

1225 (2240)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.041

1275 (2330)

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 0.048

1300 (2370)

98 / Brazing, Second Edition

Chapter 4: Base Metals and Base-Metal Family Groups / 99

react with the base metal under controlled conditions or adding base-metal powder to the filler-metal powder. One diffusion-sink filler metal is Ti-33Cr, requiring brazing temperatures of 1455 to 1480 °C (2650 to 2700 °F) (Ref 67). The other is a Ti-30V-4Be reactive filler metal that requires brazing temperatures of 1290 to 1315 °C (2350 to 2400 °F). Test results of Ti-33Cr brazed joints with D-36 niobium alloy indicate an increase in lap shear strength from approximately 17 MPa (2.5 ksi) to more than 31 MPa (4.5 ksi) at 1370 °C (2500 °F) and 7 MPa (1 ksi) at 1650 °C (3000 °F). Other niobium alloys—D43 (10W-1Zr-0.1CNb), Nb-752 (10W-2.5Zr-Nb), and C-129Y (10W-11Hf-0.07Y-Nb)—have been brazed successfully with two filler metals, B120VCA and Ti-8.5Si, whose brazing temperature was 1455 °C (2650 °F). Tantalum and Its Alloys. The filler metals for moderate-temperature applications of tantalum and its alloys are chosen on the basis of the intended application. Nickel-base filler metals (such as the Ni-Cr-Si filler metals) have been used in successful brazing of tantalum. These filler metals are satisfactory for service temperatures below 980 °C (1800 °F). Copper-gold alloys having less than 40% Au can also be used as filler metals, but gold in amounts between 40 and 90% tends to form brittle, age-hardening compounds. Copper-tin, gold-nickel, and copper-titanium filler metals also have been used in brazing tantalum and its alloys. Because most uses of tantalum and its alloys are in elevated-temperature applications (at or above 1650 °C, or 3000 °F), there is a lack of high-temperature filler metals for use with tantalum. Investigations have examined conventional brazing, reactive brazing, and diffusionsink brazing concepts for new tantalum filler metals. The reactive brazing concept uses a filler metal containing a strong melt-temperature depressant. The depressant is selected to react with the base material or powder additions to form a high-melting intermetallic compound during a postbraze diffusion treatment. As the depressant diffuses into the base material, the joint remelt temperature is increased. Successful application of this concept appears highly dependent on controlling the intermetallic compound reaction to form discrete particles. Diffusion-sink brazing with titanium and Ti30V filler metals, whose brazing temperatures range from 1675 to 1760 °C (3050 to 3200 °F), has produced remelt temperatures exceeding

2095 °C (3800 °F) in tee and lap joints. Diffusion-sink brazing with 33Zr-34Ti-33V filler metal with a brazing temperature of 1425 °C (2600 °F) has produced remelt temperatures exceeding 1760 °C (3200 °F). Therefore, these remelt temperatures indicate that service temperatures could be 1925 °C (3500 °F) for the titanium and Ti-30V filler metals and 1650 °C (3000 °F) for the 33Zr-34Ti-33V filler metal. Most of the filler metals currently available for tantalum are in powder form, which is difficult to work with at elevated temperatures. New powder filler metals, such as Hf-7Mo, Hf-40Ta, and Hf-19Ta-12.5Mo, have been used successfully but require further development and refinement. Special techniques are necessary to braze tantalum satisfactorily. All gases that have any reactivity must be removed. Oxygen and carbon monoxide must be eliminated. Tantalum forms oxides, carbides, nitrides, and hydrides very readily with these gases. A loss of ductility ensues. At high temperatures, tantalum should be protected from oxidation. One method is to electroplate the surfaces with copper or nickel, in which case it is necessary for the filler metal to be compatible with the plating. Controlledatmosphere and both hot- and cold-wall vacuum furnaces have been used for brazing tantalum and its alloys. Induction, resistance, and torch brazing are not recommended. Tungsten and Its Alloys. Tungsten has been successfully brazed in vacuum. Filler metals that have been used to braze tungsten as well as the other refractory metals are listed in Table 4.10. These filler metals and pure metals having liquidus temperatures ranging from 650 to 1925 °C (1200 to 3500 °F) are potentially useful for brazing. Tungsten can be brazed in much the same manner as molybdenum and its alloys, using many of the same filler metals. Brazing can be accomplished in a vacuum or in a dry argon, helium, or hydrogen atmosphere. To some extent, the selection of the brazing atmosphere depends on the filler metal used. For example, filler metals that contain elements with high vapor pressures at the brazing temperature cannot be used effectively in a high vacuum. Care must be exercised in handling and fixturing tungsten parts because of their inherent brittleness; these parts should be assembled in a stress-free condition. Contact between graphite fixtures and tungsten must be avoided to prevent the formation of brittle tungsten carbides.

100 / Brazing, Second Edition

Although nickel-base filler metals have been used successfully to braze tungsten, a reaction between nickel and tungsten that results in basemetal recrystallization can occur; this reaction can be minimized by short brazing cycles, minimum brazing temperatures, and the use of small quantities of filler metal. Data on brazing of tungsten are more limited than data on brazing of other refractory metals. Unalloyed tungsten has been vacuum brazed with two experimental filler metals, Nb-2.2B and Nb-20Ti. The lap shear strength of joints brazed with the Nb-2.2B filler metal was approximately 35 MPa (5 ksi) at 1650 °C (3000 °F) and 55.2 MPa (8 ksi) at 1370 °C (2500 °F). The strength of joints brazed with Nb-20Ti was somewhat lower, approximately 21 MPa (3 ksi) at both test temperatures. The flow of Nb-20Ti filler metal was more sluggish than that obtained with Nb-2.2B. Filler metals based on the platinum-boron and iridium-boron systems were developed to braze tungsten for service at 1925 °C (3500 °F). They contained up to 4.5% B and could be used for brazing below the recrystallization temperature of tungsten. Tungsten lap joints were brazed and diffusion treated in a vacuum at 1095 °C (2000 °F) for 3 h. This cycle resulted in the production of joints with remelt temperatures of approximately 2040 °C (3700 °F). A slight increase in joint remelt temperature was noted when tungsten powder was added to the filler metal. The highest remelt temperature of 2170 °C (3940 °F) was obtained when joints were brazed with Pt3.6B plus 11 wt% W powder (Ref 69). Studies have been conducted to develop and evaluate filler metals that could be used to braze tungsten for nuclear reactor service at 2500 °C (4530 °F) in hydrogen atmosphere. Butt joints have been brazed using a gas tungsten arc as the heat source and the filler metals W-25Os, W50Mo-3Re, and Mo-5Os. Furnace brazing (inert-gas or reducing atmosphere and vacuum), torch brazing, resistance brazing, or induction brazing can be used to braze tungsten. Furnace brazing operations usually are used when the parts are larger than those considered practical for the induction and resistance processes. Tungsten-copper electrode tips are materials made by P/M processes. A normal analysis of one of the materials is 80W-20Cu. These tips may be used for resistance welding electrodes that are usually brazed to the copper-chromium alloys, which constitute the main current-carry-

ing portion of the electrode. Any of the silvercopper and/or zinc series of filler metals are suitable for this application. Tungsten-copper tips usually are applied by induction or torch heating in a neutral or reducing atmosphere. The most suitable filler metal is 72Ag-28Cu. To hold parts in alignment, pins are incorporated in the design, and the filler metal is preplaced in the form of shim stock located between the mating parts. In brazing of resistance welding electrodes with pure tungsten tips, precoating of the tungsten surface aids uniform flow of the filler metal between the faying surfaces of the joint. Usually, one of the silver-base filler metals is used as a precoating. Without pretreatment, poor wetting of the tungsten by the filler metal often results, causing premature joint failure under the high pressures and temperatures of resistance welding.

Cast Iron Brazing of gray, ductile, and malleable cast irons differs from brazing of steel in two principal respects: special precleaning methods are necessary to remove graphite from the surface of the iron, and the brazing temperature is kept as low as feasible to avoid reduction in the hardness and strength of the iron. The processes used for brazing of cast irons are the same as those used for brazing of steel, which include furnace, torch, induction, and dip brazing. As with other metals, selection of the brazing process depends largely on the size and shape of the assembly, the quantity of assemblies to be brazed, and the equipment available. In recent years, brazing of ordinary gray cast irons using silver-base filler metal has become practical commercially; the development of a satisfactory surface pretreatment has opened up many design possibilities. Intricate forms can be built from simple castings or by joining castings to standard wrought forms, such as tubing and rolled shapes. Foundry work may be simplified by making intricate castings in several parts to reduce coring. Gray, ductile, and malleable irons commonly are brazed to steels or to other cast irons. In brazing of ductile or malleable irons, certain precautions are imperative. If ductile or malleable irons are heated above 760 °C (1400 °F), the metallurgical structure may be damaged; brazing thus should be done below this temperature. The effects of graphite, which is present in all gray, ductile, and malleable cast irons, are sig-

Chapter 4: Base Metals and Base-Metal Family Groups / 101

nificant. Although gray, ductile, and malleable irons all have lower brazeability than carbon or low-alloy steels, the three types of iron are not equal in brazeability. Malleable iron is generally considered the most brazeable of the three types of cast iron, largely because the total carbon content is somewhat lower (seldom over 2.70%), and, therefore, graphitization is lower. Brazeability is also enhanced because the graphite occurs in the form of approximately round nodules and thus is easier to remove or cover up (as by abrasive blasting). Also, malleable iron is lower in silicon than the other types of cast iron and thus is less graphitized, which makes it better suited for brazing. Ductile iron can have a composition nearly the same as gray iron, but the graphite particles are spheroidal rather than flake-shaped. The spheroidal shape is more favorable for brazing. Shot or grit blasting is effective in rolling metal over graphite particles that are exposed at the surface. Gray iron, which is characterized by large flakes of graphite, is the most difficult type of cast iron to braze. Until the development of electrolytic salt bath cleaning, brazing of gray iron was considered impractical. Abrasive blasting with steel shot or grit has proved reasonably successful for preparing surfaces of ductile and malleable iron castings but is seldom suitable for preparing surfaces of gray iron castings. Electrolytic treatment in a molten salt bath, alternately reducing and oxidizing, has been the most successful method for surface preparation and is applicable to all graphitic cast irons. Ordinary chemical cleaning methods, such as degreasing, detergent washing, and acid pickling, have the distinct disadvantage of not removing surface carbon, which interferes with bonding. Before any procedure for cleaning is adopted, tests should be made by cleaning samples of the iron intended for use in the castings to be brazed, fluxing the samples, and applying filler metal (preferably on a smooth, flat surface). The samples are then heated to the preestablished brazing temperature, cooled, and examined visually. If the samples show an indication that the filler metal has not uniformly wetted the testpiece, the surface is not sufficiently clean. Alkaline-based salts (sodium hydroxide) operating at 400 to 480 °C (750 to 900 °F) are extremely effective in removing surface oxides and sand from iron castings. At these temperatures, salt baths exhibit the required high chem-

ical activity. This action is further enhanced in the cleaning operations by the introduction of electrical energy. The effective use of these salts and the removal of their by-products to produce an excellent metallurgical surface has resulted in quality bonds and has permitted direct babbitting of cast iron bearings and silver brazing fittings to cast iron surfaces. With proper cleaning and preparation of the cast iron, it has been possible to produce a brazed bond between metal surfaces that exceeds the strength of the parent metal. With proper surface preparation, the usual variety of filler metals suitable for steel brazing is satisfactory for cast irons. In selection of the filler metal, the metallurgical effects of the brazing cycle on the base-metal matrix (hardening, softening, etc.) should be considered. Various filler metals in the silver-copper series are commonly used. Copper filler metal has been used successfully in furnace brazing. Some cast irons have low melting (solidus) temperatures, suggesting that care must be exercised to keep the brazing temperature below 1120 °C (2050 °F). Brass filler metals (copper-zinc series) frequently are used to braze cast irons, while the silver-base filler metals have been used to join copper to cast iron. Most brazing of cast iron is performed to join assemblies at lower cost than is possible by another process or to fabricate parts that are difficult to produce—one-piece castings, for example. In some applications, two or more cast iron components are brazed together; in other applications, one or more components of a brazed assembly are made of another metal— most often steel but also copper alloys. Localized heating methods, such as torch and induction, have the advantages of restricted heat effects and limited metallurgical phase changes in base-metal components. Furnace brazing may be desirable for production lots due to cost savings and/or consistency of results. Brazing with copper and nickel filler metals require atmosphere controls during heating, usually accomplished by furnace brazing. The use of dry hydrogen as a protective atmosphere promotes wetting by reducing oxides such as silica. Electric arc braze welding processes have successfully been used with bronze-brass filler metals on cast iron and require special preheating techniques to prevent cracking. Cast irons that contain pearlite or free carbide graphitize and decrease in strength at elevated temperatures. Because graphitization is a func-

102 / Brazing, Second Edition

tion of time and temperature, some experimentation is usually necessary to develop a brazing cycle that produces acceptable joints without excessive graphitization and decreases in strength. Two identical assemblies were heated in different plants for 1 h. In the first plant, the assemblies were heated to approximately 705 °C (1300 °F) with no significant decrease in strength. In the second plant, the assemblies were heated to approximately 790 °C (1450 °F), again without significant decrease in strength. In the latter application, however, heating was done by induction instead of in a furnace, and the heating time was much shorter than would have been required in a furnace. In some plants using furnace brazing, the furnace is operated at a considerably higher temperature than is desired for brazing (sometimes as high as 870 °C, or 1600 °F), but the time cycles used are so short that the assemblies never reach furnace temperature. For applications in which little or no decrease in strength of the cast iron can be tolerated, it is mandatory to use a filler metal with as low a flow-temperature range as possible, thus permitting a low brazing temperature, and to keep the time at brazing temperature to the minimum. The temperature required for wetting the base metal and for flow of filler metal having a melting range of 620 to 645 °C (1150 to 1195 °F) may vary from approximately 690 to 845 °C (1275 to 1555 °F), depending on the complexity of the joint design—especially the distance the filler metal must flow. Simple joints with short flow distances can be brazed at lower temperatures than more complex joints.

Low-Carbon, Low-Alloy, and Tool Steels Brazing of low-carbon and low-alloy steels is a highly developed, low-cost production process. Filler metals can be either manually or automatically applied or preplaced in the joint. The heat treatment considerations of low-alloy steels are factors in determining the specific filler metals and brazing temperatures to be used. The steels covered in this section include low-carbon (less than 0.30% C) and low-alloy steels. The low-alloy steels include the Society of Automotive Engineers/American Iron and Steel Institute (SAE/AISI) 23xx nickel steels, 31xx nickel-chromium steels, 41xx chromiummolybdenum steels, 43xx nickel-chromium-

molybdenum steels, and several other types containing less than 5% total alloy content. Leaded steels can be torch brazed using filler metal and flux combinations normally recommended for steel. Tests conducted on low-carbon free-machining steel containing 0.25 to 0.35% Pb show detrimental effects after torch brazing with silver-base filler metals; satisfactory joints have been made in these leaded steels using copper- and nickel-base filler metals. All the silver-base filler metals can be used for brazing low-carbon and low-alloy steels (Table 4.13). The silver-base filler metals containing nickel usually provide better wettability and are preferred for brazing low-alloy steels wherein joint strength is most important. Copper-base filler metals are used mainly for preplacement in controlled-atmosphere furnaces. Copper is also generally preferred because of its low cost and the high strength of the joints produced. Copper-zinc filler metals often are face fed into the joint but can also be preplaced for furnace and induction heating. The high solidus temperatures (1095 to 1150 °C, or 2000 to 2100 °F) necessary when copperbase filler metals are used often allow simultaneous brazing and heat treating of certain lowalloy steels. Nickel-base filler metals have been used for joining low-carbon and low-alloy steels when special joint requirements so dictate. Brazing usually is restricted to controlledatmosphere furnaces. Low-carbon and low-alloy steels can be brazed using virtually all known processes. Torch, furnace, and induction heating techniques are the most common. Filler metals in the form of continuous wire or strip can be automatically applied using electromechanical wire feeders; powder filler metals blended with flux and pasteforming ingredients are applied automatically with pressurized dispensing equipment. For torch brazing, the equipment would include standard oxyacetylene or similar torches. Furnaces can be of the batch or conveyor type, with or without atmosphere control; they can be electric, gas, or oil fired and should provide accurate temperature control. The principal advantage of furnace brazing over other brazing processes is that it permits the use of a variety of prepared protective atmospheres—notably, the rich exothermicbased, endothermic-based, and some prepared and commercial nitrogen-base atmospheres. These atmospheres are among the least expen-

Strip, rod, wire, powder Strip, rod, wire, powder Rod Rod

Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder Strip, wire, powder

Product form

... ... ... ...

45 50 35 30 50 40 45 50 56 30 25 40

Ag

59 48 58 58

15 15.5 26 27 15.5 30 30 34 22 38 35 30

Cu

40 41 38 39

16 16.5 21 23 15.5 28 25 16 17 32 26.5 28

Zn

... ... ... ...

... 10.0 0.5 ...

... ... ... ... 3.0 2.0 ... ... ... ... ... ...

Ni

0.6 ... 0.95 0.95

... ... ... ... ... ... ... ... 5.0 ... ... 2

Sn

Fe

... ... 0.7 0.7

... ... ... ... ... ... ... ... ... ... ... ...

Nominal composition, %

24 18 18 20 16 ... ... ... ... ... 13.5 ...

Cd

... ... 0.25 0.25

... ... ... ... ... ... ... ... ... ... ... ...

Mn

... 0.15 0.08 0.08

... ... ... ... ... ... ... ... ... ... ... ...

Si

... 0.025 ... ...

... ... ... ... ... ... ... ... ... ... ... ...

P

890 920 865 865

605 625 605 605 630 670 665 690 620 675 605 650

°C

1630 1690 1590 1590

1125 1160 1125 1125 1170 1240 1225 1270 1145 1250 1125 1200

°F

900 935 880 890

620 635 700 710 690 780 745 775 650 765 745 710

°C

1650 1715 1620 1630

1145 1175 1295 1310 1270 1435 1370 1425 1205 1410 1375 1310

°F

Solidus temperature Liquidus temperature

910–955 940–980 ... 815–925

620–760 635–760 700–845 710–845 690–815 780–900 745–845 775–870 650–760 765–870 745–860 710–845

°C

1670–1750 1720–1800 ... 1500–1700

1145–1400 1175–1400 1295–1550 1310–1550 1270–1500 1435–1650 1370–1550 1425–1600 1205–1400 1410–1600 1375–1575 1310–1550

°F

Brazing temperature

Note: AWS, American Welding Society (a) Classified for braze welding and brazing. (b) Classified for braze welding. Source: Abstracted from the mandatory and nonmandatory sections of AWS A5.7, AWS A5.8, and Ref 65

RBCuZn-A(a) RBCuZn-D(a) RCuZn -B(b) RCuZn-C(b)

Copper-zinc alloys

BAg-1 BAg-1a BAg-2 BAg-2a BAg-3 BAg-4 BAg-5 BAg-6 BAg-7 BAg-20 BAg-27 BAg-28

Silver alloys(a)

AWS classification

Table 4.13 Filler metals for torch brazing low-carbon and low-alloy steels

Chapter 4: Base Metals and Base-Metal Family Groups / 103

104 / Brazing, Second Edition

sive; they can be generated in the plant in large volume, or, in the case of commercial nitrogenbase atmospheres, they can be stored in liquid form outside the plant. They provide excellent protection against oxidation, and they can be prepared with any carbon potential in the range of approximately 0.2% to more than 1.0% C, depending on the atmosphere. This range of carbon potential is sufficient to accommodate all carbon and low-alloy steels, including those carburized before brazing. By selecting an atmosphere with a carbon potential that matches the carbon content of the work metal, brazing can be accomplished without carburizing or decarburizing the work metal. Because the protective atmospheres used for furnace brazing are sufficiently reducing to iron oxide, they usually eliminate the need for fluxes in brazing carbon steel with copper filler metal. An oxide-free surface normally promotes wetting of the workpiece by the molten filler metal. However, some low-alloy steels that contain a total of more than 2 or 3% Cr, Mn, Al, and Si form more stable surface oxides, and they require highly reducing atmospheres (such as dry hydrogen or dissociated ammonia), a flux, or nickel plating to obtain adequate wetting. Most of the limitations of furnace brazing are directly related to the high temperatures required for brazing of steels with copper filler metal. These temperatures exceed the average brazing temperature required for brazing with silver-base filler metals by 280 °C (500 °F) or more. They are high enough to cause grain coarsening in medium-carbon, high-carbon, and low-alloy steels; however, grain refinement can be obtained by subsequent heat treatment. Steels brazed with copper filler metal develop lower tensile and yield strengths and increased ductility as the brazing time or temperature, or both, are increased. These changes in properties are a result of decarburization in some types of atmospheres and alteration of grain size. Original grain size can be restored by subsequent heat treatment below the remelt temperature of the copper brazing filler metal. Loss of carbon through decarburization is generally unimportant in low-carbon steels. However, surface hardness of some low-alloy steels may be substantially lowered. Such loss of surface hardness can be very deleterious when thin-gage material is involved. For alloy steels, the filler metal should have a solidus well above any heat treating tempera-

ture to avoid damage to joints that are heat treated after brazing. In some cases, air-hardening steels can be brazed and then hardened by quenching from the brazing temperature. A filler metal with brazing temperature lower than the critical temperature of the steel can be used when no change in the metallurgical properties of the base metal is wanted. In torch brazing, a neutral or slightly reducing flame usually is preferred, because the filler metal is face fed into the prefluxed joint. Fluxcoated filler metal is often beneficial. As in all brazing processes, it is important to avoid overheating in brazing to prevent undesirable effects on the base metal, the filler metal, or the flux. Time at temperature is an important consideration, especially when the filler metal contains volatile elements such as zinc and cadmium. Excessive heat also might affect the integrity of the braze and reduce the mechanical properties of the joint. Automated torch and burner-type production equipment is available for high-production applications. These units usually use brazing fluxes and silver-base filler metals. In production applications, the filler metal (usually copper) is preplaced in, or adjacent to, the joint before the preassembled parts are charged into controlled-atmosphere batch- or conveyor-type furnaces for brazing. Induction heating for brazing is also advantageous in that maximum temperature rise is restricted to the immediate joint area by selective coil design and the use of an appropriate frequency for the induction-heating circuit. In joining certain hardenable low-alloy steels, it usually is desirable to use the lowermelting silver-base filler metals, which can be applied below the lower transformation temperature of the steel. Some local annealing, however, may occur. When postheat treatments are required, the higher-solidus filler metals are necessary to preclude the possibility of joint impairment. Low-alloy steels are sometimes brazed and heat treated simultaneously using filler metals of the copper-zinc and silver classifications. The solidus temperature of the filler metal chosen must be above the austenitizing temperature recommended for the base metal prior to quenching. In an application of this type, the joint is made at normal brazing temperature, removed from the heat source to permit a drop to the hardening temperature, and quenched. The procedure is satisfactory only for base metals compatible with rapid cooling.

Chapter 4: Base Metals and Base-Metal Family Groups / 105

Dip brazing has also been used to join carbon and low-alloy steels with silver-base, copperzinc, and other copper-base filler metals. The types of salts used in dip brazing of carbon and low-alloy steels are neutral chloride salts, neutral chloride salts plus a fluxing agent such as borax or cryolite, and carburizing and cyaniding salts, which are also fluxing-type salts. Types and compositions of brazing salts and temperatures used for brazing of carbon and low-alloy steels with various filler metals are given in Table 4.14. Neutral salts, so called because normally they do not add or subtract anything from the surface of the steel being treated, protect the surface from attack by oxygen in the air. Oxide on the workpiece, however, cannot be reduced by the salt, and a flux must generally be provided. The neutral salts are mildly oxidizing to steel when they are used at recommended austenitizing temperatures. The oxides produced by heating steel in molten salt are largely soluble; hence, the steel is scale-free after heating. The accumulation of oxide in the molten salt, however, progressively makes the salt more decarburizing, and for this reason, baths may require periodic rectification (Ref 70). Carburizing and cyaniding salts provide their own fluxing action. In addition, they supply carbon or carbon and nitrogen to the surface of the steel assembly as it is being brazed. Although silver-base filler metals have been used successfully, copper-zinc filler metal is generally preferred. Various applications involve the use of filler metals in brazing low-carbon and low-alloy steels in everyday production. These include

many components of such vehicles as automobiles, trucks, bicycles, motorcycles, snowmobiles, all-terrain vehicles, and the like. Other common brazements include window and door frames, ducts, tanks, containers of all types, perforated and expanded steel panels, steel partitions, and shelving. A great variety of tubular steel furniture is production brazed, and brazing is also used in production of cutting tools and industrial knives, hydraulic oil tanks, reservoirs, electronic chassis and supports, hand tools, honing appliances, certain instruments, and steel assemblies of all types. Tool Steels. High-carbon steels contain more than 0.45% C. High-carbon tool steels usually contain 0.60 to 1.40% C. In discussing the brazing of tool steels, it is convenient to group them in two broad classifications: carbon steels and high-speed tool steels. Brazing of high-carbon steels is best accomplished prior to or during the hardening operation. Hardening temperatures for carbon steels range from 760 to 820 °C (1400 to 1510 °F). Filler metals having brazing temperatures above 820 °C should be used. When brazing and hardening are done in one operation, the filler metal should have a solidus at or below the austenitizing temperature. Tempering and brazing can be combined for high-speed tool steels and high-carbon, highchromium alloy tool steels that have tempering temperatures in the range of 540 to 650 °C (1000 to 1200 °F). Filler metals with brazing temperatures in that range are used. The part is removed from the tempering furnace, brazed by localized heating methods, and then returned to the furnace for completion of the tempering cycle.

Table 4.14 Typical salts used for dip brazing of carbon and low-alloy steels with various brazing filler metals Brazing temperature range (a) Filler metal

BAg-1 through BAg-8, and BAg-18 RBCuZn-A

RBCuZn-D

Type of salt

Neutral Cyaniding-fluxing Neutral Neutral Fluxing Carburizing-fluxing (water soluble) Carburizing and self-fluxing Neutral Neutral Neutral

(a) Temperatures shown are those of the salt bath. Source: Ref 65

°C

°F

621–871 649–871 732–871 913–940 913–940 913–940 816–927 1038–1051 1093–1149 1093–1149

1150–1600 1200–1600 1350–1600 1675–1725 1675–1725 1675–1725 1500–1700 1900–1925 2000–2100 2000–2100

106 / Brazing, Second Edition

Localized heating for brazing may decrease the hardness of heat treated steels when the brazing temperature is above the tempering temperature of the steel. Except for thin sections, these steels must be quenched drastically during heat treatment to achieve optimal properties. Alloying elements may be added to carbon steels to impart special properties, such as reduced distortion on heat treatment, greater wear resistance and toughness, or better hightemperature properties. Such steels are referred to as alloy tool steels; they are known by various trade names and grades, and their properties are adequately covered in manufacturers’ published information and in various handbooks. The alloy steel in question should be studied carefully to determine its proper heat treating cycle, the kind of quench necessary (water, oil, or air), the best filler metal, and the proper technique for combining the heat treating and brazing operations to achieve maximum properties in service life. High-speed steels, although logically falling in the alloy steel group, are classified separately because their properties depend on relatively high percentages of such alloying elements as tungsten, molybdenum, chromium, and vanadium. Their carbon contents normally are much lower than those of carbon tool steels. Highspeed steels are widely used in industry as metal cutting tools. A common type known as 18-4-1 contains 18% W, 4% Cr, and 1% V. The choice of the filler metal to be used depends on the properties of the tool steel being brazed and the heat treatment required to develop its optimal properties. Practically all filler metals of the BAg, BCu, and RBCuZn classifications are used at various times. The best filler metal to use should be determined for the specific application. Torch, furnace, and induction heating are the three most commonly used processes in brazing tool steels. Available equipment is frequently the main factor in process selection.

Stainless Steels Stainless steels include a wide variety of ironbase alloys containing chromium that are used primarily for applications demanding heat or corrosion resistance. Tighter process controls are required than for brazing of carbon steels. These more rigorous requirements are imposed by the inherent chemical characteristics of

stainless steels and the generally more arduous service environments in which they are used. Success in the fabrication of stainless steel components by brazing depends on knowledge of the characteristics of the various types of stainless steels, and rigid adherence to certain items of process control required by these characteristics. All stainless steels are difficult to wet because of their high chromium contents. Brazing of these alloys is best accomplished in a purified (dry) hydrogen atmosphere or in a vacuum. Dewpoints of –50 °C (–60 °F) or lower must be maintained, because problems with wetting may arise following the formation of chromium oxide. In torch brazing of these base metals, fluxes are required to reduce any chromium oxides present. Austenitic nonhardenable stainless steels contain sufficient nickel or nickel plus manganese additions to stabilize austenite down to room temperature, causing these alloys to be nonmagnetic and nonhardenable by heat treatment. Stainless steels of this class possess the highest heat and corrosion resistance. They are designated as AISI 300- and 200-series alloys. One commonly used alloy is type 302, which contains approximately 18% Cr and 8% Ni. In the 200-series stainless steels, some of the nickel is replaced by manganese. For example, type 202, the parallel to type 302, contains 18% Cr, 5% Ni, and 9% Mn. The 300-series stainless steels are used widely for both torch- and furnace-brazed assemblies. These alloys have relatively high thermal expansion and low thermal conductivity. This combination of properties makes thermal distortion a major concern in furnace brazing of large or complex assemblies or assemblies in which dissimilar materials are brazed to stainless steel. In design of fixtures, heat shields, and thermal cycles, the requirement to provide uniform heating and cooling must be considered. In brazed assemblies for which corrosion resistance is important, precautions must be taken to avoid sensitization to intergranular corrosion. This problem was discussed earlier in this chapter, but briefly, it occurs when an unstabilized grade of austenitic stainless steel, such as type 302 or type 304, is held at temperatures in the range from 425 to 815 °C (800 to 1500 °F) or slowly cooled through this range. The excess carbon combines with chromium and precipi-

Chapter 4: Base Metals and Base-Metal Family Groups / 107

tates as chromium carbide along grain boundaries of the austenite. The region around the precipitate is depleted of chromium and thus becomes susceptible to corrosion. The 300-series stainless steels have been used widely for brazed components; the 200series alloys, which are only 25 years old, have not seen as much use. Due to their high manganese content, the 200-series stainless steels are more difficult to furnace braze in hydrogen atmospheres than the 300 series. Manganese forms an oxide that is not reduced easily by dry hydrogen at the furnace brazing temperatures normally used for stainless steels. It is important to start with thoroughly cleaned surfaces and to maintain a low dewpoint in the hydrogen atmosphere. All the chromium-nickel steels are subject to SCC in the presence of molten filler metals. This phenomenon occurs when the base metal is under stress—either residual or resulting from applied loads while the braze is being made. One form of applied load results from thermal gradients during brazing. The filler metal penetrates the base metal along the grain boundaries at the points of stress, producing a greatly weakened base metal. Best results, therefore, are obtained with stress-relieved material. This stress relief may be done prior to or during the brazing cycle. If it occurs during the cycle, it must be done below the solidus temperature of the filler metal (Ref 71, 72). Ferritic nonhardenable stainless steels are basically low-carbon alloys of iron and chromium in which sufficient chromium has been added to the iron to stabilize ferrite, the low-temperature phase in steels, over a wide temperature range. The more common AISI stainless grades in this category are types 405, 406, 430, and 446. Type 430 is a widely used grade that is particularly subject to a form of interface corrosion when brazed with some silver-base filler metals. This corrosion apparently is caused by electrochemical action whereby the bond between the base metal and the filler metal is destroyed. In many cases, this action has been found to occur in the presence of tapwater. It has been found that addition of small percentages of nickel to silver-base filler metal prevents interface corrosion of brazed joints in most stainless steels. The nickel-containing silver-base filler metal is not completely effective, however, with type 430, even though its use greatly reduces the rate of attack. For type 430, a special silver-base

filler metal has been developed (63Ag-28.5Cu6Sn-2.5Ni). The ferritic stainless steels (405, 406, and 430) cannot be hardened, and their grain structure cannot be refined by heat treatment. These alloys degrade in properties when brazed at temperatures above 980 °C (1800 °F), because of excessive grain growth. They lose ductility after long heating times between 340 and 600 °C (645 and 1110 °F). However, some of the ductility can be recovered by heating the brazement to approximately 790 °C (1455 °F) for a suitable time (Ref 71, 72). Martensitic hardenable stainless steels are Fe-C-Cr alloys of two basic types: the lowchromium, low-carbon grades (types 403, 410, and 416) and the high-chromium, high-carbon grades (types 440A, B, and C). These steels are closely related to the ferritic nonhardenable grades, but their alloy compositions are so balanced that they air harden on cooling from brazing; this hardening occurs at a temperature above their austenitizing temperature range. Therefore, they must be annealed after brazing or during the brazing operation. These steels are also subject to stress cracking with certain filler metals. The primary precaution in brazing components made of these alloys is that the brazing thermal cycle must be compatible with the required heat treatment. If high-temperature nickel-base filler metals are used, it is possible to reaustenitize the assembly after brazing. Although this procedure would increase costs, it may be desirable to develop optimal properties in critical components. The thermal expansion properties of martensitic stainless steels are relatively low—in the same range as those of ferritic alloys (Ref 73). Precipitation-hardening stainless steels are basically stainless steels with additions of one or more of the elements copper, molybdenum, aluminum, and titanium. Such alloying additions make it possible to strengthen the alloys by precipitation-hardening heat treatments. Some of the designations of the precipitationhardening stainless steels are 17-7PH, PH148Mo, PH13-8Mo, 15-5PH, PH15-7Mo, AM350, AM355, 17-4PH, and A-286. As in the case of the martensitic hardenable stainless steels, brazing thermal cycles used in joining these alloys must be compatible with their heat

108 / Brazing, Second Edition

treatments. Because the heat treatments vary rather widely, specific brazing procedures are required for each alloy. Precipitation-hardening alloys that contain aluminum or titanium are difficult to wet in the usual furnace brazing atmospheres. Nickel plating generally is used as a surface treatment before furnace brazing. Metallurgical Considerations in Brazing Stainless Steels. In brazing of stainless steels, base-metal inclusions and surface contaminants are even more deleterious than in brazing of carbon steels. Base-metal inclusions, such as oxides, sulfides, and nitrides, interfere with the flow of filler metal. Flow is also impeded by surface contaminants, which may include lubricants such as oil, graphite, molybdenum disulfide, and lead that are applied during machining, forming, and grinding, or by aluminum oxide particles produced by grit blasting or by grinding with aluminum oxide wheels or belts. Some filler metals in powder form contain or are mixed with an organic binder to form a paste that is subsequently applied to stainless steel for brazing; acrylics and other plastics are often used for this purpose. Although some binders form a soot residue, this residue does not usually interfere with filler-metal flow. The brazing characteristics of stainless steels can also be severely impaired by unsuitable fixturing materials, such as graphite, or by protective atmospheres with nitriding potentials. Carbon in graphite fixtures unites with hydrogen to form methane (CH4), which carburizes stainless steel and impairs its corrosion resistance. Dissociated ammonia, unless sufficiently dry and completely (100%) dissociated, nitrides stainless steel. Filler Metals for Brazing Stainless Steels. A wide variety of filler metals, including silver-base, nickel-base, gold-base, and copper, are commercially available for brazing stainless steel parts (Tables 4.15, 4.16). The factors to consider in selecting a filler metal for a particular application include the following: • Service conditions, including operating temperature, stresses, and environments • Heat treatment requirements if martensitic or precipitation-hardening steels are involved • Brazing process to be used • Cost • Special precautions, such as sensitization of unstabilized austenitic stainless steels at certain temperatures

Commercial filler metals are available that have copper, silver, nickel, cobalt, platinum, palladium, manganese, and gold as the base or as addition elements; these are grouped conveniently according to service temperature. The most widely used filler metals for brazing stainless steels are the silver-base family (BAg group). Filler metal BAg-3, which contains 3% Ni, is probably the silver-base filler metal selected most frequently, although several other silver-base filler metals can also be used successfully. Where improved corrosion resistance is needed, BAg-3 and BAg-4 are recommended. Silver-brazed joints cannot be used for hightemperature service; the recommended maximum service temperature is 370 °C (700 °F), which is the maximum temperature for BAg-13 and BAg-19 filler metals. Of the silver-base filler metals shown in Table 4.15, all except BAg-19 and possibly BAg-13 are used at brazing temperatures that fall within the effective sensitizing range for austenitic stainless steels (540 to 870 °C, or 1000 to 1600 °F). Chromium carbide precipitation occurs in the sensitizing temperature range, resulting in impairment of the corrosion resistance of the base metal. Carbide precipitation, however, depends on time as well as temperature, and exposure within the sensitizing temperature range for only a few minutes is unlikely to produce a significant amount of precipitate. Nevertheless, the lower melting temperatures of the silver-base filler metals prohibit re-solution treatment of the base metal after brazing, and if corrosion resistance in service is sufficiently critical, an extra-lowcarbon, titanium-stabilized or niobium-tantalum-stabilized austenitic stainless steel should be selected instead of a nonstabilized type. Ferritic and martensitic stainless steels that contain little or no nickel are susceptible to interface corrosion in plain water or moist atmospheres when brazed with nickel-free silver-base filler metals using a flux. Filler metals containing nickel (BAg-3) help to prevent interface corrosion. However, for complete protection, special filler metals containing nickel and tin should be used, and brazing should be done in a protective atmosphere without flux. The nickel-base filler metals usually rank second in frequency of use as filler metals for brazing of stainless steels. Nickel-base filler metals provide joints that have excellent corrosion resistance and high-temperature strength. These filler metals alloy with stainless steel,

BAg-1 BAg-1a BAg-2 BAg-2a BAg-3 BAg-4 BAg-5 BAg-6 BAg-7 BAg-8 BAg-8a BAg-9 BAg-10 BAg-13 BAg-13a BAg-18 BAg-19 BAg-20 BAg-21 BAg-22 BAg-23 BAg-24 BAg-25 BAg-26 BAg-27 BAg-28

Filler metal

44.0–46.0 49.0–51.0 34.0–36.0 29.0–31.0 49.0–51.0 39.0–41.0 44.0–46.0 49.0–51.0 55.0–57.0 71.0–73.0 71.0–73.0 64.0–66.0 69.0–71.0 53.0–55.0 55.0–57.0 59.0–61.0 92.0–93.0 29.0–31.0 62.0–64.0 48.0–50.0 84.0–86.0 49.0–51.0 19.0–21.0 24.0–26.0 24.0–26.0 39.0–41.0

Ag

14.0–16.0 14.5–16.5 25.0–27.0 26.0–28.0 14.5–16.5 29.0–31.0 29.0–31.0 33.0–35.0 21.0–23.0 bal bal 19.0–21.0 19.0–21.0 bal bal bal bal 37.0–39.0 27.5–29.5 15.0–17.0 ... 19.0–21.0 39.0–41.0 37.0–39.0 34.0–36.0 29.0–31.0

Cu

14.0–18.0 14.5–18.5 19.0–23.0 21.0–25.0 13.5–17.5 26.0–30.0 23.0–27.0 14.0–18.0 15.0–19.0 ... ... 13.0–17.0 8.0–12.0 4.0–6.0 ... ... ... 30.0–34.0 ... 21.0–25.0 ... 26.0–30.0 33.0–37.0 31.0–35.0 24.5–28.5 26.0–30.0

Zn

23.0–25.0 17.0–19.0 17.0–19.0 19.0–21.0 15.0–17.0 ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 12.5–14.5 ...

Cd

... ... ... ... 2.5–3.5 1.5–2.5 ... ... ... ... ... ... ... 0.5–1.5 1.5–2.5 ... ... ... 2.0–3.0 4.0–5.0 ... 1.5–2.5 ... 1.5–2.5 ... ...

Ni

Composition, %

... ... ... ... ... ... ... ... 4.5–5.5 ... ... ... ... ... ... 9.5–10.5 ... ... 5.0–7.0 ... ... ... ... ... ... 1.5–2.5

Sn

... ... ... ... ... ... ... ... ... ... 0.25–0.50 ... ... ... ... ... 0.15–0.30 ... ... ... ... ... ... ... ... ...

Li

... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... 7.0–8.0 bal ... 4.5–5.5 1.5–2.5 ... ...

Mn

0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15 0.15

Other elements (total)

607 627 607 607 632 671 677 688 618 779 766 671 690 713 771 601 760 677 690 682 960 660 738 707 607 649

°C

1125 1160 1125 1125 1170 1240 1250 1270 1145 1435 1410 1240 1275 1325 1420 1115 1400 1250 1275 1260 1760 1220 1360 1305 1125 1200

°F

Solidus temperature

618 635 701 710 688 779 743 774 651 779 766 713 738 857 893 713 885 766 801 699 971 707 790 801 746 710

°C

1145 1175 1295 1310 1270 1435 1370 1425 1205 1435 1410 1325 1360 1575 1640 1325 1635 1410 1475 1290 1780 1305 1455 1475 1375 1310

°F

Liquidus temperature

Table 4.15 Typical composition and properties of standard filler metals for brazing stainless steels (silver base)

618–760 635–760 701–843 710–843 688–816 779–899 743–843 774–871 651–760 779–899 766–871 713–843 738–843 857–969 871–982 713–843 877–982 766–871 801–899 699–830 971–1038 707–843 790–846 801–871 746–857 710–843

°C

1145–1400 1175–1400 1295–1550 1310–1550 1270–1500 1435–1650 1370–1550 1425–1600 1205–1400 1435–1650 1410–1600 1325–1550 1360–1550 1575–1775 1600–1800 1325–1550 1610–1800 1410–1600 1475–1650 1290–1525 1780–1900 1305–1550 1455–1555 1475–1600 1375–1575 1310–1550

°F

Brazing temperature range

Chapter 4: Base Metals and Base-Metal Family Groups / 109

Cr

Filler metal

BNi-1 BNi-1a BNi-2 BNi-3 BNi-4 BNi-5 BNi-6 BNi-7 BNi-8

13.0–15.0 13.0–15.0 6.0–8.0 ... ... 18.5–19.5 ... 13.0–15.0 ...

Nickel alloys

18.0–20.0

Cr

BCo-1

Cobalt alloys

Filler metal

BAu-1 BAu-2 BAu-3 BAu-4 BAu-5 BAu-6

Precious-metal alloys

Filler metal

BCu-1 BCu-1a BCu-2

Copper alloys

Filler metal

2.75–3.50 2.75–3.50 2.75–3.50 2.75–3.50 1.5–2.2 0.03 ... 0.01 ...

B

16.0–18.0

Ni

99.90 min 99.0 min 86.5 min

Cu

Au

Fe

3.5–4.5

W

4.0–5.0 4.0–5.0 4.0–5.0 4.0–5.0 4.0–5.0 2.5–3.5 4.0–5.0 0.5 3.0–4.0 1.5 9.75–10.50 ... ... ... 0.10 0.2 6.0–8.0 ...

Si

7.5–8.5

Si

Sn

... ... ...

37.0–38.0 79.5–80.5 34.5–35.5 81.5–82.5 29.5–30.5 69.5–70.5

... ... ...

Zn

B

P

0.7–0.9

bal bal bal ... ... ...

Cu

... ... ...

Mn

0.075 ... ...

P

... ... ... ... 33.5–34.5 7.5–8.5

Pd

0.02 0.02 0.02 0.02 0.02 0.02 0.02 0.02 0.02

S

0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05

Al

0.02

Composition, %

0.35–0.45

C

P

Composition, %

... ... ...

Ni

Composition, %

0.6–0.9 0.02 0.06 0.02 0.06 0.02 0.06 0.02 0.06 0.02 0.10 0.02 0.10 10.0–12.0 0.08 9.7–10.5 0.10 0.02

C

1.0

Fe

... ... ...

Fe

Composition, %

0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.04 0.05

Ti

0.05

S

Ni

0.01 ... ...

Al

... ... ... ... ... ... ... 0.04 21.5–24.5

Mn

0.05

Al

... ... 2.5–3.5 bal 35.5–36.5 21.5–22.5

0.02 ... ...

Pb

... ... ... ... ... ... ... ... 4.0–5.0

Cu

0.05

Ti

0.15 0.15 0.15 0.15 0.15 0.15

0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05

Zr

0.05

Zr

0.10 0.30 0.50

bal bal bal bal bal bal bal bal bal

Ni

bal

Co

Others (total)

Others (total)

... ... ...

Si

0.50 0.50 0.50 0.50 0.50 0.50 0.50 0.50 0.05

Others (total)

0.50

Others (total)

1980 1980 1980

°F

1815 1635 1785 1740 2075 1845

°F

2050

°F

977 977 971 982 982 1080 877 888 982

°C

1790 1790 1780 1800 1800 1975 1610 1630 1800

°F

Solidus temperature

1121

°C

Solidus temperature

990 890 974 949 1135 1007

°C

Solidus temperature

1082 1082 1082

°C

Solidus temperature

Table 4.16 Compositions and properties of standard filler metals for brazing steels (copper, gold, cobalt, and nickel base)

1980 1980 1980

°F

1860 1635 1885 1740 2130 1915

°F

2100

°F

1038 1077 999 1038 1066 1135 877 888 1010

°C

1900 1970 1830 1900 1950 2075 1610 1630 1850

°F

Liquidus temperature

1149

°C

Liquidus temperature

1016 890 1030 949 1166 1046

°C

Liquidus temperature

1082 1082 1082

°C

Liquidus temperature

2000–2100 2000–2100 2000–2100

°F

1860–2000 1635–1850 1885–1995 1740–1840 2130–2250 1915–2050

°F

2100–2250

°F

1066–1204 1077–1204 1010–1177 1010–1177 1010–1177 1149–1204 927–1093 927–1093 1010–1093

°C

1950–2200 1970–2200 1850–2150 1850–2150 1850–2150 2100–2200 1700–2200 1700–2000 1850–2000

°F

Brazing temperature range

1149–1232

°C

Brazing temperature range

1016–1093 890–1010 1030–1090 949–1004 1166–1232 1046–1121

°C

Brazing temperature range

1093–1149 1093–1149 1093–1149

°C

Brazing temperature range

Chapter 4: Base Metals and Base-Metal Family Groups / 111

however, and form phases with two undesirable characteristics: the phases are considerably less ductile than either the base metal or the filler metal even at elevated temperatures and thus are a potential source of rupture, and the alloys formed with stainless steel are higher-melting alloys that are likely to freeze and block further flow into the joint during brazing. Because of the relatively high brazing temperatures required for nickel-base filler metals, their use is generally restricted to furnace brazing in a controlled atmosphere (including vacuum), although there are occasional exceptions. The nickel-base filler metals normally are supplied as powders; however, they can be obtained as sintered or cast rods, preforms, plastic bonded sheet, plastic bonded wire, and tape. Many of these filler metals now are available as metallic foils produced by ultrarapid cooling from the molten state. The BNi filler metals commonly are used on stainless steels for oxidation resistance at temperatures up to 980 to 1095 °C (1800 to 2000 °F). Filler metals BNi-1, BNi-2, BNi-3, and BNi-4 (Table 4.16) tend to erode thin sheet metal because of their interaction with many base metals. Time at brazing temperature and quantity of filler metal should, therefore, be controlled carefully when these filler metals are used. Boron-free filler metals BNi-5, BNi-6, and BNi-7 are suitable for use in nuclear reactor components where boron cannot be tolerated because of its absorption of neutrons. Oxidation resistance is good up to 1095 °C (2000 °F). The BNi-5 material has the highest melting point of all the nickel-base filler metals. The gold-base filler metals (BAu group in Table 4.16) are sometimes used for brazing stainless steels, although their high cost restricts their use to specialized applications such as heat exchangers for manned space-flight vehicles. When a gold-base filler metal is used, alloying with the stainless steel base metal is minimized, and, as a result, joints exhibit good ductility (Ref 74). In addition to the gold family of filler metals, there are palladium and platinum filler metals, such as Au-Ni-Pd, Cu-Pt, Ag-Pd-Mn, and Pd-NiCr, which are useful for brazing heat- and corrosion-resistant components in the jet- and rocketpropulsion and nuclear-energy fields. The cobalt-base filler metal BCo-1 (Table 4.16) is very rarely used for brazing stainless steels, but it has been included in the table for its availability. Furnace Atmospheres for Brazing Stainless Steels. Stainless steels can be brazed with

any of the various brazing processes. A large volume of controlled-atmosphere brazing is being performed on stainless steels, and the success of this type of brazing can be attributed to the development of reliable atmosphere and vacuum furnaces. The protective atmospheres most often used in furnace brazing of stainless steels are dry hydrogen and dissociated ammonia. These atmospheres are effective in reducing oxides, protecting the base metal, and promoting the flow of filler metal. The low-cost exothermic atmospheres that are widely used in furnace brazing of low-carbon steels are not suitable for stainless steels. An inert gas such as argon, or vacuum, may be used to satisfy special requirements and provide protection in applications for which hydrogen or hydrogen-bearing gases are satisfactory. A dry hydrogen atmosphere is preferred for many stainless steel brazing applications. Hydrogen, the most strongly reducing of protective atmospheres, reduces chromium oxide and provides for excellent wetting by some filler metals without the need for flux. The principal disadvantages of hydrogen are high cost, difficulty in drying sufficiently, need for special furnace equipment, and danger involved in storing and handling hydrogen. Dissociated ammonia, when it is free of moisture and 100% dissociated, is a suitable atmosphere for brazing of stainless steel with some filler metals without the need for a flux. Dissociated ammonia is strongly reducing but less so than pure dry hydrogen. Consequently, although it promotes wetting by reducing chromium oxide on the surface of the stainless steel, dissociated ammonia may not be sufficiently reducing to promote the flow of some filler metals, such as copper oxide powders. Because of its high (75%) hydrogen content, dissociated ammonia forms explosive mixtures with air and must be handled with the same precautions as those required for handling hydrogen. Unless the atmosphere used in brazing of stainless steel is completely decomposed (100% dissociated), even minute amounts of raw ammonia (NH3) in the atmosphere nitride stainless steels, especially those containing little or no nickel. Finally, to avoid oxidation of base metal and filler metal, the dissociated ammonia atmosphere must be kept dry and pure while it is inside the furnace. Argon is occasionally used as a furnace atmosphere in brazing stainless steels to other

112 / Brazing, Second Edition

stainless steels or to reactive metals such as titanium. Argon has the advantage of being chemically inert in relation to all metals; thus, it is a useful protective atmosphere for metals that can combine with or absorb reactive atmospheres, such as hydrogen. An argon atmosphere has the disadvantage of being unable to reduce oxides; consequently, the surfaces of stainless steel components must be exceptionally clean and free of oxides when brazed in argon. The principal advantages of furnace brazing are high production rates and the means for using controlled protective atmospheres at controlled dewpoints, which often make it unnecessary to use a flux to obtain satisfactory wetting. In most furnace-brazing applications, both of these advantages are exploited. Occasionally, however, furnace brazing is selected solely on the basis of production rate, and brazing is performed without a protective atmosphere but with a suitable flux. The lower-melting filler metals are generally selected for brazing under these conditions. Vacuum brazing of many structural configurations made of austenitic stainless steels offers excellent heat and corrosion resistance for hightemperature-service applications. For stainless steels, the fundamentals of torch brazing, as well as its advantages and limitations, are basically the same as carbon steels. Because of the metallurgical characteristics of stainless steel and its requirements for corrosion resistance, however, best results are obtained when special consideration is given to the type of flame at the torch and to the filler-metal composition. To aid in reducing the oxide already present, as well as to prevent further oxidation of the work-metal surfaces, a strongly reducing flame should be used for torch brazing stainless steel to itself. A reducing flame is also satisfactory for brazing stainless steel to nickel alloys or carbon steels. In brazing stainless steel to copper alloys, however, some compromise is necessary. Although a slightly oxidizing flame is normally best for brazing copper, for brazing stainless steel to copper a slightly reducing flame usually provides a satisfactory compromise in flame adjustment. The silver-base filler metals that flow at relatively low temperatures are used almost exclusively for torch brazing of stainless steels. Filler metal BAg-3 is most often used, because it flows well in the temperature range from 705 to 760 °C (1300 to 1400 °F) and provides joints that have greater resistance to corrosion than

those brazed with filler metals such as BAg-1 or BAg-1a, although these filler metals have been used. Depending on the metallurgical and physical properties of particular stainless steels, their behavior in heating by electrical induction may differ considerably from that of carbon and lowalloy steels and from that of the more widely used nonferrous metals. In addition, depending on whether a stainless steel is magnetic or nonmagnetic at room temperature, the response of the steel to induction heating varies considerably. Differences in specific heat and electrical conductivity markedly affect response to heating by induction. Ferritic and martensitic (400-series) stainless steels are ferromagnetic at all temperatures up to the Curie temperature. The Curie temperature is the temperature of magnetic transformation below which a metal or alloy is magnetic and above which it is paramagnetic (Ref 72). Thus, given the same power input, these steels generally heat faster than austenitic stainless steels, which are nonmagnetic in the annealed condition. Stainless steels may be induction brazed in an air atmosphere, using a suitable flux, although for critical applications, induction brazing is sometimes done in a protective atmosphere or in vacuum, whereas in other applications, an inert gas such as argon may be used as a backing gas to minimize oxidation. Brazing of stainless steel by immersion of all or a portion of the assembly in molten salt offers essentially the same advantages that would apply to brazing similar assemblies of carbon steel. Similarly, the same limitations are applicable.

Maraging Steels An investigation (Ref 75) found that thin gold-nickel alloy (BAu-4, 57.5Au-42.5Ni, at.%) interlayer brazes between strong base metals were found to have very high ultimate tensile strengths as a result of the triaxial stress state within the braze (reducing the effective, or von Mises, stress) and the high strength of the filler metal. Maraging steel was joined with the previously mentioned filler metal in this study. Results of the study found that: • The transmission electron microscopy (TEM) of the gold-nickel interlayer brazes showed a very fine-grain (0.1 µm) two-phase

Chapter 4: Base Metals and Base-Metal Family Groups / 113

microstructure, with both phases coincidentally oriented. • The ultimate tensile strength of the goldnickel bonds was high, largely due to a high interlayer yield strength that resulted from the refined microstructure. • The yield strength and quasi-steady-state flow stress of the filler metal were found to be relatively strain-rate insensitive. These results explain the absence of timedependent, or delayed, tensile failures that have been observed in other braze interlayers. The fracture process appears to be one of cavity nucleation by microplasticity, followed by cavity coalescence or impingement once the required nuclei density is achieved. Substantial cavity growth does not appear to occur. This failure process contrasts with that of other interlayer metals in which the failure mechanism appears to be unstable cavity growth.

Other Stainless Steel Applications Another study (Ref 76) found that companies saved a considerable amount of money by salvaging porous stainless steel castings by using the cleaning potential of the furnace atmosphere, vacuum, and the free-flowing character of the filler metal to fill the pores by capillary action. These filler metals can also be used for building up surfaces, for example, to adjust the throat widths of gas turbine compressor stators made by precision casting or to replace metal badly machined from large stainless steel valve components used in North Sea oil submersible equipment. The nickel-base filler metals have corrosion resistance generally comparable with Inconel, while some of the filler metals have oxidation resistances that are satisfactory to 1100 °C (2010 °F). Additional research found and reported on hydrogen detector units comprising a 2.3 mm (0.09 in.) outside diameter by 1.2 mm (0.05 in.) wall thickness nickel tube helix brazed into a type 316 stainless steel pressure hub approximately 25 mm (1 in.) in diameter and 75 mm (3 in.) thick; this was an extreme case of difference in thermal mass of the components. Brazing was carried out with BNi-2 filler metal at 1020 °C (1870 °F) for 5 min. The joint had to be tested using β-dense neutron radiography as well as helium leak testing and airflow checks. Finally, the research discussed the many thousands of stainless steel/copper heat ex-

changers that have been brazed for domestic water heaters and the stainless steel tube thermometers for electric cookers and cooker hot plates. Knife handles have been brazed from stainless steel pressings, as have very much larger valve bonnets for industrial use. It is very significant that joining stainless steel assemblies for food-handling equipment requires strong ductile joints while preventing formation of pits or other imperfections that can trap bacteria and contaminate food. When developing an assembly process for a doughdivider head assembly used in a dough-forming machine, graphite-brazing fixtures were constructed for five designs of the divider. When installed, the ring-and-blade assembly divides dough into portions small enough to make dinner rolls and large enough to make bread loaves. The five designs of divider head range in number of parts from 6 to 36. Each design calls for a ring and band to hold several blades in place; all materials are of type 304 stainless steel. Nickelbase BNi-2 filler metal holds the joints. This filler metal has been used in many foodindustry applications where clean, smooth surfaces to deter bacteria formation and resist oxidation are important. Both the brazing process and the nickel-base filler metal used have been approved by the United States Department of Agriculture food safety and inspection service and the privately operated National Sanitation Foundation. Production begins by placing ring and blades in an aluminum fixture. Then, the band is placed around the assembly and tack welded to the ends of the outer blades. Operators then use gas tungsten arc welding to create joints between the ring and blades to hold them in place during furnace brazing. Filler metal is applied in paste form to each butt joint; a 36-part assembly comprises 50 such joints. Then, an operator removes the assembly from the fixture and places it in the specially designed graphite fixture for furnace brazing. Brazing the head takes 12 h in an 81.3 by 106.7 cm (32 by 42 in.) top-loading vacuum furnace. Heating, brazing, and slow-cooling cycles are strictly controlled to prevent distortion. The procedure holds the bottom of the part flat to within 0.203 mm (0.008 in.), and six assemblies are brazed per furnace run. Palladium-containing filler metals 30Pd60Cu-10Co, 30Pd-60Cu-10Ni and 30Pd-50Cu10Ni-10Co (all in wt%) were studied for use in brazing stainless steel SUS 316 L at high temper-

114 / Brazing, Second Edition

atures (Ref 77). Palladium-containing filler metals were selected because of their ductility, oxidation resistance, relatively high melting points, and lower cost than gold-base filler metals. Wettability, microstructures, and reactions were studied with experimental brazed joints between base and filler metal. Joint strengths, tensile strengths at room and elevated temperatures of 473 to 1073 K, and creep-rupture strengths at 673 to 1073 K were investigated. The joints brazed with the 30Pd-50Cu-10-Ni-10Co filler metal had superior tensile strengths at elevated temperature and creep-rupture strengths. The joints with this filler metal had tensile strengths of 327 to 429 MPa (47 to 62 ksi) at temperatures of 293 to 873 K. The solidus and liquidus of the previously mentioned filler metals are shown in Table 4.17 (Ref 77). Joints brazed with the 30Pd-60Cu-10Co filler metal had the highest tensile strength of the tested joints, and the mean tensile strength for joint clearances ranging from 0.07 to 0.15 mm (0.003 to 0.006 in.) was 536 MPa (78 ksi). Joints brazed with the 30Pd-60Cu-10Ni filler metal had the maximum tensile strength at 473 K, and joints brazed with the 30Pd-50Cu-10Ni10Co filler metal did at 673, 873, and 1073 K. Joints brazed with the 30Pd-60Cu-10Ni filler metal had the highest creep-rupture strength at 673 K, and joints brazed with the Ni-Cr-W-FeB-Si filler metal did at 873 and 1073 K (Ref 77). It was concluded that 30Pd-50Cu-10Ni-10Co filler metal is the most suitable filler metal for high-temperature brazing of SUS 316 L. A parallel research study (Ref 78) examined two types of palladium-containing filler metals, 40Pd-50Ni-10Co and 30Pd-50Ni-10Cu-10Co (all in wt%). Commercially available filler metals, palladium filler-metal braze 33Pd-64Ag3Mn, and nickel-base filler metal Ni-Co-CrMo-Fe-B were also used. The microstructures and strength properties of the brazed joints were studied as well as Table 4.17 Solidus and liquidus temperatures of filler metals studied for brazing of 316L stainless steel Solidus Filler metal

30Pd-60Cu-10Co 30Pd-60Cu-10Ni 30Pd-50Cu-10Ni-10Co Ni-Cr-W-Fe-B-Si Source: Ref 77

Liquidus

°C

°F

°C

°F

1090 1139 1130 1010

1994 2082 2066 1850

1100 1150 1160 1165

2012 2102 2120 2129

the newly developed palladium-containing filler metals 40Pd-50Ni-10Co, 30Pd-50Ni-10Cu10Co, and 33Pd-64Ag-3Mn. The joints brazed with the 30Pd-50Ni-10Cu10Co and 40Pd-50Ni-10Co filler metals showed the highest tensile strength of the tested filler metals, and the mean tensile strength of joint clearances ranging from 0.05 to 0.15 mm (0.002 to 0.006 in.) was 516 MPa (75 ksi). The 30Pd-50Ni-10Cu-10Co filler metal showed maximum tensile joint strengths at 873 K or below, and the 33Pd-64Ag-3Mn filler metal did at 1073 K. The brazed joints with the 30Pd-50Ni-10Cu10Co and 40Pd-50Ni-10Co filler metals had the highest creep-rupture strength with the time to failure of 1000 h or over at applied stresses of 425 MPa (62 ksi) at 673 K and of 125 MPa (18 ksi) at 873 K, respectively. The joints with the Ni-Co-Cr-Mo-Fe-B filler metal had good creeprupture strength with the time to failure of 1000 h at applied stress of 50 MPa (7.3 ksi) at 1073 K. The 30Pd-50Ni-10Cu-10Co filler metal for SUS 316 L was finally found to be the most suitable filler metal for high-temperature brazing, because this filler metal provided superior strength properties at room and high temperatures (Ref 79).

Titanium and Zirconium and Their Alloys Titanium and zirconium combine readily with oxygen and react to form brittle intermetallic compounds with many metals and with hydrogen and nitrogen. Parts must be cleaned before brazing and brazed immediately after cleaning. Titanium and Its Alloys. Silver and silverbase filler metals were used in early brazing of titanium, but brittle intermetallics were formed, and crevice corrosion resulted. Type 3003 aluminum foil as a filler metal joins thin, lightweight structures, such as complex honeycomb sandwich panels. Electroplating various elements on the base-metal faying surfaces lets them react in-situ with the titanium/titanium alloy base material during brazing to form a titanium alloy eutectic. That transient liquid phase flows well and forms fillets, then solidifies due to interdiffusion (Ref 69, 80). Other filler metals with high service capability and corrosion resistance include Ti-Zr-NiBe, Ti-Zr-Ni-Cu, and Ti-Ni-Cu filler metals. The best braze processing is obtained in high-

Chapter 4: Base Metals and Base-Metal Family Groups / 115

vacuum furnaces using closely controlled temperatures in the range of 900 to 955 °C (1650 to 1750 °F). Selection of filler metals for use in brazing of reactive metals is critical to avoid formation of undesirable intermetallic compounds. Selection of filler metals and brazing cycles that are compatible with the heat treatment required for (α-β) and β-titanium base metals may present some difficulty. Ideally, brazing should be conducted at temperatures from 55 to 83 °C (100 to 150 °F) below the β transus, because the ductility of (α-β) base metals may be impaired if this temperature is exceeded. The β transus can be exceeded when β-titanium base metals are brazed; however, if the brazing temperature is too high, base-metal ductility after heat treatment may be impaired by considerable interaction between the properties of the heat treatable titanium alloys and may be affected adversely by brazing, unless the assembly can be heat treated afterward. For example, the (α-β) titanium alloys must be solution treated, quenched, and aged to develop optimal properties. It is not easy to select a filler metal that permits brazing and solution treating in a single operation. Similarly, it is not always possible to quench a brazed assembly at the desired cooling rate, and certain configurations (e.g., honeycomb sandwich structures) do not lend themselves to rapid quenching without distortion. Brazing at the aging temperature is impractical, because few filler metals melt and flow at these temperatures. The possibility of galvanic corrosion must be considered when filler metals are selected for brazing titanium base metals. Although titanium is an active metal, its activity tends to decrease in an oxidizing environment, because the surface undergoes anodic polarization in a manner similar to that of aluminum. Thus, filler metals must be chosen carefully to avoid preferential corrosion of the brazed joint. When titanium is brazed, precautions must be taken to ensure that the brazing retort or chamber is free of contaminants from previous brazing operations. As a precaution against any existing contaminants in the brazing furnace, a loose cover of pure titanium foil (0.03 to 0.08 mm, or 0.001 to 0.003 in., thick) should be put over the workpieces. This acts as a getter of any remaining contaminants. Care should be exercised to ensure that the foil does not come into direct contact with the workpieces. The choice of materials to be used in fixtures

must be carefully considered. Nickel or materials containing high amounts of nickel generally should be avoided; nickel and titanium form a low-melting eutectic (28.4% Ni) at approximately 940 °C (1725 °F). Should the titanium workpieces contact fixtures or a retort made from nickel-base alloy, the parts may fuse together if the brazing temperature is in excess of 940 °C. If a fixture material containing a high nickel content (such as stainless steel) is used, it should be oxide coated. In most applications, coated graphite or carbon steel fixture materials are used. Filler metals initially used for brazing of titanium and its alloys were silver-base filler metals containing lithium, copper, aluminum, or tin (Ref 81–83). Most of these filler metals were used in low-temperature applications (540 to 595 °C, or 1000 to 1100 °F). Later developments produced a number of successful commercial filler metals, including Ag-Pd, Ti-Ni, Ti-Ni-Cu, Ti-Zr-Be (Ref 84), and Ti-Zr-Cu-Ni (Ref 85, 86). Additionally, these filler metals could be used at temperatures from 870 to 925 °C (1600 to 1700 °F). For joining applications requiring a high degree of corrosion resistance, the 48Ti48Zr-4Be and 43Ti-43Zr-12Ni-2Be filler metals are outstanding (Ref 84). A Ag-Pd-Ga filler metal (Ag-9Pd-9Ga), which flows at 900 to 915 °C (1650 to 1680 °F), is another excellent filler metal that fills large gaps (Ref 87). The following four filler metals are new developments now available for use in products (Ref 85): • • • •

Ti-Cu-Ni-2 (Ti-20Cu-20Ni) Ti-Zr-Cu-Ni-1 (Ti-38Zr-12Cu-12Ni) Ti-Zr-Cu-Ni-2 (Ti-38Zr-15Cu-15Ni) Ti-Zr-Cu-Ni-Pd (Ti-37Zr-12Cu-12Ni-2Pd)

Initial tensile and corrosion testing of joints with the latter filler metals shows that the joints exhibit nearly the same mechanical and chemical characteristics as Ti-6Al-4V. Brazing of titanium honeycomb sandwich structures using aluminum filler metal is an achievement stemming from supersonic transport materials and process technology (Ref 81). Such aircraft structures up to 7 m (23 ft) in length have been successfully fabricated using Al 3003 brazing foils as filler metal, which provide satisfactory strength up to approximately 315 °C (600 °F). High-strength, corrosionresistant Ti-Zr-Be and Ti-Zr-Ni-Be filler metals are recommended for applications from 540 to 595 °C (1000 to 1100 °F) (Ref 84).

116 / Brazing, Second Edition

Researchers (Ref 88) examined the effect of joint clearance as a critical parameter for maintaining joint integrity in brazed Ti-6Al-4V and commercially pure titanium (CPTi) materials. They determined that the mechanical properties of Ti-6Al-4V and CPTi joints brazed with Ti(Zr)Cu-Ni-(Pd) filler metals (Table 4.18) are strongly dependent on the microstructure of the brazing zone, especially with the formation of brittle intermetallic phases in the center of the brazing zone, which reduce the mechanical strength of the brazed joints. These phases form due to an exceeding of the maximum jointclearance size of the system. The formation of these phases can be avoided by minimizing the joint-clearance size of the joint to be brazed and by enhancement of brazing time (Ref 89). Tensile tests of the brazed Ti-6Al-4V and CPTi joints revealed that especially Ti-6Al-4V joints exhibit excellent mechanical properties, even comparable to those of the base metal, if the brazing zone is free of brittle intermetallic phases. Fatigue tests of the Ti-6Al-4V joints confirm the excellent mechanical properties. The fatigue properties of the Ti-6Al-4V joints approach that of the base metal at maximum stresses below 700 MPa (102 ksi). In contrast to the Ti-6Al-4V joints, fabrication of CPTi joints with mechanical properties comparable to those of the base metal was not possible. The reason can be seen in the much lower maximum joint clearance of the system CPTi/Ti-Zr-Cu-Ni with regard to the formation of intermetallic phases. Therefore, under the given circumstances (50 µm filler-metal foils), formation of intermetallic phases could not be avoided. This study also revealed that the induction brazing process is an interesting alternative to the vacuum furnace brazing process. Especially from the economic point of view, the induction brazing process under inert atmosphere had many advantages due to the short processing time. On the other hand, the short processing

time was responsible for much higher requirements on brazing joint clearance compared to the vacuum brazing process. However, fabrication of titanium joints with mechanical properties comparable to those of the base metal by induction heating were possible, when the maximum brazing joint clearance of the system was not exceeded. This required extremely small joint clearances of <30 µm in the case of Ti6Al-4V/Ti-Cu-Ni-Pd joints and <20 µm in the case of CPTi joints. Using an amorphous 25Ti-25Zr-50Cu filler metal, researchers (Ref 90) induction brazed Ti6Al-4V alloy using argon as a shielding gas. The brazing cycles were rather short; radio frequency inductor power was supplied for only 40 to 60 s. The tensile strength and ductility of the brazed joint were at the level of that of Ti-6Al4V base metal. It was also found that the postbrazing natural cooling rate of 12.5 mm (0.5 in.) diameter rod samples was sufficient to suppress the formation of undesirable brittle λ-Cu2TiZr Laves phase appearing in vacuum furnace brazing due to inherently low cooling rates. From the practical point of view, it was proven that induction brazing of compatible samples located in a simple closed chamber may be carried out as an effective and inexpensive process. Zirconium and Its Alloys. Zirconium-base metals of commercial importance are the pure metal and several alloys. The most commonly used zirconium alloy is Zircalloy, which contains small percentages of tin, iron, chromium, and nickel. These structural alloys are used for corrosion resistance in nuclear applications, especially in pressurized-water nuclear power reactors. Like titanium and beryllium, zirconium reacts readily with oxygen, hydrogen, and nitrogen and is embrittled. It also reacts with many metals and alloys to form intermetallic compounds. Therefore, as a result of this reactivity, zirconium must be brazed in a vacuum or in a

Table 4.18 Compositions and critical temperatures of filler metals evaluated for brazing commercially pure titanium and Ti-6Al-4V Chemical compositions, wt%

Solidus

Liquidus

Brazing temperature

Filler metal

Ti

Zr

Cu

Ni

Pd

°C

°F

°C

°F

°C

°F

TiCuNi TiCuNiPd TiZrCuNi

60 60 35

... ... 35

20 19 15

20 19 15

... 2 ...

923 923 832

1695 1695 1530

934 934 850

1715 1715 1560

950 950 870

1740 1740 1600

Source: Ref 88

Chapter 4: Base Metals and Base-Metal Family Groups / 117

dry atmosphere of argon or helium. Zirconium joint members must be cleaned carefully before brazing to remove oxides and other surface contaminants, and brazing should be done immediately after cleaning. Compared with the other reactive metals, very little research has been done to develop filler metals and brazing methods for joining zirconium and zirconium alloys (Ref 91). To a degree, this lack of research on brazing can be attributed to the availability of other joining methods. However, most commercial filler metals do not wet or flow well on zirconium base metals, nor are they metallurgically compatible with zirconium. In addition, many of these filler metals do not possess the corrosion resistance required in reactor environments. Research was conducted to develop filler metals for producing brazed joints in Zircalloy 2 that possessed good resistance to corrosion in pressurized water at 360 °C (680 °F). The data from these studies and from metallographic examinations of the brazed joints indicated that the following filler metals most nearly met the service requirements: Zr-5Be, Cu-20Pd-3In, Ni-20Pd10Si, Ni-30Ge-13Cr, and Ni-6P (Ref 91). Additional studies to develop improved filler metals for brazing Zircalloy base metals for use in water-cooled reactors were recently completed (Ref 92). Candidate zirconium-base and non-zirconium-base filler metals were formulated, used in vacuum brazing under pressures of 0.0013 to 0.00013 Pa (10–5 to 10–6 torr), and screened by wetting tests and corrosion tests in pressurized, high-temperature water (315 °C, or 600 °F). The following filler metals had acceptable corrosion resistance and mechanical strength: • Zr-50Ag (brazing temperature, 1520 °C, or 2770 °F) • Zr-29Mn (brazing temperature, 1380 °C, or 2515 °F) • Zr-25Sn (brazing temperature, 1730 °C, or 3145 °F) The Zr-5Be filler metal previously mentioned has been used extensively to braze zirconium base metals to themselves and to other metals (e.g., stainless steel). For example, zirconium sheet stock has been brazed with this filler metal using the following cycle: 10 min at 1005 °C (1840 °F), followed by 4 to 6 h at 800 °C (1470 °F). Because of its ability to wet ceramic surfaces, Zr-5Be has also been used to braze zirconium to uranium oxide and beryllium oxide.

As stated earlier, titanium and zirconium are both highly reactive metals. The brazing process, therefore, must not allow the joint surfaces to come in contact with air during heating. Induction brazing and furnace brazing in inertgas or vacuum atmospheres can be used successfully. Torch brazing of these base metals is difficult, requiring special precautions and techniques. Induction brazing of small, symmetrical parts is very effective, because the speed minimizes reactions between filler metal and base metal. For large, precise assemblies, furnace brazing is favored, because it allows uniformity of temperature throughout the heating and cooling cycle to be controlled readily. Titanium and zirconium assemblies frequently are brazed in high-vacuum, cold-wall furnaces.

Carbides and Cermets Carbides of the refractory metals tungsten, titanium, and tantalum, bonded with cobalt, have been used for years in cutting tools and dies. It is necessary to join these carbides to metal parts, particularly for cutting tools. Closely related materials called cermets have been developed and are ceramic particles bonded with various metals. Their high-temperature strengths are intermediate between those of the ceramic materials and the binder metals employed. Their greatest disadvantage is their brittleness. The term cermets generally describes titanium-base hardmetals. These cermets bridge the gap between carbides and ceramics. Cermets are a unique class of materials that could have excellent applications in areas such as saws and wear parts. They are harder and much more wear resistant than tungsten carbide. They do not have the impact resistance and toughness of tungsten carbide, so they are not recommended in highimpact applications such as sawmills or cabinet shops where material is hand fed. The cermets are composed entirely of titanium carbide (TiC) and titanium nitride (TiN). These are the same materials used to coat ordinary grades of carbide to make them more wear resistant. A cermet is not coated; instead, it is solid coating material. It is impractical to use coated carbides in saws, because the saws are reground and the coating would be lost in the first grinding. Typical carbide coatings are TiC and TiN. The XT3 grade is solid TiC and TiN, so there is no coating to wear off.

118 / Brazing, Second Edition

Cermets have been used successfully in clamped or indexable machining operations for decades. They have not been used in brazed applications because they have not been brazeable. There is now a method to successfully braze cermets to steel holders. Previous brazing attempts resulted in the material forming a glob in the center of the insert and an extremely weak bond. The new braze technology gives excellent wetting of the insert and excellent braze strength. A successful braze is one where the strength of the joint is in excess of the rupture strength of the material. The force to fracture the brazed cermet was approximately equivalent to the force required to fracture an unbrazed cermet. Additional research can improve the impact resistance of brazed cermet joints, as it has for brazed tungsten carbide joints. Cermets typically have a life of 3 to 5 times that of carbide. This is due to superior wear resistance. An XT3 grade of cermet has the approximate toughness of C7 carbide while being much harder. There is also a C50 grade of cermet that is much tougher. Cermets are now brazeable, with tensile strengths of up to 685 MPa (100 ksi), and can tolerate temperatures up to 750 °C (1380 °F). New developments permit cermets to be brazed with the same filler metals and for the same cost using the same equipment and techniques used for tungsten carbide. These newly developed processes use chemicals that are commonly available, allow for brazing using standard filler metals, and also use these at ambient temperatures. The process involves preparing the cermet body as follows: • Cleaning: A preferred method of cleaning is a cathodic electrocleaning process in an alkaline solution. • Surface etching: The part is treated in such a manner as to prepare the surface to bond with an intermediate material. Typically, this is a chemical bath and may or may not use electric current. The surface is roughened and chemically activated. The part is immersed in a solution for an etching effect. This process may be accelerated or enhanced by the use of electric current, higher temperature, or different combinations of chemicals. • Deposition of cobalt or other metal: The surface is plated with a metal or other material that bonds to the ceramic as well as forms a layer suitable for brazing with standard filler

metals. Typically, this is also a chemical bath and may or may not use electric current. It is beneficial if non-cyanide nitrogen is introduced so as to create intermediate metallic compounds. Cyanide compounds may be used, but they are unnecessary and significantly add to the cost in several ways. • Postcleaning and passivation: This treatment of the parts may be desirable. The parts are then brazed on saws using standard filler metals (such as BAg-24 and those similar to it) and are used successfully in sawing and other applications (Ref 93) (see Chapter 10, “Applications and Future Outlook”). Brazing carbides and cermets is generally more difficult than brazing metals. Torch, induction, or furnace brazing is used, often with a sandwich brazing technique; a layer of weak, ductile metal (pure nickel or pure copper) (Ref 94) is interposed between the carbide or cermet and a hard metal support. The cooling stresses cause the soft metal to deform instead of cracking the ceramic (Ref 95). Silver-base filler metals, copper-zinc filler metals, and copper are often used on carbide tools. Although it is possible to use any of the BAg-1 through BAg-7 filler metals, those containing nickel (BAg-3 and BAg-4) are generally recommended, because nickel improves wettability. The RBCuZn-D and BCu filler metals also have been used, particularly where a postbraze heat treatment is required. The BCu filler metal retains practically all of its strength up to a temperature of 540 °C (1000 °F); however, it requires a hydrogen atmosphere furnace for best brazing results. The 85Ag-15Mn and 85Cu15Mn filler metals are used where the brazed joint is subjected to elevated temperatures in service and also for wetting the titanium-base or chromium-base carbides. Additionally, the BNi filler metals with boron and a 60Pd-40Ni filler metal have been used successfully for brazing nickel- and cobalt-bonded cermets of tungsten carbide, titanium carbide, and niobium carbide. Tungsten-base carbides generally are readily wetted by the BAg and BCu filler metals. However, titanium-base carbides are somewhat more difficult to wet. Where it is necessary to mount TiCs by brazing, either the joint must be made in an inert or vacuum atmosphere or the surfaces must be specially treated by the carbide manufacturer, or by nickel plating. In each case, the filler metals just mentioned then wet the TiC surface. In specifying a filler metal for a given job,

Chapter 4: Base Metals and Base-Metal Family Groups / 119

first consideration should be given to the temperature range of the anticipated application. This and the other considerations of corrosion and mechanical properties dictate the alloy composition selected, the brazing temperature required, the equipment to be used, and the joining atmosphere. Generally speaking, the filler metals mentioned previously are considered for the majority of applications that require simple equipment, fluxes, and brazing temperatures below 980 °C (1800 °F). The carbides, in general, are not wetted as readily by filler metals as are most base metals; thus, it is preferable, when possible, to preplace shims in the joint rather than to face feed the filler metal in the form of wire. For larger surfaces, shims having a core of copper or nickel with a coating of the filler metal on both surfaces are frequently used. The core of the shim generally accounts for approximately 50% of the total thickness, with the filler metal providing another 25% on each side. Filler metals can provide a cushioning or a shock absorber effect that actually helps the carbide survive in use. This is generally accomplished in one of three ways. The first is through the use of a soft filler metal, such as 50% Ag with cadmium. This is a pure cushioning effect. The second way is through the use of manganese. Manganese is a hard material, but it has the property of absorbing impact. This is more similar to a shock absorber. The third method is the use of trimetal sandwiches. These are typically filler metal/copper/filler metal. During the brazing process, the copper anneals to a dead soft condition. The cadmium alloys work well, but the cadmium is a serious health risk; more importantly, it is a health risk that the government really targets. Cadmium fumes at low temperature escape from the filler metal as it is being heated. It is also fairly soluble and dissolves into the grinding coolants when the saws and tools are ground. Finally, cadmium is very readily detectable, so it is easy to find. Manganese makes a good alloy for filler metal. However, the manganese makes the alloys a bit brittle, which means it is harder to draw into wire or roll into ribbons. The trimetal sandwiches are laminated together, which makes them very expensive. In addition, they have to be cut and bent to make brazing preforms. The metal sandwiches have to be preplaced into the joint. Three practices that can lead to improvements are recommended (Ref 93). First, the use

of a high-silver filler metal with manganese increases the bond strength and greatly increases the impact resistance. It also helps prevent thermal stress fractures. Secondly, the development of a cleaning and surface-treatment process cleans the carbide completely and then selectively removes the cobalt, which leaves the surface rough, because it is all exposed tungsten carbide grains. Finally, this surface is extremely wettable and provides an excellent gripping surface for the filler metal while also promoting the intermetallic joining characteristics for successful bonding.

Ceramics The increasing use of ceramic-to-ceramic and ceramic-to-metal joints in industrial and developmental applications is due to the unique combination of properties of ceramic materials (Ref 96–100). The large use of ceramic-to-metal joints in vacuum tubes in the electronics industry stems from the following properties: • Ceramic tubes can be outgassed at higher temperatures than glass tubes. • Ceramic tubes can withstand higher temperatures than glass tubes of similar dimensions. • Ceramic tubes are mechanically stronger and less sensitive to thermal shock than glass tubes. • Ceramic components can be ground to the precise tolerances required for vacuum-tube construction. • Ceramic materials have very low electrical losses at high frequencies. Because of their inertness in many corrosive environments, ceramics are used as seals in fuel cells and other devices that convert chemical, nuclear, or thermionic energy to electricity. Ceramics are also used as friction materials for brakes, clutches, and other energy-absorbing devices; coatings for nuclear fuel particles; constituents in high-temperature adhesives; radomes used to enclose antennae; and ablative materials. Glass-to-metal seals have been made for many years in the vacuum-tube industry, and the experience thus obtained gives a general insight into the problem of fabricating ceramic-to-metal joints (Ref 98). More recently, the fabrication of refractory-tipped tools, vacuum tubes, and various experimental devices has added to the available knowledge and techniques. Glass-Ceramic Joining. There are two types of intermediate bonds that can be em-

120 / Brazing, Second Edition

ployed, and the selection of which to use is influenced by a number of practical features, including the ease of fabrication, the temperature-withstand capabilities required, whether or not the bond must possess insulating properties, and the cost. The bonds can be either metallic (e.g., a braze) or vitreous (glass or glass-ceramic). In this type of bond, the usual design configuration is that of a butt seal in which the glass-ceramic components are joined on their end faces to the metallic components, although other types of seal, such as sleeve or concentric seals, are possible. The glass-ceramic body components are initially prepared using glass-forming methods and are then heat treated. The butt type of seal usually requires that the end faces of the glassceramic components are machined flat and parallel, and other surfaces, which are very often inner and outer cylinder surfaces, are machined to tight tolerances to fit the metal parts. The metal parts themselves also require preparation before assembly. In butt seals, the sealing faces must be flat to ensure good mating with the glass-ceramic seal faces. The preparation given to the metal prior to sealing depends on the type of intermediate bond to be employed—a metallic braze or a vitreous bond. In the case of a braze metal bond, the metal seal parts are required to have smooth sealing surfaces and to be clean and free from oxidation and other defects. The type of braze normally employed with a glass-ceramic is an active metal braze. In the butt seal design, thin washers of titanium and copper-silver eutectic are

Fig. 4.9

Exploded view of the constituent parts for an active metal brazed glass-ceramic/metal seal. Source: Ref 101

assembled together with the major components, as illustrated in Fig. 4.9. The titanium is placed adjacent to the glass-ceramic, and the coppersilver eutectic adjacent to the metal, the whole being in suitable jigs to provide and maintain alignment. Alternatively, a specially developed Cu-Ag-Ti filler metal can be used (Ref 102). The firing is carried out in a vacuum in order to prevent any oxidation of metallic constituents, which would inhibit wetting and flow of the braze at a temperature of approximately 850 °C (1560 °F). The titanium serves to react with the glass-ceramic during this firing, metallizing the surface and permitting reaction with the braze, which in turn wets and bonds to the metal surface. This type of bond permits subsequent processing (e.g., vacuum bakeout) and operation to temperatures in the region of 800 °C (1470 °F). Higher temperature-withstand capability can be achieved by using more refractory brazes, such as those containing nickel. However, it will be appreciated that the limitation may become the glass-ceramic component itself, although glass-ceramics are available that withstand temperatures well in excess of 1000 °C (1830 °F). It should be noted that the more-refractory glass-ceramics generally have relatively low thermal expansion, and consequently, matched expansion systems are only possible with the lower-expansion metals and alloys, for example, molybdenum and tungsten. It should be borne in mind that work at elevated temperatures may require the presence of an inert atmosphere to prevent oxidation and deterioration of metal parts, including the braze. In the case of a vitreous bond (glass or glassceramic), the sealing faces of the metallic component are roughened by means of grit blasting or etching, and the components are preoxidized by heating in a suitable atmosphere. The metal sealing faces are then coated with the vitreous bond, which is applied in powder form via a suitable suspension, usually by spraying. The coating is allowed to dry and is then fired, often in a protective atmosphere to protect the metal from excessive oxidation, to form a continuous adherent layer on the metal sealing face. It is usually unnecessary to preglaze the glassceramic sealing faces, although this can be done in a similar manner, if it is considered desirable. The various components are assembled, as shown in Fig. 4.10, for a butt seal and are then fired to cause the bonding materials to complete reactions with the glass-ceramic and metal components and so form the desired bond. If a glass-

Chapter 4: Base Metals and Base-Metal Family Groups / 121

ceramic bond is employed, the firing schedule also causes the glass to crystallize. A glass bond permits subsequent processing and operation up to temperatures of approximately 500 °C (930 °F), depending on the glass used, and a glassceramic bond enables temperatures in excess of 800 °C (1470 °F) to be withstood, again depending on the glass-ceramic used. The use of vitreous bonds is advantageous where more than two components are required to be bonded together in such a manner that sequential bonding operations are required. It is necessary in these circumstances, while maintaining thermal-expansion-matching/settingtemperature criteria, that the second and subsequent glasses are used at successively lower temperatures so that the firing conditions do not disturb the preceding glass and thereby allow relative movement of the previously bonded components. Fortunately, glasses are available that permit this sequential operation to be achieved, particularly for higher thermal expansion systems. If necessary, a glass-ceramic or braze can be used to provide the initial, more refractory bond. Examples of seals produced by direct bonding methods, such as in-mold shaping and graphite jig shaping, include: • Glass-ceramic-molybdenum high-voltage feed through seal, which comprises a glassceramic of the ZnO-Al2O3-SiO2 type that is matched in expansion and bonded to the molybdenum parts to form a concentric design (Ref 103)

Metal Pre-glazed layer

Glass-ceramic

Fig. 4.10

Exploded view of the constituent parts for glass- or glass-ceramic-bonded glass-ceramic/metal seal. Source: Ref 101

• Glass-ceramic-metal leadthrough seals (e.g., in thermocouple terminations). In this type of seal assembly, the glass-ceramic is matched in expansion to the metal(s) of the leadthroughs. In a thermocouple termination, these are Chromel and Alumel (Hoskins Manufacturing Company). In other types of light-current leadthrough seals, the leadthrough metal is often a nickel-iron alloy, for example, 48%NiFe. In this type of seal, a higher-expansion metal or alloy, for example, mild steel or a stainless steel, is frequently used for the outer metal components. A fully hermetic bond is developed between the glass-ceramic and the outer metal, and, owing to the differential expansion, or more precisely contraction, of the two components, the outer joint is in compression, which increases the overall strength of the assembly and its ability to resist thermal shock. Clearly, the deliberate introduction of compressive stresses into the outer glassceramic influences the design of the seal assembly. Challenges in Brazing of Ceramics. Ceramic materials are inherently difficult to wet with conventional filler metals. Most of these filler metals merely ball up at the joint, and little or no wetting occurs. When bonding does occur, it can be either mechanical or chemical. The strength of a mechanical bond can be attributed to interlocking particles or penetration into surface pores and voids, whereas a chemical bond derives strength from material transfer between the filler metal and the base material. Discussions of bonding mechanisms can be found in the literature (Ref 101–109). Another basic problem in brazing of ceramics results from the differences in thermal expansion between the base material and the filler metal and, in the case of ceramic-to-metal joints, between the two base materials. In addition, ceramics are poor conductors of heat, which means that it takes them longer to reach equilibrium temperature than it does metals. Both of these factors may lead to cracking in the joint. Because ceramics generally have lower tensile and shear strengths, crack propagation occurs at lower stresses in ceramics than in metals. In addition, the low ductilities permit very little distribution of the stresses set up by stress raisers. Alumina, zirconia, beryllia, thoria, forsterite (Mg2SiO4), and silicon carbide and nitride are the leading ceramic materials that can be joined by brazing.

122 / Brazing, Second Edition

and that use temperatures exceeding 750 °C (1380 °F) (Ref 117). Further work (Ref 118) showed the development of strong, reliable joints containing ceramic components for applications in advanced heat engines. This work was focused on the joining of Si3N4 by brazing. The technique of vapor coating ceramics to circumvent wetting problems that was developed for brazing zirconia at low temperatures was applied to brazing Si3N4. An acceptable porosity-free joint was obtained with 50Au-25Ni-25Pd filler metal. The joints made with the 50Au-25Ni-25Pd (wt%) filler metal were brazed at 1250 °C (2280 °F) (approximately 120 °C, or 216 °F, above the normal brazing temperature for the Au-Ni-Pd filler metal). This work also showed that the 50Au-25Ni25Pd (wt%) filler metal is particularly well suited for brazing titanium-vapor-coated Si3N4. Figure 4.11 shows that for this material combination, high joint strength was maintained up to 700 °C (1290 °F) before dramatic loss at 800 °C (1470 °F). Investigations continue into the apparent dependence of joint strength on the filler-metal thickness, with the tendency for improvement with reduced braze-layer thickness (from 40 to 50 µm to 10 to 20 µm). Other efforts (Ref 119) to develop filler metals for Si3N4 components for the Advanced Turbine Technology Applications Project (ATTAP) engine have included the development of filler metals for ceramic-to-metal joints, including the investigation of the different alloy systems Au-Ni-Cr-Mo-Fe-Nb, Au-Ni-Cr-Fe, Au-Ni-Cu-Cr-Mo-Fe-Nb, and Au-Ni-Cu-CrFe. The effort was to increase the high-temper500 (72.5)

Joint strength, MPa (ksi)

If the ceramic is premetallized to facilitate wetting, copper, silver-copper, and gold-nickel filler metals can be used. Titanium or zirconium hydride can also be decomposed at the ceramicmetal interface to form an intimate bond. Nonmetallized ceramics have been brazed with silver-copper-clad or nickel-clad titanium wires and other useful titanium and zirconium filler metals, such as Ti-Zr-Be, Ti-Zr-V, Zr-VNb, Ti-V-Be, and Ti-V-Cr (Ref 65, 96, 101, 110–113). Silicon Nitride (Si3N4). One process for brazing Si3N4 with metallic alloys involves vapor coating the ceramic with a 1.0 µm thick layer of titanium before the brazing operation. The coating improves wetting of the Si3N4 surfaces to the extent that strong bonding between the solidified filler metal and the ceramic occurs. Braze joints of Si3N4 are made with AgCu, Au-Ni, and Au-Ni-Pd filler metals at temperatures of 790, 970, and 1130 °C (1455, 1780, and 2065 °F) (Ref 114). A study also showed that Si3N4 surfaces are not easily wet by common precious-metal-based filler metals, but that vapor coating the surfaces with titanium prior to brazing improved their wetting characteristics and permitted filler metals to adhere strongly to the ceramic. The ability of vapor coatings, particularly of titanium, to improve the wetting characteristics of oxide ceramics has been known for some time (Ref 115, 116). Using a metallic vapor coating rather than incorporating an active element such as titanium directly into the filler metal permitted readily available commercial filler metals to be used for joining. The Si3N4 ceramics have recently been joined by an oxynitride brazing method. It was useful for large-sized parts or ceramics with complex geometries in which the high-temperature mechanical properties of the original ceramic components must be maintained. Joined parts can withstand temperatures that are 300 to 400 °C (540 to 720 °F) higher than conventionally joined parts can tolerate. Joint strength is comparable to the 400 MPa (58 ksi) of the Si3N4 ceramic material. If joint strengths do not need to exceed 200 MPa (29 ksi), then these joints can be used at temperatures as high as 1250 °C (2280 °F). These large or geometrically complex parts are used in gas turbine and adiabatic diesel engines, magnetohydrodynamic generators, pumps, and numerous other applications that require high strength and corrosion resistance

400 (58)

300 (43.5)

200 (29)

100 (14.5)

0 0

200 (390)

400 (750)

600 (1110)

800 (1470)

Temperature, °C (°F)

Fig. 4.11

Variation of joint strength with test temperature for titanium-vapor-coated SN220 Si3N4 lap joints brazed with Au-25Ni-25Pd wt% filler metal. Source: Ref 118

Chapter 4: Base Metals and Base-Metal Family Groups / 123

ature strength with a minimum loss of other properties necessary for brazing. Of those alloys, Au-Ni-Cr-Mo-Fe-Nb alloys, which melt at 1050 to 1100 °C (1920 to 2010 °F), seemed to meet the processing requirements for high-temperature ceramic-metal joints. Preliminary results from mechanical testing indicated that these alloys had good ductilities and strengths. They wetted well on Si3N4 with various coating materials in the vicinity of 1150 °C (2100 °F). Researchers (Ref 120) conducted an evaluation of brazed Si3N4 joints, including their microstructure and mechanical properties. The study involved Cusil ABA (active brazing alloy) (Wesgo Metals), an Ag-35Cu-1.6Ti filler metal with a liquidus of 815 °C (1500 °F), which was used to produce Si3N4/Si3N4 (SN/ SN) joints with room-temperature strengths similar to that of the monolithic ceramic. Brazing was achieved by holding at 850 °C (1560 °F) for 30 min under vacuum, using 50 µm thick filler-metal foil. Silicon nitride/AISI 316 steel joints were also produced by brazing with Cusil ABA when interlayers were used. Molybdenum was more successful as an interlayer than copper or niobium; its coefficient of thermal expansion matched that of the ceramic, and it had less effect on the reacting route for the joint formation. Joint strengths were higher for smaller workpiece bonded areas. Failure initiated predominantly from sites at the interfaces between the braze and the ceramic or metal workpieces. An investigation (Ref 121) reported on a joining method of ceramics-to-ceramics developed with a target goal of retaining high-temperature flexural strength with more than 80% of that of base materials. To join the Si3N4 material, the system of Si3N4/yttrium oxide/alumina was selected for the base material. The sintered Si3N4 material was joined by using the powder mixture of the grain-boundary phase composition. The paste of the powder mixture of the grain-boundary phase was screen painted on the joining surfaces of the Si3N4 material. The joining was carried out by firing them in a nitrogen atmosphere and hot isostatic pressing (HIP). Optimizing the conditions, the strength of joined specimens resulted in 800 and 640 MPa (116 and 93 ksi) at room temperature and 1250 °C (2280 °F). Silicon nitride braze joints have reportedly been produced in a single step at elevated tem-

peratures without premetallizing at the University of Illinois, Chicago. Reactive brazing with a filler-metal composition of Co-10Ti was used to join Si3N4 components; the brazing process involved a 15 min hold at 1300 °C (2370 °F) under a high vacuum. Active metal brazing with Ag-Cu-Ti alloys is currently used in the automobile industry to fabricate Si3N4-metal components, and no initial metallization of the ceramic surface is required. Nonetheless, because silver- and copper-base filler metals cannot withstand service temperatures above 500 °C (930 °F), filler metals with higher melting points must be added to really utilize the high-temperature potential of Si3N4. However, during tests with Co-10Ti filler metal, the reaction layer that developed was not uniform. While thicker areas were stable, the thinner areas showed some degradation, and regardless of thickness, gaps were observed between the braze metal and the ceramic substrate. This was caused by the premature isothermal solidification of the braze as it was depleted of titanium. Brazing with Ni-13Cr-Hf was also tested but was only possible with the Si3N4 premetallized. Hafnium from the braze metal moved toward the existing reaction layer, while copper dissolved in the matrix. Silver moved from the existing reaction layer toward the braze metal but did not dissolve; it remained as islands of pure silver (Ref 122). In a study (Ref 123) conducted to improve the joining strength of Si3N4 ceramics, pure nickel and 80wt%Ni-20wt%Cr alloy laminated interlayers were used for joining. Thin, pure nickel layers in contact with the Si3N4 ceramic were used for the following reasons. The gaps between pure nickel and Si3N4 ceramics closed easily with small joining pressure in the early stage of the joining process, because the flow stress of pure nickel is smaller than that of the nickel-chromium alloy. Then, the nickel and Si3N4 joined. Furthermore, chromium diffuses in pure nickel layers toward the joining interfaces and forms chromium nitrides there. The decomposition of the nitrides is suppressed by their interception from the argon joining atmosphere because of the closed gaps. The researchers concluded that the positive effects of thin, pure nickel layers placed in contrast with Si3N4 on the joining strength of Si3N4ceramics with nickel and nickel-chromium alloy laminated interlayers were confirmed experimentally. This joining technique can also

124 / Brazing, Second Edition

be applied for the fabrication of high-strength metal-Si3N4 joints. A team of researchers (Ref 124) assessed the mechanical properties of Si3N4 braze joints after fabrication and following exposure to elevated temperatures. These tests were conducted at room temperature and using four-point bending. The filler metal used in the investigation was palladium-base 58.2Pd-38.8Ni-3.0Ti. The Si3N4 substrate was coated with a Ag-Cu-In-Ti filler metal prior to fabrication of the braze joint. Two different sample geometries were used in the fabrication of the braze joints. Weibull statistics were applied to the interpretation of the data, but the results contradict what is expected from monolithic fracture theories. Oxidation was found to be significant when the braze joints were exposed to 800 °C (1470 °F) in air, which reduced the joint strength. An interlayer consisting of titanium, silver, and copper (0.1 mm, or 0.004 in., thick titanium foil; 0.1 mm thick silver foil; and 0.5 mm, or 0.02 in., thick copper plate) was sandwiched between two such Si3N4 rods in the symmetrical arrangement of titanium/silver/copper/silver/titanium, with copper at the center, and then treated for 60 min at 1273 K and 2 MPa (0.3 ksi) to form a SN/SN joint. The previously mentioned procedure was used to investigate