Refractory Alloy Technology for Space Nuclear Power Applications

$Refractory Alloy Technology for Space Nuclear Power Applications$

CONF-8308130 (DE84001745) January 1984 Distribution Category UC-25

DE04 001745

REFRACTORY ALLOY TECHNOLOGY FOR SPACE NUCLEAR POWER APPLICATIONS

Proceedings of a symposium held at Oak Ridge, Tennessee August 10-11,1983

Sponsored by SP-100 Project Office

Edited by R. H. Cooper, Jr. E. E. Hoffman

January 1984

Published by Technical Information Center Office of Scientific and Technical Information United States Department of Energy Conference Organization

GENERAL CHAIRMAN W. 0. Harms, Oak Ridge National Laboratory

PROGRAM ORGANIZERS AND TECHNICAL PROGRAM CHAIRMEN R H. Cooper, Jr., Oak Ridge National Laboratory E. E. Hoffman, U. S. Department of Energy, Oak Ridge Operations

ADVISORY COMMITTEES Compatibility R L. Dairies, NASA Lewis Research Center N. J. Hoffman, Rockwell International, Energy Technology Engineering Center P. Roy, General Electric Company. Advanced Reactor Systems Department S. A. Shiels, Westinghouse Electric Corporation, Advanced Energy Systems Division

Processing and Production G. C. Bodine, Combustion Engineering, Inc. R L. Heestand, Oak Ridge National Laboratory R E. McDonald, Consultant R A. Perkins, Lockheed Missiles and Space Company, Inc., Palo Alto Research Laboratory W. E. Ray, Westinghouse Electric Corporation, Advance 1 Energy Systems Division

Welding and Component Fabrication E. A. Franco-Ferreira, Consultant Services W. C. Hagel, Climax Molybdenum Company of Michigan T. A. Moss, Rockwell International, Energy Systems GiOup G. M. Slaughter, Oak Ridge National Laboratory, Metals and Ceramics Division

Mechanical and Physical Properties R L. Amman, Westinghouse Electric Corporation, Advanced Energy Systems Division W. C. Hagel, Climax Molybdenum Company of Michigan H. E. McCoy, Oak Ridge National Laboratory Effects of Irradiation M. L. Bleiberg, Westinghouse Electric Corporation, Advanced Energy Systems Division J. W. Davis, McDonnell Douglas Astronautics Company East R E. Gold, Westinghouse Electric Corporation, Advanced Energy System? Division /. Moteff, University of Cincinnati

CONFERENCE STAFF Mary Helen Owens, Oak Ridge National Laboratory Lyn Elrod, Oak Ridge National Laboratory Bonnie Reesor, Oak Ridge National Laboratory CONF-8303130 (DE84001745) January 1984 Distribution Category UC-25 Refractory Alloy Technology for Space Nuclear Power Applications

ABOUT THE TECHNICAL INFORMATION CENTER ABOUT THIS PUBLICATION The Office of Scientific and Technical Information, Early in 1983 it became apparent to staff members at the Technical Information Center in Oak Ridge, Tennessee, Oak Ridge National Laboratory and the DOE Oak Ridge has been the national center for scientific and technical information for the Department of Energy (DOE) and its Operations Office thai, because of the renewed interest in predecessor agencies since 1946. In developing and manag- high-performance space nuclear power systems, a mecha- ing DOE's technical information program, the Center nism was needed to make the results of earlier research places under bibliographic control not only DOE- originated information but also worldwide literature on on candidate refractory alloys for these systems available scientific and technical advances in the energy field and to the SP-100 project. Because of the rapid termination of announces the source and availability of this information. work on these alloys for space nuclear applications in Whereas the literature of science is emphasized, coverage is extended to DOE programmatic, socioeconomic, envi- 1972 and 1973, much of the valuable data on these ronmental, legislative/regulatory, energy analysis, and materials was only marginally documented. In many cases policy-related areas. To accomplish this mission, the the progress or topical reports received very limited distri- Ginter builds and maintain:, computerized energy- iRiormation data bases and disseminates this information bution. During the period from 1973 to 1983 many of the via computerized retrieval systems and announcement pub- refractory alloy technologists who were involved in the lications such as abstracting journals, bibliographies, and systems development efforts of the 1960s have retired or update journals. Direct access to the Center s most com- preiensive data base, the Energy Data Base, is available changed their specialty areas. to the public through commercial on-line bibliographic The conclusion was made that a publication comprised retr eval systems. The Energy Data Base and many of the Csn.er's energy-related data bases are available to DOE of review papers by experts familiar with the work done offices and contractors and to other government agencies mostly during the 1960s and including more recent work via 3OE/RECON, the Department's on-line information was needed. The focus of the plan was to achieve two retrieval system. The Center has developed ard maintains systens to record and communicate energy-related goals. resea'ch-in-progress information, to maintain a register of • To review and document- the status of refractory DOE public communications publications, to track ••eseai ch report deliverables from DOE contractors, and to alloy technology for structural and fuel-cladding test end make available DOE-funded computer software applications in space nuclear power systems. programs with scientific and management applications. • To identify and document the refractory alloy The Center also maintains a full-scale publishing capability to serve special publication needs of the Department. research and development needs for the SP-100 Pro- To e'fectively manage DOE's technical information gram in both the short and the long term. resources, the Center's program is one of continual devel- opmen and evaluation of new information products, sys- The technical program staff and the editors of these tems, and technologies. proceedings wish to thank the many individuals who contributed to the success of the symposium and to the prep- UNITED STATES DEPARTMENT OF ENERGY aration of these proceedings. We sincerely hope that the Donald Paul Hodel information contained herein is used to expedite progress Secretary in this challenging technology. Martha O. Hesse Assistant Secretary Management and Administration

William S. Heffelfinger Director of Administration

Joseph G. Coyne Manager Office of Scientific and Technical Information

Technical Information Center, Office of Scientific and Technical Information Um>ed States Department of Energy, P. O. Box 62. Oak Ridge, TN 37831 Introduction It has been my pleasure to serve as general chairman for the Symposium on Refractory Alloy Technology for Space Nuclear Power Applications. The sponsor of the symposium is the SP-100 Program, a triagency endeavor involving the Office of Aeronautics and Space Technology of the National Aeronautics and Space Administration, the Office of Nuclear Energy of the Department of Energy, and the Defense Advanced Research Projects Agency of the Department of Defense. The SP-100 designation is appropriate because the program covers two classes of nuclear devices for space applications: a 100 KW(e) class and a muitimegawatt [as high as 100 MW(e)] class. For the present time, the major effort by far is on the 100 KW(e) class, although planning is already unde1" way on muitimegawatt concepts. The purpose of this symposium is twofold: (1) to review and document the status of refractory alloy technology for structural and fuel-cladding applications in space nuclear power systems, and (2) to identify and document the refractory alloy research and development needs for the SP-100 Program in both the short and the long term. As indicated, the two key words are status and needs. Nuclear fuel systems, per se, are not included in the scope of the symposium. Only the fuel cladding has been considered, principally from the standpoints of radiation damage and compatibility with coolants and working fluids. In organizing the symposium, an effort was made to recapture the space reactor refractory alloy technology that was essentially cut off in midstream around 1975 when a substantial national space nuclear reactor program, which began in the early 1960s, was terminated. An important product of this symposium will be the identification of the R&D needs of the SP-100 Program for both the short and the long term. The six technical areas covered in the program are compatibility, processing and production, welding and component fabrication, mechanical and physical properties, effects of irradiation, and machinability. The refractory alloys considered, in order of increasing refractoriness, are niobium, molybdenum, tantalum, and tungsten. The papers have had a rather thorough presymposium review. As shown on page ii of this volume, advisory committees were established for each of the technical areas. The members of these advisory committees are selected experts, and they had the opportu- nity to review the papers. A design needs panel, made up of representatives of industrial concep design teams and the SP-100 Project Office, also reviewed the papers. Workshops were held the day before the symposium at the Oak Ridge National Labora- tory mainly for the purpose of developing a consensus concerning the R&D needs. We believe that the process of bringing the symposium to fruition has been worth- while and trust that the proceedings contained herein will serve a continuing useful purpose in this nation's space nuclear reactor programs.

W. O. Harms, Director Nuclear Reactor Technology Programs Oak Ridge National Laboratory Origin and Organization of the SP-100 Program

Judith H. Ambrus National Aeronautics and Space Administration

INTRODUCTION strong and vigorous reaction in this country against nuclear power and other technologies. The The logo of the SP-100 Program (Fig. 1) reveals its country became tired of hearing about advances in philosophy. TVie spirit of cooperation among the technology, and an era came to an end by simply National Aeronautics and Space Administration landing a man on the moon. (NASA), the Department of Energy (DOE), and the Toward the end of the 1970s, however, a group Defense Advanced Research Projects Agency of us v.ere charged with examining the need fox (DARPA) is part of its history and is evident in its power in space. What had changed since the 196°3? organization. First, the shuttle had become a reality. The potential of flying into space every couple of months BACKGROUND means we can actually go into space and do work. That open^ up quite a few possibilities about what During the 1960s America expended consider- we want to do in space. able effort in space nuclear reactor power, but in Those of us concerned about energy needs in the early 1970s that program came to an abrupt space and how to supply these needs took a look at halt. A number of technical efforts were either our future space opportunities. We first noted that redirected, put on the shelf, or forgotten; people our satellites and spacecraft had either a few hun- dispersed and efforts were directed elsewhere. The dred watts for several years or a couple of specific reason for terminating these activities is kilowatts for a few months. We then tried to imag- lost in history, but it may have been part of a ine what future space missions and their power requirements might be. Quite likely we would want to explore Saturn's rings now that we have had a look at them (Fig. 2). To go to Saturn and explore the rings by conventional propulsion, however, would take about 20 or 25 years. If today's engineers and not their grandchildren want to see the spacecraft do something, we need nuclear propulsion; energy demands for the trip are likely to be in the order of 106 kWh. Other possibilities are direct broadcast satellites (Fig. 3), which again would require an order of magnitude more energy, and a space power distribution center, a large power generation center in space that would beam power down. The Program Office i3 also composed of three members, one from each agency. At its head is the DARPA member, Bill Wright, who has two deputies—Steve Lanes from DOE and me from Fig. 1 SP-100 Program logo. NASA. Taking its direction from the steering AMBRUS

Rendezvous with Saturn's Rings Manned Mars Mission

CRUISE CONFIGURATION

SPIN AXIS 4 RPM '

AlFUOi' - i STOMM S»H 't M lil

STOWf U

• POWER: 100 kW • POWFR: 6 MW » MISSION: • ION-PROPULSION ENABLED • MISSION: • 8 YEARS TO SATURN. 3 YEARS EXPLORATION • 33-MONTH EARTH RETURN • 1 YEAR TITAN MAPPING, • 3-MAN CREW ON SURFACE 2 YEARS RING SPIRAL FOR 1 MONTH

Fig. 2 Potential mission applications: apace science.

Advanced Broadcast Space-Based Radar

100 kW • 100 kW CONTINENTAL COVERAGE

Fig. 3 Potential miMion applications: civilian/eoramerclRL ORIGIN AND ORGANIZATION OF THE SP-100 PROGRAM committee, the Program Office provides technical and two agencies were talking to each other. Soon direction on a programmatic basis in a very a third agency, which was likely also to be a user, broad-brush way. It answers such vi_:;rL;:.z as: came around. After a lot of negotiation, the three How much can we—and should we—spend on mis- agencies—DOE, NASA, aad DARPA—created a sion studies? How much can we spend on system tri-agency program to develop space reactor power studies? Can we afford this particular technology? system technology (Table 1). How important is one technology likely to be to another? TABLE 1 The power that runs the program is delegated The SP-100 Program to the Project Office at J?L. In the hands of the project manager resides complete technical respon- • A joint program of DARPA, DOE, and NASA sibility. ' Defined in an agreement 3igned 7'eb.ruary 11,1933, in which: • DOE is specified to chair the steering committee and name the deputy project manager for nuclear technology SUMMARY • PARPA is specified to provide the progr i director • NASA/JPL is specified to manage che project The SP-100 Program organization was • Defined to have the DOD, DOE, and NASA charter hr all developed to be consistent with the history and space nuclear reactor power systems technology development • Funded at $14.9 million in FY 1983 (including FY 1982 complexity of the task. We believe this organiza- carryover) tion will be effective in directing the spirit, of cooperation needed to make the SP-100 Program a success. A next step for more effectively occupying space ORGANIZATION is to build a space station. Energy demands for such a project grow larger the more often we loik Because of the complex interrelationship at them. The space station task force is now talk- between the heat source, power conversion, and ing about 65 kW for the first-generation unat- spacecraft systems and because three agencies are tended space station, with a mission life of five involved, we set up a programmatic structure years. For the second-genoration space station, to (Fig. 4) that will be responsive to technical and follow a few years later, the task force is talking about 120 to 150 kW. STEERING COMMITTEE DIRECTOR OF DARPA To provide such high power for long periods of time is likely to require a nuclear reactor. The task force therefore looked around and asked, "Is there PROGRAMMATIC a space nuclear reactor?" What they found were DIRECTION - '' -REPORTING fragments. A small effort on a new type of space PROGRAM DIRECTOR reactor, little more than a feasibility study, existed NUCLEAR DEPUTY at Los Alamos National Laboratory (LANL). There SPACE DEPUTY was also a small effort in thermoelectric conversion at the Jet Propulsion Laboratory (JFL), funded by NASA. These were the enly two activities in the co-intry that had anything at all to do with space nuclear power, and both programs were woefully underfunded. Recognizing that, to accom- PROJECT MANAGER plish something with so few funds, some specializa- tion of labor was needed, JPL and LANL started talking to each other, which is wonderful for two Fig. 4 Program structure. laboratories to do. The next step was even more unique. Two spon- administrative demands. Directing the SP-100 Pro- soring agencies started to talk io each other about gram is a steering committee consisting of the developing an interface between the technical Director of DARPA, the Associate Administrator requirements of the heat 3ourc<> and the conversion for the Office of Aeronautics and Space Technol- system. This dialogue led to questions concerning ogy,* and the DOE Assistant Secretary for Nuclear the spacecraft's requirements. At that point, the missions people became involved, and a program *At present, no one occupies this position on a permanent was born. We had a system, we had a program, basis. AMBRUS

Energy, who heads the committee. The steering to go forward and advocate this concept to committee gives programmatic policy direction and Congress." Most importantly, they are the people is likely to say, "This is the time to design a getting money for the program, ground engineering system," or, "This is the time SP-100 Program Overview

Vincent C. Truscello Jet Propulsion Laboratory

INTRODUCTION The program will probably go through three phases: phase 1, technology assessment and Born to meet the special needs of America's space advancement; phase 2, ground testing, in which we effort, the SP-100 Program testifies to the coopera- will actually build and demonstrate hardware on tion among government agencies. The Department the ground; and phase 3, flight qualification. In of Energy (DOE), the National Aeronautics and actuality, phase 3 may not take place. Other Space Administration (NASA), and the Defense nuclear space programs, such as the radioisotopic Advanced Research Projects Agency (DARPA) are thermoelectric generator program, went from working together to produce a 100-kW power sys- ground demonstration directly into a flight tem for use in outer space. At this point in the program. We do not know, however, if that is a effort, it is appropriate to review (1) the approach good idea for larger systems. Our systems contrac- to meet program goals; (2) the status of activities tor will examine the issue and advise us. Right of the Project Office, managed by the Jet Propul- now the program is in phase 1, the two- to three- sion Laboratory (JPL); and (3) because this is a year technology assessment and advancement meeting on materials, answers being developed by stage. What then are our current objectives? the Project Office to vital questions on refractory Briefly, the goals of phase 1 are (1) concept def- alloy technology. inition, (2) technical feasibility, and (3) costs and schedule development. The first goal, concept definition, includes understanding what the missions APPROACH are. A number of missions are made possible with the use of a reactor system, and a large part of our Four major milestones (Fig. 1) emerge for the program is dedicated to trying to understand these SP-100 Program. The Memorandum of Agreement missions and their requirements. A complementary between the three government agencies, which aspect is then determining which systems make the really kicked off the current phase of the effort, most sense and can best me^t those particular was completed in the second quarter (February) of requirements. Identifying the missions and the FY 1983. Approximately a year from that date the concepts provides a fair understanding of the tech- concept(s) will be selected. The final milestone is nological issues that must be answered before we tentatively set for the fourth quarter (July) of FY can enter the ground testing phase of the effort. f 1985. A that time we will recommend to the Pro- Not only do we identify the technical issues but grar Office the concept which should go forw?rd, we carry out experiments and analyses. Our second the technologies that should be utilized, and Koal is to conduct those development activities whether or not to begin ground testing. To deter- required to address and *esolve these issues enough mine if we are on track and to ensure that we can to satisfy ourselves that the technical feasibility of make a ground test decision in FY 1985, the pro- the concept will not be affected. What do we mean gram will be reexamined in mid-1984. At that time by enough9. Very simply, we mean looking hard we should be prepared to determine if we will need enough at each of these technological Issues that, more time, if we can m?ke the decision by late FY once we select a particular concept and a particular 1985, or (and this is less likely) if we can do it set of technologies, materials, or what have you, we sooner. will not embark on phase 2 and suddenly find that TRUSCELLO

FY 1983 FY 1985 -f H-H—!- MEMORANDUM OF AGREEMENT

CONCEPT(s) SELECTION V INTERMEDIATE REVIEW

GROUND TEST PHASE DECISION v

Fig. 1 Major program milestones.

we have to stop. Reasons for stopping would (Fig. 2). The Project Office includes the manager, include discovering a major "show-stopper"; the assistant manager, and two deputies—one for need for a major development effort; or that we Nuclear Technology and one for Aerospace Tech- cannot use the concept, materials, or conversion nology. A coordination team established to inte- devices initially selected. We must move into tha grate program resources and responsibilities ground demonstration phase with a high probabil- includes representatives from Los Alamos National ity of completing the engineering development. Laboratory (LANL), JPL, and the NASA-Lewis The third important goal is understanding costs Research Center (NASA-LeRC). They coordinate and schedule. Before we can get any sponsor to efforts, not only of these three laboratories but of fund the second and third phases of the program, other support organizations as well [for example, we must have a thorough understanding of costs. DOE laboratories such as Oak Ridge National In the present phase, we are talking about an Laboratory (ORNL) and Hanford Engineering effort on the order of $15 million a year, or a total Development Laboratory (HEDL) and possibly r'_ about $45 million to $60 million before we move other NASA centers]. into phase 2. Obviously the ground demonstration The program itself is divided into four major phase of the effort is going to be much more areas. The Mission Analysis and Requirements and expensive—hundreds of millions of dollars each System Definition areas are each headed by a man- year. We could be talking about a total cost of a ager. The Aerospace and Nuclear Technology areas billion, or even several billion, dollars. Before any are managed by the Project Office deputies. sponsor—or even Congress—would commit to such Another area vital to the project and its viability funding, we must have a good comprehension of is Nuclear Safety. Obviously, to get launch the costs and schedule for completing these next approval, quite a few safety issues will have to be phases. Will development take two or three years surmounted. Though the Nuclear Technology man- and $0.5 to $1 billion, or will it take five or six ager is responsible for this important area, we also years and $5 to $6 billion? Our system contractors have a Safety Advisory Committee (Table 1), and in-house efforts are going to be aimed to a which has begun evaluation of our activities large extent at trying to understand this question. (Table 2). We are under pressure to generate these kinds of The first major area, Mission Analysis and numbers, because both NASA and DOE must begin Requirements (Fig. 3), is divided into planetary to put. them into their budgets for subsequent missions, military missions, space station activi- years. ties, and civilian and commercial missions. In- house activities at NASA-LeRC and JPL as well as contracted activities through the military agencies STATUS will help generate mission requirements (Table 3) for the wide range of possible missions. The idea is What is the status of the SP-100 Program? to get a set of integrated requirements that can be First, it is already organized and structured applied in the development of an appropriate sys- SP-100 PROGRAM OVERVIEW

Safety Advisory Coordi nation Commi ttee SP-1OO Project Team

Chai man Mgr - V. Truscello JPL - V. Truscello

Duane Sewell Asst Mgr - H. Davi s LANL - J. Hanson

DMNT - J. Hanson LeRC - R. Sovie

DMAT - J. Mondt Admi ni strati on an.l Operations^

Manager

Kirk Gerbracht

Mi ssion Safety ancT Analysis and System Aerospace Nuclear Requirements Definition Technology Technology

Manager Manager Manager Manager

Dick Wallace Jim French Jack Mondt John Hanson

Fig. 2 SP-100 Program structure.

TABLE 1 Members of the SP-100 Project Safety Advisory Committee

Name Affiliation TABLE 2 Status of Nuclear Safety Activities Duane Sewell Consultant (Chairman) > Two meetings held by advisory committee Robert Bacher California Institute • Preliminary safety plan completed and reviewed by of Technology advisory committee Garth Cummings Lawrence Livermnre • Preliminary 3afety design requirements corrpleted and National Laboratory reviewed by advisory committee;, requirements issued to 3ystem R. E. Schreiber Consultant contractors A. W. Snyder Sandia National Laboratory • Coordination wkn interagency nuclear safety 'eview panel James Lee United States Air Force (INSRP) begun Dave Okrent University of California (LA) Stanley M. Luczkowski Johnson Space Center Milton S. Plesset Califoi iia Institute of Technology TRUSCEU.0

MANNED MILITARY SPACE MISSIONS STATION CIVILIAN/ PLANETARY COMMERCIAL MISSIONS MISSIONS

INTEGRATED REQ'TS ON POWER SYSTEM

SYSTEM DESIGN

Fig. 3 SP-100 Mission Analysis and Requirements.

TABLE 3 is headed up out of JPL. The Air Force and Navy Summary of Mission Requirement are supplied monies to conduct military mission Inputs to Power System Design activities. JPL, NASA-LeRC, and various contractors are funded to do work in planetary, space sta- > Power use profile tion, and civil areas (Fig. 4). • Mission 9urvivability • Mission duration/lifetime To achieve the goals of the second major area, • Start-up/load following/shutoff System Definition, we put out a request for bid the • Dormancy latter part of 1982. Ws selected three system con- • Attitude control tractors: ^A Technologies teamed with Martin •Launch vehicle compatibility (mass and size) • Deployment Marietta Corporation, Gci eral Electric, and West- • interfaces (power/mechanical/control/data) inghouse teamed with Lockheed. They are right • Environment (radiation, thermal) now going through the initial screening of the vari- • Safety and cost implications ous technologies. As a result of this work, there are three formal reviews (Fig. 5). The first one occurred in June, the second in September, and the tem. We would like to develop a concept that meets final one in December. If hopes are realized, we the needs of a multitude of missions. That may or will select a contractor or contractors with one or may not be feasible. We may have such signifi- more concepts sometime ?.n February 1984. cantly different requirements—e.g., between military and commercial applications—that a single W" asked the contractors to review a large power plant design may not be possible. To the array of possible technologies with a number of extent it is possible, however, that is our goal. constraints, the moit important being that the Organizationally, the work is divided along these power plant weigh les<- than 3000 kg, produce at lines. The Mission Analysis and Requirements area least 100 kW of power, Lnd fit into no mure than SP-100 PROGRAM OVERVIEW

0 PLANETARY MILITARY • MILITARY 9 SPACE STATION t CIVILIAN/COMMERCIAL Fig. 4 Organizational support in the area of Mission Analysis and Requirements.

CY 1983

M A Ml j A S O N 0 -+--H—4-—+—-1—I—I- CONTRACT START

FIRST REVIEW

SECOND REVIEW v

FINAL REVIEW v

CONCEPT SELECTION V

Fig. 5 Systems concept selection schedule.

one-third of the shuttle bay. We are finding that meet the requirements of power and weight, how- meeting those constraints is by no means Ln easy ever, it appears that many, though not ail, of these task. concepts will have to use a deployable radiator. We have a large number of options with respect The contractors have narrowed down the list of to producing the electrical power from the thermal various technologies being considered (Table 4). unergy of the reactor. Figure 6 shows a hypotheti- For instance, chey initially looked at thermal, cal structural design of a concept using a dynamic epithermal, and fast reactors but, because of heat engine. The reactor constitutes only a very weight limitations, quickly zeroed in on a fast reac- small part of the volume and the weight of the tor. They have started looking at reactors cooled entire power system. Most of the volume is taken with gas, liquid-metal, and heat-pipe systems. As up by the device that rejects the portion of power for power conversion devices, the program has the not converted into electricity—the waste heat radi- various dynamic conversion options of Brayton, ator. This turns out to be quite i limiting aspect Stirling, and Rankine cycles. For static conversion, for the various design concepts being evolved. Our there are thermoelectric; thermionic; in-core and hope was to design a static radiator that would not out-of-core systems; and some new technologies have to be deployed ^nce it was placed in space. To (which are really in their infancy state but may 10 TRUSCELLO

PC ELECTRONICS CABLE SUPPORT SUPPORT RING

RADIATOR PANELS (TRUSSES

TRUSSES

PC I -RADIATOR-) Fig. 6 SP-100 Aerospace Technology: hypothetical structural design of an SP-100 concept.

TABLE 4 Here we run into problems: The sponsor needs an Candidate Subsystem Concepts annual operating plan, and the Project Office needs a cost breakdown of exactly what is going to be Reactor conversion done and why. The definitive inputs from the system contractors, however, are not due until • Heat-pipe-cooled •Brayton December. So we have to make judgments as to reactor (fast) • Stirling • Pumped liquid-metal- • Rankine what we should start doing right now. Using the cooled reactor (fast) • Thermoelectric significant data base that already exists at the var- • Gas-coolod reactor (fast) • Thermionic in-core ious DOE laboratories and the NASA laboratories, • Thermionic out-of-core such as JPL and NASA-LeRC, we set up a Technol- •AMTEC •TPV ogy Assessment Working Group (TAWG). In the past t^ee or four months, this large team of government people has worked together to assess the options, identify feasible technologies, and rank the well be good for growth versions), the alxali metal options. This tc^m has pretty much accomplished thermoelectric converter (AMTEC), and thermo- that task. photovoltaic conversion (TPV). Although the system contractors have not The system contractors are evaluating all of agreed with one another, as a community the these options. In fact, they have narrowed down systems they selected do line up with those we the list even further. Interestingly, each has a evaluated independently and ranked at the top. The somewhat different view. Our job, therefore, is to Project Office is comfortable with the direction come to grips with their several answers and to that things are going right now. Whether we will evaluate what does and does not make sen3e. To do be able to move ahead with all of these technolo- that, we are trying to define what the technology gies or whether we have to downscope to an even program is going to be over the next two years. smaller group, only the Program Office will tell. 31--100 PROGRAM OVERVEW 11

TEMPERATURE (°F) 1500 2000 2500 3000

STRESS (psi) 5 FOR 1% CREEP IN 10,000 h 2

800 1000 1200 i400 1600 1800 TEMPERATURE (°C) Fig. 7 SP-100 Aerospace Technology: high-temperature strength information for candidate alloys. In major areas 3 and 4, Aerospace and Nuclear structural and fuel cladding applications in the Technology, the TAWG used the expertise of each nuclear subsystem as well as piping, heat laboratory. LANL, for example, examined the exchanger, pump, turbine wheel, and Stirling cycle shield and reactor subsystems to determine where piston applications in the power conversion subsys- the technology is—the weight of the reactor per tem. unit of power produced and the t. chnical feasibility The area of refractory alloys is important not issues associated with those subsystems. NASA- only for the reactor but also for the power conver- LeRC, with romc support from ORNL, did the work sion system. The Project Office had to make sure on the dynamic machinery; and JPL did the work the SP-100 Program did not take off in all direc- on the static subsystems. tions, developing materials suitable for each particular power conversion, heat transport, and nuclear REFRACTORY ALLOYS reactor application. To meet that challenge, we set up a structure that meets the needs of all these Studies Dy JPL °taff and a subset of the TAWG diverse applications (Fig. 9). The key is a technical as well as inputs from the three systems design planning team that defines materials needs. Made contractors reveal that to meet system require- up of representatives from each of the major con- ments of a 3000-kg system weight and 100-kW out- tributing laboratories, this team will define the put, use of refractory alloys is imperative. Why requirements, the material development needs, and was this class of alloys identified? Clearly the lead- the costs and schedule to meet these needs. A ing candidates are thosa alloys that, can operate at steering committee made up of the Project Ofiice very high temperatures and have suitable creep and Aerospace and Nuclear Technology managers strength (Fig. 7). For that reason we cannot live will act on the planning team's recommendations with superalloys—and certainly not with stainless from a programmatic standpoint to determine (1) steels, which are applicable at very low tempera- whether or not we can afford it, (2) how it fits in tures compared to the types of systems we are with all our other needs, and (3) when those partic- working on. A review of possible applications of ular needs are important and should be imple- the candidate tungsten-, molybdenum-, and mented. A group housed at LANL will lead the tantalum-base alloys (Fig. 8) was conducted by the effort in implementing the actual development TAWG subset. As a result of this review, refrac- work through the various laboratories and contrac- tory alloy feasibility issues surfaced regarding core tors. 12

S. Candidate I \^ Alloy Major \. W MoRe T-111 ASTAR 811C TZM/TZC Application*--^^ by Subsystem ^\ NUCLEAR SUBSYSTEM Clad Yesa Yes NA* Yes* NA

Core Structure NA Yes Yes Yes NA (weldable)

POWER CONVERSION SUBSYSTEM Piping/HX NA NA Yes Yes NA Pumps

Turbine Wheels/ Pistons NA Yes NA NA Yes

Static Components NA NA Yes Yes NA

Heat Pioes NA Yes Yes Yes NA

in-core thermionic. loy was not considered for this application. Used with tungsten liner. Fig. 8 Possible application of candidate refractory alloys by major subsystem.

This planning team has identified a number of react with these impurities, they can transport feasibility issues (Table 5) that have to be them from one area to another, causing a buildup addressed between now and the end of the current and resulting in potential long-term failure of the phase of the program—which could be FY 1985. We components. Lithium, on the other hand, would be have to understand chemical compatibility of the better, because it is a sink for most of these impur- fuel, clad, and coolant; and we need a lot more data ities. We need to do enough work on all these on such matters as irradiation behavior, specifi- issues bj that, once we select a specific material cally property degradation and swelling. Because combination and start moving ahead in engineering the systems will not be operated during launch, we development, we will not run into any concerns must ensure that materials do not fracture. Their that would force termination of the effort. toughness at low temperatures is critical. Another In summarv, we are convinced that refractory major concern is the potential degradation in per- alloys will be necessary to meet the needs of the formance by refractory metals as a result of con- power systems, whether they are used as fuel clad- tamination by oxygen, carbon, or nitrogen. These ding, piping, heat pipes, turbines, or pistons. contaminants could be picked up from the fuel or Tantalum- and molybdenum-base alloys are prime by a fluid flowing through hot regions and de- candidates to meet temperature and weight con- posited in colder ones. Use of inert gases as a straints. They may also be suitable for the fuel working and transport fluid should solv« the prob- cladding, though tantalum alloys will require a lem, but it does not. Because inert gases do not barrier to the fuel. Tungsten-rhenium alloys are an SP-100 PROGRAM OVERVIEW 13

TABLE 5 • Trusce lo Steering Cornmi ttee • Mondt Feasibility Issues • Hanson Clad Technical • LANL Chemical compatibility. fuel/clad/coolant Planning • JPL Irradiation behavior (swelling, property degradation) Team • LeRC Fracture toughness/crack growth rate • ORNL Barrier integrity B HEOL Mechanical properties Inert gas compatibility Core structure Irradiation behavior Impie.nentor • LANL Fracture toughness r Mechanical propercies Thick section weld Inert gas compatibility

DOE/NASA Labs Piping/HX/pumps/heat pipes Fracture toughness r - - - Mechanical properties 1 Thick section weld 1. 1 Contractor;, Producibility of very large diameter pipe (3 to 4 in.) Inert gas compatibility Life testing Fig. 9 Materials Development management structure. Turbine btadea/piatoni High-temperature creep alternative for fuel cladding. For selected power Fatigue conversion system applications, tantalum- and Fabricability of wheel/piston molybdenum-base alloys again seem to be the best Inert gas compatibility candidates. At these high temperatures, molybdenum-TZM is an extremely t -> d candidate for turbine wheels for either a Bra., .on or a Ran- kine system. Both the tantalum and molybdenum ment, ground testing, and flight qualification. alloys could be used for the heat pipes. Weldability Currently the program is in the two- to three-year concerns, on the other hand, mean molybdenum technology assessment and advancement stage. alloys may be less suitable for the reactor struc- Goals are to identify the space nuclear power ture. We need to ascertain how weldable the system concept that best meets anticipated molybdenum alloys are going to be and hew much requirements of future space missions, assess the work is going to be necessary to prove they can be technical feasibility of that concept, and establish used for piping throughout the system or for the a cost and schedule for developing the concept. The reactor structure. Clearly we must understand SP-100 Project Office has begun the refractory metals. implementation activities needed to meet these goals, and we feel comfortable with the direction that things are going now. With regard to refrac- SUMMARY tory alloys, we feel a better data base will be required before we move ahead in the program The SP-100 Program is expected to go through from technology assessment to ground demonstra- three phases: technology assessment and advance- tion. Potential Refractory Alloy Requirements for Space Nuclear Power Applications

H. H. Cooper, Jr. Oak Ridgi' National Laboratory"

INTRODUCTION electricity conversion efficiencies, ard low mass and size) require very high operating temperatures. Going on a space mission is analogous to taking a As a result, materials are required that, retain high long ocean voyage in a small boat. Both spacecraft strength and other key properties at temperatures and boat components must be constructed of high well in excess of those for which more conventional quality materials designed for high reliability. One alloys, such as the stainless steels and the superal- of the more critical components for future space- loys, can be used. An approximate but useful mea- crafts will be a nuclear reactor to provide energy sure of refractoriness is the melting point (mp) of for (a) powering ion propulsion systems to expedite the metal upon which a system of alloys is trips to the outer planet?, and (b) defense applica- based—iron (mp 1535°C) for stainless steels, for tions such as surveillance satellites, large radar example. The refractory alloys under consideration systems, laser-based communication equipment, for space nuclear systems applications are those and laser or particle-beam weapons designed to based on niobium (mp 2415°C), molybdenum (mp destroy enemy missiles in flight. Because these 2620°C), tantalum (mp 2996°C), and tungsten (mp reactors should operate at a high temperature to 3410°C). The temperature limit for useful applica- ensure efficient power production, refractory (heat tion of an alloy from the standpoint of strength is resistant) alloys will be needed as the nuclear fuel in the range of 50 to 60% of its absolute melting cladding and structural materials. The purpose of temperature. this paper is to (1) provide an introduction to the design requirements for refractory alloys in space The relationship between specific energy (a use- nuclear applications and (2) briefly indicate the ful performance figure of merit) and reactor outlet status of refractory alloy technology and an temperature for a number of space power systems approach to refractory alloy development. is presented in Fig. 1. If, as is generally conceded, the absolute upper limit for the use of stainless steels is about 800°C, it is apparent from this plot RELATION BETWEEN ALLOf why the refractory alloys must be considered if the SELECTION AND SYSTEM desired high-performance levels are to be achieved. The relative elevated-temperature strengths of PERFORMANCE stainless steel, the superalloys, and the selected Discussions concerning requirements for struc- refractory alloys are shown in Fig. 2 in terms of tural and fuel-cladding alloys in spac? reactor sys- the stress required to produce 1% strain or defor- tf.'iis normally center on a class of materials mation in 10,000 h. The approximate temperature known as refractory alloys. The reason for this is limits for application of these alloys from the that thp high performance levels desired for these standpoint of strength alone are indicated in the systems (i.e., high power outputs, high heat-to- overlay of Fig. 1 shown in Fig. 3. Limited short time creep data indicate that the Mo-14Re alloy should be considered for space power applications.7 "Operated for the U. S. Department of Energy under contract W-7405-«ig-2*> with the Union Carbide Corporation. A development program on the Mo-Re alloy system

14 "OTENT1AL REFRACTORY ALLOY RECXHRQiffiNTS 15

REACTOR OUTLET TEMPERATURE (°F) REACTOR OUTLET TEMPERATURE !"F) 1OOO 1^00 180O 2CD0 2600 3000 1000 1400 1800 2200 2600 3000

10" NASA ADVANCED "SSPACE REACTOR . 5 •|^H_ANL HEAT PIFc ' a >SPft-4 S 2 1OC £

X 10'

15 in' -STAINLESS IT 'O STEELS liJ z , LJ O 2 a. 1O( 600 800 1000 1200 1400 160C 1800 400 600 800 1000 1200 1400 1600 1800 REACTOR OUTLET TEMPERATURE CO REACTOR OUTLET TEMPERATURE CO Fig. 1 Relation between specific energy and reactor outlet Fig. 3 Overlay for Fig. 1 indicating approximate tempera- temperature for a number of space power systems. ture limits of application, from the standpoint of strength, for stainless steels, supcralloys, and refractory alloys.

TEMPERATURE (°F) 2000 2500 3000 2. Availability in the product forms required (3.g., piping, tubing, forgings for pumps and other components, sheet, and plate; on a time scale that is consistent with the project schedule. 3. Weldability. STRESS (psO 4. Ductility. FOR 1% CREEP 5. Acceptable friction and wear properties. IN lO.OOOh 6. Resistance to irradiation damage.

REFRACTORY ALLOY DESIGN REQUIREMENTS 800 (000 1200 1400 16CU 1800 TEMPERATURE (°C) Although a variety of space nuclear power con- Fig. 2 Relative high temperature strengths of stainless steel [Type 316 SS (Ref. 1)]; superalloys based on nickel cepts are currently being considered, the major [Hastelloy X (Ref. 2)] and cobalt [US-188 (Ref. 3)]; and system components can be identified with two sub- refractory alloys based on niobium [Nb-l%Zr (Ref. 4)] and systems: (1) the nuclear subsystem, which princi- molybdenum tTZM (Ref. 5)], tantalum [T-lll (Ref. 5) and pally includes the reactor and radiation shielding, ASTAR 811" tlef. 6)], and tungsten [W-25%Re (Ref.4)]. and (2) the power conversion subsystem including heat transport, power conversion, and waste-heat is under way at the Los Alamos National Labora- radiator components. Reactor concepts being con- tory. sidered for the nuclear subsystem include: In addition to the strength criteria just cited, 1. Alkali metal cooled fuel pin concepts. candidate alloys for these high-temperature appli- 2. Helium or helium-xenon cooled concepts. cations must meet other requirements as well. 3. Thermionic concepts using an alkali metal These requirements include: coolant. 1. Compatibility with fuel materials and with 4. Heat pipe reactor concepts with an alkali reactor coolants and working fluids. metal working fluid. 16 COOPER

The operating temperatures for these concepts TAB1LE1 range from 625 to 1525°C and are summarized in Coolants and Operating Temperatures Table 1. The primary application of refractory for Candidate Reactor Concepts alloys in these reactor concepts are for (1) in core structural supports and fuel cladding for either gas Operating liquid metal fuel pin concepts or (2) fuel temperature range. or Reactor concepts Coolant °C support for heat pipe reactors vrhere the fuel is on the outside of the heat transport system. Design Liquid metal cooled Alkali metal 625 to i52f> requirements for refractory alloys used in either Gas cooled He or He/Xe 825 to 1325 application include effective load carrying capabil- Thermionic Alkali metal 1225 to 1525 ity at the proposed operating temperature, good Heat pipe Alkali metal 625 to 1525 fabricability, resistance to irradiation damage, and acceptable compatibility with alkali metals or inert gases and fuels (for fuel cladding). TABLK 2 The power conversion subsystems include com- Working Fluids and Operating Temperatures ponents for specific power conversion, heat transfor Candidate Power Conversioni Concepts port, and waste heat rejection. Power conversion Operating systems f^Ii into two groups—dynamv and static. Po\»er temperature Dynamic subsystems include Brayton, Stirling, and conversion Working range, Rankine cycles; static subsystems include ther- concepts fluid °C moelectric, alkali metal thermoelectric conversion (AMTEC), and th --malphotovoltaic (TPV). The Dynamic cycle Brayton He or He/Xe 875 to 1375 possible working i.uids and range of operating Sterling He 825 to 1375 temperatures anticipated for these candidate sys- Rankine Alkali metal 675 to 1325 tems are summarized in Table 2. Possible applica- Static cycle tion for refractory alloys in the power conversion Thermoelectric _ 1100 to 1325 subsystems include piping, pumps, and heat AMTEC' Alkali meta1. 925 to 1025 exchangers. With regard to specific conversion TPV+ - 1725 technologies, refractory alloys will also be required for turbine blades for Brayton and Rankine sys- 'Alkali metal thermoelectric conversion. tems and heater heads and pistons for Stirling +Thermalynotovoltaic. cycle applications. Desihrn requirements that will Currently, there is nc annotated bibliography of be imposed upon the refractory alloys in these this technology which integrates the past and cur- applications will include effective load-carrying rent alloy development accomplishments of NASA, capability, good fabricability, acceptable compati- DOD, and DOE. Also, many reports on refractory bility with working fluids, good wear resistance, alloy activities of the 1960s and 1970s were not and adaptability to being coated with high- published. Moreover, in some cases insufficient emissivity materials. technical detail was reported to allow interpretation of the data. REFRACTORY ALLOY APPROACH TO REFRACTORY TECHNOLOGY STATUS ALLOY DEVELOPMENT A review of the status of refractory alloy tech- Successful application of refractory alloys will nology provides both good and bad inputs. The pos- require a strong and continuing dialogue between itive aspects are the significant accomplishments material scientists and space power systems design made with refractory alloys during the 1960s and specialists. This dialogue typically begins with the 1970s that are directly applicable to the SP-100 design community identifying the materials Program. The negative information is the fact that requirements. The materials community then gen- refractory alloy technology for space nuclear reac- erates the engineering data needed to assess the tors has been dormant since the termination of effectiveness of candidate alloys with respect to the these development activities in the mid-1970s. As a design requirements. Because alloys rarely measure result the refractory alloy nuclear technology up to all of the designer's requirements, the inter- available today is not adequate to meet the SP-100 action between the design and the materials com- Program needs. munities is by necessity an iterative process. POfENTlAL REFRACTORY Al-LOV REQUIREMENTS 17

This meeting provides a forum for initiating the cant challenge to both the materials and the sys- development of refractory alloy nuclear technology tems design communities. by involving both the refractory alloy experts of the 1960s and 1970s and the current experts. Through the cooperation of these experts, we hope to develop an assessment of the current status and needs of a refractory alloy nuclear technology. REFERENCES

1. ASME Boiler and Pressure Vessel Code, Section III, Code SUMMARY Case N-47 (1982). 2. H. E. McCoy, Jr., Creep Behavior of HasteUoy X, 2HCr-lMo Steel, and Other Alloys in Simulated HTGR Helium, In reviewing design requirements for refractory ORNL/TM-6822, Jun; 1969. alloys for space nuclear applications, several key 3. Manufacturer's Data, Stellite Division, Cabot Corporation. points are identified. First, the successful utiliza- 4. R. E. Gluyas and G K. Wa^on, Materials Technology for c.n tion of refractory alloys is considered an enabling Advanced Space ?ower Nuclear ReacUf Concept, Program requirement for the successful deployment of high Summary, NASA TN D-7909, p. 39, March 1975. 5. K. D. Sheffler and R. R. Ebert, Generat'rm of Long Time efficiency, lightweight, ard small spaca nuclear Creep Data on Refractory Alloys at Elmated Tt mperatures, systems. Second, the recapture of refractory alloy NAS CR-134481, September 1973. nuclear technology developed in the 1960s ari 6. W. D. Kloop, R. H. Titran, and K. D. Sheffler, Lon-j Time early 1970s appears to be a pacing activity in th? Creep Behavior of the Tantalum Alloy ASTAR- !UC. wASA successful utilization of refractory alloys. Third, TP 1691, September 1980. 7. \V. D. Kloop and W. R. Witzke, Mechanical Properties of the successful ppplication of refractory alloys for EHectron-Beam-Melted Molybdenum and Dilute Molyb- space nuclear applications will present a signifi- denum-Rhenium Alloys, NASA TMX-2576, June 1972. Refractory Alloy Component Accomplish iments from 1963 to 1972

E. E. Hofttnnn U. S. Department of Energy

INTRODUCTION was abruptly t"rrw'.,;.ucu.* The General Electric portion.0 of the overall NASA-Lewis program con- Many advanced technology capabilities must be h. centrated on component development and testing, hand in order to successfully develop nuclear elec- heat transfer studies, compatibility studies, and tric power systems for space application. Many~~Df~ other tasks,, such as bearing and seal development, the rsactor/power conversion systems currently that were requirerKto support component develop- being considered for SP-100 (100 kWe) application ment. would operate at temperatures sufficiently high to The overall NASA-Lewis materials program require the use of refractory alloys. A technology during this period was a well conceived ar.d exe- element necessary to build sysiems either for cuted program that made maximum use of the var- ground tests or launch is the ability to construct ious and complimentary talents of a number of complicated refractory alloy components and to industrial contractors and laboratories. The exten- join these components into zero-leakage, durable sive alloy development and welding research heat transport systems of extremely high integrity. activities of Westinghouse Astronuclear, the very high quality creep program at TRW-Cleveland, and The purpose of this paper is to summarize the : accomplishments of a number of refractory alloy the in-house resear- : program at NASA-Lewis all component development and testing programs con- contributed to the success of the component ducted during the 1963 to 1972 time period. This development program carried out by GE-Space experience base is well documented, and effective Power and Propulsion personnel. utilization of this information should assist in reducing the cost and risk of future developments. DESCRIPTION OF REFRACTORY During the 10-year period, 1963 to 1972, General Electric, Space Power and Propulsion Section (later ALLOY COMPONENTS Nuclear Systems Programs) located in Cincinnati, The specific component developments to be Ohio, designed, built, operated, and evaluated a described in this paper are listed in Table 1. All of large number of refractory alloy components as these components or systems were tested and part of a long-term NASA-Lewis Research Center evaluated with the exception of the Nb-lZr Solar space power program. This NASA program to Brayton Heat Receiver. The experience and data develop advanced systems was focused primarily, obtained on these systems are most relevant to but not exclusively, on nuclear powered Rankine advanced Rankine applications, but much of this systems. This paper will provide a brief overview of a few of the most significant component developments oi the General Electric program in *The author of this paper led a group of materials specialists at General Electric which was responsible for the fabrication the period from 1963 until 1972 when the national and/or testing of many of the systems described in this paper program to develop advanced space power system during this time period.

18 REFRACTORY ALLOY COMPONENT ACCOMPUSH*SNTS

TABLE 1 "Refractory Alloy Component Fabricator" by Refractory Alloy Components and Test W. R. Young in these proceedings. Systems Which Will Be Described A brief d^,;.iption of the various refractory Mo-TZM and Mo-TZC potassium turbine components alloy components and test systems listed in T-111/lithium valve test loop Table 1 is given below. Nb-lZr Rankine system corrosion test loop T-lll Rankine system corrosion test loop Electromagnetic (EM) pump development Mo-TZM and Mo-TZC Potassium Nb-lZr solar Brayton heat r<"rc--•:.-. T-lll boiler development t>«r system Turbine Components One of the most ambitious undertakings of the inf-"-mgition, particularly the materials experience NASA-Lewis program at General Electric was the And such special areas as lithium compatibility, development, construction, and operation of a has wide applicability to many of the systems 3 MW potassium turbine and the associated boiler being currently considered by SP-100 contractors. and condenser system.'"4 The bulk of this two- Several of the systems that will be cited here are phase potassium system was constructed of stain- also discussed in the paper by DeVan et al. less steel and superalloys, but the turbine rotors included in these proceedings. The Nb-lZr Solar and many of the rotor blades were constructed of Brayton Heat Receiver and the T-lll Boiler the molybdenum alloys, Mo-TZM and Mo-TZC. The Development Test System, both listed in Table 1, turbine portions ~f the test facility are illustrated were the most significant component accomplish- in Figs. 1 and 2. Nichols, Fink, and Zimmerman,1 ments during the 10-year NASA program at Gen- in particular, give detailed guidance regarding the eral Electric. A detailed description of the fabrica- fabrication of the molybdenum alloy turbine blades tion of these large and complex systems is given in and rotor wheels.

UH4U5T SCROLL

PIVOTED PAD BEARING 2ND STJGf ROTOR

1ST STAG! ROTOR

DUPUX BALI BURING

FlOW

Tli BOLT

Fig. 1 Cross section of potassium test turbine. Selected rotor blades made of Mo-TZM and Mo-TZC alloys. (Three-stage version of this test system had rotors made of Mo-TZM alloy.) 20 hOFFMAW

Fig. 2 Sectioned view of potassium test turbine showing rotor and blade configuration.

The entire potassium turbine test program was conducted over a period of about six years. Both two- and three-stage turbine configurations were tested and evaluated. Potassium turbine inlet temperatures during the 12,000 hours of testing were ¥»CUUM till - mostly in the 750 to 830°C temperature range. Damage to the molybdenum alloy blades was found ! coi to be directly related to the oxygen concentration of the potassium with significant corrosion/erosion damage occurring with oxygen levels of 100 ppm or \U,

greater in the potassium and essentially no attack Hfli!ING YUYE with oxygen concentrations below 20 ppm. ISOLATION V1LVI flOH It tTU

ru»r DUTLET T-ll 1/Lithium Valve Test Loop The primary purpose of this T-lll alloy pumped lithium loop test was to evaluate the performance of refractory alloy metering and isolation valves in high velocity lithium.5 The test loop showing the

location of the valves is illustrated in Fig. 3. The DIAIM *HD Fill UNI - / various refractory metals and alloys utilized in the construction of these valves are shown in Fig. 4 prior to assembly of the valves. The valves were exercised extensively during the 5000 h test at valve temperatures as high as 1038°C. All the refractory alloy valve parts shown in Fig. 4 were subjected to rigorous evaluation following completion of the test, and no evidence of any degradation Fig. 3 Schematic view of T-lll alloy high temperature was noted on any of the valve parts wetted by lithium loop used to evaluate refractory alloy metering and lithium. Son i lling of the Mo-TZM gears which isolation valves. REFRACTORY ALLOY COMPONENT ACCOMPUSHWeNTS 21

GUt ML. WM ft'

• - -bi •

, (1 >. M I.' J N ' Mi '• -' •

•Ik • • .1 M '.*

t ••.•<•• Mi '.'•

tMP Pis 't Nb 1 Z-

=— iF . . ft

spin aiHC

}{ SUM BAS1 I <•' h

PUUil BODY III!

- SEA T BtTilHiR T1H

- INLET BODY T 111

Fig. 4 Exploded view of various refractory alloy parts used in the fabrication of the lithium system test valves. operated in the high vacuum environment was tric power systems was to evaluate candidate sys- noted, and these were subsequently replaced with tem materials under conditions as prototypic as gears made of tool steel. Based en the successful was practical. For this reason a prototype corro- performance in this test, valves of this design were sion test loop system6 was developed for use in the used in other test facilities built later at General evaluation of refractory alloys in boiling and con- Electric. densing potassium . avironments which simulated projected Rankine system conditions. The first test Nb-lZr Rankine System Corrosion system developed to evaluate these alloys in a Test Loop two-loop system designed to simulate conditions in both the heater loop (reactor coolant) and the One of the major thrusts of the NASA-Lewis two-phase Rankine loop is shown in Fig. 5. This sponsored program to develop advanced space elec- photograph was taken in the clean room prior to 22 HOFFMAN

Fig. 5 Nb-lZr Rankine system test loop mounted on vacuum chamber stainless steel spool piece prior to installation in the test facility. installation of the test loop and the vacuum cham- This test system, which was operated for 5000 h ber spool piece on which it was mounted on the without difficulty and was evaluated exi,cisively, ultrahigh vacuum test chamber. served as a prototype for the higher temperature The prototype tesc system consisted of a two- T-lll alloy system that will be described later in loop Nb-lZr facility; sodium being heated by direct this paper. The sodium circuit of this loop operated resistance in the primary loop and used in a heat at a maximum temperature of 1165°C and the exchanger (boiler) to boil potassium in the second- potassium circuit at a maximum temperature of ary teot loop. The method chosen to heat the boiler 1093° C during the 5000 h test. Details of the of the Nb-lZr Rankine System Corrosion Test Loop compatibility evaluations on this loop are given in was the use of a primary or heater loop in which the DeVan et al. paper in these proceedings. I2R heated sodium was to be pumped through the annulus of a tube-in-tube counterflov boiler where the required heat was transferred to the potassium T-lll Rankine System Corrosion in the secondary two-phase circuit. One of the Test Loop principal goals of this test program was to incorpo- This test, which followed the 5000-h Nb-lZr test rate as many components in the Nb-lZr corrosion described aL ve, was conducted for 10,000 hours in test loop as required to assure an accurate deter- a system constructed primarily of the T-lll alloy mination of the test conditions ard, thereby, mini- (90Ta-8W-2Hf). In this two-loon system7 shown in mize the possibility of undetected test variations; Fig. 7, lithium was circulated in the heater circuit, e.g., boiling instabilities, which might compromise and potassium was boiled and circulated in a two- loop operation or the posttest compatibility evalua- phase secondary ioop that contained turbine simu- tion. Figure 6 shows the entire test facility follow- lator test specimens of Mo-TZC and Nb-132M alloy. ing installation of the Nb-lZr test loop. Maximum and minimum temperatures in the REFRACTORY ALLOY COMPONENT ACCOMPLISHMENTS 23

Fig. 6 Ultrahigh vacuum test chamber and supporting equipment used in testing of refractory alloy Rankine test systems. lithium circuit were 1232°C and 1137°C, respec- Extensive chemical and metallurgical evalua- tively, while temperatures in the two-phase tion of the T-lll alloy containment material and potassium circuit ranged from 1171°C (superheated the Mo-TZC and Nb-132M turbine simulator vapor) at the boiler outlet to a minimum of 493°C materials indicated that these candidaLe materials in the coolest portion of the loop. have suitable compatibility with these energy During test startup, a leak developed in a butt transfer fluids for application in future Rankine weld of the potassium containment tube of the system electric power systems. The results of post- boiler. A detailed repair plan, which is extensively test evaluations on this test system are given in documented by Harrison, Hoffman, and Smith,7 the paper on compatibility lj DeVan et al. was developed and cairied out successfully with no included in these proceedings. significant contamination of the test components or the working fluids. Electromagnetic (EM) Pump The 10,000-h test was performed in a 1.2-m dia X 3.4-m high getter-ion pumped vacuum Development chamber shown in Fig. 6. Total pressure of the test Many helical induction EM pumps constructed chamber environment with the system at tempera- of Nb-lZr and T-lll were used as faciluy pumps ture was maintained at less than 2 X 10~8 torr during the NASA-sponsored proscram at General 6 2 (3.3 X 10~ N/m ) for most of the test period. As Electric. A typical pump of this :ype is illustrated for all refractory alloy system tests conducted as in Fig. 8. These pumps were heavy, relatively inef- part of the NASA program, an extensive bakeout ficient, and not suitable for utilization in srsace and outgassing procedure lasting several weeks power systems. NASA-Lewis sponsored an was employed on the test chamber and loop to pre- extensive program at General Electric to develop, vent contamination of the refractory alloy com- fabricate, endurance test, and evaluate an efficient, ponents. Posttest evaluation of test components lightweight potassium boiler feed pumr for use in indicated no significant contamination of the loop Rankine cycle space power systems.8-9 The product components by the vacuum chamber environment. of this effort was the T-lll pump illustrated 24 HOFFMAN

LOOP SUPPORT TURBINE COUNTERWEIGHT SIMULATOR

• BOILER

IRON TITANATE COATED CONDENSER

Fig. 7 T-lll Rankine system test loop prior to initiation of 10,0O0-h test in ultrabigh vacuum test chamber.

schematically in Fig. 9. A photograph of the com- ricated this component for use in a 10-KWe Bray- pleted T-lll pump duct is shown in Fig. 10. The ton system.10 The heat receiver was to function as test facility in which this pump was tested is an absorber of solar radiation incident on the shown schematically in Fig. 11. mirror-collector and as a heat, exchanger to trans- This T-lll alloy pump was performance tested fer heat into the Brayton cycle working fluid. and achieved the specified rating of 1.5 kg/sec The Nb-lZr heat receiver is shown in Fig. 12 with a developed head of 1.7 MPa at a potassium D prior to the addition of the shell/reflector assem- temperature of 540 C. The pump efficiency was bly. The 48 Nb-lZr heat storage receiver tubes 16.3% and did not degrade during the 10,000-h shown were filled with approximately 120 kg of endurance test that followed performance testing. lithium fluoride by the Oak RiHge National Laboratory prior to being welded into the heat Nb-lZr Solar Bray ton receiver assembly. Extensive welding process and Heat Receiver individual welder qualification testing was done One of the most challenging refractory alloy prior to the fabrication of the Nb-lZr Heat fabrication tasks successfully accomplished in the Receiver and all other refractory alloy components 1963 to 1972 time period was the building of the described in this paper. Typical records of these 10 Nb-lZr Erayton cycle heat receiver. Under NASA- qualification testa may be found in the Mendelson Lewis contract, General Electric designed and fab- report. REFRACTOHY ALLOY COMPONENT ACCOMPU&MENTS 25

Fig. 8 T-111 alloy parts of an electromagnetic pump duct before and following assembly and welding.

Potassium Outlet

Argon Filled '!^ Sutor Civity I ' h ,

Fig. 9 Sketch of T-111 alloy electromcgnstic (EM) potassium boiler feed pump. 26 HOFFMAN

Fig. 10 T-lll alloy electromagnetic pump duct following completion of fabrication and application of Nb-lZr insulating foil.

The component 3hown in Fig. 12, together with preheater in the potassium circuit. A schematic the surrounding shell and aperture structure, was which better illustrates the physical layout of the delivered to NASA-Lewis in 1970. A complete system is given in Fig. 14. The tertiaiy heat rejec- account of the construction of this heat receiver is tion loop of this system was partially constructed given in "Refractory Alloy Component Fabrication" of type 321 SS and used NaK as the coolant. by W. R. Young in these proceedings. T-lll alloy was the principal material of construction used to fabricate this test system; a list- T-lll Boiler Development ing of the various sizes and shapes of alloy Test System products used to build ihe system is given in Table 2. Also included in this figure is a similar This system, which wa3 designed, fabricated, list of T-lll products used to build the T-lll corro- and tested for NASA-Lewis at General sion loop described earlier in this paper. A large Electric11"14 in the 1967 to 1971 time period was number of demanding refractory alloy materials the culmination of the Rankine system component and process specifications were developed under development program. Its design and fabrication th6 General Electric/NASA-Lewis program, and benefitted greatly from the experience gained on most of them were employed in the construction of earlier refractory alloy component programs which the boiler development test system. The T-lll alloy have been described in this paper. specification for seamless tubing and pipe was used The purpose of this program was to develop and to obtain alloy product for the Boiler Development demonstrate the performance of a once-through System. This specification, which was typical of boiler in which potassium is transformed from a the refractory alloy specifications used by General subcooled liquid at the boiler inlet to a superheated Electric, may be found in Appendix C of the report vapor in a single pass through the boiler. Fig- by Harrison, Hoffman, and Smith.7 ure 13 shows the layout of the three circuits of the Several of the more important refractory alloy test system which included a 400 kW electrical process specification., which were developed by heater in the lithium heater circuit and a 30 kW General Electric and utilized in building the T-lll REFRACTORY ALLOY COMPONENT ACCOMPLISHMENTS

TABLE 2 Listing of T-lll Alloy Product Forms and Sizes Used to Build Two T-lll Alloy Test Systems

T-1I1 alloy procured by GE-NSP for T-lll alloy procured by GE-NSP for NASA NASA Contract NAS 3-6474, advanced Contract NAS 3-9426, devslopment refractory alloy corrosion loop of a single tube potassium boiler

Item and size, Weight, Item and size, Weight, in. lb in. ib

Rod Kod 0.250 dia 1 0.125 dia 4 0.500 dia 11 0.250 dia 4 0.625 dia r 0.690 dia 10 1.000 dia 40 0.750 dia 9 1.125 dia 10 1.000 dia 38 1.500 dia 13 1.375 dia 128 2.000 dia 85 1.500 dia 34 2.500 dia 93 2.000 dia 38 3.125 dia 74 2.250 dia 77 2.500 dia 303 332 3.063 dia 389 Bar 3.688 dia 121 1.0 by 1.0 30 4.438 dia 163 1.0 by 2.0 115 4.625 dia 176 145 1494 Wire Wire 0.062 dia 13 0.062 dia 10 0.094 dia 8 0.094 dia 15 0.125 dia 31 0.125 dia zr, 52 51 Foil, sheet, or plate Foil, sheet, or plate 0.005 X 3.5 2 0.005 X 3.5 1 0.009 X 3.5 1 0.005 X 5.5 1 0.035 X 1.0 1 0.040 X 20.5 55 0.040 X 12.0 29 0.063 X 6.0 3 0.125 X e.O 5 0.100 X 9.0 23 0.500 X 6.125 41 0.125 X 22.0 51 79 0.250 X 8.0 105 Tube 0.400 X 6.0 44 0.375 o.d. X 0.065 wall 66 283 1.00 o.d. X 0.065 wall 104 Tube 2.25 o.d. X 0.375 wall 40 0.250 o.d. X 0.050 wall 16 2.50 o.d. X 0.450 wall 46 0.375 o.d. X 0.065 wall 14 3.00 o.d. X 0.375 wall 50 0.500 o.d. X 0.075 wall 18 3.25 o.d. X 0.250 wall 40 0.625 o.d. X 0.008 wall 8 3.25 o.d. X 0.500 wall 73 0.690 o.d. X 0.045 wall 13 41B 0.750 o.d. X 0.04 wall 17 Total 1027 0.875 o.d. X 0.100 wall 363 1.325 o.d. X 0.100 wall 70 1.500 o.d. X 0.100 wall 190 Summary 2.375 o.d. X 0.220 wail 79 4.386 o.d. X 1.343 wall 180 4.420 o.d. X 0.537 wall 92 T-lll for Rankine System 1027 lb 4.424 o.d. X 1.362 wall 184 Corrosion Loop 4.625 o.d. X 0.225 wall 45 T-lll for Boiler Development 3261 lb 4.625 o.d. X 0.275 wall 55 Test System 4.625 o.d. X 0.409 wall 79 Total 4278 ib 1423 Total 3251 28 HOFFMAN

•a s «APOI TUP

VAPOR 'fcAP

Fig. 12 Nb-IZr solar Bray ton cycle heat receiver during fabrica- Fig. 11 Test facility used to conduct performance and endurance tion. Approximately 400 kg of Nb-IZr in this component. test on T-lll alloy potassium boiler feed pump.

RADIANT VAPOR DESUPERHEATER THROTTLE VALVE

NaKTO AIH COOLER LI" HUM ELEC1RICAL SINGLE-TUBE HEATER BOILEB 3-TUBE CONDENSER

EM EM FLOWBIETER FLOWMETER T NaK \ PRE-HEATE* EM PUMP

\ POTASSIUM LITHIUM Ef1 PUMP EM PUMP LIQUID THROTTLE VALVE

Fig. 13 T-lll alloy potassium boiler development test system. Portions of tbe low-temperature NaK cooling circuit were made of stainless steel. REFRACTORY ALLOY COMPONENT ACCOMPLISHMENTS

PREUUil IBAHSDUCEB

VACUUM (hiMIEB -

I PBfHEATtB

1 UQUID THROTTlt

I, HUTIB POWIB SUPP

NoH DUMP 11NK

— ION PUMP * HcK TEDTIARY 1OOP r," Ti 5UBUM1IION PUMP • ICHAHKA1 VKUUM PUir IURBGM01KUUR BOUGHIhG PUMP Mil IIGOD

Fig. 14 Schematic view of the potassium boiler development test rig.

Boiler Development System, were published by Major features of the primary lithium heater NASA-Lewis.15 Titles of these process specifica- loop are given in Table 3, and similar information tions are listed below. for the secondary two-phase potassium loop is given in Table 4. Two of the most challenging sub- Number Title of Specification system component design and fabrication tasks involved the lithium heater (Fig. 15) and the RM-1 Chemical Cleaning of Columbium, single-tube boiler shown in Fig. 16. The coils of the Tantalum, and Their Alloys lithium heater were made of 6.4-m lengths of T-111 RM-2 Gas Tungsten Arc Welding of pipe. The heater was designed to dissipate 400 kW Columbium, Tantalum, and at a line current of 8000 amperes with lithium Their Alloys exiting the heater coils at 1200°C. The boiler was RM 3 Electron Beam Welding of the most important development component in the Columbium, Tantalum, and test system. The single-tube configuration illus- Their Alloys trated in Fig. 16 was designed to represent one RM-4 Resistance Spot Welding of Refractory tube of a full-scale multiple tube design for an Metal Foil to Refractory Metal Advanced Rankine System. Components RM-5 Postheating of Cb-lZr and T-111 The test facility in which the Boiler Develop- (Ta-8W-2Hf) Weldments ment System was tested is illustrated in Fig. 17. The vacuum chamber shown was capable of achiev- Any future program utilizing these or similar ing a pressure of 10~9 torr. This capability was refractory alloys should consider utilizing or required to prevent contamination of the refractory upgrading these materials and process specifica- alioy system during testing. All the General Elec- tions as appropriate. A complete account of the tric test systems described in this paper, including fabrication of the Boiler Development System is the Boiler Development Test Facilities, were given in the paper by Young in these proceedings. equipped with partial pressure analyzers to moni- 30 HOFFMAN

TABLE 3 Potassium Boiler Development Test System Primary (Lithium) Loop Loop material T-lll alloy Main pipe size 1%-in. OD X 0.100-in. wall Maximum design pressure/temperature 100 psi/1204t>C Range of lithium temperature at boiler inlet 982 to 1204°C (for 1-tube boiler tests) Circulating pump Electromagnetic, T-lll alloy duct, 85 GPM at 75 psi head Loop flow rate: -capability 5lb/sec -range for 1-tube boile. tests 0.1 to 1.5 lb/sec Lithium heater: -type AC electrical resistance (3-phase) -current/voltage 8,000 amps/29 volts at 400 kW -pipe size (each of 3 coil sets) %-in. OD X 0.100-in. wall, 21 ft long

TABLE 4 Potassium Boiler Development Test SyBtem Secondary (Fotassiv a) Loop Loop material T-lll alloy Main pipe size: vapor piping to throttle and liquid piping from condenser ft-in. OD X 0.100-in. wall Maximum design pressure/temperature 300 psi/1204°C Potassium liquid temperatures at boiler inlet 371 to 982°C Potassium vapor temperatures at boiler outlet 982 to 1149°C Circulating pump Electromagnetic, T-lll alloy duct, 5 HT M at 300 psi head Loop flow rate: capability 0.5 lb/sec Potassium preheater: -type AC electrical resistance (1-phase) -current/voltage 3,000 amps/10 volts at 30 kW -coiled pipe size te-in. OD X 0.075-in. wall, 12 ft long Boiler ior initial tests Single-tube, 87 kW-thermal Condenser for 1-tube boiler tests Three-tube, 60 kW-thermal Range of potassium vapor temperatures at condenser inlet 593 to 871°C

Li TO BOILER tor the individual gas species responsible for the focal test chamber pressure. During bakeout and hot leak checking operations on the test loop prior to filling of the system with alkali metals, vrry small leaks were discovered in the T-lll alloy lithium heater and the potassium preheater. Various analytical techniques and laboratory tests confirmed that these leaks were caused by minute amounts of surface contamination by nickel picked up from welding clamps used in the fabrication of the test system. The successful recovery efforts are described in detail by Bond12 and in the paper of Young in the proceedings of this symposium.

ELECTRICAL SCHFHaTIC Detailed results of the program which involved 2750 hours of testing during the July-December 1970 time period are given by Deane.13 Hundreds of test conditions were investigated during the test campaign to verify the thermal-hydraulic perfor- MAX POWER 400 KU mance of system components. mx TEMP ?Z50°F Liquid lithium at temperatures up to 1232° C AP AT i LB/SEC 15 PSI LI HE CURRENT oono AM was used to boil potassium at temperatures up to 1174°C. The range of potassium conditions at the Fig- 15 T-lll alloy lithium heater for boiler development boiler exit ranged from wet vapor at 50% quality test system. to dry vapor with 167°C of superheat. Results REFRACTORY ALLOY COMPOTCNT ACCOMPLISHMENTS 31

INSTRUMENTED CENTERBODY 0.75" DIAMETER 0.04" HALL THICXHESS

HELICAL VANE WIRE COIL INSERT

^ \ \ \ -1.32" DIAMETER 18" RADIUS 0.1" WALL THICKNESS PR£SStBE TUBE

Li IN

K VAPOR THERMOCOUPLE WELLS ~'~'<-.-*k K VAPOR Li OUT OUT K IN

Fig. 16 T-111 alloy single-tube potassium boiler.

HIGH BAY CRANE TEST AREA

VACUUM CHAMBER BELL

INSTRUMENTATION & EQUIPMENT ROOM

CONTROL ROOM

VACUUM CHAMBER COOLANT LOOP ENCLOSURE

Fig. 17 Schematic view of the boiler development test facility including the 6-m tall by 2-m dia ultraHgh vacuum test chamber for the T-111 alloy loop. 32 HOFFMAN

2200 Lithium J

Critical Heat Fluxj 2100 - at Inflection_ Point

W =0.93 lbs/sec LI ^Initiation of W = 0.097 lbs/sec 2000 - Potassium K Superheat Q = 99.8 KW

5 O Well Thermocouples E-. 1900 - B Insert Thermocouples • Shell Thermocouples, Average of Four at End of Each Location 1440 - Heated Zone

1400 - _L 0 10 20 30 40 50 60 70 30 90 100 Length From Beginning of Heated Zone, inches

Fig. 18 Boiler temperature distribution during a typical test of the T-lll alloy single- tube boiler. obtained from measurements on the boiler test rience gained during this ten-year period totally unit included: Two-Phase-Flow Pressure Drop, justified the use of the exacting quality assurance Boiler Heat Transfer, Average Overall Boiling requirements imposed both by the customer, Heat Transfer Coefficients, Superheated Vapor NASA-Lewis, and the General Electric program Heat Transfer Coefficients, Critical Heat Fluxes, management. Detailed procurement and process and Shell-Side Liquid-Lithium Heat Transfer Coef- specifications vigorously implemented by the ficients. Results of a typical test to evaluate the program personnel were in large measure performance of the single-tube boiler are shown in responsible for the highly successful nature of the Fig. 18. The potassium conditions at the condenser component developments described in this paper. inlet ranged from 45 to 99% quality and from 593 A consistent element of the NASA-Lewis spon- to 871 °C saturation temperature. Results obtained sored programs directed at advancing the for the potassium condenser included: Condensing refractory alloy technology was that the Heat Transfer Coefficients and Shell-Side NaK-78 problems/solutions, as well as the successes, were Heat Transfer Coefficients. extensively documented in order to optimize the The program described above, which involved "lessons-learned" aspects of this very demanding the design, fabrication, and testing of a very com- technology. Many of the references included in plex three-loop system, was the high-water mark this paper contain that type of information. of the experimental program to evaluate very high Knowledge and utilization of this experience should temperature refractory alloy-liquid metal systems be most valuable to refraccory alloy system for space power applications which ended when the designers, fabricators, and test operators in the national programs were terminated in 1972. development of refractory alloy systems for space application.

SUMMARY REFERENCES

The refractory alloy components described in 1. H. E. Nichols, R. W. Fink, and W. F. Zimmerman, this paper were successfully designed, fabricated, Three-Stage Potassiuvi Vapor Turbine-Fabricaticm and and, in mo3t cases, tested and evaluated. The expe- Assembly—Final Report, NASA-CR-72501, May 1969. REFRACTORY ALLOY COMPONENT ACCOMPLISHMENTS 33

2. G. C. Wesling, Three-Stage Potassium Turbine Performance Intersociety Energy Conversion Engineering Conference, Test Summary, NASA-CR-1483, December 1969. Volume 1, Las Vegas, NV, Sept. 21-25, 1970. 3. E. Schnetzer and G. M. Kaplan, Erosion Testing of a 10. I. Mendelson, Design and Fabrication of Brayton Cycle Solar Three-Stage Potassium Turbine, ASME paper 70-Au/Sp T- Heat Receiver. Final F rjxrrt, NASA-CR-72872, July 1971. 37, presented at the Space Technology and Heat Transfer 11. J. A. Bond, Advanced Rankrne Cycle Potassium Boiler Conference, LOF Angeles, June 21-iJ4, 1970. Development Program, Volume I, Test Rig Design and 4. J. R. Peterson, J. A. Heller, and M. U. Gutstein, Status J/ Fabrication, NASA-CR-13451, November 1975. Advanced Rankine Power Conversion Technology, presented 12. J. A. Bond, Advanced Hankine Cycle Potassium Boiler at American Nuclear Society Meeting, Boston, MA, General Development Program, Volume II, Facility Checkout and Ekctric Report No. GESP-623, June 1971. Design Point Demonstration, NASA-CR-135452, November 5. R. W. Harrison and J. Holowach, Refractory Metal Valves 1975. for 19OO°F Service in Alkali Metal Systems, NASA-CR-1810, 13. C. W. Deane, Advanced Rankine Cycle Potassum Boiler April 1970. Development Program, Volume III, Potassium Boiler and 6. E. E. Hoffman and J. Holowach, Cb-lZr Rankine System Condenser: ETnerimental Results, NASA-CR-135453, Corrosion Test Loop, NASA-CR-1509, June 1970. November 1975. '.. R. W. Harrison, E. E. Hoffman, and J. P. Smith, T-Ul 14. J. A. Bond and M. U. Gutstein, Component and Overall Per- Rankine System Corrosion Test Loop. Volumes I and II, formance of an Advanced Rankine Cycle Test Rig. presented NASA-CR-134816, June 1975. at the 1971 Intersociety Energy Conversion Engineering 8. A. H. Powell and J. C. Amos, Fabrication and Test of a Conference, Boston, Massachusetts, August 3-5, 1971. Space Power Boiler Feed Electromagnetic Pump, NASA- 15. T. J. Moore, P. E. Moorhead, and K. J. Bowles, CR-1951, March 1972. Specifications for Cleaning, Fusion Welding and Postheating 9. A. H. Powell and J. P. Couch, Boiler Feed EM Pump for a Tantalum and Columh'.um Alloys, NASA-TMX-67879, July Rankine Cycle Space Power System, in Proceedings of Fifth 1971. Compatibility of Refractory Alloys with Space Reactor System Coolants and Working Fluids

J. B. DeVan,* J. R. DiStefaao,* and E. E. Hoffmont *Oak Ridge National Laboratory^ fDepartment of Energy, Oak Ridge Operations

INTRODUCTION AND OVERVIEW complexity. Thus, testing routines have progressed from an evaluation of specimens encapsulated in The principal compatibility issue affecting the static liquid metal containers through natural cir- development of refractory metals for power appli- culation looi>s, forced convection loops, loops simu- cations concerns their corrosion resistance against lating reactor circuits, and large engineering exper- the liquids and gases that are used as heat trans- iments. Inevitably, the volume of data in -.uded in port and working fluids. There is, however, a spe- this report is greatest for capsules and simple loop cial compatibility concern in the case of refractory experiments, but the value of the data for reactor metals that relates to the ambient environment design purposes is greatest for the more complex surrounding the metal. This assessment reviews loop experiments. Thus, while we have not the present state of knowledge concerning the attempted to screen or critique any of the data above compatibility issues. It should be noted that sources, our perceptions of the design status and it does not address compatibility problems arising research needs have been guided primarily by from other solids in the power plant, the most those test results obtained under conditions that notable being the nuclear fuel. are most representative of steady-state power Extensive information exists on the compatibil- plant parameters. ity of select niobium- and tantalum-base alloys We have, in this assessment, minimized report- with liquid lithium and boiling potassium. Thus, ing liquid metal corrosion test results of refractory the bulk of this report deals with compatibility metals obtained in conventional alloy loops. Some studies in tut latter two alkali metals. Substantial results are included to emphasize the sensitivity of information is also presented concerning the reac- refractory metals to interstitial contamination tivity of niobium and tantalum alloys with residual from conventional alloys. The reader should be gases in high and ultrahigh vacuum atmospheres. aware that the effects of interstitial contaminants The remaining information, which is much less in a mixed refractory metal-conventional alloy sys- extensive, covers the compatibility behavior of tem are basically different from those in an molybdenum and tungsten alloys in alkali metals unmixed system, and the test results from and a qualitative assessment of the use of refrac- monometallic systems should not be extended to tory metals for containing helium in a closed Bray- mixed systems. Because the latter systems, in the ton cycle. authors' opinion, do not appear amenable to space One of the principal objectives of this assess- reactor service, we have also omitted their consid- ment was to document the full record of space eration in formulating research needs. reactor corrosion studies insofar as they were known to the authors. The studies have tended to follow a classical screening pattern, ever-narrowing LIQUID METALS in the number of materials but increasing in test Liquid Lithium ^Operated for the U. S. Department of Energy under con- The principal corrosion reactions in refractory tract W-7405-eng-26 with the Union Carbide Corporation. metal-lithium systems are due to dissolution, mass

34 COMPATIBILITY OF REFRACTORY ALLOYS 35 transfer, and impurities. Isothermal dissolution is TABLE 1 governed by the solubilities of refractory metal ele- ORNI Determination ot Depth of Attack by Lithium ments in lithium. Temperature-gradient i.iass of Niobium-Zirconium Alloys RS a Function of transfer is controlled by the temperature- Zirconium Concentration, Oxygen Concentration, dtspendence of the solubilities and the Kinetics oi and Heat Treatment' dissolution and deposition. Dissimilar metal mass (Test Conditions: 10O h at Klfi°C) transfer results when an activity gradient for an element exists from two or more alloys being pres- Depth of attack (am) ent in the same system. Impurity reactions, espe- Concentration After After heat cially those involving interstitial nitrogen, carbon, oxidation treatment and especially oxygen, are easily the most serious Zirconium Oxygen at 1000T* at 1300°Ct corrosion problem in refractory metal-lithium systems. O:IH O.Mx} 0.05 0 IS 0 fi-l O.ti4i Oak Ridge National Laboratory (3RNL) 0.05 11.23 (ITfif 0.7fii 04 (109 During the mid-1950s, static lithium tests of D4 (1 l.s 0.50 niobium, tantalum, molybdenum, tungsten, and 0.4 0.23 0.7HT rhenium were conducted, and low velocity temper- OH ".09 0.25 0 ature gradient tests (seesaw) were performed on O.rt 0.18 0.50 U niobium and molybdenum in flowing lithium at 0.6 0.23 0.76§ 0.64| maximum temperatures of 870 and 1040° C, 1 0.9 0.09 0.25 0 respectively. No significant attack was noted in 0.9 0.18 0.50 0 any of these tests; however, in a thermal convec- 0.9 0.23 0.76§ o tion loop test conducted in a vacuum of 1 to 5 pm 1.3 0.09 0.13 0 at a maximum temperature of 816°C, gross attack 1.3 0.18 0.25 0 of unalloyed niobium by lithium was noted. These 1.3 0.23 O.fil 0 results and subsequent capsule tests performed to evaluate the effects of the oxygen and nitrogen •Specimen exposed tn lithium after oxidation at lOOOT. concentration of niobium on its corrosion resis- tSpecimen heat treated in vacuum for 2 h at 1300°C follow- tance to lithium led Hoffman1 tc conclude that the ing oxidation and prior to exposure to lithium. attack was due to oxygen in the niobium and that tOxyKen-to-zirconium atomic ratio was p-pater than 2. additions of a stable oxide former should make §Complete penetration of 1.52-m;n-thicU specimen. niobiui.i resistant to lithium attack. Further systematic studies were then conducted by DiStefano2 4 3 Oak Ridge National Laboratory investigated ana Klueh who identified the effect of various the compatibility of several other niobium alloys parameters such as time, temperature, and grain with lithium in static capsule tests at 500 and size on the process and extended the studies to 1000°C. The alloys investigated were: D-43 (Nb- other refractory metals and alloys. If oxygen in 10%W-l%Zr), B-66 (Nb-ll%Zr-5f7r Mo-i? V), FS-85 niobium exceeds —400 ppm, rapid intergranular (Nb-28%Ta-10%W-l%Zr), Cb-752 (Nb-lO^W- and/or transgranular attack by lithium occurs at 3^Zr), and Cb-7F>3 (Nb-l^ Zr-5^ V). Oxygen- temperatures from 300 to 1200°C Penetration contaminated specimens were attr.cked by lithium resulu from the formation of a terii^ry oxide on at both temperatures, but su?f .-tiMlity to attack grain boundaries or preferred crystallograpMc 3 was eliminated by a heat treatment at 1300 to planes ; this type of corrosion drastically lowers 1600°C before exposure. Resistance to corrosion the strength and ductility of the niobium.2 How- afforded by heat treating was again attributed to ever, URNL found that addition of zirconium to precipitation of zirconium oxide. niobium would tie up oxygen as ZrO2 and greatly 2 3 increase the threshold concentration for corrosion Capsule studies of tant him - and T-lll (Ta- to occur.2 There is a direct relation between the 8%W-2%Hf) (Ref. 4) in lithium revealed behavior zirconium and oxygen concentration in niobium similar to niobium and niob'um alloys except that the threshold concentration of oxygen that resulted (one Zr atom will tie up two O atoms) and lithium 3 attack (Table 1). However, it is generally neces- in attack was lower (—100 ppm). Heat treating T-lll at 1300 to 1600°C was similarly effective in sary to heat treat Nb-Zr alloys to precipitate ZrO2 before corrosion resistance is attained. Tempera- tying up oxygen as HfO2 thereby restoring its cor- tures from 1000 tr 1600°C were found to be effec- rosion resistance. tive. In other static capsule tests niobium and tantalum were exposed to lithium at 600°C that con- 36 DEVAN. CHSTEFANO, AND HOFFMAN tained oxygen concentrations of up to 2100 ppm. Two dynamic tests of tantalum alloys were con- Specimens showed no weight change, were bright ducted by ORNL. Results of a natural circulation and shiny, and the amount of metal in the lithium loop test of alloy T-222 (Ta-9.5%W-2.5%Hf) are 3 shown in Table 2. Comparison of weight change after test was less than the limit of detection (<10 8 ppm). results from this test (maximum temperature Oak Ridge National Laboratory also conducted 1350°C> with that of the 1297°C test of Nb-l%Zr several tests to evaluate mass transfer of niobium shows that T-222 has somewhat more resistance to and tantalum base alloys in nonisothermal lithium. temperature gradient mass transfer; however, some See-saw test results reported by Hoffman1 s^ ->d mass transfer of hafnium from hot to cold regions small weight changes, and no deposition was noted did occur. In addition, the T-222 that was exposed to temperatures from 1220 to 12&u°C was embrit- for molybdenum (1038°C, 445°C AT) and niobium 9 (871°C, 278°C AT) capsules after 150 and 300 h, tled due to an aging reaction. respectively. As a culmination of their refractory metal- In 3000 h natural circulation loop tests of lithium compatibility studies, ORNL built and Nb-l%Zr and D-43 (Table 2) at 1200°C, the small operated an engineering-scale forced-circulation weight changes observed were the result of trans- loop fabricated from alloy T-lll (Ta-8%W- c 10 port of zirconium and nitrogen between the heated 2 /< Hf). Tne loop is shown schematically in Fig. 1, and cooled regions.5-6 Transfer was greater in and design information and operating conditions Nb-l%Zr and resulted in a light zirconium nitride are shown in Table 3. Results are summarized as mass transfer deposit on cold-leg surfaces. In the follows. 1300°C test of Nb-l%Zr, transport of niobium 1. Mass transfer rates were small, and maxi- accounted for a large fraction of the weight mum dissolution rate was <1.3 /im/y. changes measured, and niobium deposits contain- 2. Hafnium, nitrogen, and carbon were ing carbon and nitrogen were found on hot-leg transferred from hot to cold regions. specimens. A test of the niobium alloy FS-85 3. Weight changes were similar to those in nat- operated for 500 h (Table 2) before a leak forced ural circulation loop test of T-222 (Table 2). termination of the test. However, results were 4. No significant microstructural changes other similar to those for Nb-l%Zr tested for 3000 h than grain u-rowth in T-lll were found. under similar conditions.7 5. Room-temperature tensile strengths of-T-lll One dissimilar alloy test was conducted as specimens were lower than control specimens after shown in Table 2. A Nb-l%Zr loop with both TZM equivalent vacuum exposure, which was attributed and Nb-l%Zr specimens .vas operated for 3000 h to increased grain size and changes in interstitial under the conditions shown. Weight changes in impurities (primarily decrease in oxygen content). both Nb-l%Zr and TZM weie higher than in the Pratt & Whitney Aircraft (PWAC-CANEL) Nb-l%Zr monometallic loop test run under similar conditions. Chemical and metallographic analyses Concurrent with the early ORNL studies PWAC showed that niobium and zirconium had also investigated the compatibility of refractory transferred to the TZM, and Nb-l%Zr picked up metals with lithium for compact reactor systems,11 molybdenum, carbon, and nitrogen. Solubility measurements12 indicated dissolution of

TABLE 2 Natural Circulation Loop Tests of Niobium and Tantalum Alloys in Lithium Conducted at OkNL

Temperature, °C Flow Weight change, mg/cm2 Time, Maximum Minimum velocity TUf tronof Alloy h hot leg cold leg (mm/sec) Maximum gain Maximum loss deposits

Nb-l%Zr 3000 1190 1080 53.8 +0.78 -0.34 ZrN Nb-l%Zr 3000 1297 1118 23.9 +3.8 -2.6 Nb, ZrN, Nb(C,N) D-43 3000 1200 1030 38.1 +0.21 -0.15 None FS-85 500 1220 980 22.9 +0.79 -0.13 ZrN T-222 3000 l.?50 H50 35.6 +2.1 -0.70 Hf(N) Nb-l%Zr/TZM 3000 1200 1090 53.8 +2.7 -4.49 (TZM) Nb, Zr +0.9 -1.6(Nb-l%Zr) Mo, C, N COMPATIBILITY OF REFRACTORY ALLOYS 37

tests showed that carbon and nitrogen in lithium / . T-111/SS ; TRANSITION react with niobium, whereas oxygen in lithium does not. In unpublished studies of niobium in lithium, Leaven worth, Geary, and Bratton12 showed that the solubility of niobium increased LCONOMlZKR from less than 20 ppm to over 600 ppm when nitrogen in lithium increased from 38 to 260 ppm. Mixtures of niobium nitrides (Nb2N and NbN) and carbide (NbN) were also identified in the heated regions of several forced convection loops. Nitride and carbide layers (usually <13 ^mj were found in Nb-l%Zr loops with stainless steel cladding. This is the same type of dissimilar metal mass transfer reported by ORNL where carbon and nitrogen from stainless steel transfer to niobium and niobium alloys in sodium and NaK systems (see section on Sodium and Potassium) The PWAC studies also confirmed the oxygen effect on corrosion of niobium and Nb-l%Zr alloys by lithium. Although the initial oxygen content of Nb-l%Zr alloy test loops did not exceed the FILL, DRAIN, i uND EXPflNSIONF* threshold oxygen level for lithium attack, tests were run in chambers containing purified helium. Impurities from desorption of insulation and metal Fig. 1 ORNL T-lll alloy forced-circulation liquid lithium surfaces resulted in heavy contamination in some loop schematic.10 areas. Chemical analyses of one Nb-l%Zr alloy loop material showed that it contained as much as 1900 ppm oxygen, and since the material did not TABIE 3 receive a ZrO2 precipitation heat treatment, Engineering Design Information fur T-ill lithium ai.Uck was noted. Forced-Circulation Liquid Lithium Loop Test at ORNL'° Preliminary corrosion evaluations were also made by PWAC on binary alloys of niobium with General information Al, Cr, Fe, Hf, Mo, Ni, Ta, Ti, V, W, and Zr by testing in the 982°C hot zone of a Nb-l%Zr thermal Material of constriction T-lll (Ta-8%W-2%Hf) convection loop for 250 h. Results for binary Heat-transfer flu'd Lithium Flow rate 141 to 160 g/sec (~ 5.8 gpm) alloys with molybdenum, tantalum, and tungsten Maximum Reynolds number 99,000 at heater outlet were similar to those of unalloyed niobium, but those with the other alloys were somewhat Pressure drop in i).40-mm-ID (0.370 in.) tubing 20 kPa/m (0.9 psi/ft) improved over unalloyed niobium. Autogenous Maximum temperature 1370°C welds of a number of niobium alloys were also Minimum temperature 1200°C exposed to flowing lithium in the 982° C hot zone of Number of test specimens a Nb-l%Zr thermal convection loop. Alloys of in loop 93 niobium with molybdenum, tantalum, and tungsten exhibited slight attack in fusion, heat affected, and base metal regions (probably oxygen related); other both niobium and molybdenum by lithium was very alloys contained strong oxide formers (Ti, Zr, Y, low to 1000°C. No significant mass transfer was Al) and generally showed good corrosion resist- found in Nb-l%Zr in pumped loop tests for 2000 h ance. at maximum temperatures of 1200°C with about Forced convection loop tests of Nb-20%Ti to 200°C AT. In 500-h pumped loop tests, 50 ^m of 1093°C and Nb-5%Mo-2%Ti-2%Zr to 871°C showed solution attack was seen in the 1330°C hot zone both alloys have good resistance to lithium attack region, but no deposition was observed in the even with high oxygen levels in the alloys (to 900 1100°C cold zone. Extensive mechanical properties ppm). tests demonstrated no stress dependent corrosion It is significant to note that during the course effects at levels to 40,000 psi. Capsule and loop of the PWAC study more than 100 tons of Nb- 38 DEVAN, CKSTEFANO, AND HOFFMAN l%Zr in a variety of mill forms were obtained of candidate alloys utilizing capsules to low- according to CANEL specifications CS-1830 to CS- velocity thermal convection test loops and, finally, 1838. to large high-velocity two-loop pulped systems involving a primary iiquid heater circuit and a sec- NASA-Lewis Research Center (LeRC) ondary two-phase boiling alkali metal circuit. During the 1970s much work on liquid metal Only a very limited amount of special capsule compaiibility with refractory metals was carried testing was done at GE, because by 1962 there was out by NASA at LeRC in studies that supported an extensive data base o& candidate refractory nuclear space power systems. Corrosion of oxygen- alloys in alkali metals on which to base alloy selec- doped tantalum by lithium at 1115 and 1390°C was 13 tion for system application. The major goal of the studied in capsule tests. Results showed that NASA/GE program was to conduct two-loop Ran- lithium getters oxygen from tantalum of low oxy- kine system tests under conditions an prototypic as gen concentration. In contrast, lithium was found was practical. to attack oxygen-doped tantalum both trans- granularly and intergranularly in a manner similar As indicated in the introduction, the authors to that leported previously by FiStefano2 and have chosen to use each of the various liquid Klueh3 at ORNL. A study14 was also made to deter- metals as the major heading and discuss the com- mine the effects of exposure of T-lll to low pres- patibility of the respective refractory alloys under sure air on its resistance to corrosion by lithium in that heading. Therefore, the results obtained in capsule tests at 980 and 1260°C. Capsules exposed the lithium primary loop of the GE T-lll Rankine at 980°C and 2 X 10~4 torr failed, but it was System Loop will be presented in this section, and reported that no lithium corrosion occurred under the results obtained in the boiling potassium sec- any conditions even though the tantalum contained ondary loop will be discussed in the potassium sec- oxygen concentrations as high as 3500 ppm. Weld tion of this paper. zone failure of one capsule after 24 h was attrib- The various lithium compatibility tests con- uted to intergranular oxidation, but photomicro- ducted by GE are discussed below. All the GE graphs of the failure region showed features that alkali metal tests used specially purified alkali more resemble lithium attack than oxidation. The metals and were conducted in getter-ion pumped purpose of this study was to determine the tolera- chambers at pressures in the 10~7 to 10~9 t^-r ble vacuum for ground testing of T-lll systems. range. Although the study concluded that 1-mm-thick T- 111 could be safely tested at a residual air pressure Lithium Capsule Tests. The purpose of these of 2 X 10~5 torr for 96 h, microstructural capsule tests16 was to determine the effect of oxy- interpretations upon which the conclusions are gen contamination (typical of that encountered based are open to question. during hot isostatic pressing in high pressure helium) on the corrosion resistance of T-lll and Isothermal capsule compatibility tests were con- W-30%Re-30%Mo in high temperature lithium. ducted rn T-lll with uranium nitrite or TZM for 2800 h at 1040°C in lithium.15 All combinations The T-lll specimens contained about 2300 ppm were concluded to be compatible except for T-lll oxygen before exposure to 1425°C lithium and 24 in direct contact with uranium nitride. There was ppm after a 100 h lithium exposure. No attack was no evidence of carbon transport from TZM to T-lll detected; this is attributed to the oxygen gettering in these tests. action of the hafnium in the T-lll alloy prior to exposure to lithium. General Electi^ (GE-Cincinnati, Ohio) The W-30%Re-30%Mo (a/o) specimens picked Compatibility studies covering a wide range of up very little contamination during autoclaving refractory alloys in contact with lithium, sodium, and were unattacked by the 1425°C lithium after a and potassium were carried out in the period from 100-h exposure. 1962 to 1972 at General Electric. Essentially all of A series of experiments17 were conducted to these studies were performed under contract to determine the leval of welding environment purity NASA-LeRC us a part of the overall technology and the postweld heat treatment required to avoid base being generated by NASA to support the lithium attack. In these tests specimens of Nb- development of advanced space electric power ^ys- l%Zr, FS-85, and T-lll were exposed to 760°C for tems. 8 h. It was determined that the welding environ- The result of this compatibility testing will be ment should contain 200 ppm or less air and that presented in order of increasing severity of the postweld heat treatment was required to avoid testing proceeding from static, noncirculating tests lithium attack. GvwvlPATlBlLJTY OF REFRACTORY ALLOYS 39

Lithium Thermal Convection Loop Test A nat- Hydrogen transfer was not detected. Essentially no ural circulation (thermal convection) loop test carbon transfer from th" ASTAR alloys was was conducted18 on T-lll, ASTAR-811C (Ta- detected, which indicates that degradation of the 8%W-l%Re-l%Hf-0.025%C) and ASTAR-1211C strength properties of these alloys due to possible (Ta-12%W-l%Re-l%Hf-0.025%C) in lithium for carbon losses to flowing, high-temperature lithium 5000 h at a maximum temperature of 1370°C with should not be a problem. a 200°C temperature gradient. The primary objective of this test was to determine if ASTAR-811C Lithium Forced Convection Loop Tests. A T- would decarburize in this environment. The loop 111/lithium valve test loop was conducted by Gen- was designed, fabricated, and operated by the Gen- eral Electric for NASA-LeRC.19 The purpose of this eral Electric Co., but results were analyzed and experiment was two-fold: (1) to determine the per- evaluated by NASA-LeRC. No gross corrosion formance of refractory metal valves and the effects were noted, but some evidence of mass lithium corrosion resistance of the various refrac- transfer was noted. Weight changes are shown in tory metals and alloys (Re, W-25%Re, T-lll, and Table 4. Some mass transfer is observed. There Nb-l%Zr) used to construct the wetted parts of the was a direct relation betwet t u ..ogen change and valve and (2) to determine the compatibility of the ' eight change. Very thin depobK with high haf- various refracto-y alloy tensile/corrosion speci- ni m contents were found in the region of high mens (T-lll, T-222, ASTAR-311, ASTAR-811C, nitrogen contents on ASTAR-811C and ASTAR- ASTAR-811CN; W-Re-Mo Alloy 256). Figure 2 1211C. Specimens were gentrnllv lower in oxygen. illustrates the various components of the T-lll

TABLE 4 Weight Change and Chemical Analyses of Tested Specimens from Lithium Thermal-Convection Loop Test* Conducted by General Electric and Lewis Research Center18

Distance from Chemical content, ppm heater inlet. Temperature, Weight change, Specimen °C* mg/cm2 Oxygen Nitrogen Carbon Hydrogen

T-lll T-T-2 6.0 1240 -0.28 T-T-3 20.0 -0.52 4 2 9 T-T-4 34.0 -0.18 4 <2 8 T-T-5 48.0 -0.82 2 <2 20 T-r-l 60.0 -0.87 13 3 61 1.0 T-C-2 62.5 ' -0.80 T-C-3 65.0 -0.93 T-C-4 77.0 -0.39 14 1 61 1.2 T-C-5 79.5 -0.69 T-C-6 82.0 -0.82 T-C-7 93.5 -0.77 12 3 62 1.2 T-C-8 96.0 1.330 -0.80 T-C-9 98.5 -0.86 T-T-C 109.5 -0.75 2 2 18 T-T-7 124.0 1265 -0 55 4 <2 10 T-T-S 138.0 -0.51 T-T-9 152.0 122U -0.34 5 4 11 T-T-10 166.0 1205 -0.26 T-C-10 178.0 -0.30 12 16 55 0.8 T-Cli 180.5 -0.30 T-C-12 183.0 -0.30 T-C-13 194.0 -0.0.21 10 14 50 1.2 T-C-l-J 197 0 -I.-.34 T-C-15 200.0 -0.23 T-C-16 210.5 -0.17 16 16 26 0.6 T-C-17 213.5 -0.25 T-C-18 216.0 1165 -0.32 9 10 18 0.2 T-T-l 229.0 -0.13 8 9 17 (Table continues on the next page.) 40 DEVAN, DISTEFANO, AND HOFFMAN

TABLE 4 (Continued)

Distance from Chemical content, ppm heater inlet, Temperature, Weight change. Specimen cm °C* mg/cm2 Oxygen Nitrogen Carbon Hydrogen

ASTAE-811C A-T-2 12.5 0.44 16 39 195 A-T-3 27.0 -0.13 6 2 183 A-T-4 41.5 -0.18 5 <2 147 A-T-5 55.5 1360 -0.27 5 <2 172 A-T-6 116.5 -0.14 <2 <2 195 A-T-7 130.5 -0.07 3

ASTAR-1211C A-C-l 68.0 1350 -0.42 11 3 348 2.1 A-C-2 71.0 1.45 A-C-3 74.0 -0.32 A-C-4 84.5 -0.30 8 2 326 0.8 A-C-5 87.0 -0.41 A-C-6 90.5 -0.3! A-C-7 101.5 -0.35 . 10 1 353 1.4 A-C-8 105.0 -0.36 A-C-10 185.5 1185 5.3G 26 108 367 1.6 A-C-ll 188.0 1185 5.51

A-C-12 191.0 5.89 A-C-13 202.0 5.13 26 105 330 1.6 A-C-14 205.0 5.04 A-C-15 208.0 5.35 A-C-16 218.5 5.30 23 117 398 ?..O A-C-17 221.5 5.62 A-C-18 224.0 5.49 18 142 245 0.4

'Measured temperatures shown. Continuous temperature gradient between measured locations.

valve test loop. A section through the valve is During the 5000-h test the isolation vahre was given in Figure 3, which shows the portion of the cycled for 95 cycles between the full open and full valve in contact with the flowing lithium. The closed position, and the torque to open and close refractory metal valves were developed and per- the valve on each cycle was measured. The meter- formance tested to satisfy the need for both metering valve was cycled between lA and % travel during and isolation valves of this type for subsequent ing this same period. The W-25%Re plugs and use in both liquid lithium and potassium Rankine rhenium seats in each valve exhibited excellent systems at temperatures in excess of 1040°C. Only performance with no sign of wear, abrasion, bond- the refractory alloy portions of the valves are ing, or lithium attack observed in either valve. The shown in Fig. 3. The design of the metering and T-lll bellows and other refractory metal com- isolation valves was done by Hoke, Inc. General ponents of the valves performed satisfactorily dur- Electric selected and procured the materials of con- ing the test. struction and performed all the joining operations. Both corrosion and tensile specimens of T-lll The material initially selected for the valve seats (Ta-8% W-2%Hf), T-222 (Ta-10.4% W-2.4%Hf- was TiC plus 10% Nb, but a lithium corrosion test 0.01 %C), ASTAR-811 (Ta-8%W-l%Hf-l%Re), AS- conducted to confirm this choice revealed extensive TAR-811C (Ta-8% W-0.7%Hf-l%Re-0.025%C), AS- lithium attack. Unalloyed rhenium was then TAR-811CN'Ta-8%W-I%Hf-Re-0.012%C-0.012%N), selected and tested in lithium at 1150°C for 100 h and W-Re-Mo Alloy 256 (W-29%Re-18%Mo) that to confirm this choice before use in vaive construc- were located in the portion of the loop circuit, tion. which operated at a maximum temperature of COMPATHJTY OF REFRACTORY ALLOYS 41

VACUUM UN«

— COPPEB IUS BAB

BESISTAHC! HEATED COILS

ISOLATION VALVE

Fig. 2 High temperature T-Ill/lithium vaK •-: test loop.

IM PUMP PUMP IhLET ^' ;/ / v; >'

PBESSUBE TBANSDUCEBS i

GAS PSiSSUSIZAIION UHl-i

DRAIN AND FILL LINE —

/ /

1150°C with a 95°C temperature drop along the T-lll clad fuel specimens, tensile specimens of W- length of the specimen containment, were Re-Mo Alloy 256, ASTAR-811CN, and ASTAR-811C evaluated following the test, and no lithium attack and corrosion test specimens of T-lll, ASTAR- was detected. The results of room temperature and 811C, and ASTAR-811CN were exposed to 1040°C 1090°C tensile tests following the 5000-h exposure flowing lithium. showed no changes in strength or ductility that The fuel element specimens were examined by could be attributed to the lithium exposure. NASA-LeRC staff members, and these results as A T-111/fuel element pumped lithium loop well as other compatibility information relevant to test20 was conducted (1) to determine the compati- the use of T-lll clad UN fuel in lithium are bility of bolh sound- and defective-clad fuel speci- reported in NASA-LeRC reports.2123 No corrosion mens and (2) to determine the compatibility and of the T-lll fuel clad by the lithium or contamina- posttest mechanical properties of corrosion and tion of the T-lll by the UN was detected. The fuel mechanical property specimens of T-lll, AcJTAR- specimen with the machined defect* (to simulate a 811C, ASTAR-811CN, and W-Re-Mo Alloy 256. A schematic of the loop and the fuel section of the *A slot 0.076 mm wide hy b.4 mm long cut through the loop is shown in Pig. 4 In addition to the three T-* 11/tungsten cladding to the UN fuel pellet surface. DEVAN, rxsTEFANO, AND HOFFMAN

Stem Fig. 3 Cross section of the T-lll valve showing the various refractory metal components in contact with flowing lithium.

:j Stainless Steel VASCO Hypercut

T-lll Cladding

Fuel Fletwnt Test Section

Corrosion Specimen Fig. 4 Pumped lithium Assembly loop to evaluate advanced refractory metal alloys and T-lll clad UN fuel specimens.

EM Pump

LOOP SCHEMATIC Flou

FUEL ELEMFNT TETT SECTION COMPATHUTY OF REFRACTORY ALLOYS

1890

Turbine Simulator

Fig. 5 T-1I1 Rankine system corrosion test loop. (Note: all temperatures are in

Condenser

cladding crack) showed a very slight loss of UN in NASA Rankine space electric power development the immediate area of the defect but no evidence of program. The test loop consisted of a T-111/lithium any compatibility problem between the UN and the primary heater loop operating at a maximum flowing lithium. No corrosion of the Mo-TZM fuel lithium temperature of 1230°C and a mini-num of capsule spacer inserts (see Fig. 4) was detected. 1135CC. The potassium secondary two-phase loop Detailed evaluation of the 38 tensile and corro- circuit will be discussed in the GE/potassium sec- sion test specimens exposed in the 7500-h test tion of this paper. A schematic of the test loop, revealed no attack of the specimens, no loss of car- including the temperatures of the various portions bon or nitrogen from the ASTAR-811C or ASTAR- of hoth the lithium and potassium circuits of the 811CN specimens, and no lithium-related degrada- loop, is shown in Fig. 5. A photograph of the test tion of the tensile properties of the alloys. loop mounted on the vacuum chamber spoolpiece A two-loop T-lll Rankine system corrosion test prior to transfer to the ultrahigh vacuum test loop24 was the culmination of the NACA-sponsored chamber is given in Fig. 6. The test loop was GE compatibility test program in support of the operated without interruption and terminated on DEVAN. DBTCFANO, AND HOFFMAN

Fig. 6 T-lll RanMne system corroeion test loop and vacuum chamber spool piece prior to installation in vacuum test chamber. completion of the planned 10,000-h run on mentioned earlier, the results of the evaluation of March 18, 1970. The lithium flow rate in the the two-phase potassium circuit are included in the heater circuit wa3 106 kg/h (234 lb/h). potassium section of this paper. After the test was completed, an extensive eval- The excellent compatibility of the T-lll alloy uation was performed on all loop components. Only with lithium at elevated temperatures is attested the results of the evaluations on T-lll components to by the fact that detailed metallurgical and in the lithium circuit will be discussed here. As chemical evaluations revealed no compatibility COMPATBBJTY OF REFRACTORY ALLOYS 45

VACUUM AND ARGON VACUUM AND ARGON PRESSURE TRANSDUCER

RESERVOIR

PUMP

SUMP

Fig. 7 High velocity lithium system for Nb-l%Zr test by Jet Propulsion Laboratory.26 problems. The oxygen concentration of the T-lll The Jet Propulsion Laboratory evaluated the components exposed to lithium was reduced from compatibility of Nb-l%Zr in high velocity, high initial levels of —20 ppm to posttest values of 5 to temperature lithium for application in magnetohy- 10 ppm. On the lithium side of the potassium drodynamic power systems in space.26 A schematic boiler tube (see inset, Fig. 5) a very thin film of of the experimental system used in this study is hafnium nitride (maximum thickness less than shown in Fig. 7. Lithium was pressurized and 0.025 mm) was detected in the very high heat flux heated in an electromagnetic induction pump with region of the boiler where boiling of potassium was a helical Nb-l%Zr duct and flowed through an occurring on the other side of the T-lll tube wall. electromagnetic flow meter to the test section This surface film had no deleterious effect on the where acceleration to higher velocities occurred. metallurgical reliability of the boiler tube. The primary flow circuit of the test system was The conclusion drawn from the GE lithium fabricated of Nb-l%Zr and was operated within a compatibility studies was that tantalum alloys of vacuum chamber. More than 20 niobium weldments the T-lll type have excellent compatibility to high were required for final assembly, and these were velocity lithium at terperatures to at least 1230°C. performed in an inert-gas weld ehamho- or by electron-beam methods. Nb-l%Zr and Y2O3 plates Others were mounted in the high velocity test section to Many other organisations have also contributed measure corrosion/mass transfer effects. The sys- to the compatibility data base on refractory metals tem was operated for two periods at constant tem- in lithium. One significant study was earned out perature. For 109 h the temperature was main- by DeMastry25 at Battelle Memorial Institute, tained .it I143°C and 48.5 m/s maximum lithium Columbus, Ohio. Corrosion tests of W-0.9%N'o, W- velocity. Examination of the test section indicated 10%Re, W-25%Re, TZM, Mo-50%Re, and Re were that complete dissolution of the Y2O3 plate had carried out in TZM capsules containing lithium for occurred. After reinstallation of the test section 100 and 1000 h at 1370, 1540, and 1650°C. At and piping, the system was operated an additional 1370°C all materials were judged to be resistant to 391 h at 1073° C. Macroscopic examination showed attack. After 1000 h at 1540°C, varying degrees of that Nb-l%Zr material removal was minor in the surface dissolution and grain boundary penetration high-velocity flow regions, but material removal of the tungsten-base alloys occurred. Chemical was heavy downstream of a cavity left by dissolved analyses of the lithium showed large increases in Y2O3. tungsten content (to 4300 ppm), but analyses of the A pumped capsule test of W-25% Re-30% Mo TZM indicated no dissimilar metal mass transfer was conducted by Lawrence Radiation Laboratory had occurred. At 1650°C there was slight dissolu- as part of a study for liquid-metal-cooled space tion of the tungsten-base alloys after 100 h which power reactor applications.27 The pumped capsule became rather severe after 1000 h. The molybde- consisted of a closed tube 1.3 cm-OD X 0.1 cm wall num alloys were resistant to attack under all con- thickness X 25 cm long containing a 0.4 mm thick ditions. Rhenium was not attacked by the lithium, plate that separated the tube into two channels. but there was evidence of mass transfer (surface (Fig. 8). Lithium was circulated by an electromag- layers) of molybdenum from the TZM capsules to netic pump at 0.3 m/sec at 1400°C maximum tem- the Re samples. perature with a AT of 150 to 200°C. The test was 46 DEVAN, DISTEFANO, AMD HOFFMAN

Transverse ring from capsule J—h hot end — Metallography

Longitjdinal strips from capsule hot section containing butt weld - Mechanical properties

Specimen from capsule- jurfoce at hot end — Micrography Longitudinal specimen- A thru butt weld — • Electropolished spot " 7 - Metallography Surface micrography and electron microprobe analysis Electropol ished spot * 1 — Surface micrography and electron micropiobe analysis Capsule Splitter platt"

LonqltudinaT l strips from capsule Longitudinal strips froTm midsection — Mechanical properties splitter plate midsection Mechanical propeities r o • Current - carrying electrodes

~ Electropol ished spot ' " - \ Surface micrography an^ electron microprobe analysis " Electropcloshed spot * 4 - Surface metallography and electron microprobe analysis I.

L Transverse ring from capsule cool end— Metallography

Fig. 8 Pumped capsule lithium corrosion test system conducted by Lawrence Radiation Laboratory.27 COMPATIBILITY O Rt-FRACTORY ALLOYS 47 operated in a vacuum of 10 8 torr for 1000 h. No of the Nb-l%Zr pipe wall by prior lithium expo- gross corrosion effects were noted; surface sure, additions of strong oxiae formers (such as Ca microprobe analysis, however, showed mass trans- or Y) to the lithium working fluid, and the pres- fer from high temperature to low temperature ence of small amounts of yttrium as an alloying regions. Unpublished data showing higher solubil- addition in tantalum. ity of molybdenum in lithium compared with tung- Tungsten and W-26%Re heat pipes, on the sten was presented to explain the higher rate of other hand, could be operated effectively with mass transfer of molybdenum that was found com- lithium at 1600°C for as long as 10,000 h with no pared with tungsten and rhenium. Grain boundary special requirements for oxygen scavenging.28'31 attack or intrusions to 25 nm were noted. One Busse and coworkers29'30 conducted posttest explanation postulated was that grain boundaries examinations of failed Nb-l%Zr and tantalum heat may be enhanced in molybdenum due to kinetic pipes to evaluate the oxygen distribution in the consideration, and, because solubility of molybde- heat pipe wall and the microatructure of the fail- num is higher in lithium, preferred attack in these 25 ure area. On the basis of these examinations, they regions will occur. The I esults cited by DeMastry, concluded that the corrosion problem was directly however, do not agree with such an explanation related to the initial oxygen content of the wall because molybdenum alloys were resistant to grain material (i.e., oxygen was transported by the boundary attack by lithium after 1000 h at 1370 to refluxing lithium from the condenser wall to the 1650°C. Tungsten alloys showed evidence of surface heating zone where local increases in oxygen con- dissolution and grain boundary attack at 1540 and tent in the liquid lithium could result from the 1650°C. Further compatibility studies of these evaporation process). He postulated that attack of materials in lithium are required for more com- the wall then occurred through the dissolution of plete understanding of the corrosion mechanisms niobium or tantalum by lithium containing high involved. oxygen levels. DiStefano and DeVan32 have ques- The mechanical strength of the W-25% Re-30% tioned the feasibility of the latter step and have Mo alloy was unaffected by the test exposure, but argued that attack of the wall can occur only by ductility of both hot and cold end samples was direct chemical reaction with lithium oxide accom- greater after the test. panying dry out of liquid lithium in local areas of the evaporator. Nevertheless, the observations and Lithium Heat Pipes oxygen control procedures of Busse and his coworkers at Ispra provide the most extensive and Because the boiling point of lithium is high useful background studies to date for the operation (1342°O) relative to space power, most of the of refractory metal-alkali metal heat pipe systems. extant corrosion data for lithium relate to convec- tive rather than evaporative heat transport. How- The Heat Pipe Design Handbook,™ compiled by ever, there has been a long-standing interest in Dynatherm Corp., lists four other heat pipe tests lithium as a two-phase working fluid for high tem- with lithium in the 1500 to 1600°C temperature perature heat pipes, and performance testing of range conducted prior to 1972. Four lithium heat refractory metal wick and containment materials pipe tests with Mo-TZM tubes were operated by for this application has been conducted by at least RCA Electronics Components Division. Three were four different investigators. Although there have terminated after approximately 10,000 h by weld been no parametric corrosion studies of refractory failures, and a fourth developed an evaporator leak metal-lithium heat pipes to date, metallurgical after 4600 h. examinations were performed following certain Lithium heat pipes operating at 1000 to 1350°C long-term (up to 10,000 h) performance tests. have shown fewer corrosion problems than those Prior to 1972 the most extensive testing of operating at 1500 to 1600°C. Ranken and lithium-filled heat pipes was reported by the Eura- Kemme's33 examination of an Nb-l%Zr pipe with a totn Ispra Establishment.28 Eusse and screen wick revea'ed no visual evidence of corro- coworkers29'30 were the first to document corrosion sion after 4300 h at 1100°C; also, an Nb-l%Zr/Li problems in Nb-l%Zr and tantalum heat pipes heat pipe operated at 1350°C by RCA completed operating at 1500 tc 1600°C. In initial tests of these 2300 h without evidence of macroscopic corrosion.28 materials, the evaporator wall was perforated in Finally, Ispra workers obsarved33 that the same times less than 10 h unless procedures were Nb-l%Zr pipe design that was rapidly perforated employed to deal with the mass transport of oxy- by lithium at 1600°C (discussed above) operated gen from the condenser region to the evaporator without failure for 3570 h at 1000°C. However, region. Effective procedures included deoxidation examination of the 1000°C pipe did reveal dissolu- DEVAN. DISTEFANO. AND HOFFMAN

..ion and corrosion product buildup in the evaporator zone within grooves and capillaries used for wicking. m Since 1972 published data on lithium heat pipes Q _ 2 E have narrowed to molybdenum and TZM containment systems with operating conditions representative of a UO2-fueled heat pipe reactor. These tests, conducted by Thermacore, Inc.3* and Los Alamos National Laboratory,35 have investigated the hydraulic and heat transfer performance of various O wick geometries and pipe configurations. Only lim- O ited corrosion information has been extracted to 10 ? 5 10 ? 10- ? '- '0 date from these tests, although operating times in CONCENTRATION Or OXYGEN IN SODIUM (ppm) excess of 25,000 h at 1230 to 1430°C have been reached without significant changes in heat trans- Fig. 9 The effect of oxygen on the solubility of niobium in fer performance. sodium at 600°C.! Sodium and Potassium In contrast with the results from niobium tests, The compatibility of refractory metals with tantalum exposed to sodium with increasing oxygen levels gained weight, picked up oxygen, and sodium and potassium is characterized by the same 38 general type of reactions as occur with developed progressively blacker scales. Further- lithium—dissolution, mass transfer, and impu- more, the amount of tantalum in sodium did not rities—but there are some important differences. show as muc^ dependency on oxygen in sodium (up Unlike lithium, oxygen in sodium and potassium to 2000 ppm) as had been noted with niobium. increases corrosion Several studies have shown However, at higher levels (to 12,000 ppm) weight that the effect of oxygen on the apparent solubili- losses were noted, and tantalum in sodium ties of refractory metals is critical. Another impor- increased markedly. tant consideration results from the relatively Results from studies at 600°C of tantalum higher vapor pressures of sodium and potassium exposed to potassium containing increasing oxygen compared with lithium. At temperatures where levels were similar to those for niobium, and it was refractory metal containment of sodium and potas- concluded that the apparent solubility of tantalum increases with the concentration of oxygen in sium is required, Rankine cycle systems utilize 39 two-phase boiling liquids. The condensing vapor in potassium. In contrast with niobium, oxygen such a system is extremely pure, and rate of disso- redistribution is the result of the formation of a lution can be a principal factor in corrosion and ternary oxide which either flakes off during expo- mass transfer. sure or is dissolved when the specimen is cleaned after the test. Oak Ridge National Laboratory (ORNL) In tests where oxygen in the niobium or tan- The compatibility of niobium and tantalum talum exceeded a threshold level, sodium and potassium penetrated the refractory l.ietals37"39 as .vith liquid sodium and potassium has been 16 18 reported in a number of studies conducted at was noted earlier with lithium. " As the oxygen ORNL. Klueh36 determined niobium solubility at concentration in the refractory metals is increased, 600°C as a function of oxygen concentration in the depth and amount of attack are thought to pro- sodium (Fig. 9). Similar results for the ceed by formation of a ternary <™ide that forms niobium-oxygen-potassium system at 400 and along grain boundaries or certain crystallographic 600°C were also reported by Litman.37 The increase planes. in solubility noted was proposed to be the result of Dissimilar-metal interactions between niobium interaction between solutes (oxygen and niobium) or Nb-l%Zr and type 316 or 318 stainless steel in in liquid sodium or potassium solution rather than sodium or NaK were investigated by ORNL in the formation of a ternary-oxide phase as had been static capsule tests40 at temperatures from 816 to previously proposed. The effect was characterized 982°C. The principal interaction was the transfer by use of thermodynamic interaction parameters of carbon and nitrogen from the stainless steel to that empirically describe the effect of oxygen in the niobium or niobium alloy. Complex thp alkali metal on the activity coefficient of carbide-nitride layers formed on the surface of the niobium in the alkali metal. niobium. Subsequent analysis revealed that only COMPATBSJTY OF REFRACTORY ALLOYS

, THERMOCOUPLE FEEDTHROUGH

TO VACUUM (SEALED AFTER BAKE-OUT) A

GET ;t R 'ON

CONTAINER WALL NSERT • £ SPECIMENS

REFRACTORY METAL CAPSULE

:nLr-.VOCO!JPLE Ff ' DiHROUGt

Fig 10 Schematic of ORNI. refluxing capsule test Bystem."

nitrogen had diffused into the niohium. The sub- mined u.\. a known condensing rate; (2) natural- surface nitrogen pickup greatly increased tensile circulation loop tests (Fig. 11), which provided dis- strength and decreased ductility. The amount of solution and deposition rates in a condenser and carbon and nitrogen transfer increased with tem- subcooler, respectively, and (31 forced circulation perature, exposure time, and amount of stairless loop tests (Fig. 12), designed tn investigate mass steel present. transfer and corrosion effects with ihe condenser Compatibility studies of refractory metals in at a much lower temperature than the boiler, and boiling potassium were carried out by ORNL in the corrosion-erosion resistance of potential nozzle three types of facilities: (1) refluxing-capsi;1^ tests and turhine-biadt materials. Alloys examined (Fig. 10), in which dissolution rates were deter- included Nb-1 *Zr, D-43. FS-85, C-129 Y, T-lll, so DEVAN, DtSTEFANO, AND HOFFMAN

COILED To WIRE HEATERS ON VAPOR CARRY-OVER LINE

FOUR LAYERS Or To RADIATION FOIL

O97b-in. OD x 50 mil WALL CONTAINING 16 INSERTS, EACH I in. LONG I 30 in.

TYPICAL INSERT SECTION

- COILED Tr WIRE PRE-HEATF.R

Fig. 11 Schematic of ORNL natural circulation, Nb-l%Zr boiling loop." (Note: all dimensions are given in inches.)

TZM, W and W-26%Re. Results have been Based on their resistance to relatively pure reported41"44 and are summarized in Table 5. (One potassium, the alloys tested in refluxing capsule refluxing capsule test was conducted in boiling tests fell into two general categories. One class sodium and data from this test are also included.) consisted of niobium, tantalum, or molybdenum All tests were conducted under vacuum at 10~7 to with minor additions of zirconium, hafnium, or 10~9 torr, and initial oxygen content of purified titanium. These alloys showed negligible surface potassium ranged from 20 to 50 ppm (amalgama- attack but did undergo changes in interstitial tion method of analysis) or 70 to 200 ppm impurity concentrations. The expectation of (gettering-vacuum-fusion method of analysis). weight losses in the condensing region, reflecting COM>A1BUTY OF REFRACTORY ALLOYS 51

TABLES Summary of Boiling Sodium or Potassium Metal-Structural MeUlJ Compatibility Testa Completed by ORNL

Temperature, °C Number Time Condennlng weight of Maxi- Mini- per rate. change, 1 tests Material Fluid mum mum test, h g/min '/cm' rag/cm Results

Refiuzing Capsules (55,000 h)* Nb K 1200 5000 0.3 to 0.4 t Penetration of niobium by potassium. Nb-l%Zr K 1100 to 600 to 0.2 to 0.4 1.1 to -2.0 No evidence of penetration 1200 5000 by potassium. Oxygen pickup and migration from top to bottom of capsule. Nb-l%Zr Na 1250 5000 0.17 +0.7 to +0.1 Mo evidence of penetration. D-43 K 1200 *o 950 to 0.3 to 0.6 + 1.7 to -0.002 No evidence of penetration. 1400 5000 Oxygen transferred from top to bottom of capsule. Decarburization occurred in areas of high oxygen pickup. 1 C-129 Y K 1200 5000 0.3 +0.5 to 0 No attack except to 0.03 mm on one specimen. 2 TZM K 1200 to 5000 0.3 + 7 to -2 Increase in concentration of Ti 1300 and Mo in K but no other evidence of attack. 1 T-lli K 1200 5000 0.3 +0 to -1.7 No evidence of penetration. Some oxygen pickjp by T-lll. 2 W-26%Re K 1200 to 5000 0.3 No attack of powder product 1250 capsule; mass transfer deposits of tungsten in arc melted capsule test

Natural Circulation Loops (13,000 h)* Nb-l%Zr 1200 650 1200 to 0.2 +0.7 to -0.2 No evidence of penetration. 2800 Oxygen pickup in specimens from subcooler region. D-43 1200 835 3000 C 35 +0.6 to +0.2 No evidence of penetration. Oxygen pickup in specimens from subcooler region. T-lll K 1250 1040 3000 0.53 -0.4 to +0.7 D-43 Loop/ K 1250 1040 3000 0.53 +0.3 (Av) No evidence of corrosion. TZM Inserts ZrO2 transferred to subcooler. TZM inserts decarburized.

Forced-Circulation Loops (3000 h)' 1100 355 3000 270 0.03 mm erosion, second stage blade. Oxygen ; jkup in boiler and condenser.

*Total test hours. tNot determined. 62 DEVAN. DBTB=ANO. AND HOFFMAN

DRYER

Nb-1%Zr ALLOY TO TYPE 316 STAINLESS STEEL JUNCTION

I ZIRCONIUM MHOT TRAP

AIR

FARADAY PUMP

Nb-l%Zr ALLOY TO TYPE 3(6 STAINLESS STEEL JUNCTION WATER-COOLED ELECTROMAGNETIC OIL TRAP PUMP

12-in VACUUM BLC ZK VALVE

• MECHANICAL VACUUM PUMP -VACUUM CHAMBER

4-in DIFFUSION PUMP

TRAP

Fig. 12 Schematic of OSNL boiling potassium, forced circulation loop test system.14 (Note: all dimensions are given in inches.)

dissolution of the alloy by the condensing vapor, and molybdenum were found in higher than initial vas not realized. Both losses and gains were concentrations in the potassium after the tests. observed, depending on specimen location in the Unalloyed niobium and W-26% Re comprised a condenser, but the total gains outweighed the total second class of materials that proved less resistant losses. In those tests with niobium-base ahoys, to attack. Niobium was heavily attacked at 1200° C changes in oxygen content were of the same magni- by potassiuiu initially containing <30 ppm oxygen tude as the weight changes, so the weight changes and virtually disintegrated in potassium to which appear to be symptomatic of migration of oxygen 300 ppm oxygen had been added. These results rather than metal dissolution. Weight changes in contrast with the behavior of unalloyed niobium in a TZM capsule (+7 to -2 mg/cmz) were larger static (nonfluxing) potassium.36-37 In all liquid than those found for niobium-base alloys and were systems, attack resembling that produced under more indicative of dissolution; there; was very little iefluxing conditions was observed only when the evidence of interstitial migration. Both titanium before-test niobium contained >500 ppm oxygen COHPATBtlTY OF REFRACTORY ALLOYS 53

temperatures from 780 to 1015° C was reported by and occurred irrespective of the oxygen level of the 46 potassium. It was concluded that attack in niobium LeRC. The data are described as being apparent refluxing capsules is the result of oxygen entering solubility because results were affected by oxygen the boiler wall.41 Furthermore, transfer of oxygen in niobium and tantalum. From Lhis study to the niobium from potassium appears to be solubilities are described by the relation peculiar to a boiling system, because under static conditions the opposite was observed. log wt ppm Ta = 4.75 - 3048/T ) In two otherwise similar W-26% Re refluxing fnr T = 1055 to 1287K capsule tests for 5000 h at 1250° C, no measurable log wt ppm Nb = 5.23 - 3739/TJ corrosion was noted in a powder-metaliurgically produced capsule while a ring of tungsten crystal- An extensive test program to evaluate the line deposits formed in the condenser section of an resistance of refractory metals to refluxing potassium at 980 to 1315°C was also carried ^ut at arc-melted capsule. The arc-melted W-26% Re cap 47 sule originally contained about th -ee times more LeRC. Refluxing capsule tests were carried out on oxygen (—60 ppm) than the deposit-free powder the following niobium- and tantalum-base alloyF. product capsule (—20 ppm oxygen). B-33 (Nb-5%V) Oak Ridge National Laboratory also conducted SCb-291 (Nb-10%Ta-S%W) a series of refhixing capsule tests with Nb-l%Zr at Nb-1 Zr (Nb-l%Zr) 1200°C to compare the relative corrosiveness of 41 42 As-55 (Nb-5% W-17%Zr-0.3% Y-0.07%C) sodium, potassium, rubidium, and cesium. - Abso- B-6P 'Nb-5%V-5%Mo-l%Zr) lute weight changes in Nb-l%Zr were small in all D-4S (Nb-10%W-l%Zr-0.1%C) tests, and there was nc evidence of chemical inter- FS-85 (Nb-25%Ta-10%W-l%Zr) action other than migration of oxygen. The results Cb-752 (Nb-10%W-3%Hf) indicated little difference in the relative corrossive- D-14 (Nb-5% Zr-0.02%C) ness of the alkali metals at 1200°C. 43 C-129Y (Nb-10%W-10%Hf) Natural-circulation loop results correlated in Ta-10%W (Ta-10%W) all respects with the effects observed in T-lll (Ta-8%W-2%Hf) corresponding refluxing-capsule tests; i.e., weight T-222 (Ta-10.4%W-2.4%Hf-0.01%C) changes, when corrected for the migration of interstitial elements, showed negligible dissolution of Results are summarized in Table 6. The most the condenser walls, and there was no evidence of noticeable trend was that the ungettered alloys mass transfer of metallic elements to the sub- (those containing no strong oxide former such as cooier. 44 Hf or Zr) were less resistant to corrosion by One Nb-l%Zr forced-circulation loop with refluxing potassium than were the gettered alloys. Nb-l%Zr nozzle and turbine-blade specimens in Each alloy was placed in one of three groups, in the test section has been operated under conditions order of decreasing corrosion resistance. shown in Table 5. Maximum attack was 0.03 mm Group 1: T-lll, T-222, C-129, D-14, Cb-752. where a 915-m/sec stream of 86% -quality vapor D-43, B-66, AS-55, FS-85 impinged on a simulated second-stage turbine Group 2: Nb-l%Zr blade. No evidence of attack or deposits wa3 found Group 3: B-33, SCb-291, Ta-10% W in the boiler, condenser, or subcooler, although chemical analyses of the loop wall showed a slight The corrosion resistance of Group 1 was judged increase in interstitial oxygen concentration. sufficient for consideration in space power systems In the lat» 1960s, ORNL completed an analytical as containment materials. Nb-l%Zr in Group 2 comparison of cesium and potassium as working was considered marginal because one heat of mate- fluids for NASA. One phase of this work was a rial was attacked and another was not. The review of turbine biaue erosion in operating tur- Group 3 (ungettered) alloys were judged unsatis- bines and a review of analytical and experimental factory because of the large amount of attack investigations of erosion and related phenomena. noted. 5 Results of this work have been reported* , and a Brookhaven National Laboratory (BNL) series of tables and curves summarizes the data obtained. Sodium refluxing capsule studies were conducted with seventeen refractory metal alloys as NASA-Lewis Research Center (LeRC) shown in Table 7 for Jmes up to 10,000 h and tem- A study of the solubility of niobium and tan- peratures to 1320°C.48 Comparison of several talum in potassium containing <15 ppm oxygen at materials tesied at 1315°C in boiling sodium is 54 DEVAN, DBTEFANO, AND HOFFMAN

TABLE 6 Summary of LeRC Refluxing Capsule Currosion Test Results"

Corrosive attack Tempera- Time, ture, °C h Alloy Run Type Depth, M"a Location Comments

980 110 Ta-10 W 6 Intergranular 300 Weld 1000 3Cb-291 6 Intergranular 38 Weld 2000 B-33 1 None 2000 Nb-1 Zr 1 Intergranular 175 Liquid-vapor interface Intergranular 20 Liquid section General attack 4000 T-222 7 None 1200 8 Ta-10 W 6 Intergranular 400 Weld 123 Nb-1 Zr 1 Solution Condensing General attack section Intergranular 113 Liquid-vapor 38 fim deposit interface 1000 SCb-291 6 Solution 50 Condensing section General attack Intergranular 400 Liquid section General section 1000 B-33 4 Solution 25 Condensing section Solution 75 Liquid-vapor Highly localized interface 2000 D-43 3 None Film at liquid-vapor int rface 4000 Cb-7E2 2 None Niobium and zirconium ioxide film at liquid-vapor interface D-14 2 None Niobium and zirconium dioxide film at liquid-vapor interface AS-55 2 None Heavy etching, about 5 iu~ B-66 2 Solution 5 Condensing section Niobium and rrconium dioxide film at liquid-vapor interface C-129 2 None Film at iiquid-vapor interface T-lll 2 None Film at liquid-vapor interface FS-85 7 None Niobium or tantalum film at liquid-vapor interface T-222 7 None K2Ta40n film at Iiquid-vapor interface 1250 380 B-33 5 Intergranular 25 High stress area Very slight solution attack at in condensing liquid-vapor interface section 2000 D-43 None Amber film over entire inner surface 2000 FS-85 None Film at liquid-vapor interface 2000 T-lll None Film at liquid-vapor interface 1 13 5 2000 FS-85 3 None Film at liquid-vapor interface 2000 T-lll 3 None Film at liquid-vapor interface 4000 T-222 7 None KjTa^n film at liquid-vapor interface COMPATBiUrY OF REFRACTORY ALLOYS 55

TABLE 7 Nominal Composition of Alloys Tested in BNL Test Program*

Alloying element, wt % Alloy Nb Ta Mo W V Zr Ti Hi* C Others designation

AS-55 Bal 5 1 0.06 0.2 Y B-66 Bal 5 5 1 Cb-753 Bal 5 1 Cb-Ti Bal 10 D-ll Bal 1 D-14 Bal 5 D-43 Bal 10 0.1 F-48 Bal 5 15 1 0.05 FS-60 Bal 10 FS-85 Bal 27 10 1 S-291 Bal 10 10 T-lll Bal 8 2 T-222 Bal 10 2.5 0.01 TZC Bal 0.35 1.4 1.15 TZM Bal 0.38 0.5 W-26 Re Bal 26 Re X-34 Bal 5 0.38 C.5 0.1

TABLE 8 the D-43 capsule because the D-43 gettered oxygen Corrosion Results" on High-Strength from the sodium and maintained it at a low level Refractory Metal Alloys Tested by BNL in throughout the test. Refluxing Sodium at 1315°C In order to determine if prestraining the refractory metai alloys would have any effect on Tested as insert in corrosior, several alloys wers tested in Nb-l%Zr Alloy Monometallic capsule, D-43 capsule for capsules with purified sodium for 2000 h at 1000°C, designation test time 6271 h 3COOL and results aio shown in Table 9. No effect of pre-

D-43 No corrosion No corrosion FS-85 0.05-mm G.B. attack in 0.05-mm G.B. liquid TABLE 9 weld, both liquid region Corrosion Results of Refractory Metal Alloy Prestrained vapor regions Inserts Tested by BNL in Sodium at 1000°C for 2000 h'8 FS-60 No corrosion T-Ul No corrosion 0.025-mm G.B. attack vapor region Alloy designation Results of metallographic examination TZM 0.01-mm G.B. attack ho corrosion of weld in vapor FS-85 0.02-mm grain boundary attack. No region cracking. X-34 No corrosion T-lll Slight transgranular rougl uning of TZM polished surfaces. No cracking. S-291 shown in Table 8. Initially, each material was AS-55 No corrosion detected on any of these evaluated as an insert in a D-43 capsule, and then B-66 inserts. No cracking. the tests were repeated in a monometallic capsule Cb-75? system. D-43 did not show any evidence of attack, Cb-Ti D-ll while the other alloys showed some slight evidence D-14 of corrosion as indicated. The authors noted that D-43 FS-60 (Ta-10%W) was not tested as a monometal- F-48 lic capsule because previous results49 at BNL had FS-60 T-222 shown severe corrosion of this alloy. Speculation W-26 Re was that FS-60 was not attacked when exposed in 56 DEVAN, DIST5FANO, ANO HOFFMAN

TABLE 10 Summary of Results at BNL on Relative Aggressiveness of Alkali Metals in Refractory Metal Alloys" Nb-l%Zr and D-43

Capsule material: Nb-l%Zr Capsule material: D-43 Alkali Temperature: 1150°C Temperature: 1235°C metal Time: 6000 h Time: 9437 h

Cs Slight attack at liquid vapor No corrosion interface Rb Slight attack at liquid vapor No corrosion interface K No corroaion No corrosion Na Slight attack at liquid vapor No corrosion interface Li Slight attack at liquid vapor No corrosion interface

straining was noted, and only the FS-85 alloy other the capsule was Nb-l%Zr. There was a sig- showed evidence of attack as had been noted previ- nificant weight loss in wrought and welded No and ously (Table 8). Nb-l%Zr whsn they were exposed to potassium at Comparison tests were run with Nb-l%Zr and 790 to 1090°C in Hastelloy X capsules. In addition, D-43 in Li, Nc, K, Rb, and Cs as shown in Table 10 severe intergranular ~)enetration was noted in the to measure the relative corrosiveness c* the alkali welded specimens ana X-ray diffraction tentatively metals. Although very little attack occurred, identified NbN and NbC as the corrosion product. Romano et al.,48 judged that Cs, Rb, Na, and Li This is a dissimilar metal effect due to testing were comparable, with K being the least aggres- these materials in a Hastelloy X capsule. When sive. tested in Nb-l%Zr capsules at 1090 to 1000°C, both BNL also operated several natural convection niobium- and molybdenum-base alloys were judged sodium boiling loops of Nb-l%Zr and D-43. No cor- highly compatible with even impure potassium. No rosion was reported in either of two Nb-l%Zr reaction layers were found, and weight changes loops49 under test conditions as follows. were small. A Haynes 25/type 321 stainless steel boiling potassium loop was operated50 for 50 h with the Temperature, Time, Mass flow rate, boiler at 980°C and the condenser at 730°C. Sam- °C h g/mia ples of Nb, Nb-l%Zr, Mo, ana Ta were suspended in the boiler, and Nb and Nb-l%Zr were suspended 1095 8007 10 at the exit of the nozzle (vapor velocity — Mach 1094 to 1150 1330 75 2.5). Results are shown in Table 11. Examination of the boiler specimens indicated the decreasing order of corrosion resistance to be Mo, Nb-l%Zr, Nb, and Four additional natural circulation loop tests were conducted by BNL, one each of Nb-l%Zr (1100°C maximum) and D-43 (1200°C maximum) with liq- TABLE 11 uid sodium and one each of Nb-I%Zr (1260°C max- Rocketdyne Low-Temperature Boiling imum) ar " D-43 (1350°C maximum) with boiling Potassiumi Loop Results50 sodium. However, the program wa" terminated while tho tests were in progress, ana no results Approximate Approximate Weight-change were obtained. Boiler tempera- exposure rate, samples ture, °C time, h mg/cinVday Rocketdyne Two types of capsule tests were conducted by Rocketdyne to evaluate refractory alloys in Nb 980 50 -77 50 Nb-l%Zr 980 50 -66 potassium. In one case the capsules were made of Mo 980 50 -7 a Ni-base alloy, usually Hastelloy X, and in the COMPATBUTY OF REFRACTORY ALLOYS 57

Ta. The Nb and Nb-l%Zr samples downstream experienced large oxygen losses as a result of expo- from the nozzle wer<* not found after tests indicat- sure to NaK. ing very severe corrosion/erosion had occurred. Sodium Thermal Convection Loop Test This test loop52 operated at a hot leg temperature of General Electric (GE-Cincinnati, Ohio) 1305°C and a cold leg temperature of 760°C at a sodium velocity of 0.1 m/sec. This loop was the Extensive sodium and potassium compatibility first of a series of three Nb-l%Zr sodium thermal testing was donr at General Electric (GE) under convection and pumped loop tests to qualify com- tht, sponsorship of NASA-Lewis. As a result, a ponents and materials for a 10,000-h T-lll Rankiii- review of GE's sodium and potassium compatibility System Test Loop. The primary purpose of this +est test results will be summarized in separate sec- was net only to determine the performance of the tions. 1^ (autoresistance) heater design but also to qual- The principal purpose of most of the sodium ify instrumentation procedures and insulation tech- loop testing to be described in this section of tho niques and to establish the vacuum requirements paper was to confirm the adequacy of the many necessary to prevent contamination of the external 2 test components, such as pumps, I !! heaters, pres- surfaces of the test loop. sure transducers, valves, flow meters, thermocou- No significant corrosion of the Nb-l%Zr by the ples, etc., prior to the use of these proven com- high temperature sodium was observed. Oxygen ponents and procedures in the 10,000-h T-lll Ran- migration from the hotter to the cooler regions of kine System Corrosion Test Loop which is covered the loop was noted. Due to the very high tempera- in the lithium and potassium sections of this ture in the loop heater coil, the Nb-l%Zr tube paper. Nb-l%Zr and sodium were selected as the experienced critical strain growth resulting in one prototypic container material and coolant, respec- or two grains across the entire 1.7-mm tube wall. tively, because of the advanced state of knowledge Another significant observation was the large of these materials at the time the GE program was hydrogen pickup and associated embrittlement of Initiated in 1963 under NASA contract NAS 3-2547. the Nb-l%Zr tubing during the alcohol stripping The only capsule tests conducted at GE on refrac- procedure used to remove residual sodium from the tory metals were sodium-potassium alloy (NaK) ioop tubing. The hydrogen embrittlement was read- tests in support of the SNAP-8 program, and these ily eliminated by vacuum annealing of the tubing. tests are described below. Alkali metal removal procedures were subsequently developed by GE and ORNL which prevented Sodium Capsule Tests. Earlier work at Oak hydrogen pickup during alkali metal removal Ridge National Laboratory (ORNL) had indicated operations. that unalloyed tantalum contaminated with oxygen was subject to rapid attack by NaK at tempera- Sodium Forced Convectitnt, Loop Test The Nb- tures of 260 to 315°C. The SNAP-8 reactor utilized l%Zr Pumped Sodium Test Loop^ operated with a NaK as the reactor coolant to boil mercury in the 1130°C hot leg and 1075°C cold leg and L .low rate secondary Rankine loop and because of the corro- of 188 kg/hr. The loop was built and operated to siveness of mercury to the nonrefractory alloys of verify the performance of additional loop com- the reference design, tantalum was proposed as the ponents beyond those evaluated in the Nb-l%Zr material of construction for NaK-heated mercury thermal convection loop experiment described boiler. For these reasons Harrison51 conducted a above. The major additional components incor- series of tests on tantalum specimens contaminated porated into this system, which is illustrated in to precise oxygen levels and subsequently welded in Fig. 13, included a helical induction electromag- helium environmeiits with varying air contamina- netic SEM) pump, metering and isolation valves, tion levels. Other uncontaminated specimens were pressure transducers (%o types), and a flowmeter. also welded in helium containing varying amounts The test was operated for 2650 h. After the test of air. Harrison's tests used 22%Na-78%K a3 a the loop was cleaned, and the various components coolant and were operated at 650 and 730°C for were evaluated. Some loss of oxygen from the hot- both 100- and 1000-h tests. test regions of the loop by the sodium was The results of these experiments showed that detected, and a modest increase (159 ppm before tantalum specimens with oxygen concentrations of test to 404 ppm after test) in the oxygen concentra- 270 ppm or greater were attacked by NaK at 650°C tion of the uninsulated tube wall was noted despite and 730° C. Uncontaminated tantalum specimens the fact that the chamber was operated at 1 X welded in helium containing up to 250 ppm air 10"8 torr for 2200 h of the 2650 h. It is assumed were not attacked by NaK. All tantalum specimens that the contamination occurred during the first 68 DEVAN, DBTEFANO, AND HOFFMAN

Vacuum Tatilt - Loo-1 Pressure TransducersI Tempo ratures

T/C Resistance No. °F ' Hetted Colls 1 1975 2 585 3 800 4 580 5 1965 6 1950 7 1985 8 2020 9 2 000 10 2035 11 2065 12 2065 13 2050 l4 2045 15 650 - Valve BypaBfi IS 2030 17 2015 18 670 19 Surge Tank 1985 20 1880 21 18s0 22 505 23 475 24 •155 25 1555 26 1575 27 1720 28 1580

Fig. 13 Temperatures of the various components of the Nb-l%Zr pumped sodium loop during the 2650-h test.

400 h when the chamber pressure was in the 2 X 111 in the T-lll Rankine System Corrosion Test 10~8 to 2 X 10~7 torr pressure range. Detailed Loop wliich is discussed in the lithium and potas- metallographic e~

PKES5LJKL IRANSliuCt SLACK PIAPHhAGM

VACUUM UNk

TUR3INF SIMULA IOr. 1 N - • i.-t

SECONT'ARi [M Pl:W

SODIUM POTASSIUM ARGON ELFCTRICAL PCWER [ FLEX I B LE DRIVE CABLE

Fig. 14 Isometric drawing of the Nb-l%Zr RanJrine system eorronon test loop and vacuum test chamber. potassium was vaporized. Results obtained on and 1095° C on tube tensile specimens machined materials in the potassium circuit will be covered from the boiler inlet tube (oxygen concentration, 36 later in this paper. ppm), and the boiler outlet tube (oxygen concentra- The temperatures maintained in the sodium tion, 730 ppm) showed essentially no change from portions of the loop during the 5000-h test are the before-test material. Deoxidation of the Nb- illustrated in Fig. 15. All portions of the sodium l%Zr tubing from the hotter regions of the loop circuit in contact with flowing sodium were made was noted while oxygen increases were observed in of Nb-l%Zr. The pressure in the vacuum chamber the cooler regions of the sodium circuit. The most was maintained at 2 X 10~8 torr during the 5000-h significant oxygen increases were measured in the test to prevent oxygen pickup by the Nb-l%Zr loop regions of the boiler where the heat flux to boiling components. potassium section was very high resulting in a Extensive metallographic evaluation of the ftb- rapid drop in the sodium temperature. Figure 16 l%Zr specimens taken from all portions of the shows the tube-in-tube boiler and indicates the sodium circuit revealed no measurable corrosion. oxygen concentrations of the Nb-l%Zr following Tensile tests were performed at room temperature the test. The boiler plug detail (inset) in this figure 60 DEVAN, D1STHFANO. AND HOFFMAN

1780

Turbine • Turbine Simulator SimulalOf

Test Chamber Pressure D ..mg 5000 Hour Tesf 2-5 x 10's Torr

— Condenser

(Temperature in °F)

K Pump Surye T.inks

Fig. 15 Nb-l%Zr Rankine system corrosion test looping Bhowing temperatures (°F) and flow rates of sodium and potassium during 5000-h test. COMPATIBILITY OF REFRACTORY ALLOYS 61

' TURBINE __ Sll-'JLATOR (STAG," NO. 1)

12130 F / _ HEATER

21]0°F WALL (?n'jO°KK 148 Na

TUBF-l.'l-VJBE WALL (1990°F;, 1270 — - -BOILIR OUTEK 0.020" - L1'5 MIODI E IJ.020V - 533 "JMR 0.0/0" - '.76

BOILED PLUG DETAIL

PLUG SECTION OF BOILER — WALL_- 7^0 OUTER""0.020" - 330 MIDDLE 0.020" - 455 INNER 0.020" - 1460 NUMBERS INDICATE OXYGEN 3/8" CONCENTRATIONS IN PPM - Na TO EH PUMP BEFORE TEST ANALYSIS

177O°F WALL -J62 3/8" TU3E - 260 PPM 1" TUBE - 245 PPM

K FhOM I PREHLATER

Fig. 16 Oxygen concentration of various regions of the boiler of the Nb-l%Zr Rankine system corrosion test loop following 5000-h operation. and the oxygen concentrations of various sections inventory of refractory alloy/alkali metal boiling of the tube wall separating the sodium from the systems at as low a 1 el as is practical. boiling potassium show the large oxygen pickup The potassium compatibility testing done at (2445 ppm) on the sodium (outer) side of the boiler General Electric (GE) was spor.sored by NASA- tube wall. Figure 17 consists of two photographs of LeRC in support of Rankine system technology the straight region of the boiler. The very sharp development. The T-lll Ra"kine System Test Loop transition in appearance from the deposition that will be discussed in the last part of this sec- region to the clean tube surface results from the tion was the culmination of an eight-year program sharp transition from wet to dry potassium on the to document ^he compatibility of selected Rankine other side of the tube wall and the associated System materials for application in space power change in heat flux through the wall in this loca- plants. Long-term refluxing potassium capsule tion. Figure 18 shows metallographic cross sections tests conducted during this period are d:scussed of the Nb-l%Zr tube wall in the region oi depori- below tion shown in Fig. 17. The X-ray image, logether with other supporting analysis, strongly suggests Potassium Capsule Tests. An extensive bearing that the crystals are ZrO. These very thin deposits materials evaluation program was carried out at are of no engineering significance, but they illus- GE55 to identify materials suitable for use in trate the importance of keeping the overall oxygen potassium lubricated journal bearings over the DEVAN, DISTEFANO. AND HOFFMAN

i POTAS :TV , t us F fiM

Overall view of sodium side of ?r boiler tube and enlarged how the very thin deposits of impound on the tube surface.

NL-lZr

temperature range from 205°C to 870°C. Test base material (Grade 7178) lost carbon resulting in materials included Nb-l%Zr, Mo-TZM, W, the formation of elemental potassium on the sur- TiC + 5%W, TiC + 10% Mo, TiC + 10% Nb, face. The conclusion of the study was that the Grade 7178 (84%W-8%Mo-2%Nb-6%C). In addi- refractory metal bonded carbides TiC + 10%Nb tion to potassium corrosion evaluation, all the and Grade 7178 offered promise for 650°C bearing additional studies performed on candidate bearing applications. materials included friction and wear in potassium, The Mo-TZM alloy was a candidate turbine dimensional stability, hot hardness, thermal expan- blade material for application in Rankine space sion, ccmpressive strength, modulus of elasticity, power systems in which the turbine blades would and potassium wetting characteristics. The corro- operate in the 1095°C temwrature range. There- sion test results on the candidate materials cited fore, refluxing capsule tests were conducted to above indicated that the Mo-TZM, W, determine the corrosion resistance of the Mo-TZM TiC + 5%W, TiC + 10%Mo, and TiC + 10%Nb alloy in a refluxing potassium environment.56 The had excellent corrosion resistance but that the WC test facility, which was located in an ultrahigh vac- COMPATfBOJTY OF REFRACTORY ALLOYS 63

Fig. IK Cryptals of Zr-O found OD the sodium side of the potassium boiler tube of the Nb-I%Zr RanMne system loop.

uum chamber operated at 10" to 10 torr, is observed in these capsule tests was slightly higher illustrated in Fig. 19. As indicated in this figure, than would be predicted based on uniaxial creep the tubular insert specimens of the Mo-TZM alloy data. were located in the condensing potassium region of In addition, the test facility described above a Nb l%Zr capsrle. Both the 2500- and 5000-h cap- was used to assess the response of two biaxially sule tests operated at potassium condensing ratos stressed capsules in a refluxing potassium environ- 2 of l.n of the D-43 alloy ih'tijiw,} Lx>p 7V.sf.s. Both Mo-TZM and Mo-TZC when the stess is sufficiently 'arge to produce sub- turbine blades were utilized in the turbine of the stantial p counts of creep. The test facility and the Pnfnxsiiim Turbine Test Facility as shown in D-43 capsule design are illustrated in Fig. 20. Two Fig. 21. The test facility was constructed on type separate capsule tests were conducted.5" In the first 31

Thermistor

Optical Sight Port Stainless Steel H"oaL Sink

—TZM Allov Inserts

-Nb-llr Alloy Capsule

W-Re Thermocouple .,,

K Liquid-Vapor Interface Split Tantalum Strip Heater

— Tantalum Radiation Shields Tantalum Base Support |{{ Nb-lZr Alloy Spacer Wire

W-Re Thernocuuple

Fig. 19 Nb-l%Zr refluxing potassium capsule test system used ;o evaluate corrosion resistance of Mo-TZM alloy tube inserts.

Mo-TZM and Mo-TZC alloys have excellent simulator nozzles and blades made of the Rankine resistance to potassium liquid and vapor. system candidate turbine blade alloy, Mo-TZM. The A general description of the Nb-l%Zr Rankine location of the turbine simulator teat sections were System Corrosion Test Loop was given in the shown earlier in Figs. 14 and 15. Figure 23 shows sodium section covering the GE compatibility the appearance of the ten nozzle and blade combi- work. Figures 14 and 15 illustrate the general nations after the 5000-h test exposure. The nozzle features of this system and give the temperatures diameters and the heat rejection conditions were and flow rates in the boiling potassium circuit. The selected to achieve about 88% quality vapor im- results of the extensive evaluation performed on pinging on the nozzle blades at velocities in the 275 the Nb~l%Zr loop and the components located in to 410 m/sec range. The blade and nozzle speci- the potassium circuit are given in the report of mens shown in Fig. 23 were essentially identical in Hoffman and Holowach.60 A brief summary will be appearance to the before-test specimens. Cross sec- given of the most significant compatibility observations of several of the nozzles are shown in Fig. 24, tions in this 5000-h test. ant; thp metallographic appearance of the nozzles One of the major reasons for conducting this attests to the excellent resistance of the Mo-TZM test was to determine the corrosive/erosive effect alloy. No change in nozzle diameters occurred dur- of wet high velocity potassium vapor on turbine ing the test. All the blade specimens were exam- COMPATHUTY OF REFRACTORY ALLOYS 86

-Nb-lZr Alloy

W Caps

D-S3 Vapor Capsule Linear Variable I.VDT's Differential Transformers

Split Tan ta 1 uir Tantal um, Had i a t1 on Heater Liquid Shu-Ids \

Boil'ng '''ad ea tor

Fig. 20 Biasially stressed reflating potassium capsnle test facility/D-43 (Nb-base) alloy capsule.

ined at the impingement point, and no attack was (13 Mm) was noted in the high heat flux region of detected under the3e vigorous test conditions as the boiler tube. Extensive grain growth was noted illustrated in Fig. 25. in the hottest (1095cC) regions of the boiler tube Extensive metallographie examinations of all wall, but this was not related to the potassium (or Nb-l%Zi regions of the potassium circuit were sodium) environment. performed. No corrosion of either Nb-l%Zr base The results of this test confirmed the exoellent metal tubing or weldments was observed in any corrosion resistance of Nb-l%Zr to high tempera- portion of the loop. Very slight surface roughness ture sodium and the resistance of both Nb-1% Zr 6b

Fig. 21 Potassium turbine, housing, and exhaust scroll showing second stage (right side) where Mo-TZM and Mo-TZC blades were located.

F13. 22 Mo-TZM alloy turbine bladss in the second stage of potassium turbine test system following completion of a 3000 -h endurance test.

Test conditions at second stage: Effects of 3000-h test: Temperature, 138S°F Weight change, -0.029% K vapor velocity, 535 fps Maximum corrosion, < 1 mil K v&por velocity, 95% COMPATBUTY OF REFRACTORY ALLOYS 67

Fig. 23 Mo-TZM alloy turbine simulator nozzle and blade specimen* froe Nb-l%Zr Ranidne system corrosion loop following 5000-h exposure to high-velocity potassium vapor. Potassium Vk vor flow path can be seen on some of the blades.

CROSS SECTION AND PHOTOMICROGRAPHS OF NOZZLES FOLLOWING EXPOSURE TO POTASSIUM FOR 5000 HOURS IN THE Cb-lZr CORROSION TEST LOOP

Fig. 24 Cross-section viowa and photomicrographs of turbine simulator nozzles following exposure to potassium for 5000 h in the Nb-l%Zr corrosion test loop., 68 DEVAN, DtSTEFANO. AND HOFFMAN

Mo-TZM TURBINE BLADE SPECIMENS FOLLOWING 5000 HOURR?S EXnOSURE TO POTASSIUM IN THE Cb-lZr CORROSION TEST LOOP

.i?t';!"E'l AND HOLDING 'A BLADE SPECIMEN '10. 1, ^r.:' "I0Z7LE I'.'LE" TLHP. , 2rj " ' : :MPV-rME.,T vELnr;:.-, 11 :-=

flLADE SPECIMEN NO. 2, REGION® •''V '-L"::TI r^ RLADr SDrc;ur'l NOZZLE INLET TEMP., 1785OF !V?I'ir,EMENT VELOCITY, 1350 FT/SEC

Fig. 25 Mo-TZM turbine simulator blades following 5000-h exposure to high-velocity potassium in Nb-l%Zr Rankine system loop.

and Mo-TZM to both boiling and condensing potas- The potassium circuit of this loop included tur- sium. bine simulation test sections identical in design to The T-lll Rankine System Corrosion Test Loop,61 those illustrated earlier in Fig. 23. In this section which had a lithium primary heater circuit and a the potassium flow rate was 18 kg/h with a vapor boiling potassium secondary circuit (1175°C maxi- nozzle exit velocity of 305 m/sec at a vapor quality mum and 515°C minimum operating temperature) at the exit of 88%. However, in this test the alloys was previously illustrated in Figs. 7 and 8. This Mo-TZC and Nb-132 M (Nb-base) nozzle and blade test was the culmination of an eight-year NASA- specimens were substituted for the M

Vapor Carryover Line

• Total Wall + Inner 0.020" A Outer 0.020"

Oxygen Analysis Shown in ppra Pre-Test Analysis: 3/8-inch Tube - 16 .-Inch Tuba - 17

1101. 596

Fig. 26 Results of oxygen analyses of various T-Hl alloy components following the lG,CO0-h test Note the high oxygen concentrations on the potassfam side of the T-111 wall in the boiler plug section and the very low oxygen concentrations on the lithium side.

Studies have generally been conducted in In 1967 Klueh and Jansen66 at ORNL completed association with applications such as a working a literature survey on compatibility of structural fluid in heat engines or to dissipate the space material with cesium as part of the Analytical charge in thermionic generators. Many of the early Comparison of Cesium and Potassium on Rankine- results on cesium compatibility were affected by Cycle Working Fluids for Space-Power Plants. Few oxygen contamination, and most often impurity additional compatibility studies with cesium have concentrations were not determined. The same cor- been conducted since this report; therefore, it has rosion mechanisms that occur with the other alkali been used as the basis for the summary presented liquid metals also generally occur with cesium. here. COMPATIBILITY OF REFRACTORY ALLOYS 71

Chandler and Hoffman67 tested niobium, Nb- lamellar carbide or carbonitride precipitates after l%Zr, molybdenum, tantalum, and tungsten in 600 h at 1540 and 1700°C. The source of the carbon static capsules containing ce3ium liquid for 300 h and nitrogen was thought to be the cesium and the at 870°C and found no metallographic evidence of TZM container. The tungsten-base alloys tested attack. Similar results were found by Winslow68 for were unattacked after 1000 h at 1700°C, but they molybdenum and tungsten after 500 h at 400°C. experienced surface dissolution and grain boundary British workers,69 on the other hand, found a sur- attack after 1000 h at 1870°C. face scale on niobium and tarnishing of tantalum Tables 12 and 13 summarize the results of the and molybdenum after 10,000 h at 570°C. These refractory alloys tested in the cesium vapor.67"69'73'76 latter tests, however, were conducted in stainless More recently, a performance test of a thermionic steel capsules. converter with a tungsten emitter was conducted Tepper and Greer70'71 at Mine Safety Appliance by GA Technologies.77 The tungsten emitter was (MSA) tested the following dissimilar metal cou- exposed to cesium vapor at a pressure of 5 to 10 ples in cesium liquid. torr for nearly five years with no evidence of deg- : 1. Nb-l%Zr/"TD" Nickel (500 h at 980DC) radation in performance. ' 2. Nb-l%Zr/Haynes alloy 25 (500 h at 1370°C) Several investigators41'67'71'78'79 have conducted 3. Nb-l%Zr/Mo-0.5%Ti (500 h at 137O°C) refluxing capsule experiments with cesium. 4. Mo-0.5%Ti/"TD" Nickel (500 h at 980°C) Chandler and Hoffman07 found surface dissolution 5. Mo-0.5%Ti/Haynes alloy 25 (500 h at and severe attack of niobium and Nb-l%Zr after 1370T) 720 h at 980 and 1370° C. These results, however, 6. Mo-0.5%Ti/Zirconium (725 h at 1370°C) have been contradicted by more recent tests of niobium alloys.41'71'78'79 Inserts removed from the For couples (1) through (5), considerable mass condenser section of an Nb-l%Zr capsule -^sted by transfer of both metallic and nonmetallic constit- 41 DiStefano at OKNL for 5000 h at 1200°C showed uents was observed. No detectable changes were 8 only slight weight changes on tiie order of 10~ noted in the Mo-0.5%Ti/Zr couple after 725 h at 2 mg/cm . The only mass transfer noted was a 1370°C; but when oxygen and carbon additions change in oxygen concentration of the inserts. were made to the system, measurable changes were Analyses of the cesium following the test gave noted, especially in the zirconium which gettered niobium and zirconium concentrations of 6 and 1 both species. Conclusions drawn from this investi- ppm, respectively, indicating little dissolutive cor- gation were that the utilization of dissimilar metal rosion. systems can often lead to premature failure if one 78 79 is not cognizant of the effect of interalloying. Similarly, BNL and M3A found no attack of btudies on refractory metals with cesium vapor Nb-l%Zr after 6000 h at 1200°C and 886 h at 1150°C, respectively. Two other niobium-base have been of three types: (1) capsules containing 79 cesium with liquid and vapor in equilibrium, allo'jS that have been tested under refluxing con- (2) capsules containing only cesium vapor, and dition, D-43 (9000 h at 1200°C) and FS-85 (818 h (8) test specimens in stainless steel reservoirs con- at 1150°C)> were also unattacked. nected to an outside source of cesium vapor. In the Rocketdyne67 found surface and dissolutive first two methods, test specimens may have been attack on refluxing capsules of tantalum tested for suspended in the vapor, or the capsule itself may 720 h at 970 and 1370°C. Mine Safety have served as the specimen. Appliance,71-79 on the other hand, found no mass When the capsule and test specimen were of the transfer or attack of Ta-10%W after 528 h at same material, there was generally no metallic evi- 1170°C. Finally, severe dissolutive corrosion of dence of attack.67-72-73 DeMastry and Griesenauer,74 molybdenum and Mo-0.5%Ti (Refs. 71, 79) has been in what effectively were dissimilar metal tests, observed. Table 14 summarizes the refluxing cap- used TZM as a container material for various sule tests conducted to date. refractory metal test specimens and, in some There are three references to cesium loop tests instances, found evidence of corrosive attack and conducted for corrosion studies. Results have been mass transfer. B-66 experienced slight surface reported by Westinghouse Astronuclear Labo- dissolution after 1000 h at 1370°C, and T-lll, ratory80 (a two-phase natural circulation loop of which was resistant, to attack for 1000 h at 1370°C, T-lll), Brookhaven National Laboratory,78 and exhibited surface dissolution after 100 h at 1540°C. aerojet General Nucleonics.81 (The latter two were After 100 h at 1370°C, Ta-12%W was unattacked two-phase forced-circulation loop of Nb-l%Zr.) by cesium vapor, but, showed surface dissolution The natural-circulation (two-phase) T-11I loop after 1000 h and showed surface dissolution and conlpined three tubular inserts, two in the boiler 72 DB/AN, DISTEFANO, AND HOFFMAN

TABLE 12 Corrosion of Refractory Alloys in Cesium Vapor in TZM Containers74 (Static Capsule Tests)

Tempere- Time, Tempera- Time, Material ture, "C h Results Material ture °C h Results

B-66 1370 100 No attack W-0.9%Nb 1370 100 No attack 1000 Slight surface 1000 No attack dissolution 1540 100 No attack 1540 100 No attack 100 No attack 170U 100 No attack Ta-12%W 1370L 100 No attack 1000 No attack 1000 Surface dissolution 1870 100 No attack 1540 100 Carbide precipitate 1000 Surface dissolution 1700 100 Carbide precipitate W-15%Nb 1370 100 No attack Ta-8%W- 1370 100 No attack 1000 No attack 2%Hf (T-lll) VAti 100 No attack 1000 No attack 1000 No attack 1540 100 Slight surface 1700 100 No attack dissolution 1000 No attack TZM 1370 100 No attack 1870 100 No attack 1000 No attack 1000 Slight surface and grain 1540 100 No attack boundary dissolution 1000 No attack W-l0%Re 1370 100 No attack 1700 100 No attack 1000 Nc attack 1000 No attack 1540 100 No attack 1870 100 No attack 1000 No attack 1000 No attack 1700 100 No attack Mo-50%Re 1370 100 No attack 1000 No attack 1000 No attack 1870 100 No attack 1540 100 No atUck 1000 Grain boundary attack 1000 No attack and surface dissolution 1700 100 No attack W-25%Re 1370 100 No attack 1000 No attack 300 No attack 1870 100 No attack 1000 ' No attack 1000 No attack 1540 100 No attack Tungsten 1540 100 No attack 1000 No attack 1000 No attack 1700 100 No attack 1700 100 No attack 1000 No attack 1000 No attack 1870 100 No attack 1870 100 Slight surface 1000 Slight surface di 'solution dissolution and 1000 Grain boundary grain boundary penetration penetration

*J. A. DeMastry and N. M. Griesenauer, Investigation qf High-Temperature Refractory Metrte and Alloys for Thermionic Con- verters, APAPL-TR-65-29, Battelle Memorial Institute, April 1965.

and one located in the condenser section. The test Condenser area, cm2 56 was terminated due to a high-temperature Tubing sizes, mm 13 OD X 0.9 wall excursion that caused a failure in the boiler. Condi- Test duration, h at °C 246 at 1315°C and tions during operation are given below. 175 at 1090°C Temperature, °C (max) 1370 Posttest analyses indicated a slight increase (26 Temperature, °C (min) 1060 to 40 ppm) of oxygen at the boiler region and a sig- Vapor pressure, MPa 4.2 nificant increase (20 to 110 ppm) at the condenser Power input, kW 3 location. A decrease of oxygen in the cesirm was Flow, kg/h (max) 10 also noted. No evidence of corrosive attack was Vapor velocity, m/s 0.7 found in the loop. COMPATESJTY OF RffRACTORY ALLOYS 73

TABLE 13 Corrosion of Refractory Alloys in Cesium Vapor"

Temperature, Time, Material °C h Remarks Results Reference

Niobium 570 10,000 Stainless steel capsule Surf8.ce scale 69 870 300 No attack 73 870 720 No attack 67 1370 300 No attack 73 Nb-l%Zr 870 300 No attack 73 870 720 No attack 67 1370 300 No attack 73 D-43 J70 300 No attack 73 1370 3C0 No attack 73 FS-82 900 281 Stainless steel container Reaction zone at surface 76 PWC-33 870 300 No attack 73 1370 300 No attack 73 Tantalum 450,600 100 Stainless steel capsules No attack 75 570 10,000 Stainless steel capsules Slight surface tarnish 69 26u 430,650 48 Stainless steel container Surface tarniah 76 870 300 No attack 73 870 720 No attack 67 1370 300 No attack 73 900 281 Stainless steel container Internal pores and pre- 76 cipitate near surface Ta-in%W 870,1370 300 h Hack 73 T-11I 870, 1370 300 •>• ack 73 Molybdenum 260, 430, 650 48 Stainless steel container No attack 76 400 500 No attack 68 Molybdenum 450, 600 100 Stainless steel capsule No attack 75 570 10,000 Stainless steel capsule Slight surface tarnish 69 870 .300 No attack 73 870 720 No attack 67 900 281 Stainless steel container No attack 76 1370 300 No attack 73 Mo-0.5%Ti 870, 1370 300 No attack 73 Tungsten 260, 430, 650 48 Stainless steel container No attack r' j 400 500 No attack 68 870, 1370 300 No attack 73 W-3%Re 1370 300 No attack 73 W-5%Re 1370 300 No attack 73 W-25%Ke 1370 300 No attack 73

Operating conditions of the Brookhaven forced- The loop was constructed of Nb-l%Zr and con- circulation cesium loop,78 Fig. 27, are listed below. tained a nozzle-blade test assembly with Nb-l%Zr Test section material nozzles and TZM or TZC blade specimens. Nozrles Nb-l%Zr Posttest visual examination of the nozzle-blade Blades TZM, TZC test section revealed a vortexing of the vapor on Number of stages 2 the underside (opposite to the impingement side) of Temperature, °C the blades. Only a darkening of the high-velocity Boiling 960 vapor impingement region was detected on the Impingement (max) 840 TZM or TZC alloy blade specimens. Results of Vapor quality, % 80 metallographic examination of the blades indicated Vapor velocity, m/s (max) 245 no evidence of corrosion or loss of materials but, Mass flow rate, kg/h 40 rather, an adherent metallic layer (identified as Test duration, h 1100 niobium) 1 mil thick. 74 DEVAN, DISTEFANO, AND HOFFMAN

TABLE 14 Summary of Cesium Refluxing Capsules'*

Temper- ature, Time, Refer- Material °C h Results ence

Niobium 980 720 Surface dissolution 67 1370 720 Severe attack 67 Nb-l%Zr 980 720 Severe attack 67 1370 720 Surface dissolution 67 1200 5000 No attack 41 1205 6000 No attack 78 1150 886 No attack 79 D-43 1205 9000 No attack 78 FS-85 1150 818 No attack 79 Tantalum 980 720 Surface dissolution 67 1370 720 Severe attack 67 Ta-10%W 1150 528 No attack 71,79 Molybdenum 980 720 Severe attack 67 1370 720 Severe attack 67 Mo-0.5%Ti 1150 292 Surface dissolution 71,79 1000 Surface dissolution 71,79 1370 255 Severe attack 71,79

TURBINE SIMULATORS (NB-ZR NOZZLFS) SEO'ON »:TJM SLIDE; SECTION B'TZC BLADE

Fig. 27 Brookhaven forced-circulatk. .. loop operated with VAPOR 18 EXPANSION boiling cesium. TANK RAOiANT CONDLNSER COIL ( 695'C) AIR COOLED HEAT EXCHANGER AIR OUTLET

TRIM COOLER

PRESSURE TRANSMITTER

INLET (B7O-C)

HELICAL INDUCTION ELECTROMAGNETIC pUMp I ft« »/hr) COMPATOIUTY OF REFRACTOR* ALLOYS 76

The Nb-l%Zr forced-circulation loop containing at high temperature show that the total residual boiling cesium that was operated by Aerojet Gen- gas pressure must be held to below 1.3 X 10~5 Pa 7 eral Nucleonics81 was stable for 2014 h with the (1 X 10~ torr) to avoid serious deterioration of cesium boiling and condensing at 1010° C. The per- mechanical properties in periods as short as tinent operating conditions were months at 600°C and weeks at 1000°C. Contamina- tion rates are dependent, of course, on the residual Boiling temperature, °C 1010 impurities present, but it can be reasonably argued Condensing temperature, °C 1010 that the impurity ratios in these vacuum exposures Vapor quality, % 10 + 3 approximate those ratios that would be attained in Cesium flow rate, gpm (lb/h) 0.84 (693) a helium system purified by chemical gettering. Cold-leg temperature, °C 370 However, to achieve and measure an equivalent Power, kW pressure at 1.3 X 10~5 Pa of residual impurities in To preheater 6.3 helium is currently impossible. For example, in a To boiler 8.5 to 9.6 helium circuit at 5 MPa (50 atm) absolute pressure, Test duration, h 2014 a partial pressure of 1.3 X 10~5 Pa would represent a volumetric concentration of (1.3 X 5 6 Oxygen conLent of the cesium increased from 10" )/(5 X 10 ) or 0.0026 ppl. <10 ppm before the test to 57 ± 10 ppm at the Because of the obvious contamination problem, conclusion of the test (method of analysis was not very few experiments have actually been conducted specified). to evaluate the reaction rates of refractory metals The loop did not contain nozzle and companion in helium with controlled purity levels. Testing of refractory metals was conducted at Battelle blade specimens but was designed to provide high- 82 velocity vapor in various regions. Posttest exami- Northwest Laboratories in the mid 1960s using a nation revealed no corrosion of the loop piping and conventional alloy, recirculating helium loop with a no mass transfer effects. mass flow of 225 kg/h, fLw rate of 40 m/s, and pressure of 2.1 MPa (300 psi). Before each test run, the loop was charged with h!gh purity helium (<1 ppm total impurities), was evacuated, and was GAS COOLANTS charged again with high purity helium. Specimens of molybdenum, tungsten, tantalum, and niobium Compressible gases of interest to high- were exposed for a total of 504 h at 1150°C. Most temperature Brajton power cycles include hydro- specimens showed some tracer of iron and chro- gen and certain of the noble gases. Although com- mium on the surface, anil niobium and tantalum patible at high temperatures, hydrogen does not specimens showed a thin nitride case. The results appear to be a potential working fluid for closed were summarized by the Battelle workers as fol- Brayton refractory metal systems due to its rapid lows. diffusion rate in b.c.c. metals and reactions to form hydrides ai low temperatures. Of the noble gases, Niobium and tantalum readily pick up impuri- helium and helium/xenon mixtures appear best ties from (helium) loop atmospheres pure enough to suited to closed cycle Brayton operation. Since they permit the evaporation of superalloys, whils molyb- are chemically inert, these elements present no denum and tungsten remain comparatively unaf- inherent compatibility problems. In practice, how- fected. TI:e mechanical properties of Nb and Ta can be drastically altered by contaminant absorption ever, difficulties do arise because of impurity during dean run conditions (i.e., without purposeful atoms that enter the gas from the various system impurity additions). components Juat it contacts. At the detection limits reached by present-day analytical methods Migge83 addressed the relative thermodynamic (~-l ppb), the chemical activities of nitrogen and stability of oxide and carbide Dhases in equilibrium oxygen impurities in helium ar^ still far above with niobium or molybdenum in helium gas con- those needed to form the lower oxides of tantalum taining CO, CO2, H2, H2O, and CH4 impurities. Not or niobium in equilibrium with their respective surprisingly, he showed that, for virtually any gas metals. chemistry that has equilibrated with graphite, car- The difficulty of maintaining helium purity lev- bides (NbC and Mo2C) rather than the oxides are els commensurate with niobium or tantalum purity the stable phases above 400° C. (However, it should requirements is shown by the following simple cal- be noted that the equilibrium oxygen concentra- culation. As discussed in the next section, studies tions in solution in the Nb can be quite high if CU, 84 of the contamination of Nb-l%Zr in high vacuum CO2, or H2O is present.) Kohl examined the effect 76 DEVAN, D1STEFANO, AND HOFFMAN of helium pressure on the absorption rate of oxy- with the absorption of the same impurities by gen into pure vanadium. The result, which should refractory metal components contacting the be applicable to niobium and tantalum, was that at helium. Rather, the approach to the use of helium a constant oxygen partial pressure of 97.5 Pa and as a coolant must be to rid the circuit of accessible temperatures from 600 to 900°C, the oxidation rate impurities prior to their reaction with refractory (in static gas) decreased by a factor of 4 as the metals (i.e., vacuum degassing) and then to limit helium pressure increased from 8.7 X 10 ~3 MPa the ingress of impurities (from leaks and make-up (0.09 atm) to 8.7 X 10"2 MPa (0.9 atm). The oxida- gas) to acceptable levels. Such an approach will tion rates were constant at all helium pressures place severe limitations on the nature of materials and were directly proportional to oxygen pressure that can be tolerated in the circuit, particularly (i.e., adsorption controlled). thermal insulation and neutron moderators. Also, Japanese workers85-86 have more recently inves- it is doubtful that conventional alloys could be per- tigated the oxidation rates of TZM, molybdenum, mitted in the same circuit at temperatures where and Nb-l%Zr in commercially purified (>99.995%) diffusion processes could bring carbon and nitrogen helium and the same heHum containing 13 ppm to the alloy surfaces. oxygen. The oxygen, nitrogen, and water vapor concentrations in the former gas were 0.1, 5, and <1 ppm (by volume), respectively. Specimens were AMBIENT ENVIRONMENTAL exposed to the flowing gas at 0.12 MPa (1.15 atm) EFFECTS at temperatures from 700 to 1000°C. The Nb-l%Zr specimens exhibited Nb2N surface Iayer3 after The mechanical and corrosion properties of exposure in the commercial grade helium, but exhi- refractory meta s are strongly dependent on the bited NbO and NbO2 films in the helium with 13 concentration of interstitial impurities—carbon, ppm oxygen. The oxygen and nitrogen concentra- oxygen, nitrogen, and hydrogen. Since refractory tions of Nb-l%Zr after 250 h at 800 to 1000°C in metals are highly reactive toward carbon, oxygen, either test environment reached 1.0% and 0.1%, and nitrogen at elevated temperatures, the metals res-pectively. The oxidation rates were essentially must be maintained under ambient environments constant with time for the first 250 h at 700°C and that contain very low activities (or total amounts) 500 h at 1000°C. Molybdenum specimens showed of these elements. no measurable oxidation effects in cuuimercial Considerable analysis has been done to measure grade helium, but evaporation of molybdenum the uptake of interstitial impurities by niobium oxide occurred above 800°C in the helium contain- and tantalum alloys in dynamic vacua having varying 13 ppm oxygen. The oxygen content and ing activities (partial pressures) of oxygen-,87"96 microhardness of the TZM specimens ii:c-°ased in nitrogen-,97-98 and carbon-containing99 gases. These both environments. studios have provided data concerning the vacuum Other studies that have some relevance to the requirements and sfandards that must be applied compatibility of helium coolants with refractci-y to the heat treatment and terrestrial testing of metals are discussed in the ^ext section in the con- these refractory alloys. (The standards relating to text of protective atmospheres for refractory mechanical property testing are discussed in the metals. In the cases cited where Nb-l%Zr was paper on mechanical properties by Conway, this blanketed by helium or argon, every attempt was volume, and will not be covered here.) These find- made to continuously purify the blanket gases ings, coupled with operating experience, ultimately during use. However, the methods used proved led to the general acceptance of ultrahigh (<10~8 much less effective than what has conventionally torr) vacuum atmospheres achieved by ion pump- been achieved using ultrahigh vacuum systems, and ing "i a cold-wall chamber as the most effective the resultant contamination levels were much service environment for the extended testing of higher than acceptable for most testing purposes. refractory metal systems at elevated temperatures. Furthermore, as discussed above, the background Oxygen absorption rates in tantalum, niobium, pressures required for residual gases are impossi- and certain of their alloys at high temperatures ble to monitor except by vacuum technology, so and relatively low oxygen partial pressures (<10~3 there is no procedure for controlling the quality ot Pa) were reviewed by Inouye,100 who reported that the blanket gas. the average oxygen uptake by sheet specimens is It does not appear realistic that purification directly proportional to exposure time and oxygen methods for removing impurities from helium pressure and inversely proportional to specimen could be achieved on a scale that would compete thickness under pressure and temperature condi- COMPATIBILITY OF REFRACTORY ALLOYS 77 tions that do not produce visible surface reaction •^00 products. He concluded that the absorption rate TEST CONDITIONS HI t: NO 4 MA"Ifl"Al Nb Mr J was surface reaction limited and therefore was 400 DuTlf.T ION HV0 h- ' independent of oxyge^ mobility (diffusivity) in the PR! '..•".. ii;f ? ! > I)' loir i matrix. Inouye90 has also studied the contamina- o 0*YO? N I tion of niobium, molybdenum, tantalum, Nb-l%Zr, a C.'.WHON | O M I h\>C;t N

; FS-85, D-43, Cb-752, TZM, ?nd T-ill resulting from « H i ,i vi I:,I N their prolonged exposure to the residual gases in o Nt 1 C.-.'...t 'high vacua at pressures between 1.3 X 10~4 to 6 6 8 1.3 X 1(T Pa (10~ to 10" torr). The residual ^ too gases were mainly hydrogen, H2O, CO, CH4, nitrogen, and CO2. A comparison of the total chemistry change of Nb-l%Zr, niobium, tantalum, and molybdenum resulting from their exposure to the residual gases for 1000 h at 2.7 X 10"7 torr is made in T ZOO H'"' 900 l COO llnO l.'OO Fig. 28. Tp to about 600°C, contamination was neg- H V.'i H.M •.•'.•[ ("C ) ligible for all compositions. Above 600°C, the differences in the roauivity of these metals became ^ig. 29 Contamination of Nb-l%Zr by oxygen, carbon, apparent and were observed to decrease in the hydrogen, and nitrogen at 2.7 X 10 ' torr. order Nb-l%Zr, niobium, tantalum, and molybdenum. The contribution of the specific contaminants kinetic models, these authors concluded that the to the total contamination of Nb-l%Zr is shown in oxygen absorption rate and concentration gradient Fig. "29. In this test, oxygen contamination in T-lll were best described by a solution of Fief's accounted fci over 80^ of the total, while carbon second law using a constant oxygen flux into the contamination essentially accounted for the bal- specimen as the external boundary condition. Their ance. On the basis of these results, a background kinetic analysis also showed that, at 1000°C, the pressure below 1.3 X 10~u Pa (10~7 torr) can be diffusion constant for hafnium in the alloy is so seen to be a basic operating requirement for vac- low that this element is essentially immobile and uum atn ^spheres used to protect niobium and tan- that its only effect on oxygen transport is to talum alloys for extended periods. reduce the oxygen diffusion coefficient about an In smdies of T-lll at 1.3 X 1(T5 Fa oxygen order of magnitude below thai in pure tantalum. pressure, Carpenter and Liu92'101 showed that the From the Carpenter-Liu model it follows that, up rate of oxygen uptake by T-ill at 1L'OO°C remained to at least 1^00°C, contamination from a low- constant with time (i.e., that the average oxygen pressure o:;ygen source will produce a significant concentrator increased linearly with time;. After concentration gradient in T-lll, while pure tan- correlating experimental observations with various talum develops virtually no gradient if the temperature is above ~900°C. If the same model is applied to niobium and Nb-l%Zr and if the oxygen 500 diffusivities measured by Lauf102 are used, an cxy- gen gradient should not occur in niobium above —800°C but should be present in Nb-l%Zr up to a temperature of at least 1100°C. These predictions are in reasonable accord with the observed 300 responses of thpse refractory metals under service conditions and annealing treatments. ?00 An extensive contamination sturiv of tantalum alloys in vacua was performed by Harrison and Hoffman95 under conditions directly analogous to those used in refractory metal system and component testing. Their experiments were conducted in a large-diameter, cold-wall, ion-pumped vacuum chamber using a variable leak valve ant1 mass spec- -(00 500 600 700 800 900 1000 1100 1200 ti meter to control the oxygen inleakage to the TEMPERATURE ( °C ) chamber. After initial outgassing, the rate of oxygen inleakage was adjusted to hold the chamber Fig. 28 Contamination of Nb-l%Zr, niobium, tantalum, and 6 molybdenum Qt 2 X KT'torr. pressure at 1.3 X 10"" Pa (10~ torr) and resulted 78 DgVAN, DISTEFANO, AND HOFFMAN in an oxygen partial pressure that ranged from a controlled to either nondetectable or negligible lev- maximum of 1.1 X 10"3 Pa (8.3 X 10~6 torr) to a els. The operating procedures, vacuum histo „, and minimum of 5.7 X 10"5 Pa (4.4 X 10 "7 torr). Con- contaminant rates for these experiments have been taminrtion rates of 1-mm-thick T-lll strip exposed well documented107"110 and need only be referenced at 1316°C under these conditions averaged 47 in t' 3 discussion. The major problem in these g cm""2 sec"1 Pa"1. The contamination rate of exo' nents was to control the desorption of oxy- ASTAR-811C waa 56 g cm"2 sec"1 Pa"1 as com- Z1- n from heater surfaces, reflective insulation, and pared to a rate of 40 g cm"2 sec"1 Pa1 for pure vacuum chamber accessories so that the bulk of tantalum. Calculations of sticking factors showed this oxygen was picked up by the vacuum pumps that from 30 to 50% of oxygen molecules striking and not the refractory metal test components. In the specimens were absorbed. many cases the refractory metal components were A similar study103 using a cold wall, ion-pumped wrapped with tantalum, molybdenum, or titanium vacuum chamber was conducted by Lyon at GE foil to effect a mechanical barrier between the using Nb-l%Zr strip as the test material. Ho component and the chamber environment. The measured the sticking coefficient for oxygen on bakeout cycle and heat-up of the test system were controlled to limit the vacuum atmosphere to Nb-l%Zr at 927°C at an oxygen partial pressure of 4 2.6 X 10"4 Pa (8.57 X 10"6 torr) and found the <1.3 X 10~ Pa once the system had reached a coefficient to depend strongly on the prior anneal- temperature of 450°C. After the system operating ing treatment given the alloy. Following a 1200°C temperatures were attained, the steady-state ambient pressure in the vacuum chamber was usually vacuum aimeal, which removed any prior oxide 6 film on the alloy, the sticking coefficient was 0.6 to well below 1.3 X 10~ Pa, and hydrogen was the 0.7. Without the anneal, the coefficient was about dominant species in the residual gas. 0.16 to 0.18. The oxidation rates determined in One of the obvious contamination concerns Lyons study were 90 g cm"2 sec"1 Pa~~l and 23 2 1 1 relates to interstitial effects on mechanical proper- g cm" sec" Pa" for the annealed and uaan- ties. (See, for example, the paper by Conway, this nealed specimens, respectively. volume.) However, contaminant effects are also The earliest attempts to operate niobium and critical to the resistance of tantalum and niobium niobium alloy corrosion loops made use of helium alloys to attack by lithium and boiling potassium. as a protective environment. During the 1950s, As discussed under the subject of lithium corro- Pratt-Whitney CANEL examined various purifica- sion, initial tests of Nb and Nb-1% Zr were tion processes for large recirculating boJinjn sys- plagued by intergranular attack which was ulti- tems, but all fell short of protecting Nb-1 % Zr loop mately traced to oxygen contamination occurring components.104'105 Later Jet Propulsion Labo- during loop fabrication. Oxygen contamination has ratory106 operated a Nb-l%Zr loop under recircu- also been observed U. affect the operation of these lating argon using space-age seals ai.'d degassing alloys with boiling potassium. Each of these prob- practices. Although make-up argon was maintained lems has already been addressed in the context of below 1 ppm O2 and H2O, respectively, after 1000-h contamination "^curring during welding or during operation at 1093°C the combined oxygen p;us heat treating prior to liquid metal exposure. What nitrogen concentration of the loop walls had has not been addressed is the effect of contamina- increased to 4480 ppm. These results show the tion from supposedly protective environments dur- importance of low oxygen activity as an environ- ing the actual operation of these liquid metal sys- mental constraint in high temperature exposures of tems. niobium, tantalum, and their alloys. In the case of lithium, the pickup of oxygen by The only demonstrated method for protecting Nb and Nb-l%Zr from the ambient environment niobium and tantalum alloys from serious oxida- has led to localized penetration of the refractory tion has been the use of bakeable, cold-wall metal by lithium and the weeping or wickinf of chambers pumped by ion-sputtering, titanium sub- lithium to the outside of the piping system. Such a limation, or turbomolecular pumps. Even with this result is pioduced by the diffusion of oxygen from equipment, the bakeout cycle and heating rate of the external piping surface to the lithium-exposed tha refractory metal system are critical to the con- surface under conditions where all or a large frac- tamination problem. However, a number of large- tion of the diffusing oxygen is in solid solution in scale experiments with Nb-1% Zr and/or T-lll the refractory metal (i.e., at temperatures below were carried out in the late 1960s and early 1970s 1000°C or when the oxygen concentration has for times as long as 10,000 ii in which contamina- exceeded the concentration of active alloying addition by oxygen, as well as carbon and nitrogen, was tions such as Ti, Zr, or Hf). The penetration re&c- COMPATIBILITY OF REFRACTORY ALLOYS 79 tion is initiated at a threshold oxygen content in water- or alcohol-based solvents that produce the refractory metal and is not affected by the oxy- hydrogen when they react with alkali metals. Oak gen content of the lithium. In the case of potas- Ridge National Laboratory119 sfudies have shown sium, oxygen will Le pulled from the refractory that the susceptibility of a given alloy to embrittle- metal wall into the notassium at those surfaces ment is dependent on (1) the thermal history where condensation of potassium vapor occurs. The before decontamination, (2) the type of solvent oxygen will be swept to the evaporator region and used to dissolve the alkali metal, and (3) the sur- •••ill add to the oxygen inventory in the licjuid face condition of the sample. Alloys of tantalum potassium in that region. Given sufficient oxygen and niobium with high initial oxygen concentra- permeation through the condenser walls, the oxy- tions cracked worse than those with lower oxygen gen content of potassium in the evaporator can contents. Alloys that were oxidized, heat treated, reach a level where significant attack of the wall is and exposed to lithium at 1000°C were much more possible through oxidation reactions involving K2O susceptible to cracking than similar specimens and the refractory metal (see the Sodium and Po- exposed to lithium at 500°C. Cracking was always tassium section of this paper). associated with an increase in the hydrogen con- In addition to the environmental requirements tent of the alloy; however, the cracks were not a that must be maintained during the high- repult of hydride formation, for no hydride phase temperature operation of refractory metals, was o' served in the microstructure. Buik hydrogen another set of environmental practices must be increases of as little as 50 ppm to as high as 1200 invoked during the preparation and handling of ppm were recorded for samples that showed cracks. refractory metals at room temperature These 120 General Electric workers also found that include the avoidance of even trace contamination removal of residual sodium from Nb-l%Zr loop of refractory metal surfaces by contact with non- surfaces with ethyl alcohol led to hydrogen pickup refractory metals (e.g., vice-jaws, cutoff wheels, that affected the room temperature bend ductility etc ) and pickup of hydrogen. Among the more det- of the wall. Even though only a thin film of sodium rimental metal contaminants are nickel, copper, was present and was removed in a few seconds, the and platinum. Thus, contact of Pt/Pt-Rh thermo- bulk hydrogen content of the loop tubing increased couples with refractory metals must be avoided as much as 170 ppm. Heat treating the embrittled above ~800°C. tube sections for four hours at 1093° C completely A considerable amount of work has been restored their room temperature ductility. reported111"118 concerning hydrogen embrittlement of refractory alloys. Experimenters at The problem of hydrogen embrittlement during Rocketdyne111"114 evaluated the compatibility of alkali metal decontamination can be overcome by 119 121 Ta 0% W and alloy B-66 with hydrogen over a several techniques, ' the most successful of wioe temperature range. The investigators113-114 which is the use of anhydrous liquid ammonia, found that under certain conditions both B-66 and which dissolves the alkali metal without generating Ta-10% W could be made to absorb hydrogen at hydrogen. However, if the ammonia solution is room temperature and that this absorption often allowed to stand for a period of hours or is heated resulted in general disintegration of the sample. If above room temperature under pressure, it will a completely clean or newly made surface was form amide and generate hydrogen. This reaction exposed to hydrogen at room temperature, the is also catalyzed in the presence of certain metals. sample fragmented and created new surfaces for Oak Ridge National Laboratory experiments, how- continued fragmenting. Whether or not actual for- ever, have shown nG evidence of hydrogen pickup mation of hydride at room temperature was by tantalum or niobium alloys when continuous responsible for the cracking was not stated. How- flushing with ammonia was used to dissolve ever, several investigators112'115'117 have reported lithium residues. embrittlement of one type or another in either tan- Vacuum distillation is also a feasible technique talum or niobium alloys before formation of a visi- for removing residual films, particularly in the ble second phase. Thus, hydrogen in solution case of potassium and cesium which have relatively appears to be a sufficient condition to induce high vapor pressures. Organic solvents that react embrittlement. slowly and generate hydrogen very slowly, such as The Rocketdyne observations concerning frag- butyl cellusolve, are much less prone to produce mentation correlate with observations of cata- hydrogen embritt'ement than solvents that react strophic cracking of refractory metals during the fast, such as alcohol and water, but the use of removal of alkali metals after corrosion testing. these solvents is generally not as reliable as the Sach cracking has been associated wifeh the use of use of ammonia or vacuum distillation. 80 DEVAN, KSTEFANO. AND HOFFMAN

Room temperature embrittlement of T-111, both metals appear highly corrosion resistant to apparently related to hydrogen r

• Derive deoxidation rate expression for T-lll bine rotor assembly in potassium vapor under and ASTAR-811C in lithium and verify exper- prototypic Rankine cycle operating conditions. imentally. • Conduct simple heat pipe experiments to eval- • Conduct long-term (10,000 h) demonstration uate corrosion properties of Mo-13%Re in of ASTAR-811C forced convection heat trans- lithium under evaporative heat transfer confer system with lithium that simulates proto- ditions. typic reactor circuit operating conditions • Determine long-term (7 years) rates of oxy- (1000° to 1350°C). gen diffusion from UO2 to lithium and its effect on evaporator corrosion in molybdenum Tungsten Alloys and Mo-13%Re heat pipes. • Determine mass transfer rates of substitu- tional and interstitial elements in T- Brayton-Cycle Working Fluids 111/tungsten and T-lll/W-26%Re dissimilar Tantalum Alloys metal loop systems between 1200 and 1650°C. Conduct a small-scale feasibility loop experi- Niobium Alloys ment to demonstrate outgassing and helium- • Long-term (>10,000 h) corrosion loop tests of cleanup procedures that will achieve accept- niobium-base alloys in lithium will b" needed able contamination rates of T-lll over a if such alloys are to be qualified for auxiliary seven-year operating period at temper- (<800°C) lithium heat transfer circuits. atures up to 1200°C. • Determine susceptibility of 10% hafnium lev- • Determine mass transport rates of L ;rsiit'?.i els (as exist in C-103) to temperature gradient impurities in T-lll from hot to cold regions mass transfer in lithium and sodium. of a pressurized helium circuit. • Determine lithium corrosion resistance of • Determine coating and purity requirements Nb-l%Zr to T-lll dissimilar metal welds as a for graphite moderator and fuel assemblies to function of weld parameters and postweld achieve compatibility with T-lll in helium. heat treatment. Tungsten Alleys Molybdenum Alloys • Determine oxidation and carburization rates • Determine mass transfer behavior of Mo- of tungsten and W-26%Re in reactor-grade 13% Re in lithium at 1000 to 1350°C. helium at temperatures to 1350°C. Two-Phase Alkali Metals Niobium Alloys Tantalum Alloys • Based on test results for helium loop feasibility ex jeriments with tantalum alloys, similar • Conduct long-term (10,000 h) demonstration feasibility jxperiments should be conducted test of ASTAR-811C forced convection boiler, on niobium alloys (particularly, Nb-l%Zr and condenser, and static turbine blade assem- C-103) ai temperatures to 1000°C. blies with potassium under prototypic Rank- Molybdenum Alloys ine cj Je operating conditions. Tungsten Alloys • Determine oxidation and carburization rate3 of Mo and Mo-13%Re in reactor-grade helium • Conduct simple heat pipe experiments to eval- at temperatures to 1200°C. uate corrosion properties of tungsten and W- 26% Re in lithium under evaporative heat transfer conditions. REFERENCES Niobium Alloys 1. E. E, Hoffman, Corrosion of Materials by Lithium at Elevated Temperature, pp. 56-60, ORNL-2674 (1959), and • Evaluate corrosion resistance of C-103 to E. E. Hoffman, The Effects of Oxygen ami Nitrogen on boiling potassium and cesium in refluxing the Corrosion Resistances of Columbium to Lithium at capsule experiments. Elevated Temperatures, ORNL-2675 (1959). 2. J. R. DiStefano, Corrosion qf Refractory Metals by Molybdenum Alloys Lithium. ORNL-3551, March 1964. 3. R. L. Klueh, The Effect of Oxygen on the Corrosion qf • Conduct long-term (10,000 h) demonstration Niobium and Tantalum by Liquid Lithium, ORNL- test of small molybdenum and/or TZM tur- TM-4069, March 1973. CGMPATHJTY OF REFRACTORY ALLOYS 83

4. 0. E. Sessions and J. H. DeVan, Effects of Oxygen, Heat 24. R. W. Harrison, E. E. Hoffman, and J. P. Smith, T-lll Treatment, and Test Temperature on the Compatibility of RanHne System Corraeion Test Loop, Volumes I and II, Several Advanced Refractory Allays With Lithium, NASA-CR-134816, June 1975. ORNL-4430, April 1371. 25. J. A. DeMaatry, Corrosion Studies of Tungsten, Molybde- 5. J. H. DeVan snd C. E. Sassions, Mass Transfer of num, and Rhenium in Lithium, NucL AppL, Vol. 3, FebrtS* Niobium-Base Alloys in Flowing Nonisothennal Lithium, ary 1967. NucL AppL, 3:102-109, February 1%7. 26. L. G. Hays, Corrosion of Nwtoium-1% Zr Alloy and Yttria 6. C. fc. Sessions and J. H. DeVan, Thermal Convection Loop by Lithium at High Flow Velocities, Jet Propulsion Tests of Nb-1% Zr in Lithium at 1200 and 1300°C, NucL Laboratory, Technical Report 32-1233, December 1,1967. AppL TechnoL, 9: 250-259, August 1970. 27. G. E. Tardiff, Corrosion Damage to a Tungstetir25 AL% 7. C. E. Sessions, High-Temperature Material Program Rkeniuvi-SO AL% Molybdenum Containment Alloy After Quarterly Progress Report, pp. 42-47, ORNL-iM-1520, Exposure to Flowing Lithium, Lawrence Radiation April 30, 1966. Laboratory, UCID-15356, September 2, 1968. 8. J. H. DeVan, Fuels and Materials Development Program 28. E. A. Skrabek, Heat Pipe Design HandbooJz, Part 2, Quarterly Progress Report, pp. 127-130, ORNL-4350, NASA-CR-134265, Dynatherm Corporation, August 1972. Sept. 30,1968. 9. J. H. DeVan, Fuels and Materials Development Program 29. C. A. BuBse, Korrosion in Hochtemperatur-Lithium- Quarterly Progress Report, pp. 105-106, ORNL-4420, Warmerohren Mit Niob-lZikon Oder Tantal Als Wand- March 31,1969. .naterial, in Corros. £ei, 10: 65-84 (1970). 10. J. H. DeVan and E. L. Long, Jr., Evaluation of T-lll 30. I. Von Didie Quataert, Investigations of the Corrosion Forced-Convection Loop Teal id With Lithium ot 137n°C, Mechanism in Tantalum-Lithium High Temperature Heat ORNL-TM-4775/NASA CR-134745, April 1975. Pipes by Ion Analysis, Forsch. Ing.-Wes., 37: 37-38 (1971). 11. Corrosion of Columbium Base and Other Structural Alloys 31. C. A Busse, Heal Pipe Research in Europe, paper in High Temperature Lithium, PWAC-355, June 30, 1961. presented at Second International Conference on Ther- 12. H. W. Leavenworth, R. V,. Heary, and H. D. Brattcn, mionic Electrical Power Generation, Stresa, May 27-31, Solubility of Some Transition Metals in Lithium, 1968. PWAC-356, June 30,1961. SL J. R. DiStefano .cid J. H. DeVan, Refluxing Capsule 13. Stephan Stecura, Corrosion of Oxygen-Doped Tanta- Experiments with Refractory Metals and Boiling Alkali lum by Lithium, NASA Technical Menuvt. Mum, Metals, NucL App. TechnoL, 8: 29-44 (January 1970). NASA-TM-2989, Lewis Research Center, Cleveland, Ohio, 33. W. A. Ranken a^d J. E. Kemrae, Survey of LOB Alamos March 1974. and Euratom Hiat Pipe Ln cstigations, pp. 325-36 in 14. Randall F. Gahn. Effects of Low-Pressure Air on Oxygen IEEE Canf. Record of 1965 Thermionic Conversion Special- Contamination and Lithium Corrosion of a Tantalum ist Conf.. October 1965. Alloy, T-lll, at 980° and 12S0°C, NASA Technical Note, 34. G. Yale, Eastmrn Thermacore, Inc., private communica- NASA TN D-7588, Lewis Research Center, Cleveland, tion to J. H. I'eVan, Oak Ridge National Laboratory, Ohm, April 1974. January 1983. li). Juhn H Sinclair, Compatibility Tests of Materials for a 35. M. Merrigan, Los Alamos National Laboratory, private Lithium-Cooled Space Power Reactor Concept, NASA communication to J. H. DeVan, Oak Ridge National Technical Memorandum., NASA TN D-7529, Lewis Laboratory, August 1983. Research Center, Cleveland, O1-"?, May 1973, 36. R. L. Klueh, The Effect of Oxygen in Sodium on the Com- 16. G. P. Brandenburg, E. E Hoffman, and J. P. Smith, patibility of Sodium and Niobium, in Proceedings of the Pumped Lithium Loop Test to Evaluate Advanced Rejrac- liiternati'mal (hnference on Sodium Technology and Large tory Metal Alloys and Simulated Nuclear Elements, Fast Reactor Design, ANL 7520, Part I, November 7-9, pp. NASA-CR-134527, August 1973. 171-176. 17. R. W. Hairison, The Effects of Welding Atmosphere 1'ur- 37. A. P. Litmar, ''Tie Effect of Oxygen on the Corrosion of ity on the Lithium Corrosion Kesistance of Refractory Niobium by JJquid Potassium, ORNL-3751, July 1965. Alloys, Corrosion by Liquid Metals, J. E. Draley and J. R. 38 R. L. Klueh. The Effect of Oxygen on Tantalum-Sodiui.i Weeks (Eds.), Plenum Press, 1970. Compatibility, ORNL-TM-3590, December 1971. 18. Coulson M. Scheuermann, Mass Transfer in a IS7O"C 39. R. L. Klueh, The Effect of Oxygen on the Compatibility of (2500°F) IMhium Thermal-Ccnvection Loop, NASA-TM X- Tantalum and Potassium, ORNL-4737, November 1971. 2986, Lewis Research Center, Cleveland, Ohio, February 40. J. R. DiStef uno, Mass Transfer Effects in Some Refractory 1974. Metal-Alkali Metal-Stainless Steel Systems, ORNL-4028, 19. R. W. Harrison and J. Holovach, Refractory M^tal Values November 1MJ6. for 1900°F Service in Alkali Metal Systems, NASA CR- 41. J. R. DiStofano and J. H. DeVan, Refluxing Capsule 1810, 1970. Experiment With Refractory Metals and Boiling Alkali 20. G. B. Brandenburg, E. E. Hoffman, and J. P. Smith, Metals, NwL AppL TechnoL, 8: 29-44, January 1970. Pumped Lithium Loop Test to Evaluate Advanced Refrac- 42 J. H. DeVpn, Compatibility of Structural Materials With tory Metal Alloys and Simulated Nuclear Test Elements, Boiling PoUissium, ORNL-TM-1361, April 1966. NASA-CR-134527, January 1974. 43 D. H. Jan3i n and E. E. Hoffman, Niobivm-1% Zirconium, 21. G. K. Watson, Preliminary Evaluation of T-lll Clad UN NaturalrCirculation, Boiling-Potassium Corrosion Loop Fuel Specimens Frvm 2600-Hour, lQlfiX (19O0°F) Lithium Test, ORNL-3603, May 1964. Loop Test, NASA-TM X-529S8, March 197i. 44. D. H. Janben et al., Ntobium-1% Zr Boiling Potassium 22. J. H. Sinclair, Compatibility Tests of Materials for a Forced-Circulation Loop Test, ORNL-4301, December 1968. Lithium-Cooled Space Power Reactor Concept, NASA TN 45. A. P. Fraa.---, H. u. Young, and A. G. Grindell, Survey of D-7258, May 1973. Informatior, on Turbine Bucket. Erosion, ORNL/TM 2088, 23. R. E. Gluyas ano G. K. Watson, Materials Technology for July 1968. an Admnced Space Power Nuclear Reacltr Concept- 46. S. Stecura, Apparent Solubilituji cf Commercially Pure Program Summary. NASA TN D-7H09, March 1975. and Oxygei^Doped Tantalum and Niobium in Liquid 84 DEVAN. DISTEFANO, AND HOFFMAN

Potassium, NASA TN D-5875, Lewis Research Center, Vapor Cesium on Container Metals, Report ASD-TRD- Cleveland, Ohio, July 1970. «?°-965, Rocketdyne Division of North American Aviation, 47. C. M. Scheuermann and C. A. Barrett, Compatibility of inc., March 1963. Columbium and Tantalum Tubing Allays with Refluxing 6S. P. M. Winslow, Synopsis of Cesium Compatibility Studies, Potassium, NASA TN D-3429, Lewis Research Center, CONF-650411, p. 334, April 1965. Cleveland, Ohio, May 1966. 69. F. Hargreaves, G. T. J. Mayo, and A. G. Thomas, A Study 48. A. J. Romano, A. H. Fleitman, and C. J. Klamut, Behavior of the Long-Term Compatibility of Thermionic Converter of Refractory Metals and Alloys in Boiling Sodium and Materials with Cesium, J. Nucl Mater., 18: 212-218 Other Boiling Alkali Metal Coolants, in Proceedings qf (1966). the Symposium on Alkali Metal Coolants, Vienna, 70. F. Tepper and J. Greer, Factors Affecting the Compatibil- November 28-December 2,1966, IAEA pp. 673-674 (1967). ity of Liquid Cesium with Containment Metals, Report 49. A. J. Romano, A. II. Fleitman, and C. J. Klamut, A Sum- AFML-TR-64-327, MSA Research Corporation, November mary of tlie High Temperature Alkali Metal Corrosion 1964. Program at Brookhaven National Laboratory, AEC-NASA 71. F. Tepper and J. Greer, Factors Affecting the Compatibil- Liquid Metals Information Meeting, USAEC Report ity of Liquid Cesium with Containment Metals, CONF-350411, p. 234 (1966). CONF-650411, p. 323, April 1965. 50. W. T. Chandler, Alkali Metal Corrosion Studies at Rocket- 72. W. H. Hall and S. W. Kessler, Cesium Compatibility of dyne, NASA-AEC Liquid Metals Corrosion Meeting, Thermionic Converter Structural Materials, N65-33774, USAEC Report TIC-7626 (P..rt 1), pp. 42-62, April 1962. Radio Corporation of America, 1963. 51. R. W. Harrison, SNAPS Refractory Boiler Develop- 73. D. V. Rigney, unpublished internal memorandum sum- ment—Corrosion of Oxygen Contaminated Tantalum in marizing cesium corrosion data, Pratt and Whitney NaK, NASA-CR-1850, July 1H71. Aircraft-CANEL, July 15,1965. 52. E. E. Hoffman and J. Holowach, Cb-1% Zr Sodium Ther- 74. J. A. DeMastry and N. M. Griesenauer, Investigation of mal Convection Loop, NASA-CR-1097, September 1968. High Temperature Refractory Metals and Alloys for 53. E. E. Hoffman and J. Holowach, Cb-1% Zr Pumped Thermionic Converters, AFAPL-TR-65-29, Battelle Sodium loop, NASA-CR-1135, November 1968. Memorial Institute, April 1965. 54. E. E. Hoffman and J. Holowach, Cb-1% Zr Rankxne Sys- 75. R. G. Smith et al., A Study of the Compatibility of Ther- tem Corrosion Test Loop, NASA-CR-1509, June 1970. mionic Converter Materials with Cesium, J. Nuc. Mater., 55. B. L. Moor and R. G. Frank, Friction and Wear of Materi- 10: 191-200 (1963). als in Sliding Contact from Rcom Temperature to 1200°F 76. M, J. 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W. Harrison, The Compatibility of ity of Liquid Cesium with Containment Metals, Refractory Metals with Liquid Metals, Refractory Metal ASD-TDR-63-824, Part I, MSA Research Corporation, Alloys, Metallurgy, ami Technology, I. Machlin, R. T. Beg- September 1?<53. ley, and E. D. Weisert (Eds.), Plenum Preso, pp. 251-287 80. R. L. Ammon, R. T. Fegley, and R. L. Eichinger, T-lll (1968). Cesium Natural Coriection Ixxyp, CONF-650411, April 60. E. E. Hoffman and J. Holowach, Co-1% Zr Rankine Sys- 1965. tem Corrosion Test Loop, NASA-0r'-1509, June 1970. 81. P. F. Young and P. Achener, Operation of 1850°F Pumped 61. R. W. Harrison, E. E. Hoffman, and J. P. Smith, T-lll Loop System and Determination of Specific Heat, Density, Rankire System Corrosion Test Loop, .NASA-CR-134816, and Vapor Pressure of Cesium Between 174 and 1800°F, June 1975. Aerojet General Nucleonics, May 1962. 62. G. E. Raines, 0. V. Weaver and J. H. Stang, Corrosion 82. L. A. Chariot, R. A. Thiede, and R. E. We^'jrman, aiU Creep Behavior of Tantalum in Flowing Sodium, Corrosion of SuperaUoys and Refractory Metals in High Batte'.ie Memorial Institute Report BMI-1284, August Temperature Flowing Helium, BNWL-SA-1137, Battelle 1S58. Northwest Laboratories, Richland, Wash., September 63. R. Cygan and E. Reed, Molybdenum Corrosion by Sodium, 1967. North American Aviation, Inc., USAEC Report NAA-SP- 83. H. Migge and H. Andresen, Compatibility of Vanadium, 161, November 20, 1951. Nirbium, and Molybdenum wKh Impure Htlium, in 64. S. S. Bleckerman and J. Hodel, The Compatibility of Struc- Proceedings of 9th Symposium, on Fusion, Garisch- tural and Turbo-Machinery AUnys in Boiling Potassium, Partenkirchen (FRG), June U-18, pp. 605-609, Pergamon USAEC Report PWAC-501,1965. Press, 1976. 65. E. Kovacevich, Potassium Corrosion Studies, AiResearch 84. H. K. Kohl, The Oxidation of Vanadium with Oxygen in Manufacturing Company of Arizona, NASA-AEC Liquid Helium, J. Nucl Mater., 41: 231-234 (1971). Metals Corrosion Meeting, USAEC Report TID-7626 85. T. Noda, M. Okada, and R. Watanabe, The Compatibility (Part 1), pp. 63-68, April 1962. of First Wall M"'nllic Materials with Iu^u.. Ilchum, 66. R. L. Klueh and D. H. Jansen, Effects of Liquid and Vapor J. Nuc Mater., 85-86: 329-333 (1979). Cesium on Structural Materials, ORNL/TM-1813, June 86. T. Noda, M. Okada, and R. Watanabe, Corrosion 1967. Behaviors of Iron-Base Alloy, Nickel-Base Alloy, and 67. W. T. Chandler and N. J. Hoffman, Effects of Liquid and Refractory Me'als in High Temperature Impure Helium CCWFATBHJTY OF REFRACTORY ALLOYS 81)

Gas, J. NucL ScL TechnoL, 17(3): 191-203, Atomic Energy Pressures, TTM-901, Pratt-Whitney Aircraft, CANEL, Society of Japan, 1980. August 1965. 87. B. Bergenov-Hansen, N. Endow, and R. A. Pasternak, 105. E. A. Limoncelli, Use of Helium and Argon as Protective Sorptian qf Gases on Metal Surfaces in UUrahigh Vacuum, Environments for SNAP-50/SPUR Program, TIM-824, Annual Report, February 15, 1964, to February H, 1965, Pratt-Whitney Aircraft, CANEL, November 1965. Stanford Research Institute, March 23,1965. 106. W. M. Phillips, Some Alkali Metal Corrosion Effects in a 88. R. A. Pasternak, High Temperature Oxidation and Ni- Rankine Cycle Test Loop, pp. 197-217, Corrosion by Liquid tridation of Niobium in UUrahigh Vacuum, Final Report, Metals, Plenum Press, New York, 1970. July 1, 1963, to October 31, 1964, Stanford Research 107. D. H. Jansen and E. E. Hoffman, Niobium-1% Zirconium, Institute, November 15,1964. Natural Circulation Boiling-Potassium Corrosion Loop 89. D. T. Bourgette, High-Temperature Chemical Stability qf Test, ORNL-3603, Oak Ridge National Laboratory, Oak Refractory-Base Alloys in High Vacuum, ORNL/TM-1431, Ridge, Tennessee, 1964. <">ak Ridge National Laboratory, Oak Ridge, Tennessee, 108. E. E. Hoffman and J. Holowach, Cb-1 Zr Sodium Thermal 1J66. Convection Loop, NASA CR-1097, General Electric Co., 90. H. Tnouye, Contamination of Refractory Metals by Residual Nuclear Systems Programs. Gases in Vacuums Below 10'" Torr, ORNL-3674, Oak 109. J. H. DeVan and E. L. Long, Jr., Evaluation of T-lll Ridge National Laboratory, Oak Ridge, Tennessee, 1964. Forced Convection Loop Tested with Lithium at 1S7O°C, 91. V. C. Marcotte and W. L. Larsen, Phase Studies of High- NASA Cr-134745 (ORNL/TM-4775), Oak Ridge National Niobium-Conteut Nb-Zr-0 Alloys, J. Less-Common Met., Laboratory, Oak Ridge, Tennessee, 1975. 10: 229-236 (1966). 110. R. W. Harrison, E. E. Hoffman, and J. P. Smith, T-Ul tf2. R. W. Carpenter and C. T. Liu, Surface Reaction Con- Rankine Syste-m Corrosion Test Loop, NASA Cr-134816 trolled Oxygen Absorption in a Ta-8 W-2 Hf Alloy: Kinet- (Vol. 1), General Electric Co., Nuclear Systems Programs, ics and Concentration Gradients, Metallurgical Transac- 1975. tions A, 6A: 2235-2241 (1975). 111. R. J. Walter, The Columbium-Hydrogen System and 93. V. K. Sikka and C. J. Rosa, The Abnormal Oxidation Hydrogen Embrittlement of Colwmbium, Rocketdyne Kinetics of Niobium at High Temperatures, Z. Metallkde, Research Report 64-6, February 1964. 69: 777-780 (13?^. 112. R. J. Walter, Compatibility of Tantalum and Columbium 94. C. T. Liu and H. Inouye Internal Oxidation and Mechani- with Hydrogen, Rocketdyne Research Report 65-3, Febru- cal Properties of TZM-Mo Alloy, Metallurgical Transac- ary 1965. tions, 5: 2515-2525 (1974). 113. R. J. Walter and J. A. Ytterhus, Compatibility of Colum- 95. R. W. Harrison and E. E. Hoffman, Contamination of bium with Hydrogen, Quarterly Progress Report, Tantalum and Tantalum Alloys in Low Pressure Oxygen Rocketdyne Research Report R-6345-2, January 1966. Environments, pp. 29-57 in Transactions of the Interna- 114. R. J. Walter and W. T. Chandler, Compatibility of Tan- ikmal Vacuum Metallurgy Conference, R. B. Barrow and talum and Columbium AlloyB with Hydrogen, in AIAA J., A. Simko ' (Eds.), American Vacuum Society, New 4(2): 304-307, February 1966. York, 1968. 115. B. Longson, The Hydrogen Embrittlement of Niobium, 96. H. Inouye, Equilibrium SolM Sdittions oil Nitrogen in TRG-1035 (C), 1966. Cb-1% Zr Between 1200-180O°C, ORNL/TM-1355, Oak 116. A. G. Ingram et al., Notch Sensitivity of Refractory Ridge National Laboratory, Oak Ridge, Tennessee, 1966. Metals, ASD-TR-61-474, August 1961. 97. H. lnouye, High-Temperature Sorption of Nitrogen by 117. J. R. Stephens and R. G. Gorlick, Compatibility of Tan- Nb-1% Z: in UUrahigh Vacuum, ORNL^403, Oak Ridge talum, Colvmbium, and Their Alloys with Hydrogen in National Laboratory, Oak Ridge, Tennessee, 1969. Presence of Temperature Gradient, NASA-TN D3546, 98. R. F. Bunshah, Contamination of Reactive Metals in August 1966. Vacuum Heat Treating at Various Vacuum Levels, 118. J. R. Stephens, Role of Hf and Zr in th". Hydrogen Embrit- California University, Livermore, California, July 5,1968. Oement ofTa and Cb Alloys, NASA-TM X-68293, ir73. 99. H. E. McCoy, Carburization of Niobium- and Tantalum- 119. C. E. Sessions and J. H. DeVan, Effect of Oxygen, Heat Base Allo>s, J. Less-Common Metals, 12:139-145 (1967). Treatment, and Test Temperature on the Compatibility of 100. H. Inouye, Refractory Metal Alloys, Plenum Press, p. 165, Several Advanced Refractory Alloys with Lithium, pp. 1968. 43-46, ORNL-4430, Oak Ridge National Laboratory, Oak 101. 0. T. Liu, R. W. Carpenter, and H. Inouye, Oxygen Distri- Ridge, Tennessee, 1971. bution in Internally-Oxidized Ta-8 Pet W-2 Pet Hf Alloy, 120. E. E. Hoffman and J. Holowach, Cb-1% Zr Sodium Ther- MetalL Trans. A, 6A: 419-421 (1975). mal Convection Loop, pp. 64-65, NASA Cr-1097, General 102. R. J. Lauf, A Study of Metal-Oxygen Solid Solutions Using Electric Co. Solid Electrolytic Cells, Ph.D. Thesis, University of Illinois, 121. R. F. Gahn, Techniques for Lithium Removal from 104O"C Urbana, Illinois, 1978. Aged Tantalum Alloy, T-lll, NASA-TM X-2775, April 103. T. F. Lyon, Low-Pressure Oxidation of Cb-1 Zr Alloy, J. 1973. Vac. Sd TechnoL, 8(6): VM488SS-58, June 1971. 122. R. W. Harrison, E. E. Hoffman, and J. P. Smith, T-lll 104. J. F. Hogan, E. A. Limoncelli, and R. E. Cleary, Reaction Rankine System. Corrosion Test Loop, pp. 47-50, NASA- Rate of Columbium-1 Zirconium Alloy with Oxygen at Low CR-134816 (Vol. II), General Electric Co., 1975. A Review of Tantalum and Niobium Alloy Production

R. W. Buckman, Jr. Westinghouse Electric Corporation

INTRODUCTION refractory metal alloy developments that took place between 1955 and 1965. The second, Niobium, and Niobium Alloys in Nuclear Power, by D. C. Materials requirements for space nuclear power 2 systems as embodied in the SP-10O program are Goldberg et al. is an excellent treatise on niobium governed by a unique set of operating require- alloy development. The information is equally ments: high power density, light weight, and liquid applicable to tantalum alloy development. alkali metal coolant. The reactor will drive some The appended bibliography will serve 83 a type of power conversion system which could be source of more details relating to niobium and tan- Rankine cycle, Brayton cycle, thermoelectric, or talum alloy production. thermionic. Heat rejection will be by radiation. All of these factors inevitably lead to very high temperatures of operation which in turn restrict the REQUIREMENTS selection of structural materials to the refractory metal alloy systems based on tantalum, niobium, The performance required of refractory alloys molybdenum, and tungsten. will ultimately depend on the reactor type and Advanced nuclear space power systems were conversion system chosen. If a gas reactor concept last in vogue between 1965 and 1975 when exten- is chosen, then the structural alloys will have to be sive work was performed under the direction of the compatible with the high purity inert gas working National Aeronautics and Space Administration fluid. If the reactor is liquid metal cooled, then (NASAj. System requirements have not changed lithium compatibility is a nectssary requirement significantly since then, and the developments for the reactor system. In addition, resistance to vhich took place in materials fabrication are boiling potassium or cesium in the conversion sys- directly relevant to the SP-100 program. With the tem will be important. Regardless of the reactor contraction of the space program, and with limited concept chosei., however, efficiency and mass terrestrial aerospace application (because of the requirements dictate that maximurr operating need for reliable oxidation resistant coatings), only temperatures will be in excess of 1200°C thus a few of the refractory metal alloys were produced essentially eliminating all but refractory metal in any commercially significant quantities. alloys from consideration. This paper will concentrate on the current btat~ Conceptual representations of a compact reactor of niobium- and tantalum-base alloy production. To and u 2000 kW potassium boiler for a Rankine sys- provide an orderly discussion, the materials tem having an outlet temperature of 1150°C requirements, alloy compositions of interest, and (Ref. 3) are illustrated in Figs. 1 and 2. These are production status will be discussed. Finally, a list vintage mid-1960s concepts but are comparable in of developments needed to support the SP 100 pro- size to those being currently considered for SP-100. gram will be identified. The sizes of the SP-100 compact reactor system Two key references are worthy of mention. The components c.re auch that the mill shapes, sheet, first, A Decade of Progress in Refractory Metals, by plate, strip, tubing, bar, rod forgings, wire, and G. M. Ault1 is a thorough historical account of the foil, are well within the industry's fabrication

86 REVCW OF TANTALUM AND NIOBIUM AUOY PRODUCTION 87

III PRESSURE VESSa

CORE FUa ELEMENTS

H0NEYC0W& I

ORUMFUa : ,-'',' ELEMENTS^" ,•' TROL DRUM

MOLY»DENUM-TZM-' (

T-lll ABSORBER - NIOLYBDENUM-TZM REF:£C'Oa

1 0.58 m '

Fig. 1 Nuclear power plant reactor.5

LITHIUM INLE

POTASSIUM VAPCR EXIT POTASSIUM LIQUID INLtT

Fig. 2 200G kW thermal potassium boiler, outlet tempera- tu~* 1150°C (Ref. 3). capability. } typical product availability list is development effort occurred between 1960 and 1965 phown in Tabr.. 1. in response to Air Force aerospace needs for high temperature, oxidation resistant alloys for turbine and air frame applications. Compatibility with oxi- ALLOY COMPOSITIONS dation resista.it. coatings; high, short time strength; fabricability; and weldability were of Niobium and tantalum allo> compositions that prime importance. reache.] commercial and/or pilot &cale production are listed in Table 2. In 1959, the only niobium These programs lod to the development of a alloy available wad Nb-iZr, and the only tantalum series of niobium alleys with B-66, C-103, FS-85, alloys were the tantalum-tungsten binar!os which Scb 291, C-129ir, Cb-752, and D-43 being the most wer? used in the lamp industry.9 An intensive alloy prominent. Of these alloy compositions, C-103 and 88 BUCKMAN

TABLE 1 Typical Product Availability for Niobium and Tantalum Alloys4-7'

Thickness, Width, Length, Furm mm (in.) m(in.) m(in.)

Foil 0.025 (0.001) 0.23 (8) Coil Strip 0.027-0.483 (0.005-0.019) 0.610 (24) Coil Sheet 0.508-1.575 (0.020-0.062) 1.22 (48) 3.05 (120)» Seamless 3.175 0.0. X 0.254 wall - 3.05 (120)* tubing to 63.5 O.D. X 0.5 wall (0.125 O.D. X 0.010 wall to 2.5 O.D. X 0.5 wall)

•Limited by maximum vacuu n annealing f"rnace length.

TABLE 2 Niobium and Tantalum Alloys1"2-*"'

Element w/o Alloy designation Nb Ta V Mo W Re Ti ZT Hf Y C Status

Nb-lZr Bal. . _ _ 1.0 _ CommerciM C-103 Bal. 0.5 - 0.5 1.0 0.7 10.0 - Commercial C-129Y Bal. - - 10.0 - 0.7 10.0 0.1 _ Commercial Cb-752 Bal. _ 10.0 - 2.5 _ _ Commercial D-43 Bal. _ - 10.0 - 1.0 - 0.1 Commercial SCb-291 Bal. 10 - 10.0 - - - - Commercial FS-85 Bal. 27 _ 11.0 _ 0.9 _ Commercial B-66 Bal. „ 5.0 5.0 _ _ 1.0 _ _ Commercial Ta-2.5W - Bal. - 2.5 - - - - Commercial Ta-7.5W - Bal. - 7.5 - - - - Commercial Ta-10W - Bal. - 10.0 - - - - Commercial KBI-Alloy 40 40 60 ------Commercial Till _ Bal. 8.0 _ 2.0 _ Commercial T-222 - Bal. - 9.6 - - 2.4 - 0.010 Pilot plant prod. ASTAR-811C - Bal. - 8.0 1.0 _ 0.7 0.025 Pilot plant prod. ASTAR-1211C - Bal. - J2 1.0 - _ 0.7 - 0.025 Pilot plant prod. ASTAR-1511C - Bal. - 15 1.0 - 0.7 0.025 Pilot plant prod.

Nb-lZr are still in production for aerospace (C-103) properties.8'12 Although significant quantities of T- and sodium vapor lamp applications (Nb-lZr). ill were produced in the 1960s, very little is in production today. Limited quantities are being pro- The tantalum alloys T-lll and T-222 were also duced for radioisotope encapsulation. The primary developed in this period under Navy sponsorship demand today is for the Ta-W binary alloys, which for aerospace applications 1(M1 However, T-lll with are used extensively in the chemical industry its demonstrated compatibility with liquid alkali because tantalum has outstanding corrosion resis- metals and combination Oi strength, fabiicability, tance to strong acids such as sulphuric, nitric, and i'-i ^eldability was selected as the baseline refer- hydrochloric.5"7 'X.cf: alloy by N^SA for the space nuclear power * or SP-100 concepts utilizing liquid alkali metal in the mid 1960s:3 ASTAR-811C was coolant, Nb-lZr, FS-85, T-lll, and ASTAR-811C all to extend the operating temperature have demonstrated a combination of corrosion r::.X\S-A of T-lll without compromising its resistance to alkali metals, adequate temperature -iyuv..;'// and/or liquid alkali metal corrosion creep strength, fabricability, and weldability.9'13"15 REVEW OF TANTALUM AND NIOBIUM ALLOY PRODUCTION 89

ELECTRON BEAM PURIFICATION AND MELTING ALLOY ADDITION

DOUBLE VACUUM CONSUMABLE ELECTRODE FINAL ALLOY ADDITION ARC MELTING AND COMPOSITION CONTROL

PRIMARY BREAKDOWN INGOT 900-1300°C J

ROUND EXTRUSION SHEET BAR FORGING TUBULAR IN PROCESS ANNEAL (1100-1650°C) SECONDARY BREAKDOWN (R.T. to 1300°C)

EXTRUSION RCU.ING } SWAGING, ROCKING | iN PHOCESS ANNEAI !1100-1650°C) PLATE, ROD, TUBE REDRAW STOCK, PIPE

FINAL WORKING (R.T.)

DRAWING ROLLING Oti PILGERING S'VAGING iN PROCESS ANNEALING (VARIABLE) ROD, BAR, SHAPE (11OO-165O°C) WIRE SHEET, STRIP ROD j TUBING FOIL WIRE FINAL ANNEALING (1100-1o50°C)

Fig. 3 Process flow diagram. PRODUCTION OF NIOBIUM melting for purification and consumable electrode AND TANTALUM ALLOYS vacuum arc melting for final alloy addition and composition control. A typical ingot melting sequence for T-lll is illustrated in Fig. 4. Tan- Production of niobium and tantalum alloy mill product is accomplished following the process flow talum starting powders are produced by sodium diagram illustrated in F;.g. 3. The production tech- reduction of . xide or by the hydriding/dehydriding niques developed in the 60s are still in general use of metal scrap. After one or more electron beam witn only minor modifications. The following melts, the as-melted ingot is analyzed for impuri- description of the production of niobium and tan- ties and alloy content. Necessary alloy additions talum alloys will be brirf because the information are GTA welded into the electron beam melted is well documented elsewhere.2-3'11'13"14'16 ingot to form the electrode for the first of two consumable vacuum arc melts to ensure a homogeneous final ingot composition. For niobium, ingot sizes of 208 to 406 mm (8 co 16 in.) diameter are Melting common; for tantalum, the Ingot size ranges fror" nominally 178 to 254 mm (7 to 10 in.). Generally, Melting of niobium and tantalum alloy ingots is ingot size is not limited by melting capacity but by accomplished by a combination of electron beam primary working equipment capability. Vacuum arc 90 BUCKMAN

76.2 mm DIAMETER PRESSED r!<\FNIUM POWDER 2ND ADDITION ELECTRODE 1ST ELECTRON ELECTRON BEAM MELT

102 mm DIAMETER TANTALUM- CONSUMABLE 102 mm DIAMETER INGOT ELECTRODE • VACUUM CONSUMABLc ARC MELT

152 Tim DIAMETER CONSUMABLE 650 ELECTRODE VACUUM CONSUMABLE <=• ARC MELT

254 mm DIAMETER x 1.27 m LENGTH INGOT 2400 lb

Fig. 4 Typical ingot melting sequence for producing T-lll (Ref, 14).

melting of niobium and tantalum is accomplished Primary Working 3 at pressure (1 X 10 torr). Primary breakdown of the ingot microstructure Aschoff17 estimated that the melting capacity of is accomplished either by direct forging or extru- the refractory metals industry (circa 1961) was sion and depends on the practices developed and approximately 5.5 X 106 kg'y (12 X 106 lb/y). used by a particular vendor. Nb-lZr ingot can be The melting capacity could have easily been tripled worked by forging or extrusion in the temperature by improving the vacuum capabilities of the then range of 800 to 1000°C, but Ta-lOW, T-lll and existing reactive metal arc melting furnaces, which ASTAR 811-C are normally heated to ~1300°C for generally operate at 50 to 500 fim pressure. This is forging or extrusion. To prevent interaction of the just- as true today as then. An estimated 1810 kg billet with the heating medium, billets are pro- (4000 lb) of refractory metal alloy is required for tected by an evacuated mild steel hermetically one space power system; therefore the ability of sealed or purged cladding.14 Also hafnium silicide, the industry to supply niobium and tantalum alloy molybdenum, or tungsten coating can be applied to ingots will not be taxed by the SP-100 demands. protect the billet during heating for forging and/or REVSEW OF TANTALUM AND HSOSUM ALLOY PRODUCTION 91

TABLE 3 Specifications for Niobium and Tantalum Alloys

Form

Alloy Ingot Sbeet Strip Plate Bar Rod Wire Tubing Specifications

Nb ASTM B-391-78 ASTM B-°92-80 ASTM B-393-78 Nb-lZr ASTM B-394-70 Nb-lOHMTi ASTM B-652-79, AoTM B-654-79, ASTM B-655-79, AMS-7852-80, AMS-7857-80 Nb-10W-2.5Zr AMS-7851A-67. AMS-7855A-67 Ta ASTM-B-364-82, ASTM B-365-82 Ta-lOW ASTM-K-708-82, AMS-7847A-76, AMS-7848A-75 Ta-HW-2Hf (T-lll) B50YA325-S1"

'General Electric specification number for "Seamless Tubing and Pipe," April 21, 1971 (Ref. 13 Appendix C, p. 281). extrusion.18"19 Before proceeding to secondary with niobium or tantalum foil. Currently, refrac- working operations, billet conditioning to remove tory metal producers have vacuum annealing fur- contaminated materials and/or defects is accom- naces operating at pressures of —1 X 10 ~5 torr at plished by mechanical and chemical means. This is temperatures up to 1600°C with a size limitation of an extremely important step to assure maximum -0.25 m wide X 3.05 m long (10 in. X 120 in.). yield of high quality material in subsequent work- This area of technology is similar to the capability ing operations. established in the mid-sixties.

Secondary Working Specifications Secondary working operations are generally Currently available ASTM and AMS specifica- carried out at room temperatures to ~260°C. tions for niobium and tantalum alloys are listed in Working can be accomplished readily in this tem- Tatle 3. During the 60s, us jr-generated specifica- perature range with minimal concern for oxidation tions were used to procure niobium and tantalum of the billet. The Nb-lZr, T-lll, and ASTAR 811C 26 alloys for the space power program. These specifi- can be readily worked at room temperature to all cations, generated and used by GE in the 1965 time the common mill shapes. period under the NASA program, were generally more stringent than the commercial specifications Heat Treatment in use at that time. As an example, Table 4 com- Proper annealing conditions are absolutely pares the limits for impurities found in the ASTM essential in order to obtain high quality niobium and GE specifications to the actual chemical analy- and tantalum alloy product. In-process and final sis of as-melted ingot product. The purity of the annealing should be accomplished at pressures of T-lll ingot material was maintained well within <1 X 10~5 torr. Cleanliness of the product prior specification. to annealing is as important as the vacuum tight- A similar summary for the Nb-lZr alloy is ness of the annealing system. The interaction of shown in Table 5. Although the interstitial levels niobium and tantalum alloys with oxygen at low 20 25 are well within the limits established by the W64 pressure is well documented. " Steps to minimize revision of the ASTM specification for reactor this interaction include wrapping the work piece grade material, the latr 't revision (1978) is not 92 BUCKMAN

TABLE 4 Chemical Analysis Results of T-lll Ingot Produced During 1960 to 1970 Specification limits Element Current Prior GE» Experience!

ASTM-B-364-82 ASTM B-364-70 C 200 300 50 22 ± 17 0 150 300 100 38 ± 16 N 100 150 50 18 ± 5 H 15 100 10 5 ± 3 Ni 100 200 50 <60 Co 100 _ 50 <60 Fe 100 200 50 <60 Cu 100 _ _ 20 Si 50 200 - <60

•D. W. Miketta and R. G. Frank, SPPS-22-R1, Tantalum Alloy T-lll Specification, Dec. 23,1964, Contract NASA-2547 (Ref. 26). fF^ m time period 1960 to 1970, data from commercial size ingot {6Vi- to 10-in. dia), average of 13 different heats for interstitials, one heat from Ref. 14 for metallics.

TABLE 5 Chemical Analysis Results of Nb-lZr Ingot Produced During 1960 to 1970 Specification limits (WPPM) Element Current Prior Experience!

ASTM B-391-78* ASTM B-391-64* C 100 100 39 + 7 0 150 300 154 ± 82 N 100 300 50 ± 25 H 10 20 3 ± 1 Ni 50 200 <20 Co 20 30 Not reported Fe 50 500 <50 Cu 40 100 Not reported Si 50 300 <50 ^Reactor grade. fFor nine different heats for interstitials, metallic^ from Ref. 2 typical for E.B. melted Nb ingot

met, specifically for oxygen. It will be necessary to ditioning during processing, contamination pickup work with the vendors in establishing acceptable in the final product can be expected. Chemical and realistic specifications that will allow material analysis of niobium and tantalum alloy tubing, to be produced on a commercial basis. The specifi- showing interstitial content variations from ingot cation limits should be consistent with the require- through final product, is shown in Table 6. Pro- ments of the end item hardware service. cessing down to 9.5 mm dia X 1.57 mm wall (0.375 in. X 0.002 in.) produced very little change The specification limits discussed were for the in interstitial content. However, significant pickup as-melted ingot and thus represent the lowest of carbon and oxygen was observed after process- expected level of impurities. Processing to final ing to 0.38 mm (0.015 in.) wall. Significantly lower mill form requires exposure to elevated tempera- pressures during annealing would be required to tures, lubricants, etc., and thus, depending on final prevent this contamination. Current commercial form and degree of the chemical and chemical con- annealing furnaces do not yet have this capability. REVHEV OF TANTALUM AND NIOBIUM ALLOY PRODUCTION 93

TABLE 6 Interstitial Pickup During Processing" of B-66 (Cb-5Mo-5V-lZr) and T-lll (Ta-8W-2Hf)

Analysis, WPPM

B-66 T-lll Product form O N O N

Ingot-203 mm (8 in.) dia 40 95 70 10 25 20 (E.B. + vac. arc melt) Tube extrusion billet, 35 90 60 15 80 30 102 mm (4 in.) O.D.* 9.5 mm (0.375 in ) O.D. X 35 110 60 15 30 20 1.57 mm (0.062 in.) wallf 6.35 mm (0..250 in.) O.D. X 65 430 80 30 125 20 0.38 mm (0.015 in.) wall*

*As-extruded condition (B-66 extruded 1600°C) and annealed 1 h @ 1600°C. fFive in-process anneals @ <5 X 105 torr, wrapped in Ta foil; B-66 @ 1040 to 12CO°C for 10 min to 1 h; T-lll @ 1500°C for 10 min to 1 h. JSeven in-process anneals @ <5 X 10^ torr, wrapped in Ta foil; B-66 @ 1040 to 1200°C for 10 min to 1 h; T-lll @ 1500°C for 10 min to 1 h.

Supply Cautions There has been no domestic mining of niobium Some of the following caveats are worth noting or tantalum ores since 1959 (Ref. 27). Niobium ore and were a part of the Earning experience during is imported from Brazil, Canada, and Thailand; the 1960s. tantalum ore is imported from Thailand, Malaysia, • The production of high quality niobium and Australia, Canada, South America, and South tantalum alloy forms for space nuclear power Africa. The United States has approximately applications will require development of 360 X 106 kg (800 X 106 lb) of niobium resources 6 6 stringent process control and materials and 1.5 X 10 kg (3.4 X 10 lb) of tantalum specifications. Although the situation today is resources in identified deposits which were consid- 27 29 much improved over the 1960s, there is still ered uneconomic at 1981 prices. " Current price no totally integrated refractory metal pro- estimated for tantalum ingot is about $240/kg ducer who can perform all the necessary fab- ($110/lb), and for niobium ingot it is about $77/kg 18 rication operations. The number of knowl- ($35/lb). The formation of an organization named edgeable persons intimately familiar with the the Tantalum Information Center in Brussels has subtleties of refractory metal production has helped to stabilize the price of tantalum which significantly decreased over the past 15 years experienced instability during the late 1970s (Refs. due to lack of activity in the space nuclear 28 and 29). Between 1978 and 1980, the price of power field. tantalum oxide more than tripled in price.29 • Niobium and tantalum alloys, particularly those containing zirconium or hafnium, are Producers extremely sensitive to contamination by iron, copper, nickel, and cobalt. Control must There are four major suppliers of tantalum and therefore be exercised to assure that this does niobium alloys in the United States. Three of the not occur. For example, ingots were vacuum firms are completely integrated from raw material arc melted in furnaces normally used to melt processing through primary fabrication of the 28 29 alloy steels and superalloys. Material was niobium and tantalum ingot. " No single com- annealed in vacuum furnaces used for pro- pany is totally integrated through secondary fabri- cessing other nickel base materials and/or for cation. These four suppliers, however, do have dedi- brazing. The end result was defects and cated facilities for melting, primary and secondary failures attributed to irace amounts of con- working, and annealing of refractory metal alloys. tamination by these critical elements. 94 BlICKMAN

• Particular attention must be given to the in Refractory Alloys for Space Power Systems, mechanical and chemical condition of ingots NASA-SP-245, June 26,1969. 4. Anon., Columbium and Tantalum Alloys, Technical Informa- and billets, before working and or heat treat- tion, Teledyne Wah Chang Corporation. ment to assure removal of defects or contam- 5. Anon., Columbium by KBI, Bulletin 313-PD1, Kawecki- inants such as lubricants from prior working Beryllco, Reading, PA. operations. 6. Anon., Fansteel Production Bulletin, Fansteel Metals, North o Vacuum annealing must be accomplished in Chicago, EL. 7. Anon., Looking into Tantalum, NRC, Newton, MA. furnaces dedicated solely to refractory metal 8. R. W. Buckman, Jr., and R. C. Goodspeed, Considerations alloys. in the Development of Tantalum Base Alloys, Metallurgy • Specifications will reyuire overview by knowl- and Technology, pp. 373-394, I. Machlin, R. T. Begley. and edgeable persons to assure compliance with E. D. Weisert (Eds.), Plenum Press, 1968. 9. J. R. Lane and G. Mervin Ault, The Refractory Metals nv rial and process control requirements. Sheet Rolling Program, Metals Eng. Quart, Vol. 5, No. 3, August 1965. 10. R. L. Ammon, A. L. Field, A. J. Lewis, and L. S. Richard- DEVELOPMENTS NEEDED TO son, Research and Development of Tantalum and Tungsten SUPPORT SP-100 Base Alloys, Final Report, NOw-58-852C, May 26,1961. 11. R. L. Ammon and R. T. Begley, Pilot Production and Evaluation of Tantalum Alloy Sheet, Summary Phase Report No major developments are necessary in the Part 1, NOw-62-06562. June 1963. areas of nicbium and tantalum alloy production 12. R. W. Buckman, Jr., and R. T. Begley, Development of High and fabrication. Some consideration should be Strength Tantalum Base Alloys, in Recent Advances in given, however, to upgrading production capabili- Refractory Alloys for Space Power Systems, pp. 19-38, ties and enhancing material quality control. Specif NASA-SP-245, June 26,1969. 13. R. W. Harrison, E. E. Hoffman, and J. P. Smith, T-lll ically, the following items should be considered: Rankine System Corrosion Test Loop, NASA-CR-134816, June 1975. • Increasing vacuum annealing capability and 14. D. R. Stoner and R. W. Buckman, Jr., Development of large imposing more stringent pressure limits. Diameter T-lll (Tar8W-2Hf) TvJAng, Final Report, NASA- Existing capability limits product size to CR-72869, December 1970. 1.5 m X 3.05 m (60 in. X 120 in.) at pres- 15. W. R. Young, Fabrication of T-lll Test Loop System, in sures of ~1 X 10~3 Pa (1(T5 torr) at tem- Recent Advances in Refractory Alloys for Space Power Sys- tems, pp. 165-220, NASA-SP-245, June 2fi, 1969. peratures to 1600°C. Consideration should be 16. A. W. Buckman, Jr., and J. L. Godshall, Final Technics given to limiting annealing pressures to Report on Development jf Hig'1 Strength Columbium ant 5 7 <1 X 10~ Pa (1 X 10~ torr) for thin- Tantalum Alloy luting, WANL-PR-(N)-004, Contraci walled products. AT(30-l)3108, May 1,1964. 17. W. A. Aschoff, Current Melting Practices for Refractor} • Development of acceptable material and pro- Metals Columbium, Tantalum, Molybdenum and Tungsten cess control specifications written with due High Temperature Materials, II, G. M. Ault, W. F. Barclay consideration to the end use, i.e., space power and H. P. Munger (Eds.), Interscience Publishers, 1961. applications. 18. R. Marsh, Teledyne Wah Chang Corp., private communica tion. 19. R. MacDonald, OP.NL, private communication. ACKNOWLEDGMENT 20. H. Inouye, The Contamination of Refractory Metals ii Vacuum Below 10"* Tor" Refractory Metals and Alloys III Applied Aspects, pp. 871-883, Robert I. Jaffee (Ed.) The critical review and helpful comments pro- December 1963. vided by co-workers R. L. Ammon, Chris Bagnall, 21. R. W. Buckman, Jr., Operation of Ultra High Vacuun and Dr. Stu Shiels during the preparation of this Creep Testing Laboratory, in Transactions of the Vacuum paper are gratefully accepted and deeply appreci- Metallurgy Conference, New York City, June 1966, pp. 25-31 ated. M. A. Orehoski and R. F. Bunshah (Eds.), 1967. 22. R. F. Bunshah, Contamination of Reactive Metals i; Vacuum Heat Treating at Various Vacuum Levels, i: Transactions of the Internationcl Vacuum Metallurgy Confei ence, Beverly Hills, CA, June 1968, pp. 611-619, R. B. Bai REFERENCES row and A. Simkovich (Eds.), 1968. 23. E Fromm, Effect of Inert Gas Atmosphere on Refractor; Metal Losses During Evaporation, No. 499233 AIX-1C 1. G. Mervin Ault, A Decade of Progress in Refractory Metals, 444471, EDB-79: 073333 (in Russian), International Sympo Gillette Memorial Lecture, ASTM, Philadelphia, PA, 1965. sium on Interaction of Metals with Gases and Carbor 2. D. C. Goldberg, G Dicker, and S. A. 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Research Center, Report No. BM-RI-6521, p. 16, February Snyder, W. B., Jr., Effects of Orientation on the Rolling and 1964. Recrystallization Behavior of Tantalum Single Crystals, No. Ortner, H. M., and D. Hirschfeld, Round-Robin Results on the 230518 INS-77: 008046, EDB-77: 088389, Thesis (Ph.D.), Univ. Determination of Hydrogen in Tantalum and Niobium, No. of Tenn., Knoxville, pp. 228 (1976). 855988 AIX-12: 618055, EDB-82: 030827, Metallwerk Plansee , Effects of Orientation on the Rolling and RecrystaUization A.G., Reutte/Tyro' (Austria); Krupp (Friedr.) Q.m.b.H., Behavior of Tantalum Single Crystals, No. 217552 ERA-02: Essen (German), Forschungsinstitut, Mikrochirn. Ada 025020, INS-77: 006311, EDB-77: 055354, Thesis, Report No. (Austria) 2(5-6): 475-490,1980. Y-2062, Oak Ridge Y-12 Plant, Oak Ridge, TN, p. 118, De<\ Palatnik, L. S., V. R. Kopach, I. B. Skatkov, Yu. L. Pozdeev, 1976. and I. V. Mancheno, Method of Investigation of Interface Topology in Laminated Ta-Nb-Ta Sheet Metal 2, No. 302112 Stephan, H., Present Position of Electron Beam Melting and AIX-08: 337258, EDB-77: 140544 (in Russian), Zavod. Lab., Casting Technology, No. 355730 ERA-03: 021837, EDB-78: (USSR), 42(10): 1211-1212 (1976). 044358, Elektrmvaerme Int. Ed. B (Germany), 35(3): 150-153, Palme, R., The Importance of Vacuum Technology in the Pro- June 1977. duction and Processing of Refractory Metals, M03-26196 Tantalum Ingots and Flat Mill Products, No. 909001 INS-82: (German), Vahuum Tech., 13(8): 229-232, December 1964. 008040, EDB-82: 083847, Anna. Book ASTM Stand. (U.S.), 45: Perryman, J. T., The forming of a Cb-lOW-lOTa Alloy, in 160-163 (1981). Proceedings c/ Conference on Refractory Metals and AUoys, Tantalum Rod and Wire, No. 909002 INS-82: 008011, EDB-82: 3d, Los Angeles, 1963, Refractory Metals and AUoys III: 083848, Annu. Book ASTM Stand. (U.S.), 45: 164-167 (1981). Applied Aspects, pp. 67-72, Robert I. Jaffee (Ed.), 1966. Torti, M. L., Arc-Melting Procedures for Refractory Metals, in Peters, J. Irwin, How to Work Tantalum, M08-14710, Amer. Transactions of the Vacuum Metallurgy Conference, New York Mack/Metalwork Manuf., 108(23): 105-107, November 1964. University, June 1959, pp. 1-4, Rointan F. Bunshah (Ed.), Pirch, I. I., A. I. Pugin, and A. A. Erokhin, Enthalpy and Aver- 1960. age Metal Bath Temperature During Plasma-Arc Melting, Physical Properties and Fabrication Techniques for the No. 873270 AIX-12: 629090, EDB-82: 048111 (in Russian), Fix. t Tantalum 10% Tungsten Alloy, in Proceedings of High Khim. Obrab. Mater. (USSR), 1:150-153, Jan.-Feb. 1980. Temperature Materials II, Metallurgical Society O- •ferences, Precision Forging, 310-F, Metal Treatment and Drop Forging, 27: 18: 161-169, Cleveland, OH, April 26-27, 1961, G. M. Ault, 363-365, September 1960. W. F. Barclay, and H. P. Munger (Eds.), 1963. Quadt, R. A., Precision Cold Extrusion of Metals, 164, 249-G, Vandermeer, R. A., and W. B. Snyder, Jr., Recovery and Recrys- Met Prog., 77: 121-124, May 1960. tallization in Rolled Tantalum Single Crystals, No. 539029 Randall, R. N., The Planetary Ball Swager—New Tool for thr ERA-04: 053430, EDB-79: 118495, MetalL Trans., A (U.S.), Production of Refractory Metal Tubing, p. 9, M07-48339, 10(8): 10-1-1044, August 1979. ASTM, Dearborn, MI, Tech. Paper No. SP66-46 (1966). Rostoker, W., Conversion of Refractory Metals, Armour Vereschchagin, L. F., and N. S. Fateeva, Multing Temperatures Research Foundation, Chicago, IL, in Refractory Metals and of Refractory Metals at High Pressures, No. 667442 AIX-il: AUoys II, Proceedings of a Technical Conference. 17: 379-394, 546535, EDB-S0: 106968, High Temp.-High Pressures (United Chicago, IL, April 12-13, 1962, M. Semch;shen and Kingdom), 9(6): 619-628, 1977. I. Perlmutter (Eds.), 1963. Von Ardenne, M. and S. Schiller, Advances in Vacuum Metal- Schmidt, F. F., D. J. Maykuth, and H. R. Ogden, The Effects of lurgy, in 1965 Transactions of the Third International Vacuum Heal-Treating and Testing Environments on the Properties of Congress, Invited Pavers, 1: 137-153, Long Island City, New Refractory Metals, p. 28, M17-13014, Defense Metals Infor- York, Pergamon Press, Inc., 1966. mation Center, Battelle Memorial Institute Memos & Wahlster, M., and F. Sperner, Vacuum Technology and Use of Reports, Report 205, Aug. 20,1964. Important Reactive Metals, No. 328033 AIX-08: 340149, Schmitt, B. F., H. U. Fusban, G. Kraft (Eds.), Photon Activa- EDB-78: 016660 (in German), MetalL (Germany), 31(7): tion Analysis of Carbon, Nitrogen and Oxygen in Refractory 781-784, July 1977. Metals, No. 903160 EDB-82: 078005 (in German), in International Conference tm the Analysis of Non-Metals in Weisert, E. D., T'.ie Present Status of Refractory Metals and Metals, Berlin, Germany, June 1980. Alloys, Met. Eng. Q. 1S9-A (American Society for Metals), 2. Schumacher, J. F., and W. A. Fischer. Levitation Melting 3-15, February 1962. Installation with a Large Working Scope as Regards Load Wong, J., S. S. Christopher, and S. A. Worcester, Jr., Capacity and Temperature, and also Melting of Iron, Consolidation of Refractory Metals, Wah Chang Corporation, Niobium, Molybdenum and Tantalum, No. 468651 AIX-09: Albany, OR, in Refractory Metals and AUoys II, Proceedings 41<*141, EDB-79: 042750 (in German), Stahl Eisen (Germany), of a Technical Conference, 17: 351-377, Chicago, IL, April S3(8): 410-413, April 1978. 12-13,1962, M. Semchyshen and I. Perlmutter (Eds.), 1963. Schwirtlich, I. A., and H. Schultz, Quenching Experiments on Zedler, E., Dieter Knaut, and Karl Schlaulitz, On the Properties Niobium, No. 773154 AIX-12: 591271, EDB-81: 081416, of Electron Beam Melted Metals, pp. 187-197, 92556C (Ger- Stuttgart Univ., Germany, Inst. fuer Theoretische und man), Freiberger Forschungshefte, No. B76, 1963. Angewandte Physik; Max-Planck-Institut fuer Metall- forschung, Philos. Mag., (Part) A (United Kingdom), 42(5): Zlatin, N., M. Field, and J. Gould, Machining Refractory Metals, 613-620, November 1980. in Proceedings of Conference on Refractory Metals and AUoys, Shaping Space Age Metals in Argon, 179-F, Am. Mach., 104: 3d, Los Angeles, 1963, Refractory Metals and AUoys III: 120-121, June 1960. Applied Aspects, pp. 483-496, Robert I. Jaffee (Ed.), 1966. Smith, H. R., Jr., Design and Performance Parameters of Com- Zotov, Yu. P., N. N. Kroshkina, and V. E. Bukharin, Determina- marcial and Laboratory-Scale Electron Beam Furnaces, in tion of Composition and Melting Points of Eutectics in Transactions of the Vacuum Metallurgy Conference, New York Refractory Metal (Alloy)-Silicon Systems, No. 749543 AIX- University, June 1960, pp. 255-265, Rointan F. Bunshah (Ed.), 12: 584460, EDB-81: 057801 (in Russian), Izv. Akad. Nauk 1961. SSSR. Neorg. Mater. (USSR), 16(5): 842-844, May 1980. Processing and Production of Molybdenum and Tungsten Alloys

W. C. HageU J. A. Shields, Jr., and S. M. Tuominen Climax Molybdenum Company of Michigan

13 13 16 INTRODUCTION products with K and Si or Al, K, and Si, ' respectively, yields products with higher recrystal- Alloy designers have relied on strain hardening, Ii7-\tion temperatures and higher creep strength than undoped material due to potassium bubble solution hardening, and strengthening by dispersed 23 phases to achieve improved strength levels, pinning of grain boundaries. Doping is used in increased low temperature ductility, and increased the lamp industry to improve the sag resistance of recrystallization temperatures in molybdenum and filaments and for sheet products to be used at high tungsten alloys. Because molybdenum and tung- temperatures. Thoriated tungsten alloys, which sten are body-cefltered-cubic metals which do have higher electroemhaivities, recrystalliza- not undergo allotropic transformations, the tion temperatures, elevated temperature tensile strengths,24 and lower creep rates at elevated strengthening that can accompany such transfor- 25 mations is unavailable for these metals. Table 1 temperatures than unalloyed tungsten, are also summarizes the strengthening mechanisms manufactured for lamp and electron emission exploited by a number of commercial and promis- applications. ing experimental alloys.1 ~20>22>34 Strain hardening is The high melting points of Mo and W make not listed explicitly, though it is nearly universally them candidates for applications involving high employed to improve the low-temperature ductility temperatures. Alloying is necessary to decrease the of Mo and W alloys. Solution hardening is ductile-brittle transition temperature (DBTT), employed in Li« Mo-W1-2 Mo-Re,3 and W-Re15"16 increase creep strength, and increase recrystalliza- systems. However, the primary reason for alloying tion temperatures. The effectiveness of strengthen- Mo with 30% W is to obtain corrosion resistance ing and ductilizing mechanisms is highly depen- to liquid zinc equivalent to that of unalloyed dent on the production and processing techniques tungsten.2 Rhenium additions to Mo and W used to prepare the alloys. For example, the pri- increase the elevated temperatme yield strength,20 mary difference between alloys produced by but the primary reason for alloying with rhenium vacuum-arc casting (VAC) and powder-metallurgy is to take advantage of the well-known rhenium- (PM) techniques is the oxygen content. According ductilizing effect21 which leads to improved low- to ASTM Standard B387-74 the maximum allow- "ernperature ductility. The family of Mo alloys able oxygen content in VAC-'TZM is 30 ppm while -:riOTrn as TZM,5-6 TZC,7"10 MHC,7-8-11-12 and the W- that of PM-TZM may exceed 300 ppm. The higher Hi'C alloys all employ hardening by dispersed oxygen levels in PM-TZM are attributable to the •»a/:tive metai carbides. The tungsten-rhenium- tendency of reactive metals in the PM product to '..if.'i'jm carbide and molybdenum-rhenium- form oxides du-ing sintering.26 These oxides can be ".A/-;-J.T. ".arbide alloys noted in Table 1 (Refs. 11, expected to decrease the volume fraction of Ti and '..'-. '.' J ; are based on the same dispersed carbide Zr available to form carbides. Vacuum-arc-cast or s!v»naf.;--:fi-ning mechanism but make use of a electron-beam-melted Mo-Re alloys have lower fimu." addition to obtain ductile behavior at low DBTTs than powder-metallurgy Mo-Re alloys, with r.r.r»a. Doping powder metallurgy Mo or W the difference in DBTT believed to be due to lower

98 PROCESSING AND PRODUCTION OF Mo AND W ALLOYS 99

TABLE 1 Strengthening Mechanisms in Mo end W Alloys

Alloy and production techniques Mechanism employed Molybdenum-base ulloys Tungsten-base alloyB

Solution hardening Mo-30WU(VAC, PM) W-3 to 50Re164718 (PM, VAC) Mo-3 to 50Re (VAC. PM)14 Intrinsic dispeision TZM: Mo-0.5Ti-0.lZr-0.01 to W-HfC'-^fPM, VAC) hardening CMC5* (VAC, PM) TZC: Mo-1.0 to 1.5Ti-0.20 to 0.35Zr-0.0fi to 0.20C" " (PM, VAC) MHC: Mo-0.5 to 2.0Hf-0.04 to 0 20C'*1IJZ(PM, VAC) 13 16 Extrinsic dispersion Doped Mo with K and Si (PM) W-ThO2 (PM) hardening Doped W with Al, K, Si13aii(PM) Combinations of the Mo-Re-Hf-C14 (VAC) above Mo-30to45W-Hf-C16(PM) W-Re-Hf-C1I16(VAC, PM) interstitial contents in the cast alloys.42 This dif- standard compositions have not been developed for ference in DBTT for Mo-Re alloys is very impor- these alloys. tant because of the need to handle components at Producers of commercial W-base alloys, a3 room or lower temperatures during fabrication. given in the 1982 Thamas Register, are listed Before discussing future needs, including alphabetically in Table 3. Commercial alloys are detailed discussions of alloy systems and potential prepared using only powder metallurgy techniques improvements in properties, the technological ana are available only in thin sections which means to produce and process Mo and W alloys include rod, wire, sheet, and foil. must be summarized because for many Mo and W alloy systems the mechanical properties can be optimized only by thermomechanical processing PRODUCTION CAPABILITIES requiring production and processing capabilities that are not widely available. First, Lhe producers For the purposes of this paper, we will limit our of commercial Mo and W alloys will be presented consideration of production and processing along with currently available product forms. capabilities to those facilities located in the United Second, currently disclosed standard capabilities of States. The production capabilities of the commer- producers and processors in the United States will cial facilities in the United States listed in Table 4 be presented. The capabilities of firms th*t pro- are of primary importance because th^.re are duce special products for internal use, such as Gen- currently maximum standard sizes and tempera- eral Electric's production of wire for lamps, are not tures for many of the processes involved in refrac- included because capabilities of such firms are mn tory alloy production. The first step in production generally disclosed and the facilities are not avail- of Mo and W alloys is reduction of oxides to form able for other uses. metal powders. Both metal powders are produced by hydrogen reduction of oxides at elevated temperatures. Molybdenum is usually reduced from PRODUCERS OF Mo AND MoO3 or from ammonium dimolybdate which forms W ALLOYS MOO3 when heated for the first stage of reduction to MoO2 (Ref. 27). Tungsten powder is reduced The producers of commercial Mo alloys, as from ammonium paratungstate which is also given in the 1982 Thomas Register, are listed alpha- heated in air or hydrogen to form WO3 which is 17 betically in Table 2. Teledyne Wah-Chang Hunts- reduced to WO2 in hydrogen. Both MoO2 and WO2 ville is not included in Table 2 because they pro- are reduced to metal powders using higher tem- 17 27 duce only unalloyed molybdenum products in a peratures " than are used for the formation of wide variety of forms. The alloys which are avail- the dioxides. The commercial U. S. suppliers of Mo able upon special request, PM-TZC and PM-MHC, powder, ASMC, GTE-Sylvania, and Teledyne Wah do not have fixed compositions in Table 1 because Chang Huntsviile, all control the reduction param- 100 HAGEL, SHIELDS, AND TUOMtNEN

TABLE 2 Producers of Commercial Mo Alloys

Production Source technique Alloys Available forms

ASMC VAC TZM Wide variety* Mo-30W Billets, forgings, bar, rod PM TZM Billets, forgings and bar TZC Available upon special request MHC Available upon special request H. Cross, Inc. PM Mo-47Re Thin sectionst and tubing GTE-Sylvania PM TZM Wide variety, and pressed and sintered shapes Mo-30W Billets, forgings, bar, and rod MHC Available upon special request Metallwerk Plansee PM TZM Wide variety TZC Available upon special request MHC Available upon special request Mo-5 to 50Re Thin sections Doped Mo Thin sections Philips Elmet PM Doped Mo Thin sections Re Alloys, Inc. PM Mo-10 to 48Re Thin sections and tubing Ultramet CVD Mo-Re Any shape Mo-W Any shape

"Includes billets, forgings, bar, rod, plate, sheet, foil, and custom fabr'cated shapes but not thin wall tubing. •("Includes rod, wire, sheet, and foil.

TABLE 3 Producers of Commercial W Alloys

Production Source technique Alloys Available forms

H. Cross, Inc. PM W-26Re Thin sections* Metallwerk PM W-3 to 50Re Thin sections Plansee W-l to 4 ThO2 Thin sections Doped W Thin sections Philips Elmet PM W-50Mo Rod and wire W-l to 3 ThOz Rod and wire Doped W Wire Re Alloys, Inc. PM W-5 to 26Re Thin sections W-5 to 25Re Thin sections plus up to 2 ThOz Ultramet CVD W-Re Any shape

'Includes rod, wire, sheet, and foil.

eters for Mo to yield powders with controlled parti- capabilities. Information on in-house isostatic cle sizes, compactibility, and bulk density. The press sizes are available from the producers (see commercial suppliers of W powders, GTE-Sylvania Table 5). In addition, larger isostatic presses out- and Teledyne Wah Chang Huntsville, also reduce side the refractory metals indastry are available W powders with controlled particle sizes, compact- for toll usage by refractory alloy producers. All ibility, and bulk density. ASMC is also capable of producers of PM alloys have elevated temperature co-reducing other powders along with molybdenum sintering capability in hydrogen atmosphere fur- powder to form alloyed powders. Philips Elmet naces. GTE-Sylvania can produce as-sintered bil- reduces Mo and W powders but does not currently lets up to 940 mm (37 in.) diameter.15 The only sell the powders. tungsten shape sintered by ASMC is sheet bars for All domestic producers of PM Mo and W alloys rolling of flat products at ASMC's Cleveland flat or unalloyed products have isostatic compaction rolling facility. Teledyne Wah-Chang Albany has PROCESSING AND PRODUCTION OF Mo AND W ALLOYS 101

TABLE 4 Production Capabilities of Commercial V. S. Facilities

Production process Isostatic Vacuum Electron Chemical Powder compaction arc beam vapor Source reduction from oxides Sintering easting melting deposition

ASMC Mo Mo, W Mo, W Mo _ _ H. Cross, Inc. - Mo, W Mo, W - - - GNB Corp. - - - - Mo, W - GTE-Sylvania Mo, W Mo, W Mo,W - - - Philips Elmet Mo, W Mo, W Mo, W - - - Re Alloys, Inc. _ Mo, W Mo, W _ _ _ Teledyne Wah-Chang - Mo, W Mo, W - Mo, W Albany Teledyne Wah-Chang Mo, W Mo, W Mo, W - - Huntsville Ultramet - - - - - Mo, W

TABLE 5 Production and Processing Sources

Firm Location Telephone

AMAX Specialty Metals Corp. (ASMC) Parsippany, NJ (201) 884-2900 H. Cross, Inc. Weehawken, NJ (201) 863-1134 GNB Corp. Hayward, CA (415) 537-4722 GTE-Sylvania Towanda, PA (717) 265-2121 Metallwerk Plansee Holliston, MA (617) 429-6801 ^Schwarzkopf Development Corp.) Philips Elmet Lewiston, ML (207) 784-3591 Rhenium Alloys, Inc. Elyria, OH (216) 365-7388 Teledyne Wah-Chang Albany Albany, OR (503) 926-4211 Teledyne Wah-Chang Huntsville Huntsville.. AL (205) 837-1311 Thermo Electron Woburn, MA (517) 933-7610 Ultramet Pacoima, CA (213) 899-0236 furnaces for sintering of tungsten up to 2300°C Solid solution strengthened molybdenum and using a hydrogen atmosphere or vacuum. These tungsten alloys can be produced from elemental furnaces are available for toll usage. The notable powders directly tc a wide variety of shapes by feature of these furnaces is that they do not con- chemical vapor deposition. This technique is used tain large masses of refractory insulating brick for producing tungsten tubing. Oak Ridge National and thus will cool relatively rapidly. Laboratory (ORNL) experience is that duplex The two melting techniques available for pro- extrusion is the best method for producing tung- duction of Mo alloys are vacuum-arc casting (VAC) sten alloy tubing. and electron-beam (EB) melting. Commercial VAC Mo alloys are produced using press-sinter-melt (PSM) equipment which has been described in PROCESSING CAPABILITIES detail by Timmons and Yingling28; this equipment is suitable for casting some tungsten alloys. The capabilities of U. S. processors are sum- Electron-beam melting units can be used to pro- marized in Table 6. Commercial VAC molybde- duce Mo or V/ alloy ingots, but the electrodes must num alloys are currently always processed initially first be prepared using powder-metallurgy by extrusion because the alternative technique that techniques. Double electron-beam-melted tungsten was used from 1946 to 1956, forging with a V- ingots have been reported to have lower impurity 17 shaped lower die block, resulted in lower yields due levels than arc-cast tungsten ingots. to cracking and oxidation.30 Laboratory studies are 102 HAQEL, SHELDS, AND TUOMNEN

TABLE 6 Proceailng Capabilities of Commercial U. & Facilities

Procea* Round Flat Wire Tube Source Extrusion Forging* roiling Swaging rolling drawing drawing

ASMC Mo Mo Mo Mo Mo,W _ _ H. Cross, Inc. - - Mo,W Mo,W Mo.W Mo GTE-Sylvania Mo, W Mo Mo, W Mo, W Mo, W - Philips Elmet - - Mo, W Mo Mo, W - Re A'.loys, Inc. _ _ Mo.W Mo,W Mo, W Mo Teledyne Wah-Chang - Mo Mo, W Mo, W Mo, W _ Huntsville Thermo Electron - Mo Mo, W - Mo Mo

•Forging services available to sources.

seeking ways to avoid the need for extrusion while thicknesses ranging from 2.5 to 0.05 mm (0.10 in. minimizing material losses. Extrusion press capa- to 2 mils). Lengths up to 7.6 m (25 ft) are avail- bilities, including liner sizes, loads, preheat tem- able except in very small diameters. peratures, and feasibility of preliminary laboratory The alloy producers and processors listed in extrusion studies, are available from ASMC (see Tables 2, 3, 4, and 6 are listed in Table 5 in Table 5). The in-house forging capabilities of U. S. alphabetical order along with locations and producers are suitable only for production of small telephone numbers as of June 1983. It should be forgings, but the producers have access to outside noted that these firms may be capable of extending forging facilities. their standard processing capabilities if the Molybdenum alloys, but not tungsten alloys, are customer is willing to pay additional costs, if the rolled as rounds. After rolling to diameters near economics of scale permits the purchase of addi- 25 mm (1 in.), round stock is swaged to smaller tional equipment, or if the customer is willing to diameters. All producers are capable of swaging pay on a best-effort basis. alloys to about 3.2 mm (V6 in.) diameter. Smaller diameters are obtained using wire drawing facilities. EFFECTS OF PRODUCTION AND Practically all producers have the capacity to PROCESSING ON MECHANICAL flat roll Mo and W alloys (Table 6): Philips Elmet PROPERTIES does not produce flat W alloy products. Standard rolling capabilities, including starting and final Processing of molybdenum and tungsten alloys thickness for molybdenum or tungsten, widths, and can critically influence the mechanical properties preheat temperatures of each producer, are availa- of the finished product. Both consolidation tech- ble from the sources in Table 5. The maximum niques and thermomechanical processing of consol- width of sheet for the processors listed is 610 mm idated products must be considered in any analysis (24 in.). Molybdenum and tungsten with of the effects of processing. thicknesses of less than 0.13 mm (5 mil) and For unalloyed molybdenum and tungsten, the 1.0 mm (40 mil), respectively, are available only in technique used to consolidate the material has lit- narrower widths. Minimum standard thicknesses tle influence on mechanical properties. The for molybdenum and tungsten aie 0.013 mm mechanical properties of vacuum-arc-cast material (0.5 mil) and 0.05 mm (2 mil), respectively. and powder-metallurgy material are quite close to Three firms listed in Table 6 are capable of one another. The two routes offer different advan- drawing Mo alloy tubing; none of these firms tages. VAC material is lower in interstitial impuri- currently draws tungsten tubing. Standard tubing ties and is normally processed initially by extru- sizes, including diameters, wall thicknesses, sion. Currently the maximum ingot size is limited lengths, and alloys that are normally drawn, are by available mold sizes. Powder-metallurgy tech- available from the sources in Table 5. Standard niques produce material which is generally less tubing outside diameters range from 32 to 0.05 mm pure, but which can be forged and rolled without (1.26 to 0.02 in.) with corresponding wall intermediate extrusion because of the fine grain PROCESSING AND PRODUCTION OF Mo AND W ALLOYS 103 size o? the sintered products. Powder metallurgy alloy are compared with those of powder metal- can also yield billets which are limited in size only lurgy TZM, TZC, and MHC. Also shown are proper- by the capacities of available compacting and ties for powder metallurgy Mo-^ORe alloy. All the sintering equipment. alloys are in the wrought or wrought and stress- Alloys which contain reactive metal additions relieved condition. The scatter bands shown in (and these comprise the majority of the alloys of Fig. 1 arise because data are included on materials having a variety of compositions and processing interest) result in a different situation. For these 0 materials, the VAC process is clearly advantageous historic . Both of these parameters strongly affect because of the low oxygen content resulting from the properties of molybdenum alloys. For example, very efficient carbon deoxidation during the vac- thin sheet materials, which are highly cold worked, uum melting process. Powder-metallurgy products can have much higher strength at low tempera- usually contain much higher levels of oxygen, tures than those given by the scatter bands. At resulting in microstructures containing reactive temperatures above -~800°C, these materials have metal oxides. These oxides are detrimental from properties which fall with the bands. The effects of two standpoints. The inclusions themselves can act processing on strength are par lcularly evident for as nuclei for fracture, and they remove the reactive the VAC MHC alloys. In a study of the effects of processing and composition on properties of MHC, metal from solution, making it unavailable for 29 solution strengthening or dispersion strengthening. Raffo measured properties for VAC-MHC which Thermomechanical treatment (TMT) also plays fell withii the PM material scatter band. However, an important role in developing good mechanical for processing which resulted in strain-activated properties in tungsten and molybdenum alloyt. The precipitation during processing, the alloy possessed properties whicn fell within the "optimum" MHC importance of deformation in improving the DBTT 11 of the materials was recognized early.35 Solution band. Klopp et al. at NASA-Lewi3 Research treatment of VAC29'31"32 and PM7 material has been Center also studied the effects of processing on the shown to improve the strength, creep resistance, tensile strength of VAC-MHC of various composi- and recrystallization behavior of alloys employing tions. The processing which yielded by far the carbide-forming elements. highest tensile strength at 1315°C included extrud- The techniques used to develop strength should ing at 2205°C and swaging at 1370°C. These probe specified with a mind to the final application. cess parameters, however, may not yield the same properties in materials larger than i-he smal1 rods Processing routes which rely on deformation to l develop strength are less resistant to long-term about 3 mm ( A in.) diameter processed by Klopp creep at higher strength levels than at lower et al. Another potential limitation is cnat the strengths.25'33 Strain-activated precipitation of strength levels shown in Fig. 1 for powder carbides during processing can produce outstanding metallurgy TZC and MHC were obtain^ uy forging after solution treating at 2200°C with rapid tunsile and short-term rupture properties, but the 7 resultant materials may be inferior to "lower cooling from the solution treatment temperature. strength" materials under conditions of low-stress The cooling rates obtained in the laboratory-size creep.29 Precise control of TMT is critical to forging blanks are unlikely to be obtainable in obtaining the appropriate mechanical properties in larger blanks, which may limit the size of products tungsten and molybdenum alloys and is likely to of these alloys with the indicated strength levels. be even more important in the future, as alloy design becomes more sophisticated. Specific exam- Consolidation technique (PM vs. VAC) makes ples of the influences of production and processing little difference in tensile strength for those on properties of experimental alloys will be materials not containing reactive elements (Mo, presented in the following sections and related to Mo-W, Mo-Re). Problems with processing of VAC TZC and MHC ingots have limited the availability the feasibility of commercial production and pro- 7 cessing. of these alloys as VAC material to small sizes. They are available upon special request as PM products. Our laboratory is currently working to STRENGTH AND DUCTILITY optimize the composition of MHC for VAC production. Figures 1 and 2 illustrate the effects of alloying The effect of alloying is shown in Fig. 1. on the ultimate tensile strength of several Rhenium provides a dramatic increase in the low- 3 5 7 29 35 37 17 19 38 41 molybdenum - ' - ' " and tungsten ' - " alloys, temperature strength of molybdenum, but the ben- respectively. In Fig. 1, the tensile strengths of efit is greatly reduced at high temperatures. (The vacuum-arc-cast molybdenum, TZM alloy and MHC alloy illustrated in the figure received 50% reduc- 104 HAGEL, SHELOS, AND TUOMNEN

TEMPERATURE,°F 1000 2000 3000 1500

-200

VAC-MHC (Optimal Processing) o a. 2 - 150 to 1000 o z tu tr. (- at in 111 ui - 100

i- LU ui 500 - < 2

Jo 500 1000 1500 2000 TEMPERATURE,°C

Fig. 1 Effect of temperature on the ultimate tensile strength of molybdscum and molybdenum-base alloys.

TEMPERATURE,°F 1000 2000 3000 4000 -i 1 r- —I—' T 200

150 z UJ CE PM W-Hf-C I-

:oo _J

PM W-2ThO2 (As-sintered)

500 1000 1500 2000 2500 TEMPERATURE,°C Fig. 2 Effect of temperature on the ultimate tensile strength of tungsten and tungsten-base alloys, All alloys tested in the wrought and stress-relieved condition, except as noted. PROCESSWG AND PRODUCTION OF Mo AND W ALLOYS 105

TEMPERATURE,1* 1000 2000 3000 600 TZM Mo-O.6Hf-O.5C Mo-35 Re --3T-111 500 - —aCb-752 W-3.9Re-0.41Hf-0.51C W-HfC 400

ID 300 J5 a

500 1000 1500 2000 TEMPERATURE,°C Fig. 3 Tensile strength/density ratio for several refractory metal alloys as a function of temperature.

tion during processing.) 'E * strength advantages tory alloys. All alloys were compared in the conferred by alloying with <.,tanium, zirconium, and stress-relieved condition. Altnough this is not the hafnium are smaller at low temperature but per- most common structural condition for such alloys sist to high temperatures. All the alloys app"oach as T-lll, which are normally used in the recrystal- one another in strength at the highest tempera- lized condition, comparisons between alloys are tures as recrystallization occurs, and precipitate simplified because all are in the sams microstruc- coarsening occurs in the carbide-strengthened tural condition. At temperatures below ~1200°C, alloys. Greater amounts of precipitate-forming ele- • TZM is preferred over a range of tungsten, tan- ments provide a larger strength increment than do talum, and niobium alloys. Above this temperature, lower levels (compare TZC and MHC with TZM). W-Hf-C alloys are superior to TZM. The high Similar observations can be made for the W strength of Mo-35Re drops fairly rapidly, and this alloys shown in Fig. 2, although fewer alloys are alloy is less desirable than TZM above —700°C. illustrated than for the Mo-base systems. Solution The Mo-0.6Hf-0.5C alloy investigated by Raffo is hardening is effective over a relatively wide tem- the strongest of all for temperatures below perature range, as illustrated by the curves for ~1600°C. Figures 4 and 5 illustrate the ductilities tungsten-15 molybdenum and tungsten-25 rhenium of several of the alloys shown in Fig. 3. The ductil- alloys. Carbide-strengthened alloys offer the ity of Mo and W alloys at temperatures below greatest improvement in strength over the widest ~1000°C is lower than the ductility of Ta and Nb temperature range. The addition of thoria24 to alloys but is still adequate for engineering applica- tungsten yields wrought and rtcrystallized prod- tion. Ductility is also important in determining an ucts with higher strength than pure tungsten over alloy's fabricabiiity and It3 resistance to fracture temperatures ranging from 800 to 2400° C. during use. Elevatod-temperature ductility is not a In applications where weight is important, the problem for any of the W or Mo alloys considered strength of the material is legs important than its here, either in the worked or recrystallized condi- strength/density ratio. Figure 3 illustrates the tion. However, there are significant differences in strength/density ratio for several different refrac- the DBTT behavior of various alloys. Figure 6 106 HAGEL. SHIELDS. AND TUOMNEN

TEMPERATURE ,°F 1000 2000 3000 TOO so- > T-111 (Stress Relievad) . Cb-752(Stress Relieved) 80 - 70 60 50 40 30 20 10 0 500 1000 1500 2000 TEMPERATURE,°C

Fig. 4 Tensile elongation of tantalum and niobium alloys AS a function of temperature.

TEMP£SATURC,°F 1000 2000 3000 100 90 35Re - 80 r^TZM 1st™.. o 70 » -.W-3.9Re-0.4lHf-0.51C As-Swaged 60 3 50 UJ 40 30 20 10

500 1000 1500 2000 TEMPERATURE,°C

Fig. 5 Tensile elongation of molybdenum and tungaten alloys as a faction of temperature.

illustrates the DBTT behavior of vacuum-arc-cast incre?~es as reactive metal levels are increased. I molybdenum35 following working and either stress is difficult to compare the work of different inves relief or recrystallization. The advantage to be tigators, because many different tests and transi gained by using stress-relieved materials is clearly tion criteria hn_ve been used to define DBTT. Com illustrated. The effect of alloying on the DBTT monly employed tests are IT bend, 4T bend, tensile depends upon the alloying element. Increasing the notched Charpy, and unnotched Charpy tests carbon level in any alloy increases the DBTT, Rhenium has a dramatic effect42 on the DBTT c either for pure Mo37 or for alloys containing reac- 25 molybdenum, as can be seen in Fig. 7. Thi tive metals. Individual lots of TZM material may rhenium ductilizing effect is not a new discovery/ possess lower DBTTs than individual lots of pure 35186 but it offers the promise of alloys which are resis molybdenum, but in general the DBTT tant to fracture at low temperatures. As noted pr< PROCESSMG AND PRODUCTION OF Mo AND W ALLOYS 107

TEMPERATURES marizes data from a number of investigators on 17 40 80 120 160 200 220 the D'JJTT of several tungsten alloys. Similar conclusions may be drawn for these materials as well. Mechanical working followed by stress relief sharp- ly lowers the DBTT as compared to recrystallized material. The effect of carbide-forming elements is not so clear, although the W-4Re-Hf-C alloy appears to have an advantage over the W-2 to 3Re alloys. The ductilizing effects of rhenium are also apparent. Arc melting produces material with lower DBTT values in the recrystallized condition than either EB-melting or powder metallurgy, although not quite as low as those of CVD material.

CREEP RUPTURE -40 -20 0 20 40 60 80 100 Figures 8 and 9 summarize creep-rupture data TEMPERATURE,^ for molybdenum and its alloys7'14'34'42"45 and tung- 43 45 46 Fig. 6 Ductile-brittle transition behavior of vacuum arc- sten and its alloys, - ' respectively. The cart molybdenum. Larson-Miller parameter, based on rupture life, assumes a constant of 13.42 and 12.28 for molybdenum alloys and tungsten alloys, respectively. No TOO -200 attempt was made to optimize the constant for each alloy. Also shown as an alternate abscissa is the temperature at which failure occurs in 75,000 hours, for a given Larson-Miller parameter value. This axis can provide an estimate of the maximum working stress to which the alloys can be subjected at a given temperature in a high performance application such as a space nuclear power system. Alloying can dramatically improve the creep strength of these materials. The data for molybdenum shown in Fig. 8 suggest that alloys have much broader scatter bands than pure molybdenum, a result of processing and composition variations. Low-stress tests, which are normally performed for long times or at high temperatures, are probably associated with material which has - -300 recrystallized during testing. The leaner-alloyed 10 20 30 40 50 TZM material shows its greatest advantage in tests RHENIUM CONTENT, Wt.% at intermediate stress levels, while TZC and MHC Fig. 7 Effect of rhenium on the 4T bend transition temper- are significantly more creep-resistant than TZM at ature of recrystallizt r vacuum-arc-cast and powder- high stresses. The MHC and TZC data exhibit metallurgy molybdenum. some curvature, probably the result of a tendency for the precipitate structure to coarsen at higher temperatures. It is interesting to note that, although the optimized processing employed by viously and illustrated in Fig. 7, there is a large 29 difference between powder-metallurgy and arc-cast Raf^o leads to dramatic improvements in tensile materials in the effect of rhenium on DBTT. In properties, the creep properties of these alloys still this respect, the VAC material responds much fall within the band of Fig. 8. As can be noted, PM materials also fall within the band for aix-cast more strongly than PM material to the ductilizing 7 effects of rhenium. materials. The DBTT behavior of tungsten alloys is analo- Figure 9 includes several tungsten alloys, in gous to that of molybdenum alloys. Table 7 sum- addition to pure VAC tungsten. The improvement 106 HAQEL. SHELOS, AW3 TUOMNEN

TABLE 7 4-T Find Transition Temperature for Tungsten and Tungsten Alleys*

Approximate Alloy Production technique Microstructure DBTTt, °C

PureW Powder metallurgy Stress-relieved 100 PureW Powder metallurgy Recrystallized 390 PureW Arc-cast Recrystallized 340 PureW E8 melted Recrystallized 345 PureW Chemical vapor deposited 210 W-5Hf SB melted Rec.>

'Data obtained by numeroub investigEtors and summarized in Ref. 10. tAverage of several values rounded to nearest ;i°C. ^Tensile elongation transition temperature.

75,000 h RUPTURE TEMPERATURE I0OO 2000 3000 °F 75,000 h RUPTURE TEMPERATURE 2000 3000 4000 500 1OOO 1500 1000 2500 °C

500 -

100 -

30 40

T(12.28 + log tr)

Fig. 8 Creep-rupture behavior of molybdenum and Fig. 9 Creep-rupture behavior of wrought tungsten and molybdenum-base alloys. (All materials tested in the tungsten-base alloys. (All materials tested in the wrought wrought and stress-relieved condition.) and stress-relieved condition, except as noted.) PROCESSING AMD PRODUCTION OF Mo AND W ALLOYS 109

TABLE 8 Nominal Compositions of Several Refractory Metal Alloys

Chemical composition, wt% Alloy Mo Nb Tft W Re Ti Zr Hf C

VAC-TZM Bal. _ _ _ 0.50 0.08 0.020 PM-TZC Bal. - - _ - 1.5 0.35 - 0.15 T-111 _ - Bal. 8 _ _ _ 2.4 - 1-222 - - Bal. 9.6 - 2.4 0.01 ASTAR-811C - - Bal. 3 1.0 - - 0.7 0.02E NAS-36* 0.7 _ Bal. 5.7 1.56 0.13 0.25 0.015 NWZr - Bal. - - - - 1.0 - - FS-85 - Bal. 28 10 _ 1.0 - - W-25Re - - - Bal. 25 - - - -

*NA3-36 also contains 0.015N. in creep strength offered by carbide-strengthened for 1% creep in 10,000 h for a variety of refractory alloys is ag ,;n apparent as is the scatter in proper- metal alloys.43-56 Table 8 lists nominal compositions ties resulting from chemistry and processing varia- for the alloys. The data indicate that for tempera- tions. The thoriated PM material also has a slight tures above ~1100°C a variety of materials per- advantage over pure tungsten. form better than the standard TZM alloy A word Density-normalized creep strengths are a useful of caution is in order, however. All creep data are measure of an alloy's creep strength for subject to scatter, especially when comparison is weight-sensitive applications. Figure 10 shows made between lots of material which differ from direct comparison of the denf'ty-normalized stress one another in chemistry and processing. This is especially true for the refractory metal alioys, which are very sensitive to the thermomechanical processing used to develop strength. This is well TEMPERATURE,°F 45 56 1500 2000 2500 3000 3500 illustrated by the two curves for \V-25Re. - Only when more complete data sets are available, based on statistically significant numbers of tests cf material representative of best commercial practice, will it be possible to accurately determine properties which the designer can use with confi- 50 dence. '10.0 When strength/density ratio is used, alloys that 4 are attractive on a raw strength basis may be NAS-36 found less desirable. The higher densities of the tantalum and tungsten alloys compared to T-222 molybdenum alloys serve as an example. On the 10.0 <" ASTAR 811C basis of strength/density ratio considerations, TZC, Y,-25Re(Ref. 56) W-25Re, ASTAR 811C, and NAS-36 all show promise for high-temperature application.

0.5 RECRYSTALLIZATION Because a significant contribution to both the strength and the ductility of tungsten and molybdenum alloys depends upon the development of a

1200 1600 2000 stable deformation structure, the recrystallization TEMPERATURE,°C behavior of these materials is quite important. In Fig. 11, data are summarized for molybdenum and Fig. 10 Density-normalized stress for 1% creep in 10,000 h. several of its alloys.5-7-35 The carbide dispersions of 110 HAOEL.SHELDS.Aff

1-H ANNEALING T£MPERATURE,°F dispersions in increasing the material's resistance 2000 3000 to recrystallization. The hafnium carbide alloy data plotted in Fig. 12 were selected from data from among several heats.48 The scatter bar! in the figure includes the best of the heats investigated. Note that the data given in Figs. 11 and 12 are for 1-h anneals. Recrystallization data are useful for developing processing schedules because they can be used to set annealing and stress-relief cycles. This information, however, is not useful for the designer, who must know about the long-term stability of the microstructures and strengths of ths alloys under consideration. Kinetic information on recrystallization is needed for the normally processed commercial alloys, as well as for alloys subjected to thermoruechanieal processing variations. Doping powder-metallurgy tungsten billets with aluminum, potassium, and silicon increases the 1000 1500 2000 49 1-H ANNfcALING TEMPERATURE,°C recrystallization temperature of tungsten wire. Figure 13 shows the comparison of the mean Fig. 11 Recrystallization behavior of molybdenum and boundary spacing of doped and undoped tungsten molybdenum-base alloys. wire, as determined by transmission microscopy, as a function of the annealing temperature.49 Doping can be seen to effectively reduce the mean bound- these alloys are important in retaining the worked ary spacing after annealing. It is employed to structure. Even the TZM alloy, which has a much increase the recrystallization temperature of highly lower alloy content than TZC and MHC, enjoys a worked, unidirectionally deformed refractory 300 to 400°C advantage in recrystallization tem- metals. For example, the temperature for complete perature over unalloyed molybdenum. The carbides recrystallization can be increased from ~1200°C for pure molybdenum to —1700°C for doped which form in TZM are effective in improving both 50 the strength and recrystallization resistance of molybdenum. molybdenum. Thc recrystallization temperature of VAC-TZM can be increased by solution treating31 at temperatures as low as 185O°C. The TZC and 1H ANNEAl ING TEMPERATURE,' MHC alloys, which contain higher alloy contents 3000 and rely to a greater extent on dispersion harden- zooo ing, have even higher recrystallization temperatures. However, the higher recrystallization temperatures were obtained in the powder metallurgy en 500 versions of these alloys after solution treating7 at 2200°C with rapid cooling. As was observed with tensile data, MHC and TZC recrystallization data fall in approximately the same scatter band. Less information is available for tungsten and 400 its alloys,45-47'48 partly because there are fewer tungsten alloys than molybdenum alloys produced commercially and partly because measurements of the recrystallization in tungsten alloys have emphasized hardness testing to a lesser degree. Re- 100 L^ 2000 3000 4000 crystallization behavior of tungsten alloys, sum- 1H ANNEALING TEMPERAT'JRE.t marized in Fig. 12, is consistent with that of molybdenum and molybdenum-based alloys and Fig. 12 Recryrtallization behavior of tungsten and again illustrates the effectiveness of carbide tungsten-base alloys. PROCESSING AND PRODUCTION OF Mo AND W ALLOYS 111

1 H ANNEALING TEMPERATURE, F investigate ductilizing analogues to rhenium, such .1200 2000 3000 DOp— r 1 i 1 i— as osmium, iridium, and platinum, and to make use of multiple alloying additions. Work at our own

50 - laboratory, employing combinations of several reactive plloy additions, has identified composi-

19 tions possessing superior strength at elevated tem- z peratures. GTE-Sylvania is developing a powder o metallurgy Mo-30 to 45W-HfO alloy.15 Investiga- 2 UNDOPED / tions are also necessary to determine the utility of i

TABLES Research Needs for Tnnfiten and Molybdenum Allo;ri

Consolidation techniques Lowered oxygen content in PM alloys with reactive metal additions Grain-refinement of arc-melted and EB-melted ingots Development of "interstitial-free" Mo and W alloys to maximize low temperature ductility Alloy development aud characterization Synergistic combinations of strengthening mechanisms High-strength alloys with good room-temperature ductility in recrystallized conditions Alloy compositions optimized for the consolidation technique to be used Development of kinetic data on recryatallization of Mo and W alloys Development of statistically significant data bases fo> physical and mechanical properties Thermomechanical treatment Processing schedules optimized for requirements of the end application Effects of TMT on kinetics of recrystallization Development of statistically significant data bases on the effects of TMT on mechanical and physical properties

form powders of high tungsten content, is being 2. R. Eck, Corrosion of Powder Metallurgy Molybdenum- studied by Allied Corporation.55 Conventional Tungsten Alloys in Liquid Zinc, Preprint of 5th Eur. Symp. powder metallurgy, while extremely valuable as a Powder Metallurgy, 3:187-194 (1979). 3. CRM Molybdenum-50 Rhenium, Alloy Digest, Filing Code tool in producing large alloy billets, suffers from Mo-11, Engineering Alloys Digest, Inc., Upper Montclair, oxide inclusions in alloys containing reactive NJ, 1970. metals. Research is needed to develop techniques 4. F. Fo.'ie, Rhenium Alloys, Inc., personal communication. for handling, blending, compacting, and sintering 5. Janet F. Briggs and Robert Q. Earr, Arc-Cast Molybde- these alloys which minimize or eliminate oxygen num-Base TZM Alloy: Properties and Applications, High contamination. These requirements are summa- Temperatures-High Pressures, 3: 363-409 (1971). rized in Table 9. The list contains several items 6. L. P. Clare, J. S. Smith, A. J. Kegel, and R. L. Koss, Prop- erties of a 1300 kg P/M TZM Billet, preprint of Tth Plansee which have been goals of refractory metal technol- Seminar, Vol. I, No. 30 (1971). ogists for many years. A great deal of progress has 7. S. M. Tuominen and V. Biss, Properties of PM been made toward thoss goals. With continued Molybdenum-Base Alloys Strengthened by Carbides, Modern effort we can expect to close the gap even further. Developments in. Powder Metallurgy, 14: 215-230, H. Haus- ner, H. W. Antes, and G. D. Smith (Eds), Metal Powder Industries Federation/American Powder Metallurgy Inst., ACKNOWLEDGMENTS Princetor.,NJ, 1981. 8. L. P. Clare and R. H. Rhodes, Superior P/M Molybdenum The authors wish to express their appreciation Die Alloys for Isothermal Forging, High Temperatures- to the following individuals wlio provided informa- High Pressures, 10(3): 347-348 (1978). tion on standard production and processing capa- 9. W. H. Chang, Effect of Titanium and Zirconium on Micro- bilities: R. W. Burman, R. S. Fusco, J. P. Linteau, structure and Tensile Properties of Carbide-Strengthened and R. C. Rhoades of ASMC; T. Baker of H. Cross, Molybdenum Alloys, Trans. ASM, 57: 527-553 (1964). Inc.; R. Misata of GNB Corp.; F. Nair of GTE- 10. M. Semchyshen, R. Q. Barr, -nd S. Kalns, Arc Cast Molybdenwm-Base Alloys (1962-196J,), ONR Report Contract Sylvania; S. Van Savage of Philips Elmet; Nonr-2390(00), Project NR 039-002. S. McCrossan of Schwarzkopf Development Corp. 11. W. D. Klopp, P. L. Raffo, and W. R. Wiizke, Strengthening (representing Metallwerk Plansee); F. Foyle of of Molybdenum and Tungsten Alloys with HfC, J. Met, 23: Rhenium Alloys, Inc.; R. Marsh of Teledyne Wah- 27-38 (1971). Chang Albany; R. Pan'in of Teledyne Wah-Chang 12. M. Semchyshen, U. S. Patent No. 3,169,860. Huntsville; J. DiMascio of Thermo Electron; and 13. R. Eck, Production and Properties of Powder Metallurgy R. Tuffias of Ultramet. Doped Molybdenum and Tungsten Wires, Metal, 33: 819-823 (1979). 14. W. R. Witzke, Compositional Effects on Mechanical Proper- REFERENCES ties of HafniumrCarbide-Strengthened Molybdenum. Alloys, NASA TM X-3239, National Aeronautics and Space 1. Exploring Molybdenum and Tungsten, Product Brochure Administration, Washington, DC (May 1975). AMAX Specialty Metals Corp-, Refractory Metals Division. 15. V. Nair, GTE-Sylvania, personal communication PROCESSING AND PRODUCTION OF Mo AND W A'J-OYS 113

16. R. Eck, Powder Metallurgy of Refrictory Metals and Appli- 37. H. Braun, M. Limchyshen, and R. Q. Barr, Recent Develop- cations, Int J. P/wder MetalL Powder TechrwL, 17: 201-211 ments in the Technology of Arc-Cast Molybdenum, (1981). Proceedings of 5th Plansee Seminor—Metals for the Space 17. Stephen W. H. Yih and Chun T. Wang, Tungsten: Sources, Age, F. Eenesovsky (Ed.), pp. 351-352 (1965). Metallurgy, Properties, and Applications, Plenum Prea9, New 38. M. Semchyshen, E. Kalns, D. J. Kalphaak, and R. Q. Barr, York. 1979. Development of High-Strength Molybdenum- and Tu'igsten- 18. 3. McCrossan, Schwarzkopf Development Corporation, per- Base Alloys, Final Report, Bureau of Naval Weapons Con- sonal communication. tract No. NOw 65-0385-c (1966). 19. R. L. Ammon, Consolidation anH Processing of W-0.35% 39. P. L. Raffo and W. D. Klopp, Mechanical Properties of HfC Alloy, in Proceedings of 12th Naticnai SAMPE Confer- Solid-Solution and Carbide-Strengthened Arc-Melted Turnsten ence m Materials, 1980. AUoys, NASA TN D-3248, National Aeronautics and Space 20. R. L. Ammon and R. W. Buckman, Jr., Vacuum Arc Melt- Administration, Washington, DC (February 1966). ing of Tungsten-Hafnium Carbon Alloy, J. Vac Sci Tech- 40. "Oremst HPC 85W15Mo," Alloy Digest, Filing Code W-9, noL, 11: 385-388 (1974). Engineering Alloys Digest, Inc. Upper Montclair, NJ, 1965 21. G. A. Geach and J. E. Hughes, The Alloys of Rhenium with 41. Wah Chang W-25Re, Alloy Digest, Filing Code W-15, Molybdenum or with Tungsten and Ha\dng Good High Tem- Engineering Alloys Digest, Inc., Upper Montclair, NJ, 1970 perature Properties. Plansee Proc 1955, pp. 245-253, 42. I,. B. Lundberg, An Evaluation of Molybdenum and Its F. Bene9OV9ky (Ed.), Pergamon Press, Oxford, 1956. Allays, Paper No. AIAA-81-llOft, Am. Inst. of Aeronautics 22. William D. Klopp, A Review of Chromium, Molybdenum, and Asti nautics, New York, NY (1981). and Tungsten AHoys, J. Less Common MeL, 42: 261-278 43. J. B. Ccnway and P. N. Flagella, Creep-Rupture Data for (1975). the Refractory Metxds to High Temperatures, Gordon and 23. R. Eck, Production and Properties of Powder Metallurgy Breach Science Publishers, New York, 1971. Molybdenum and Tungsten Doped Wires, Wire Industry, 47: 44. J. A. Houck, Physical and Mechanical. Properties of Com- 523-527 (1980). mercial Molybawm-Base Alloys, DMIC Report 140, Defense 24. G. W. King and H. G. Sell, The Effect of Thoria on the Materials Infornation Center, Battelle Memorial Inst., Elevated-Temperature Tensile Properties of Recrystallized Columbus, OH (November 30,1960). High-Purity Tungsten, Trans. AlME, 233: 1104-1113 (1965). 45. M. Semchyshen and R. Q. Barr, Extrusion and Mechanical 25. G. W. King, The Stress-Dependence of High Temperature Properties of Molybdenum and Tungsten AUoys, ASD Tech- nical Report 61-193, Aeronautical Systems Div., Air Force Creep of Tungsten and a Tungsten-2 Wt Pet ThO2 Alloy, Met Trans., 3: 941-945 (1972). Systems Command, USAF, Wright-Patterson AFB, OH, 26. R. Eck, VorgSnge beim Sintern von technischen MolybdSn June 1961. legierungen insbesondere mit Ti, Zr, Cr und C und Einfluss 46. L. S. Rubenstein, Effects of Composition and Heat Treat- auf die Eigenschaften der Fertigprodukte, MetsU., 28: ment on High-Temper ture Strength of Arc-Melted 1147-1151 (1974). Tungstev-Hafniuni-Carbou Alloys, NASA TN D-4379, 27. S. M. Tuoimnen and K. H. Carpenter. Powder Metallurgy National Aeronautics and Space Administration, Washing- Molybdenum: Influence of Powder Redaction Process on ton, DC, February 1968. Properties, J. MeL, 32: 23-26 (1980). 47. W. D. Klopp and W. R. Witzke, Mechanical Properties and 28. G. A. Timmons and R. G. Yingling, Arc Melting Molybde- RecrystaUization Behavior of Electron-Beam-Melted num, The Metal Molybdenum, ASM, pp. 80-108 (1958). Tungsten Compared vrith Arc-Melted Tungsten, NASA TN 29 P. L. Raffo, Thermomechanical Processing of Molybdenum- D-3232, National Aeronautics and Space Administration, Hafnium-Carbon Alloys, NASA TO D-564o, National Washington, DC, January 1966. Aeronautics and Space Administration, Washington, DC, 48. W. D. Klopp and W. R Witzke, Mechanical Properties of February 1970. Arc-Melted Tungsten-Rhenium-Hafnium-Carbon Alloys, 30. G. A. Timmons, consultant for ASMC, personal communica- NASA TN D-5348, National Aeronautics and Space tion. Administration, Washington, DC, July 1969. 31. E. Kalns, R. Q. Barr, and M. Semchyshen, Effect of In- 49. D. B. Snow, The Recrystallization of Commercially Pure Process Solution Heat Treatment? on the Properties of the and Doped Tungsten Wire Drawn to High Strain, MetalL TZM Alloy, High Temperature Refractory Metak Part 2, Trans., A, 10A: 815-821 (1979). Metallurgical Society Conferences, VoL Si, R. W. Fountain, 50. Molybdenum, General Survey, Product Information Bro- Joseph Moltz., and L. S. Richardson (Eds), Gordon and chure, Metallwerk Plansee, Reutte, Austria. Breach Science Publishers, New York, 1968. 51. N. E. Ryan and J. W. Martin, Hardening of Some 32. W. H. Chang, The Effect of Heat Treatment on Strengt.i Molybdenum-Base Alloys by Precipitation of Nitride and Properties of Molybdenum-Base Alloys, Trans. ASM, Carbide Phases, Proceedings of Plansee Seminar, 56: 107-124 (1963). Reutte-Tyrol, Austria, June 1968. 33. R. L. Stephenson, The Eftect of Fabrication Variables on 52. N. E. Ryan, W. A. Soffa, and R. C. Cr.-wford, Orientation the Creep-Rupture Properties of Molybdenum-Base Alloys, and Habit Plane Relationships for Carbide and Nitride Pre- Trans. AIME, 245: 997-1001 (1969). cipitates in Molybdenum, Metallography, 1: 195-220 (1968). 34. J. Wadsworth.. A Reevaluation of the Mechanical Properties 53. N. E. Ryan and J. W. Martin, The Formation and Stability of Molybdenum- and Tungsten-Based Alloys Containing of Group IVA Carbides and Nitrides in Molybdenum, Hafnium and Carbon, Met Trans., 14A: 285-294 (1983). J. Less-Common Met, 17: 363-376 (1969). 35. M. Semchyshen and R. Q. Barr, Mechanical Properties of 54. W. J. Boesch, J. K. Tien, and T. E. Howsen, Progress in Molybdenum and Molybdenum-Base Alloy Sheet, Vacuum Melting, from VIM to VADER, Met Prog., 122(5): Symposium on Newer Metals, ASTM Special Technical Pub- 49-56, October 1982. lication No. 272, pp. 1-24, Americ., Society for Testing and 55. M. Skrypa, Allied Corp., personal communication. Materials, 1959. 56. R. Gluyas and G. Watson, Materials Technology for an 36. M. Semchyshen, G. D. McArdle, and R. Q. Barr, Develop- Advanced Space Power Nuclear Reactor Concept, Program ment of Molybdenum-Base Alloys, WADC Technical Report Summary, NASA TN D-7909, National Aeronautics and 59-780, October 1959. Space Administration, Washington, DC, March 1975. CVD Refractory Metals and Alloys for Space Nuclear Power Application

L. Yang, T. D. Gulden, and J. F. Watson GA Technologies, Inc.

INTRODUCTION TECHNOLOGY STATUS CVD Electron Emitter for In space nuclear power systems, many key com- Thermionic Space Power System ponents have to -psirie at temperatures higher CVD Tungsten Emitter for than 1200°C for long periods of time. To ensure 1 high reliability in performance and life, they have In-Core Thermionic System to be made of refractory metals and alloys. In an in-core thermionic system, electrons are Although conventional metallurgical processes can used as the working fluid for the direct conversion be used to fabricate most of these components, of fission heat to electrical energy. Materials of low there are special cases when other methods are vapor pressure and high mechanical strength are better suited for meeting the needs. Chemical needed as electron emitters operating in contact vapor deposition (CVD), which was used exten- with nuclear fuels at high temperature for long sively from 1963 to 1972 for the fabrication of the periods of time. Extensive investigations of the electron emitters in in-core thermionic fuel ele- high temperature compatibility between various ments, represents one of these special cases. carbide and oxide fuels and various refractory In CVD, metal atoms are produced by chemical metals and alloys during 1961 to 1963 led to the reactions and deposited on the surface of a man- conclusion that tungsten is the best cnoice for the drel maintained at the desired temperature until electron emitter in a thermionic fuel element. The the required thickness of the deposit is attained. reference ,' sign for the emitter is in the form of a The mandrel is then removed either mechanically cup of 2.8 cm O.D., 6.4 cm length, and 1 to 2 mm or by using an appropriate solvent, leaving the wall thickness. Fission heat generated by the deposit in the shape of a replica of the mandrel. nuclear fuel contained inside the cup heats the This method is thus specially suited for the fabri- emitter to Loil off electrons from its external sur- cation of items with complicated shapes and rela- face. Successful fabrication of the tungsten emitter tively thin walls without welded or brazed joints; cup of the desired purity, microstructures, and this can be accomplished with relatively inexpen- electron emission properties is the key to the suc- sive equipment. By controlling the purity of the cess of in-core thermionics. Fabrication by drilling reactan*3 and the deposition parameters, products a cylindrical tungsten billet is obviously impracti- of desired purity, microstructures, and properties cal. Fabrication by welding a bottom to a tungsten can be prepared for specific applications. Further- tube runs the risk of failure at the weld. CVD more, the CVD technique is specially suited for the appears to be the most attractive method for fabri- deposition of thin coatings as high temperature cating the tungsten emitter cup. diffusion barriers for compatibility consideration. During the period of 1963 to 1970, extensive Major efforts applying CVD techniques in the fab- efforts were made at GA Technologies Inc. (GA), rication of refractory alloy components for space Oak Ridge National Laboratory (ORNL), Lawrence nuclear power systems are described in this paper. Radiation Laboratory (LRL), General Electric

114 CVD METALS AND ALLOYS FOB SPACE APPLICATION 116

Company (GE), and San Fernando Laboratory the NaOH traps to remove HF and any unreacted (SFL) to study CVD tungsten deposition and the WFb. properties and microstructure9 of CVD tH-^oten The deposition rate of fluoride tungsten deposits. The information that was generated increases with the temperature of the mandrel and helped to lay the foundation of CVD tungsten the pressure in the deposition chamber. Figure 2, emitter technology, which was used for the fabrica- taken from a paper by Holman and Huegel,2 tion of the emitters in thermionic devices tested illustrates this point. The deposit tends to be out "pile and in-pile at GA during the period rough and dendritic when the temperature of the ii&3 to 1972. mandrel is above 550°C; therefore the mandrel Fluoride Tungsten Deposition. Tungsten depos- temperature cannot be too high. For the prepara- its prepared by the hydrogen reduction of tion of cylindrical fluoride tungsten emitter, the following conditions have been used: mandrel tem- WF6(WF0 + 3H2 — W + 6HF) are referred to l as fluoride tungsten. All the emitters tested prior perature, 540°C; WF6 flow rate, h liter per minute; to 1968 were made of fluoride tungsten. Figure 1 hydrogen flow rate, 2 liter per minute; chamber shows schematically the deposition apparatus for pressure, % atmosphere. Figure 3 shows the as- fluoride tungsten. The hydrogen used is 99.999% deposited emitter and the emitter after it has been pure and is further purified by diffusion through a ground to the required dimension. Pd-Ag membrane. The WF6 used is purified by Typical fluoride tungsten deposits contain less cooling the gas cylinder to dry ice temperature and than a few ppm of nitrogen or hydrogen, 5 to 10 pumping off the volatile impurities. The deposition ppm oxygen, 10 to 15 ppm carbon, and very low chamber is water-cooled and can be made of Pyrex metallic impurity contents. Fluoride tungsten has a or Teflon. Various mandrel materials have been columnar grain structure which is stable at 1800°C used; molybdenum is favored because it can be out- (Fig. 4). Studies carried out at ORNL3-4 and GE5 gassed at high temperature and it has a thermal showed that such grain structure stability is due to expansion coefficient closely matching that of the presence of fluorine-containing gas bubbles at tungsten. The mandrel is heated by induction, the grain boundaries. When the fluorine content of although resistance heating can also be used. Dur- the deposit is high (>20 ppm), the deposit swells ing the deposition, the chamber is continuously when heated to high temperatures (Fig. 5). On the evacuated with a pump, and the pressure in the other hand, wher the fluorine content is low chamber can be adjusted to the desired value by (<~5 ppm), grain growth occurs readily at high varying the opening of the valve connecting the temperatures (Fig. S). A fluorine content of 10 to pump to the chamber. The exhaust passes through 20 ppm is considered as optimum for stable grain

TO ROUGHING PUMP

THERMOCOUPLE 'GAUGE SIGHT THERMOCOUPLE GLASS MOLYBDENUM / MANDREL FOR 'UMGSTEN DEPCSITICN

Fig. 1 Schematic diagram of fluoride tunggten deposition apparatus. 118 YANG, GULDEN. AND WATSON

TEMPERATURE (deg C)

1000 850 750 650 550 450

0.050

0.020 _

| 0.5 0.010

Z z o 0.005 ° PRESSURE FLOwRATE (rwr ) o H2 WF6 UJ o 0.1 ' 5 500 70 B 10 1000 70 D 10 500 7C 0.002 • 10 1500 XO a 20 500 70 0.05 o 40 500 70 VBO 500 70 0.001 ° 400-500 2400 40-300

0.8 0.9 1.0 1.1 1.2 1.3 1.4

Fig. 2 Effects of plating variables on fluoride tungsten deposition rate (from Ref. 2).

Fig. 3 CVD cylindrical fluoride tungsten emitter. Left, as-deposited; right, after grinding. CVD SiSTALS MS) ALLOYS FOR SPACE APPLICATION 117

: 1t '

•». V

(b) 200 urn 200 ym

AS-DEPOSITED AFTER 113 HOURS AT 1800°C IN VACUUM

Fig. 4 Microslructures of CVD fluoride tungsten.

0.03G . a

0.032 0.7 All Samples Between 0.450- and 0.500-in. o.d. 0.028 inoo h O.d E at 195irc E 0.024 / 0.5 i

0.016 '/, 0.4 (5 c 0.012 r o.:i | 0.2 -H 0.008 10(10 h at 1750°C 0.00-1 0.1

0 o 20 40 60 80 100 120 Fluorine Content, ppm

Fig. 5 Fluorine content vs. diametral change (from Roff. 5). 118 YAN&. GULDEN, AND WATSON

formed by partial reduction of WF6 with hydrogen, which is responsible for the observed bubble formation in fluoride tungsten at high temperatures.5 1 Ii has been shown that oxygen additives to WF6 (2 to 5 vol %) lower the fluorine content of fluoride tungsten without significantly increasing the oxygen impurity content in the deposit, presumably by connecting the WF4 to the more volatile W0F4. By exercising careful control of the purity of the WF6 and the deposition conditions, it was possible to produce fluoride tungsten emitters with fluorine content in the range of 5 to 20 ppm. In the latter part of 1972, fluoride tungsten emitters of fluorine content equal to 15 ± 5 ppm have been produced consistently.

Chloride Tungsten Deposition. In the fluoride tungsten deposit, the crystal grains are oriented preferentially with the (100) plane parallel to the substrate. Emitters prepared from such deposits have an electron work function of 4.5 eV. Early in 1966, efforts were made to prepare CVD tungsten deposits having (110) preferred orientation. Because the (110) plane is more densely packed with tungsten atoms than the (100) plane, it is expected that the (110) oriented emitter should "•S- - have a higher electron work function, which can improve the electrical output of the converter by reducing the cesium pressure required for optimum operation. It was found that (110) oriented tungsten cnuld be obtained by the hydrogen reduction of tungsten chloride under certain deposition condHions.8 The tungsten chloride used can be c WC(6 powder or tungsten chloride formed hy the chlorination of CVD fluoride tungsten chips. Figure 7 shows schematically the arrangement for preparing chloride tungsten deposits by the hydrogen reduction of tungsten chloride formed by the chlorination of fluoride tungsten chips at the top of the rig. 6 Effect of fluorine content on the microstructures of S series tungsten after annealing 1 h at 25C0°C. deposition chamber. Extensive efforts were made a. S OT-0 (1-ppm F); b. S 82-1 (6-ppm F); c. S 90-1 in defining the deposition conditions for achieving (30-ppm F); 125X (from Hef. 4). a high degree of (110) preferred orientation in the deposit and a high electron work function, usin^ structure and dimensional stability. Control of the planar molybdenum mandrel. Some of the results fluorine content of fluoride tungsten has received a are shown in Fig. 8. Chloride tungsten deposits of gre^t deal of attention during the latter part of the 4.9 to 5.0 eV electron work function can be obtained at a mandrel temperature of 1100°C, a H /W ratio 1960s. Deposition at low pressure (5 to 10 torr) in 2 of 2.2 to 2.8, a C1 /H ratio of 1, and a deposition chamber cooled with water jacket reduces the 2 2 3 chamber pressure of 4 to 6 torr. Typical chloride fluorine content. Increase of deposition tempera- tungsten deposits have impurity contents (ppm) as ture to 600°C, and H2/WF6 ratio (>50/l) also 6 follows: N < 1; 0, 2 to 7; C, 6 to 10; F, <3; Cl, 6 to lowers the fluorine content. Study carried out with 14; total metallic impurities, <10 ppm. Figure 9 time-of-flight mass spectrometer showed that th • shows the microstructures of as-deposited chloride gaseous species released from fractured CVD tungsten. fluoride tungsten could be a subfluoride (e.g., WF4) CVD METALS AT© .MXOYS FOR SPACE APPLICATION 119

Pressure Chlorine Gauge—> Thermocouple Supply Threeway Needle ^Needle Stopcock Valve Valves Ground Joints

•Flowme Lers Tungsten Chips Needle Valve Nichrome *• Resistance Furnace Stopcock Hydrogen -Quartz Fibers Supply ..Liquid Nitrogen Cold Trap • Mo Mandrel Induction Coil 8 •Hohlraum .Quartz Jacket To Vacuum Pump -McDanel Tube .Stainless Transition - McDanel Tube

Mercury — Manometer

Ground T '"int

Absoxute Mercury — Manometer .Optical Flat Glass Prism

Fig. 7 Chloride tungsten deposition apparatus using chlorination of tungsten chips to provide tungsten source material. 120 YANG, QULDBi AND WATSON

. (mrtlk.ay

jiioKJ.oi)

. (UOH--55) • (U0)(k.B7)

1200 1300 CiEoaltlco Toptratur* (°C)

Fig. 8 Preferred orientation and vacuum work function in chloride tungsten as a function of deposition conditions. The numbers in parentheses represent the lattice planes of preferred orientation and the vacuum work functions.

MOLYBDENUM 50X MANDRELSIDE

Fig. 9 Microstructure of a cross section (110) oriented chloride tungsten deposit. CVD METALS AND ALLOYS FOR SPACE APPLICATION 121

CHLORIDE TUNGSTEN

FLUORIDE TUNGSTEN

(a) AS-DEPOSITED 200 ym

CHLORIDE TUNGSTEN

FLUORIDE TUNGSTEN

(b) AFTER 1000 HR AT 1800°C 100 ym

Fig. 10 Microstructures of duplex tungsten.

Because chloride tungsten has a strong ten- tantalum transition piece which is then electron dency for grain growth, the cylindrical tungsten beam welded to other niobium components. emitters in the thermionic devices are composed of Figure 10 shows the microstructures of a duplex a chloride tungsten emitting layer (—0.3 mm tungsten emitter before and after heating at thick) on a fluoride tungsten substrate. The chlo- 1800°C for 1000 h and illustrates the excellent ride tungsten is used as the emitting layer because bonding between the fluoride tungsten and the of its high electron work function which provides chloride tungsten. While the grain structure of the better electrical output. The fluoride tungsten is fluoride tungsten layer remains stable, grain used as the substrate because of its stable grain growth has occurred in the chloride tungsten layer. structure which provides mechanical stability. An However, the electron work function of the chlo- emitter with such a two-layer design is called the ride tungsten emitting surface did not show am duplex tungsten emitter. In a thermionic fuel ele- significant change. These results obtained on ment, the tungsten emitter is diffusion bonded to a planar samples were subsequently used as a guide 122 YANG, GULDEN, AND WATSON in the deposition of the chloride tungsten layers of tron work function:13 chlorine flow rate 200 cylindrical duplex emitters. The duplex tungsten cc/min., rhenium pellet temperature 800°C, man- emitter is the reference emitter design for in-core drel temperature 1200° C. Figure 12 shows the thermionic fuel element and was the one tested in- microstructures of some CVD rhenium deposits. pile during the late 1960s and early 1970s. The impurity contents of these deposits in ppm are: C, 9 to 25; O, 25 to 48; N, 3 to 5; Cl, <1; Ca, 1 to 3; CVD Rhenium Emitter Na, 2 to 5. The conditions established were used in the preparation of cylindrical rhenium emitters on The (0001) face of a rhenium single crystal has 9 tantalum .nandrel 11.7 cm in length and 2.2 cm in an electron work function of 5.59 eV. A cesiated diameter. To improve the uniformity of the as- converter containing such a rhenium emitter 10 deposited thickness of the rhenium layer over the showed outstanding thermionic performance. entire length, the rhenium chloride generator was Because the use of a single crystal emitter in moved from the top of the deposition chamber to engineering devices is impractical from the point of its side so that the rhenium chloride vapor stream view of cost and the cylindrical configuration was perpendicular to the cylindrical axis of the required, efforts were therefore made to develop mandrel, i.e., a cross flow configuration. In addi- (0001) oriented polycrystalline CVD rhenium tion to the rotary notion at 2 rpm, the mandrel emitter. A CVD polycrystalline rhenium emitter was also moved axially across the rhenium chloride exhibiting (0001) preferred orientation was 11 vapor stream during the deposition. Rhenium prepared by the pyrolysis of rhenium chloride layers of strong (0001) preferred orientation and and yielded a vacuum electron work function of 5.1 12 uniform thickness could be deposited over the eV. Detailed investigations of the deposition entire 11.7 cm mandrel length at a rate of 0.1 to parameters were carried out at GA in the early 0.13 mm per h. 1970s in order to establish the conditions for achieving porosity-free rhenium deposits of a high In spite of the high electron work function of degree of (0001) orientation and high electron work (0001) oriented rhenium, the use of rhenium function. emitter is limited to out-of-core thermionic systems because rhenium is a strong neutron The deposition apparatus used (Fig. 11) is simi- absorber. lar to that for chloride tungsten deposition. The rhenium chloride used was prepared by the chlorination of rhenium pellets contained in a quartz reservoir located at the top of the deposition appa- CVD Tungsten Diffusion Barrier ratus. These pellets were prepared by taking 200 Between Fuel and Cladding mesh rhenium powder procured from Cleveland Refractory Metals and cold pressing to 70 MPa T-lll and Nb-lZr clad uranium nitride fuel pins (10.000 psi) and sintering the pressed pellets at were promising candidates as the heat source for 1400°C in vacuum for 15 h. The sintered pellets liquid metal cooled reactors for space power appli- were 6.4 mm in diameter, 1.6 mm in thickness and cations. To prevent fuel-clad interaction and loss of 80% dense. Prior to the deposition operation, the nitrogen, a 0.13 mm thick CVD tungsten barrier pellets were baked in situ in flowing hydrogen for was deposited on the I.D. of the cladding. This bar- a few hours at 900° C to rid them of any oxide pres- rier worked successfully at 1200°C cladding tem- ent, and the molybdenum mandrel was also cleaned perature during in-pile irradiation to 2% burnup.14 in hydrogen at 1200°C. The chamber was then The adherence of CVD tungsten coating to various flushed with helium to remove the hydrogen. The metal and alloy substrates has been discussed by mandrel and the rhenium pellets were brought to Bryant 15 and by Federer and Poteat.16 Good adhe- the desired temperatures and the chamber was sion is achieved if the substrate surface is free of evacuated before chlorine was admitted to the contaminants and does not react with the coating rhenium pellet reservoir to initiate the deposition. gas (e.g., WF6) at the deposition conditions. The rhenium chloride formed was decomposed at Because of its columnar grain structures and the hot molybdenum mandrel surface to form the associated grain boundary diffusion, CVD fluoride rhenium deposit. The chlorine released was tungsten is not expected to be as good a barrier for pumped out of the chamber to maintain a residual fuel component transport as CVD chloride tung- pressure of <1 torr. A thorough evaluation of the sten. CVD chloride tungsten develops equiaxia! deposition parameters established the following grain structure similar to that in recrystallizec1 conditions for obtaining porosity-free (0001) wrought materials after heat treatment at high oriented rhenium deposits of 5.17 to 5.2G eV elec- temperatures (e.g., 1500°C). CVD METALS AND ALLOYS FOR SPACE APPLICATION 123

Thermocouple Monel Pressure Gauge

Chlorine Supply Ground Joints

Needle Absolute Valves Mercury Thermocouple Well Manometer Nichrome Resistance Stopcock Furnace Quartz Reservoir

Liquid Nitrogen Cold Trap Rhenium Chips Quartz Fibers

To Vacuum Hohlraum Pump Mandrel 3* Induction Coil

Quartz Jacket Mullite Ceramic Tube Absolute Glass Mercury Wool- Manometer Trap Stainless Stenl Tube Rubber Gasket Stainless Ste?l Liquid Grooved Plate Nitrogen Cold Trap • Rubber 0 Ring Seals Motor •Wilson Seal Optical Flat Rotational Worm Glass Prism Gear Drive

Fig. 11 Rhenium deposition apparatus.

CVD Tungsten Heat Pipe flow heat exchanger involve both circular and rectangular cross sections (Fig. 13). Tungsten-lithium Tungsten is one of the most promising wall heat pipe has been operated up to 2000°C for materials for lithium heat pipe for high tempera- 5000 h without failure.17 ture operation because of its high temperature CVD techniques were used to fabricate the strength and resistance to oxygen-induced corro- tungsten heat pipe shell for two designs, the sion in lithium systems. Some of the heat pipe smooth wall heat pipe utilizing autoclaved tungsten designs for transferring heat from the reactor core screen wicka in both the evaporator and the con- to the conversion system involve complicated denser (Fig. 14a) and the channeled wall heat pipe shapes. For instance, heat pipes designed for cross with channels in the condenser wall and a swaged 124 YANG, GULDEN. AND WATSON

Rhenium Deposit

/••

*Uuw f Molybdenum ~ T Mandrel 100X

Fig. 12 "^icrostructures of CVD rhenium sample, deposited at a chlorine flow rate of 200 c.c./min, a rhenium pellet temperature of 800°C, and a mandrel temperature of 1200°C.

REACTOR evaporator, 10 cm condenser, 2.2 cm adiabatic HEAT PIPES length), a wall thickness of 0.1 cm, a diameter of 0.955 cm for the evaporator, and a condenser cross r POWER LEVELING spcuon of 0.955 cm X 1.91 cm. Fabrication of '\ HEAT PIPES these heat pipes was carried out at SFL for NASA Lewis Research Center in 1970.18 Both heat pipes have operated to the design temperature of 1575°C. r- CONVERTER ,'i HEAT PIPES CVD Tubing FUEL -T. CVD tubing has the potential of being used as fuel cladding. Both tungsten and tungsten alloy tubing has been fabricated by CVD techniques during the 1960s and early 1970s.

CVD Tungsten Tubing CVD can produce high-purity tungsten tubing Fig. 13 Cross flow heat exchanger (from Ref. 17). of various • diameters and wall thicknesses in lengths up to 1.2 m. A technique using resistance heated thin-wall stainless steel mandrels to pro- tungsten screen wick in the evaporator (Fig. 14b). duce long tubings of uniform as-deposited wall The smooth wall heat pipe had a total length of 32 thickness was described bv Martin et al.19 These cm (18 cm evaporator, 12 cm condenser, 2.2 cm adi- authors also concluded that for small diameter abatic length), a wall thickness of 0.1 cm, a diame- (—1.9 cm) thin wall (<1.3 mm) tubing, fabrication ter of 1.0 cm for the evaporator, and a condenser by CVD technique should have an economical cross section of 1.3 cm X 2.0 cm. The channeled advantage over fabrication by conventional metal- wall heat pipe had a total length of 25 cm (13 cm lurgical processes. CVD METALS AND ALLOYS FOR SPACE APPLICATION 125

3. Provisions (not shown in Fig. 15) for rub- bing the deposit surface during the deposition to break up the columnar grain structure of the deposit and to change the mechanical properties of the deposit. The ranges of the deposition parameters studied are: temperature, 600 to 1000°C; pressure, 10 to 100 torr; K2/MF6 ratio = 1 to 14; MF6 flow rate 30 to 530 cc/min and concentration of ReF6 in WF6 = 6.8 to 25%. It was found that lower temperature and pressure led to high rhenium concentration in the deposit. It appears that the rhenium deposition rate was supply limited, with a deposi- L ••,-•:•.. [£,• -|r • tion efficiency nearly 100% in the temperature range studied, while the tungsten deposition rate was limited by a surface process and therefore was Fig. 14a Smooth wall CVD tungsten heat pipe (from Raf. 17). temperature dependent. The overall deposition rate of the alloy increased with temperature and for each temperature was a maximum at the H2/MF6 ratio of 3. The microstructures of the tungsten- rhenium alloy deposited on the rotating mandrel were characterized by a chevron-like pattern super- imposed on the normal grain structure. It is believed that those markings were caused by fluctuations in composition as the area under observation traveled through regions with different composition, with the side facing the inlet gas having higher rhenium concentration tnan the side 180° from the gas entrance. When the rhenium content exceeded 18% to 20%, a very hard meta- stable phase with the (3-tungsten A-15 structure appeared.20 High temperature (1400 to 1600°C) annealins is needed to achieve compositional homo- Fig. 14b Channel wall tungsten heat pipe (from Ref. 17). geneity, to restore the equilibrium phase relationship, and to improve the mechanical properties. The apparatus shown in Fig. 15 could produce tubes of only 2.5 cm in length. It was later modi- CVD Tungsten-Rhenium Tubing fied to allow the deposition of homogeneous W-25 21 CVD tungsten-rhenium alloy deposition by Re alloy tubing of 0.3 m in length. This was hydrogen reduction of a mixture of WF6 and ReF6 accomplished by using a rectangular cross-section has been studied both at ORNL and LRL. Holman flue that could direct the coating gas across a and Huegel2 used the apparatus shown in Fig. 15 10.1 cm length of the mandrel and a cam driven to study the effect of temperature, pressure, ratchet mechanism for slowly translating ti.e H2/MF6 ratio (MF6 is WF6 + ReFP), concentration rotating mandrel along its axis through the plaiing of ReF5 in MF6, and hydrogen flow rate on the region for a distance of 0.4 m. A schematic draw- composition, m .crostructures, and mechanical prop- ing of the apparatus is shown in Fig. lfi. erties of the deposit. The special features of this 22 apparatus are: Federer, Schaifhauser, and Leitten deported tungsten-rich alloy (7% and 22%, Re) tubings 1. Accurate control of the flow rate and (1.9 cm O.D.) having uniform composition in both WFg/ReFe ratio by positive displacement the axial and the radial directions in a moving hot meiering of premixed liquid fluorides into a zone by the hydrogen reduction of (WF6 + ReF6). flash evaporator. A (WF6 + ReF6) injector was used to minimize 2. Cross-flow of the coating gas with respect deposition in a temperature gradient. Figure 17 to the rotating mandrel to improve the uni- shows the deposition apparatus. The 22% Re alloy formity of deposit thickness. was deposited under the conditions: 900° C, 10 torr, 126 YANG, GULDEN, AND WATSON

MASS FLOW METERS (HASTINGS)

H2 GAS BLANKET DISCONNECT (FOR OPTIC PROTECTION) LOW POWFR MICROSCOPE EXHAUST COLD TRAPS A MECH. PUMPS O-RING SALL VALVE AH LOCK PACKING THERMOCOUPLE GLAr-.C .SLIP RING FLASH CHAv5i?^\

PRESSURE GAUGE

INDUCTION COIL MANDREL (THERMOCOUPLE ON INSIDE Su \CE)

Fig. 15 Chemical vapor deposition apparatus, for tungsten and tungsten- rhenium alloys (from Ref. 2).

/Blankef gas {Hj

Gas flow

exhaust (HP,excess H Reactonts (Ho, WF,,ReF WF,, ReFT zoo 6 6

Fig. 16 Long-tube fabrication apparatus (from Ref. 20). CVD METALS AND ALLOYS FOR SPACE APPLICATION 127

WATER-COOLED WF.-ReF, INJECTOR (STATIC ^Y)

DEPOSITION FURNACE (STATIONARY)

DEPOSITION MANDREL IMOV^BLE)

VACUUM PUMP HF SCRUBBER' DRAIN

Fig. 17 Tungsten-rhenium deposition apparatus (from Ref. 21).

and H2/(WF6 + ReF6) ratio of 15:1. The 7% Re For the liquid metal cooled fuel pin using ura- alloy was deposited under the conditions: 800°C, 5 nium nitride as fuel and T-lll and Nb-1 Zr as clad- torr, and H2/(WF6 + ReF6) ratio of 20:1. In ding, a lungsten barrier possibly produced by CVD both cases, the mandrel traveled at a rate of 1.8 methods is essential to the fuel-cladding compati- in./h. The rhenium concentration uniformity along bility at the designed operating temperature. CVD a distance of 22.9 cm in the axial direction was technology is well established for fabricating tung- ±0.5% for the 7% Re alloy and ±1% for the 22% sten and tungsten-rhenium tubes which can be Re alloy. Deposition on the outer surface of used as fuel cladding for space reactors. mandrels in a static hot zone also showed promise Reactors for space nuclear power systems may for obtaining uniformity in composition. use heat pipes to transfer heat from the reactor core to the conversion system. CVD technology has been used for fabricating the heat pipe used as SUMMARY OF TECHNOLOGY cross-flo'.v heat exchanger, including the built-in STATUS channels on the condenser wall for liquid lithium return. CVD represents the best means for CVD technology has made significant contribu- fabricating heat pipes, when the condenser and the tions to the development of space nuclear power evaporator have different shapes. systems during the period 1962 to 1972. CVD capability for fabricating refractory metal For the in-core thermionic concept, CVD tech- and alloy components is available in a number of nology is essential to the fabrication of the tung- laboratories. These include ORNL, Los Alamos sten electron emitter. The CVD fluoride tungsten National Laboratory (LANL), LRL, GA, Thermal process provides the emitter substrate for mechan- Electron Corporation (TECO), Battelle Memorial ical strength and structural stability. The CVD Institute (BMI), SFL, and ULTRAMET. There chloride tungsten process provides the emitting should be no difficulty in getting the required work layer of high electron work function and superior done if need arises. thermionic performance. The controlling parameters for these processes have been thoroughly investigated during the past ten years. Various TECHNOLOGY NEEDS LIST physical and chemical methods have been 6 23 24 developed for quality control. ' " These processes CVD of refractory metals and alloys has a are ready for scaleup if needed. broad technological base to support the develop- 128 YANG. GULDEN, AND WATSON ment of space nuclear power systems. To meet REFERENCES future needs, CVD technology should be applied directly to resolve the critical problems controlling 1. L. YBng and R. G. Hudson, Evaluation of Chemically Vapor the life and performance of these systems. Deposited Tungsten as Electron Emitters for Nuclear For the thermionic space power system, the Thermionic Application, in Proceedings of the Conference on process for fabricating the duplex tungsten Chemical Vapor Deposition of Refractory Metals, Alloys, and emitters is well established. However, some of the Compounds, pp. 329-348, A. C. Schaffhauser (Ed.), American material proptrties of CVD tungsten remain to be Nuclear Society, Hinsdale, Illinois, 1967. 2. W. R. Holman and F. J. Huegel, CVD Tungsten and studied. The first is the creep behavior of fluoride Tungsten-Rhenium Alloys for Structural Applications. tungsten and duplex tungsten as a function of Part I: Process Development, in Proceedings of the Confer- fluorine content. Excellent work has been carried ence on Chemical Vapor Deposition of Refractory Metals, out daring the 1960s and early 1970s at ORNL25"27 Alloys, and Compounds, pp. 127-148, A. C. Schaffhauser in this area. These studies should be extended to (Ed.), American Nuclear Society, Hinsdale, Illinois, 1967. stress level below 1000 psi and to the biaxial creep 3. A. C. Schaffhauaer and R. L. Heestand, Effect of Fluorine Impurities on the Grain Stability o* Thermochemically of duplex tungsten in contact with nuclear fuel Deposited Tungsten, in 1966 IEEE Conference Record of the such as UO2. Such information is needed for the Thermionic Conversion Specialist Conference, pp. 204-211. viscoelastic analysis of the dimensional stability of 4. K. Farrell, J. T. Houston, and A. C. Schaffhauser, The tungsten emitter during in-pile operation and the Growth of Grain Boundary Gas Bubbles in Chemically life of the thermionic fuel element. The second is Vapor Deposited Tungsten, in Proceedings of the Conference the fast neutron effect on the dimensional stabil- on Chemical Vapor Deposition of Refractory Metals, Alloys, ity, thermal conductivity, and electrical conductiv- and Compounds, pp. 363-390, A. C. Schaffhauser (Ed.), 28 American Nuclear Society, Hinsdale, Illinois, 1967. ity of CVD tungsten. Some efforts were initiated 5. J. V. Festa and J. C. Danko, Some Effects of Fluorine Con- at ORNL in 1972 ir. this area but were not contin- tent on the Properties of Chemically Vapor Deposited ued because of the termination of all space power Tungsten, in Proceedings of the Conference on Chemical programs early in 1973. Work in this area is Vapor Deposition of Refractory Metals, Alloys, and Com- urgently needed for the design of the thermionic pounds, pp. 349-361, A. C. Schaffhauser (Ed.), American fuel element. Nuclear Society, Hinsdale, Illinois, 1967. 6. J. Chin and J. Horsley, Chemical Vapor Deposition of For the liquid metal loop or heat pipe cooled Tungsten, in 1968 IEEE Conference Record of the Therm- reactor used as the heat source of various conver- ionic Conversion-Specialist Conference, pp. 51-59. sion systems, CVD technology should be extended 7. F. A. Glaski, The Use of Oxygen Additive to Control Re- to the following two areas. The first is the fabrica- sidual Fluorine in Chemical Vapor Deposited Tungsten, in 1970 IEEE Conference Record of the Thermionic Specialist tion of long W-Re tubing as fuel cladding. The fab- Conference, pp. 72-73. rication techniques developed at LRL and ORNL 8. R. G. Hudson, T. Tagami, and L. Yang, Chemical Vapor have produced W-Re tubing of 0.3 m length. The Deposition of Tungsten Emitters of (110) Preferred Crystal process should be scaled up for the fabrication of Orientation, in Proceedings of the Second International tubing of the lengths needed in current reactor Conference on Thermionic Electrical Power Generation, pp. 565-573, Stresa, Italy, May 27-31 (1968). designs and of uniform composition! and structure. 9. R. Wichner and T. H. Pigford, Work Function of Mono- The second is the development of CVD Mo-Re crystalline and Polycrystalline Rhenium, in 1966 IEEE technology for the fabrication of heat pip" of low Conference Record of the Thermionic Conversion Specialist ductile-to-brittle transition temperature, utilizing Conference, pp. 405-412. the experience gained by LRL and ORNL on CVD 10. T. H. Pigford and B. E. Thinger, Performance Characteris- tica of a 0001 Rhenium Thermionic Converter, in 1969 IEEE W-Re system. It is realized that cost, geometric Conference Record of the Thermionic Conversion Specialist configuration (length, diameter, and shape), and Conference, pp. 34-38. material quality (purity, composition, and 11. F. A. Glaski, The Formation of (0001) Oriented Rhenium microstructural uniformity) are the important fac- Surfaces by Chemical Vapor Deposition, iu 1970 Conference tors for determining whether CVD or conventional Record of the Thermionic Conversion Specialist Conference, pp. 128-129. metallurgical processes should be used as the fabri- 12. R. G. Hudson and L. Yang, Diffusion and Electron Emis- cation technique. These factors should be taken sion Properties of Duplex Refractory Metal Thermionic into consideration in the planning of future CVD Emitters, Metallurgical Society Conferences, 41, Refractory programs for supporting the development of Metals and Alloys IV, Research and Development, Volume II, space nuclear power systems. pp. 1253-1268 (1969). 13. L. Yang, R. G. Hudson, and J. J. Ward, Preparation and For the general purpose of using CVD tungsten Evaluation of Chemically Vapor Deposited Rhenium Therm- as a corrosion resistance barrier, surface glazing ionic Emitters, in Proceedings of the Third International Conference on Chemical Vapor Deposition, pp. 253-269, F. A. by means of laser or electron beam should be Glaski (Ed.), American Nuclear Society, 1972. investigated as a mean3 of surface modification to 14. Private communication between L. Yang of GA and reduce defects and enhance corrosion resistance. J. ScotL of ORNL, June 1983. CVD METALS AND ALLOYS FOR SPACE APPLICATION 129

15. W. A. Bryant, The Adhesion of Chemically Vapor Deposited International Conference on Chemical Vapor Deposition, pp. Tungsten Coatings, in Proceedings of the Second Interna- 171-191, J. M. Blocker, Jr., and J. C. Withers (Eds.), The tional Conference on Chemical Vapor Deposition, pp. 409-421, Electrochemical Society, 1970. J. M. Blocker, Jr., and J. C. Withers (Eds.) The Electro- 22. J. I. Federer, A. C. Schaffhauser, and C. F Leitten, Jr., chemical Society, 1970. Thermochemical Deposition and Evaluation of Rhenium and 16. J. I. Pederer and L. E. Poteat, A Study of the Adherence of Tungsten-Rhenium Alloys, in 1966 IEEE Conference of Tungsten and Molybdenum Coatings, in Proceedings of the Thermionic Conversion Rvecialists, pp. 229-228. Third International Conference on Chemical Vapor Deposi- 23. L. Yang, R. G. Hudson, T. Tagami, and J. W. R. Creagh, tion, pp. 591-599, F. A. Glaski (Ed.), American Nuclear Preferred Crystal Orientation Evaluation and Vacuum Society, 1972. Work Function Measurements on Chemically Vapor Depos- 17. D. M, Ernst, He»t Pipe and Radiator Design, Direct Conver- ited Tungsten Emitters, in 1968 IEEE Conference of sion Nuclear Reactor Space Power System, Appendix II, Thermionic Conversion Specialists. Report to Wright Patterson Air Force Laboratory by Rasor 24. L. Yang, R. G. Hudson, and J. W. R. Creagh, Evaluation of Associates, 1981. Cylindrical CVD Tungsten Emitters of (110) Preferred Crys- 13. R. J. Bacigalup, Fabrication and Testing of Tungsten Heat tal Orientation for Thermionic Application, in Proceedings Pipes for Heat Pipe Cooled Reactors, in 1971 IEEE Confer- of the Second International Conference on Chemical Vapor ence Record of the Thermionic Conversion Specialist Confer- Deposition, pp. 817-837, J. M. Blocker, Jr., and J. C. Withers ence. (Eds.), The Electrochemical Society, 1970. 19. W. R. Martin, R. L. Heestand, R. E. McDonald, and G. A. 25. H. E. McCoy and J. 0. Stiegler, Mechanical Behavior of Reimann, Application of Chemical Vapor Deposition to the CVD Tungsten at Elevated Temperatures, in Proceedings of Production of Tungbten Tubing, in Proceedings of the the Conference on Chemical Vapor Deposition of Refractory Conference on Chemical Vapor Deposition of Refractory Metals, Alloys, and Compounds, pp. 391-425, A. C. Metals, Alloys, and Compounds, pp. 303-314, A. C. Schaffhauser (Ed.), American Nuclear Society, Hinsdale, Schaffhauser (Ed.), American Nuclear Society, Hinsdale, Illinois, 1967. Illinois, 1967. 26. R. E. Stephenson and J. I. Federer, Creep Rupture Proper- 20. J. I. Federer and J. E. Spruiell, Formation and Character- ties of CVD-Tungsten, in 1970 IEEE Conference of Therm- ization of an A15-Type Structure in Chemically Vapor De- ionic Conversion Specialists, pp. 90-91. pos'fed Tungsten Rhenium Alloys, in Proceedings of the 27. K. Farrell, Creep Properties of CVD Tungsten, in Conference on Chemical Vapor Depositions of Refractory Proceedings of the Fourth Internationil Conference on Chem- Metals, Alloys, and Compounds, pp. 443-458, A. C. ical Vapor Deposition, pp. 439-455, G. I Wakefield and J. M. Schaffhauser (Ed.), American Nuclear Society, 1967. Blocker, Jr. (Eds.), The Electrochemical Society, 1973. 21. I. A. Huegel and W. R. Holman, CVD Tungsten and 28. F. W. Wiffen, Radiation Damage to Refractory Metals as Tungsten-Rhenium Alloys for Structural Applications. Part Related to Thermionic Applications, in 1ST: IEEE (Xvftir- III: Recent Developments, in Proceedings of the Second ence of Thermionic Conversion Specialists, pp. 156-165. Machining Refractory Alloys-An Overview*

John D. Christopher Metcut Research Associates, Inc.

INTRODUCTION the molybdenum TZM and the niobium D31, can be machined at approximately 300 fpm. However, the The machining of refractory metals has generally tungsten and the tantalum both show lower cutting been characterized by low metal removal rates and speeds. The tungsten was machined at 200 fpm, but poor cutting tool life. These manufacturing prob- only a 15-min tool life was obtained. The tantalum lems have somewhat limited the widespread usage alloy was machined at <100 fpm. These cutting and application of this category of engineering speeds generally compare with other high tempera- materials and their alloys. A literature search has ture materials that are nickel base or with high revealed little, if any, new published machining strength steels at hardness levels above 55 Re. data on traditional machining operations such as A comparison with the same group of alloys on turning, milling, drilling, and tapping. Even grind- the basis of feed rate used for a 30-min tool life is ing of these materials is difficult or tricky because shown in Fig. 2. Only three of the materials shown of the susceptibility of surface damage such as exhibit the need for a low feed rate. Two of those heat checks or cracks. materials are high strength stsels above 55 Re, and Nontraditional machining is a generic term for the third is the niobium D31. This material can be those material removal processes that differ dras- turned only at 0.005 in./rev for the 30-min tool tically from the historic operations such as turn- life. The other refractory materials, tantalum alloy, ing, milling, drilling, tapping, and grinding. The tungsten, and molybdenum, can all be machined at use of primary energy modes other than mechani- feed rates as high as .009 ipr. cal, such as thermal, electrical, and chemical, sets The cutting speed used for 60 linear in. of travel these operations apart and reinforces their non- per tooth in face milling is shown for a variety of traditional label. Several of these newer processes materials in Fig. 3. The molybdenum alloys are have been very successful in machining close toler- machined at the highest speed of any of the refrac- ance parts from refractory materials. tory materials, approximately 300 fpm. The This paper will provide a general overview of niobium alloy can be face milled at approximately both traditional and nontraditional aspects of 200 fpm and the tungsten alloys at lower speeds machining refractory materials. with lower tool life levels than 60 ipt. The tantalum alloy was machined at the slowest cutting DISCUSSION speed of any of the refractory materials, approximately 75 fpm for the 60 ipt tool life level. Traditional Machining Methods The feed rate to produce a 60 ipt. tool life in face milling for the various materials is shown in Figure 1 shows a comparison of the cutting Fig. 4. The refractory materials are grouped at the speeds used to obtain a 30-min tool life for four lower part of the figure. In this example, the two refractory materials: molybdenum TZM, tungsten molybdenum alloys were face milled at the lowest at 93% density, nioDium D31 and tantalum-10 feed rate of .005 ipt; the tungsten, niobium, and tungsten alloy. Two of the refractory materials, tantalum alloys were all machined as high as .010 ipt feed rate. With the exception of 302 stainless •Note: Metric units have not been usH in this paper. steel, which is very easy to machine, none of the

130 4340 (340 BHN)

MOI.Y-. 5Ti or TZH 4340 (220 BHN) (52 R

D6AC (56 "-) NIOBIUM D31 (207 DH1!) D6AC (58 R

4340 302 SS (52 He) (170 BHN)

RENE 41 Ti-6A1-4V (365 BHN) (312 BHN) WASPALLOY P S S TUNGSTEN (15 HIN LIFE IIAX. ) (3yo BHN) (93% DENSITY) INCO 700 (302 Bill!) D6AC (56 "el UDIMET 500 (340 BHN) Ta-lOW (220 BHN) Ti B-120VCA

Ti 6A1.-4V INCO 700 (312 BHN) (302 BHN) MOLY - TZM (220 BHN) D6AC (58 Rc) HOLY - .5 Ti (220 BHN) UDIMET S00 RFFKAC'TOFY ALLOYS (340 BHN) PiS TUNGSTEN (931 DENSITY)

RENE 41 Ta-lOW (365 BHN) (220 BUN)

IIIOBIIJM .031 (2"7 BHII) I 200 .007 .011 CUTTIIIG SPEED - FT/KIH nxu -

Fig. 1 Carbide turning aerospace alloy*, relative cutting ipeed for 30-nun tool life. Fig. 2 Carbide turning aerospace alloys, feed for 30-min tool life. 4340 4340 (340 BUN) (340 BUN) ~i 4340 302 SS (52 Re I (170 BUN) ID D6AC HOLY TZM (56 Rc) ZD (220 BUM) 302 SS HOLY - .5 Ti (170 Bllll) :••;•-. •::•• ,.|: (220 BHN)

KENE 41 II .s.s. 4340 (365 BHN) 1 (52 Re)

11.s.s. NIOJI'JM D?L IHCO 700 (302 HHN) m (207 M<1)

UDIMET 500 11.s.s. PSS TUNGSTEN (30 IN./700TK) (340 BUN) (93* DENSITY) i] Ti A-110 FORGED TUNGSTEN (un I;I./TOOTH) (96* DENSITY) (110 BUM) ZT

Ti G.M.-4V Ti B-120VCA (312 BHN) (2B5 BHN) TI B-120 VCA Ti B-120VCA (285 BUN) (400 BUN) Ti B-J::3 VCA Ti 6A1-4V (365 BHN) (312 BHN) m

D6AC PSS TUNGSTEN (56 "c) (

Ti B-120 FORGED TUNGSTEIi (961 DENSITY) 1 (36S BHN)

1II0BIU" 1)31 REFRACTORY ALLOYS ' 1NC0 700 1 (302 BHN) ] (207 BHii) HOLY -.5 Ti RENF. 41 HIGH SPEED STEEL CUTTER (220 BUN) (365 BHN) ] D MOI.Y -TZM UDIMET 500 (220 BHN) (340 BUN) ] Z] Ta-lOW Ta-lOW (220 BUN) s.s. •r ,.-;;,- • |. (220 BHti)

• •sr-•••"'r i''"4 1 1 .005 .010 .015 .020 300 450 600 ITED - IN./TOOTH Cl'TTIilC, SPEED - IT/IUil

Fig. 3 Carbide face milling aerospace alloys, relative catting speed for Fig. 4 Carbide face milling aerospace alloys, relative feed for 60 in./tooth : ; : ; >« 1- U..1L >M) Itf. tool life. : ' .•: •" ' " ' '' :":-. -: ;v:. :,;.> :,.':V'\'i MACHNNG REFRACTORY ALLOYS—AN OVERVEW 133

HOLY - TZM (220 BHNi

HOLY - .5 Tl (220 BHN)

302 SS 1

IIIOBIU!! D3.1 (207 SKI) 1 Ta-lOW (220 BHN) 1

4340 (340 BHN) 1

Ti 6A1-4V (312 BHN) 1

Ti A-110 (310 BHN) 1

4340 152 B=) 1

INCO 700 1 (30 HOLES) (302 BUN)

UDIMET 500 1 (30 HOLES) (340 BHN)

Ti B-12OVCA (285 BHN)

Ti B-120VCA (365 BHN)

RENE 41 (365 BHNi

DfaAC SOLTD CARBIDE DRILL (70 HOLES) j (56 "c)

P&S TUNGSTEN SOLID CARBIDE DRILL (14 HOLES) | (93% DENSITY)

FORGED TUNGSTEN SOLID CARBIDE DRILL (96* DENSITY) (15 HOLES) | 1 1 1 1 1 25 50 75 100 125

CUTTING SPEED - FT/-1I1I

Fig. 5 BLS& drilling aerospace alloy?—Vi-in.~dia X '/4-in.-deep thru hole, relative catting speed for SO- to 100-hole drill life.

other alloys shown in this figure were machined at approximately 125 to 135 fpm. However, the tool higher feed rates than .010 ipt. Most of the nickel life obtained in these tests even with carbide drills and titanium alloys were machined at .005 ipt. was quite low, only 14 to 15 holes. The feed rates The relative cutting speed used to drill a variety used in the same drilling tests on the same group of alloys with both high speed steel and carbide of materials are shown in Fig. 6. The refractory drills is shown in Fig. 5. The two molybdenum alloys are shown at the bottom portion of the fig- alloys were machined at approximately 100 fpm, ure. Molybdenum and niobium alloys were drilled compared with 75 fpm with the niobium D31 mate- at a feed rate of .005 ipr which would be the rial. Both of these were machined with high speed expected feed to use on a W-in. drill. This was the steel drills along with the tantalum alloy, which feed used on the 4340 steel at 340 BHN, the 302 was drilled at approximately 50 fpm with the high stainless, and on two of the titanium alloys. How- speed steel drills. Both of the tungsten alloys were ever, the tungsten alloys at 93% and 96% density drilled with solid carbide drills at speeds of and the tantalum alloy were both drilled at .002 134 ipr. Again it is not noted on Fig. 6, but the low TZM alloy machined as the best. The tantalum tool life of 14 and 15 holes obtained with the two alloy was machined at the lowest cutting speed, tungsten alloys was produced with solid carbide approximately 50 fpm for 50 holes tool life. drills and not high speed steel drills as w^re the In Fig. 11, the tool life curves for tapping the other refractory, alloys. -.::: -';:: ~r - ip-5: ~' " refractory alloys ate shown. The tungsten material Tapping is one of the most difficult of all con- was tapped at the slowest cutting speed with the ventional machining operations. Figure 7 shows lowest tool life, approximately 14 to 15 holes. This the cutting speeds used in tapping the variety of operation was performed with the tungsten mate- alloys shown in the previous figures. As in the rial preheated to 600°F to improve the ductility. turning, milling, and drilling tests, the The molybdenum TZM alloy shows a strange tool molybdenum alloys exhibit the capability of the life response to cutting speed, reaching a peak of highest cutting s speeds, f approximately 50 fpm. approximately 120 holes but then dropping back as However, the niobium D31, the tungsten at 96% the speed was reduced. The tantalum alloy and the density, and the tantalum alloy were all tapped at niobium alloy were tapped at a speed range of 5 to cutting speeds of [10 fpm or less. The tungsten 10 fpm. alloy was even preheated to •iOO°F because of its Tables 1 through 5 show recommended machin- extremely low ductility at room temperature. ing and grinding conditions for tungsten, niobium Figure 8 shows a series of tool life curves de- D31, TZM molybdenum, molybdenum-0.5 titanium picting the relationship between tool life and cut- alloy, and tantalum-10 tungsten alloy, respectively. ting speed for four of the refractory alloys and All these data were produced * in U. S. Air Force other reference materials such as 302 stainless, report No. ASD-TRD-81. Table 6 shows recom- 4340 steel and Rene" 41. The molybdenum TZM mended conditions for machining niobium FS-85 alloy machined the best of the refractory materi- alloy. This'•alloy contains 28% tantalum, 10% tung- als, even better than 302 stainless, cutting between sten, 1% zirconium with the balance niobium. It 400 and 500 fpm for an acceptable tool life. The was heat treated in vacuum at 2400°F for one h. poorest machining of the four refractory alloys was This work was performed by C. E. Glynn at Gen- the tungsten material because of the relative eral Electric Corporation, Evendale, Ohio. insensitivity to reduced cutting speeds. The tan- Table 7 presents several grinding recom- talum alloy shows the tendency of increased tool mendations3 on: various V refractory materials. It life with reduced speed, and the tool life on the should be noted that some of the alloys require tungsten material seems to level off. lowerthan the normal 6000 fpm wheel speed in The relationship between cutting speed and tool order to maintain suitable surface integrity. These life in the face milling operation on the four considerations particularly involve the zirconium refractory materials and other referenced materi- material and its alloys. als is shown in Fig. 9. In this operation, the 302 stainless machines at higher cutting speeds than Nontraditional Machining of the molybdenum alloy but not in the turning oper- Refractory Alloys ation. Again, the tungsten and the tantalum alloys are the poorest machining, with the tantalum Two of the most popular nontraditional machin- showing better tool life response to cutting speed ing operations that have proven quite successful in than the tungsten material. The tool life curve machining refractory alloys are electrochemical with the tungsten alloy reaches a maximum of machining (ECM) and electrical discharge machin- about 14 or 15 in. of work travel and then drops as ing (DM). Both of these processes have been very the cutting speed is reduced. The niobium D31 successful in producing high tolerance, high quality alloy was machined satisfactorily at cutting speeds parts from difficult-to-machine alloys, including between 150 and 200 fpm. refractory materials as well as nickel-base alloys. The cutting speed and tool life relationship in Kotora4 gives some general guidelines on the appli- drilling is shown in Fig. 10. Drilling is generally a cation of wire EDM plutonium and depleted more difficult operation than either turning or face uranium as well as formed electrode cutting of milling, and the shapes of the tool life curves tungsten, beryllium, and tungsten-uranium sheet reflect this response. Notice the relative insensitiv- material to produce tensile specimens from" Vu-ih. ity of the tungsten material to changes in cutting thick sheet. speed. A tool life of approximately 10 to 15 holes Detailed summaries of both the ECM and the was produced, regardless of the speed, which was EDM processes are available in Sections 11.5 and varied between 90 and 150 fpm. The molybdenum (Text continues on page 145.) 434CJ (340 niu:)

434U (52 "cl KOI.Y - TZM D6AC (.!?II mini (56 "cl

302 SS HUI.Y - .5 TI (220 DUN) U70 BI1NI

1NCC 700 Ti A-110 (310 BHN)

Ti B-120 VCA UDIMET 500 (•100 BUN) (340 BUN) II B-120 VCA RENE 41 CB5 BHN) (365 WIN I Ti 6A1-4V Ti 6A1-4V (312 BUN) (312 BHN] A-286 Ti A-110 (320 BUM (310 BHN) 4340 Ti B-120VCA (52 Re) (285 BHN) 1 11-11 Tr D-120VCA (52 "c) (365 WIN) 1 INCO 700 HOLY TZH (302 WIN) (220 BUM) 1 UDIMET 500 HOLV-.5 Ti (340 BHN) 13 (220 BHN) 1 RENE 41 NIOBIUM D31 REFRACTORY (365 BHN) (207 BUM) 1 ALLOYS NIOBIUM D31 P S S TUNGSTEN V-l-28 THREAD 1 (14 HOLES) (207 BH1I) (93% DENSITY) PSS TUNGSTEN 5/16-24 THREAD, STUB TAP, 14 HOLES (400°F) FORGED TUNGSTEN (96% DENSITY) .95% DENSITY)

fa-1OW Ta-lOW 1/4-28 THREAD (220 BHN) Zl (220 BHN) , 1 1 1 1 I .002 .003 .004 .005 20 30 40

FEED-IN./REVOLUTION CUTTING SPEED-F7/VIN

Fig. 6 HAS. drilling aerospace alloys—Vi-ln.-dia X H-in.-deep thru hole, feed Fig. 7 Tapping aerospace alloys, relative cutting speed for tap life of 60 to for SO- to 100-hole drill life. 100 holes, Vn-18 NC Up, 75% thread. TABLE 1 Reoommended Conditions for Machining and Grinding Unalloyed Tungsten Nominal Chemical Composition, % 0 N C Mo Mn Si Nb W

Pressed and sintered .005 .002 .010 .02 .001 .01 .01 Bal. Forged and reaintered .005 .001 .010 .02 .001 .01 .01^ Bal. Arc cast .002 .001 .004 .50 BaL

Depth Width Cutting Wear- Tool Tool Tool of out) of cut, speed,; Tool land, Cutting Operatior. material geometry used in. in. Feed ft/min life in. fluid

Turning C-4 BR: -15°; SCEA: 15° H-in.-square .050 .009 in./ 200 ISmin .030 Soluble oil Pressed and sintered carbide SR: 0"; ECEA: 15° brazed tool bit rev (1 : 20) 93% density Relief: 5° 32-34 R, Nfo'/atin. Face milling C-4 AR: 0°; RR: 0° 4-in.-dia .080 .012 in./ 100 70 in./ .016 Highly chlo- Pressed and sintered carbide CA: 45" single-tooth tooth tooth rinated oil i 85% density Clearance: 15° face mill 90Bb Face milling C-4 AR: -15°; RR: 0° 4-in.-dia .060 1.250 .010 in./ 78 27 in./ .030 Highly chlo- Pressed and sintered carbide CA:45° single-tooth tooth tooth rinated oil

93% density Clearance: IS" face mill ; ~i "w '''•'•' 26 Rc Face milling C-2 AR:-15°; RR: 0° 4-in.-dia .060 1.5 .009 in./ 142 39 in./ .080 Soluble oil Forged and reBintered carbide CA:45° single-tooth tooth tooth (1 : 20) 96% density Clearancs : 15° face mill 35 V End mill slotting C-3 AR: 0°; RR: 0" lK-in.-dia .125 1.250 .003 in./ 200 45 in. .030 Soluble oil Pressed and sintered carbide CA: 45° X .060 4-tooth carbide tooth (1 :20) 93% density Clearance: 12° tipped end mill 34 R,. Peripheral C-3 AR: 0"; RR: 0° lW-in.-dia .125 .250 .004 in./ 140 110 in. .030 End milling end milling carbide CA: 45° X .060 in. 4-tooth carbide tooth : done with Pressed and sintered Clearance: 12" tipped end mill workpiece tern 93% density, peratureof 35R, 800°F End mill Blotting C-3 AR: 0°; RR: 0° lW-in.-dia .125 1.250 .003 in./ 200 26 in. .030 Soluble oil Forged and resintered carbide CA: 45° X .060 in. 4-tooth carbide tooth (1 : 20) 96% density, Clearance: 12° tipped end mill 35 R. Drilling C-2 118V90" notched .213-in.-dia W-in. .002 in./ 125 14 holes .030 Highly chlo- Pressed and sintered carbide point 29s helix angle thru rev rinated oil 93% density 7° clearance solid carbide 34 R, drill Drilling C-2 U8V90" notched .213-in.-dia fc-in. .002 in./ 150 15 holes .030 Highly chlo- Forged and resintered carbide point 29° helix angle thru rev rinated oil 96% density 7° clearance solid carbide 35 R, drill Drilling C-2 118^90° notched .213-in.-dia W-in. .002 in./ 150 9 holes .030 Highly chlo- Arc east carbide point 29T helix angle thru rev rinated oil 99% density 7° clearance solid carbide 31 Re drill -'"'•••

Tapping M-10 4-flute special V16-24NF H-in. 5 14 holes Tapping done Pressed and sintered HSS stub type plug thru \ with work- 96% density tap, 76% thread piece temper- 34 R, ature of 400-F

Surface Grinding Wheel apeed, Table speed, Down feed, Crow feed, Wheel grade Grinding Hold ft/min ft/mln in./pasa in./pau Gratia

32A46N5VBE KNOjjBolution 2000 40 .0005 .050 5.0 32A46N5VBE Highly sulfuri-ied 2000 40 .0005 .050 2.6 oil TABLE 2 Recommended Conditions for Machining and Grinding Niobium D-31 Alloy, 207 BHN (Nominal chemical composition: Ti, 10.0%; Mo, 10.0%; Nb, balance)

Depth Width Cutting Wear- Tool Tool Tool of cut, of cut, •peed, Tool land, Cutting Operation material geometry used In. in. Feed ft/min life In. fluid

Turning M-2 BR: 0°; SCEA: 0° H-in.-square .030 .005 in./ 60 40+ .030 Soluble oil HSS SR: 20°; ECEA: 5° solid HSS rev min (1:20) Belief: 5° NR:'/«in. Turning C-2 BR: 0°; SCEA: 0° H-in.-square .030 .005 in./ 300 40+ .010 Soluble oil carbide SR: 20°; ECEA: 5° brazed tool bit rev min (1:20) Relief: 5°

Face milling Super AR: 0°; ECEA: 5° 4-in.-dia .030 IK .010 in./ 135 50+ in./ .016 Highly chlo- HSS RK:20° single-tooth tooth tooth rinated oil CA:45° face mill Clearance: 10° Face milling C-2 AR: 0°; ECEA: 10° 4-in.-dia .030 2 .010 in./ 150 90 in./ .016 Highly chlo- carbide RR: 10° single-tooth tooth tooth rinated oil CA: 45° face mill Clearance: 10° End mill T-15 Helix angle: 30° Vi-in.-dia .060 .500 .003 in./ 100 200+ in. .008 Highly chlo- slotting HSS RR: 10° 4-tooth tooth rinated oil Clearance: 10° HSS end mill CA:45° Drilling M-l 118° plain point 125-in.-dia tt-in. .005 in./ 75 175+ .008 Highly chlo- HSS 7° clearance drill thru rev holes rinated l'A-in. long hole oil Reaming M-2 10° RH helix .213-in.-dia fc-in. .010 .005 in./ 125 105 .012 Highly HSS CA: 45° 6-flute chucking thru depth rev holes Rulfurized Clearance: 10° reamer hole on hole oil radius Tapping M-10 2-flute chip V. in. 28 NF V4-in. - - 12 50 holes - Highly chlo- HSS driver tap, tap thru rinated 75% thread hole oil

Surface Grinding Wheel ipeed, Table ipeed, Down feed, Cross feed, Wheel grade Grinding fluid ft/min ft/min in./pau in./pas8 G ratio

32A46K8VBE 5% KNO8 solution 2000* 40 .005 .025 7.5 32A46K8VBE 5% KNOj. solution 4000 40 .001 .050 4.5 32A46K8VBE Soluble oil (1 : 20) 4000 40 .001 .050 3.5

•If wheel Bpeed of 2000 ft/min is not available, use conditions for wheel speed of 4000 ft/min. MACHNNG REFRACTORY ALLOYS—AN OVERVEW 138

60. -i

56. -

i/l 40. A ill 39. H

10. -

0, 200. 400. 600. 800. CUTTING SPEED-FT/M1N

Fig. 8 Relationship between tool life and cutting speed in the turning operation.

125. -i m Jr 100. -

O / 302 STAINLLJS STEEL 75. h^ : \ HI 1 50. MOLYBDENUM TZH i LU

25. - o o 4340 52 Be

0. 200. 400. 600. 800, CUTTING SPEED-FT/MIN

Fig. 9 Relationship between cutting speed and tool life in the face milling operation. TABLES Recommended Conditions for Machicing and Grinding TZM Molybdenum, 217 BHN (Nominal chemical composition: Ti, 0.50%; C, .015%; Zr, 0.08%; Mo, balance)

Depth Width Cutting Wsar- Tool Tool Tool of cut, of cut, speed, Tool land, Cutting Operation material geometry used in. in. Feed ft/min life in. fluid

Turning C-2 BR: 0°; SCEA: 15° %-in.-Bquare .030 - .009 in./ 450 25min .010 Soluble oil carbide SR: 20°; ECEA: 15° brazed tool kit rev (1 :20) Relief: 5° NR: Varin. Turning C-2 BR: 0°; SCEA: 15° %-in.-square .060 .009 in./ 350 20min .010 Soluble oil carbide SR: 20°; ECEA: 15° brazed tool bit rev (1 : 20) Relief: 5° NR: Vsrin. Face milling T-15 AR: 0°; ECEA: 10° 4-in.-dia .060 2 .010 in./ 100 70 in./ .015 Soluble oil HSS RR:20° single-tooth tooth tooth (1:20) CA:45° face mill Clearance: 15° Face milling C-2 AR: 0°; ECEA: 5° 4-ia.-dia .060 2 .005 in./ 350 100 in./ .015 Soluble oil carbide RR:0° single-tooth tooth tooth (1 :20) CA:45° face mill Clearance: 10° End mill T-15 Helix angle: 30° «-in.-dia .125 .750 .004 in./ 190 78 in. .012 Soluble oil slotting HSS RR: 10° 4-tooth tooth (1:20) Clearance: 10° HSS end mill CA:45° End mill M-3 Helix angle: 30° K-in.-dia .125 .750 .004 in./ 190 142 in. .012 Soluble oil peripheral HSS RR: 10° 4-tooth tooth (1 :20) cut Clearance: 10° HSS end mill CA:45° Drilling M-33 118° plain point .250-in.-dia 14-in. .005 in./ 150 68 holes .015 Highly chlo- rev rinated oil HSS 7° clearance angle drill, thru " 2V4 in. long hole Reuming M-2 Helix angle: 0° .272-in.-dia .010-in. .015 in./ 60 51 holes .012 Highly chlo- HSS CA:45° 6-flute thru depth rev rinated oil Clearance: 10° chucking reamer hole or. hole radius

Tapping M-10 4-flute plug, V16in.24NF V4-in. - - 70 100+ - Highly chlo- HSS 75% thread tap thru holes rinated oil hole Surface Grinding

Wheel speed, Table speed, Down feed, Cross feed, Wheel grade Grinding fluid ft/min ft/min in./pass in./pass G ratio

32A46N5VBE 5% KNOZ solution 2000* 40 .001 .050 25 32A46L8VBE 5% KNO2 solution 4000 40 .001 .050 13

•If wheel speed of 2000 ft/min is not available, use conditions for wheel speed of 4000 ft/min.

TABLE 4 Recommended Cutting Conditions for Machining and Grinding Mo-0.5 Ti Molybdenum Alloy (Nominal chemical composition: Ti, 0.45%; C, .020%; Mo, balance)

Depth Width Cutting Wear- Tool Tool Tool of cut, of cut, speed, Tool land Cutting Operation material geometry used in. in. Feed ft/min life in. fluid

Turning C-2 SR: 20°; SCEA: 15° W-in.-square .060 - .009 in./ 300 25min .010 Soluble oil carbide BR: 0°; ECE.': 15° brazed tip rev (20:1) Relief: 7° tool bit Face milling C-2 AR: 0°; ECEA: 10' 4-in.-dia .060 2 .005 in./ 225 120 in. .012 Soluble oil carbide RR: 0°; Cl: 15° single-tooth tooth (20: 1) CA: 45° face mill Drilling M-l 2-flute, 118° ,193-in.-dia .500 - .005 in./ 100 100 .012 Highly chlo- HSS plain point. drill, thru rev holes rinated oil 7° clearance 2'A-m. long hole Reaming M-2 10° RH hehx 6-flute straight .500 .010 .015 in./ 85 45 .010 Highly chlo- HSS 45° CA shank chucking thru depth rev holes rinated oil 10° clearance reamer hole on hole radius Tapping M-10 2-flute chip M-in. 28 NF .500 56 100+ Tap still Highly chlo- HSS driver tap, thru holes cutting rinated oil 80% thread hole

S. rface Grinding

Wheel speed, Table speed, Down feed, Cross feed, Wheel grade Grinding fluid ft/min ft/min Ln./pass in./pass G ratio

32A4W8VBE Soluble oil (40 : 1) 4000 40 .001 .050 3.3 142 OffBTOMER

125. T

50. 100. 159. 200. CUTTING SPEED-FT/MIN

Fig. 10 Relationship between cutting apeed and tool life in the drilling operation.

125. -i

MOLYBDENUM TZM

100. -

LU _l O 75. X I Ui LL

a. 50. a

25.

0. 0. 20. 10. 60. 80, CUTTING SPEED-FT/MIN

Fig. 11 Relationship between cutting gpeed and tool Ufa in the tapping operation. TABLES Recommended Conditions for Machining and Grinding 90 Ta-10 W Tantalum Alloy, 207-241 BHN (Nominal chemical composition: Ta, 89%; W, 10%)

Depth, Width Cutting We»x- Tool Tool Tool of cut, of cut, speed, Tool land, Cutting material geometry used in. in. Feed ft/min life in. fluid

Turning M-2 BR: 0°; SCEA: 0° %-in.-square .030 - .009 in./ 50 44min .030 Soluble oil HSS SR: 20°; ECEA: 5° Bolid HSS rev (1:20) Relief: 5° NR: VM-in. Turning C-i. BR: 0"; SCEA: 0" %-in.-square .030 .009 in./ 75 27min .010 Soluble oil carbide SR: 20°; ECEA: 5° brazed tool bit rev (1:20) Relief: 5° NR: VM-in. Face milling Super AR: 0°; ECEA: 10° 4-in.-dia .030 1.125 .010 in./ 80 53 in./ .016 Soluble oil HSS RR:20° single-tooth tooth tooth (1:20) CA: 45° face mill Clearance: 10° End mill T-15 Helix angle: 30° V4-in.-dia .060 .500 .002 in./ 70 80 in. .012 Soluble oil slotting HSS RR: 10° 4-tooth tooth (1:20) Clearance: 15° HSS end mill CA: 45° X .040 in. End mill M-2 Helix angle: 30° V^-in.-dia, .060 .500 .002 in./ 65 70 in. .012 Soluble oil peripheral HSS RR: 10° 4-tooth tooth (1:20) cut Clearance: 6° HSS end mill CA: 45° X .040 in. Drilling M-l 118° plain point .125-in.-dia tt-in. .002 in./ 50 125 .015 Highly chlo- HSS 7° clearance angle drill, thru rev holes rinated oil 2%-in.-long hole

t TABLE 6 Recommended Condition* for Machining Niobium FS-86 AUoy* (90 Rb hardened)

Patting Tool Tool Tool Took Depth, Feed, •peod, Tool wear, Operation material geometry oied in. ipr •fm life in. Fluid Eenurka

Turning Carbide SR 15"; BR: 0° Toolholder .030 .008 140 27 .005 Trim Higher speeds 883 SCEA: 15° KSDR85C sol, up to 400 Bfm; Relief: 6°; NR: .06 20 : I poor finish Turning HSS SR: 15"; BR: 0° Toolholder .030 .005 110 24 .015 Trim Higher Bpneds M42 SCEA: 15° KSDR 85C .020 sol, up to 150 Bfm; Relief: 6°; MR: .06 20: 1 poor finish End HSS 30° helix 4-fluta .020 .001 80 50 .002 — No surface scratch milling M42 25-30° axial 1.0-in. dia C/T sa with 10* axial 0° radial W-in. radius Drilling HSS 30° helix NAS907 M-in. .002 39 35 .002 Trim Rings in holes; M42 30° axial TypeJ thru 600 rpm uniform sol, finish may require W-in.-dia .004 mist flooded cutting oil local or reaming Drilling HSS 30° helix NAS907 .089 .0002 .18 16 Trim Tool broke on M42 30° axial TypeJ thru 2200 rpm holes sol, 17th hole Vffl-in. dia miBt Drilling HSS 3G° helix NAS907 .089 .0005 29 61 .001 Trim Stopped test; M42 30° axial TypeJ thru 1800 rpm holes sol, chattered; %,-in. dia wist chip welding Grinding A1A 97A60I6VFM 10-in. dia X .0002 .050 3120 - 4.0 Oil, 30to50rms 1 in. X 3 in. D/F C/F G ratio Gl&S

•Work performed by Glynn,3 Conventional Machining Laboratory, General Electric Company, Evendale, Ohio, 1973. MACHNNG REFRACTORY ALLOYS—AN OVERVEW 146

TABLE 7 Grinding Refractory Alloy*

Wheel Infeed or Cross Wheel sneedor downfeed, feed. Table speed, Grinding Surface finish, Operation Alloy grade work speed in./pan in./paw fpm fluid RA

Cylindrical Mo-05Ti A60H9V 6000 fpm .001 .030R/.010F 50 Water base 13R/7F Cylindrical 50Mo-50W A60H9V 6000 fpm .001 .030 70 W/B, det. 7/20 Centerless Pure Mo C36IB 650 rpm - - 55 ipm - 90 Centerless Pure Mo C320L8E 370 rpm - - 30ipm - 5 Centerless 70 Mo-30 W C60H11V 120 rpm - 60 ipm Water base 60/90 - - C80H11V 200 rpm .0005 - 60 ipm Water base 8 Suiface Pure W C3U4V 8000 .001 .030 50 ipm W/B, det., 1: 40 14 Cylindrical Pure W C60I8V 5300/ .025 - - 2.5% nitride 20/40 150 rpm amine Centerless Pure W C60J4V 6500 .001 - W/B, det. 17 Surface 99 Nb-1 Zr A60H8V 6000 - .032 50 W/B or dry 30/90 Surface Zr & alloys C60K5V 1500/3500 - - - Oil 25 Surface Zr & alloys C100N7V 16C0 _ _ _ Oil _ Surface Zr & alloys A80M8V 3500 10% nitride - amine Cylindrical Zr & alloys C60K4V 1500/3500 - - - - -

5 12.5 of the Machining Data Handbook By follow- Report No. ASD-TRD-581, USAF Contract 33(600)-42349, ing the guidelines in these two process summaries Wright Patterson Air Force Materials Laboratory, July and using the correct equipment, high tolerance 1963. parts can be machined using these two nontradi- 2. C. E. Glynn, Machining of Columbium (FSS5 Alloy), FPD tional processes from virtually any refractory alloy General Electric Corporation, Evendale, Ohio. that is electrically conductive. [The editors have 3. Machining Difficult Alloys, Chapter 20, pp. 229-244, spon- chosen not to reprint these two sections of the sored by the U. S. Air Force, ° American Society for handbook. Interested readers are encouraged to Metals, 1962. review this information.] 4. John Kotora, Jr., Machining Exotic Materials with EDM, Rpt. SME (ASTME) MR68-901, Argonne National Labora- REFERENCES tory. 1. N. Zlatin, M. Field, and J. Gould, Technical Report an 5. Machining Data Handbook, 3rd Ed., Vol. 2, Metcut Research Machining Refractory Materials, Technical Documentary Associates Inc., Cincinnati, Ohio, 1980. Welding of Refractory Alloys

G. G. Leasmann Westinghouse Research and Development Center

INTRODUCTION creep strength (primarily) were resolved, and emphasis was given to tantalum-base alloys which This review primarily summarizes welding evalua- combined fabricability, toughness, and high tem- tions supported by the National Aeronautics and perature creep strength. The evaluation of tungsten Space Administration-Lewis Research Center in (and the W-25Re alloy) continued quite indepen- the 1960s. dent of fabricability considerations because of its A literature search run in preparation for this high temperature potential. A special aspect of this review indicates that more recent work is modest work was the screening of these materials for by comparison. Hence, this review restates thesa longtime structural stability (10,000 h at 1315°C). accomplishments briefly and addresses opportuni- Special tools useful to the welding engineer ties which have evolved in welding technology were also perfected in these programs. Examples (such as lasers) in the intervening decade. include adaptation of the Varestraint test to Emphasis in this review is given to tantalum- refractory metal welding for sheet thickness; and niobium-base alloys. These were perceived conci3e methods for using the bend test as an inex- then, as now, as the most promising structural pensive screening tool for quality and process con- materials for Rankine cycle space power systems trol as well as a sensitive weldability test; an using liquid metal working fluids. Considerable appreciation for weldability trade-offs in alloy work was also done to assure that a consistent development including creep strength as related to comparison was made with tungsten. Thus, tung- multipass welding; and methods for control of sten was recognized for its potential for very high welding atmosphere and the influence of materials temperature operation despite practical fabrication and component designs in inert gas handling sys- problems associated with its high ductile-to-brittle tems. transition temperature. The bibliographies of this report and others These evaluations were extremely thorough but presented at this meeting represent substantially in a focused and practical manner. A wide variety the basis for the current technology status for of candidate alloys derived primarily from develop- welding refractory metal alloys. ments directed at aircraft propulsion applications were available. Early efforts by NASA were directed at screening studies to select promising ALLOY TECHNOLOGY STATUS structural alloys for the space power application. This objective required fine tuning of welding procedures, e.g., the demonstration of stvingent stan- Background f dards or control of welding atmosphere to assure It is doubtful that one would characterize the good co.-rosion resistance in liquid alkali metals. refractory metal alleys as commercially available. These guidelines remain today as definitive meth- Most likely, they would be supplied on a single-lot odology for welding react'.ve and refractory metal basis with close technical following to substitute alloys. for the confidence attributed to the more common Eventually, the trade-off requirements for use use of well-established specifications. The designer of these materials in terms of fabricability and and fabricator can be expected to play a more inti-

146 WELWNQ OF REFRACTORY ALLOYS 14T

TABLE 1 tantalum alloys. These form complex systems with Alloy IdttttifUatteii the other elements but are alloyed at levels below their equilibrium solid solubility limit Hence, AUoy* Nominal eompocitioti, wt% wrought structures are single phase, while cored weld structures may be multiphased. Nb-lZr Nb-lZr Strengthening is realized through both solid solu- AS-55 Nb-5W-lZr-O.06C -1- Y B-66 Nb-6Mo-W-lZr tion and dispersion because hafnium and zirconium C-129Y Nb-lOW-lOHf + Y tend to form very stable precipitates with the Cb-752 Nb-10W-£5Zr residual interstitials. Reactive metal additions D-43 Nb-lOW-lZr-O.lC enhance corrosion resistance in alkali liquid FS-85 Nb-27Ta-10W-lZr metals. SCb-291 Nb-lOW-lOTa D43 +Y Nb-10W-lZr-0.1C + Y Several alloys also contain intentional carbon T-Hl Ta-8W-2Hf additions. These alloys respond to thermal treat- T-222 Ta-9.6W-2.4Hf-O.01C ment during processing and realize their strength ASTAR 811C Ta-8W-lBe-lHf-0.02C in part from carbide dispersions. In this respect a Ta-lOW Ta-lOW W-25Re W-2SRe knowledge of the probable phase relations is W Unalloyed important, and many of these relationships have W-25Re-30Mot been investigated. Additions of minor amounts of yttrium in sev- •All alloys from arc-cast and/or electron eral niobium alloys provide an interesting modifi- beam melted material. t(a/o), all other (w/o). cation of mechanical properties resulting primarily in improved ductility. Yttrium is essentially insoluble in niobium and is very reactive with oxygen. Hence, the most probable mechanisms for mate role than normal in following material from improved ductility are an effective reduction in ingot to finished product. Because we can antici- matrix oxygen level by preferential combination pate alloy modifications in the future, the user can with yttrium, purification during melting and gain some insight into the future based on the met- welding by slagging of the oxide, and grain refine- allurgical background of the many alloy systems ment resulting from the presence of the highly sta- given early consideration. ble oxide. The grain refining influence of yttrium is An early grouping of available alloys is given in evident in as-received material. AS-55 and C-129Y Table 1. This list at first appears formidable but are recrystallized at the highest temperature simplifies very easily because of similarities iu among the niobium alloys yet they have the small- metallurgical behavior. The list simplifies even fur- est grain size. ther because the tantalum-base alloys have the Rhenium in tungsten is an unusual solid solu- most promise based on high temperature creep tion addition which improves ductility. Rhenium strength and fabrieability. Among the niobium- added to ASTAR-811C is a substitute strengthener base alloys, only the most fabricable and simple, allowing reduction in the hafnium for improved such as the Nb-lZr alloy, are of interest and only high temperature creep strength. for lower temperature application. The two tungsten alloys remain aloof—very difficult and risky A simplified summary indicates'that the most in fabrication. promising alloys are the family of alloys T-lll, T-222, and ASTAR-811C for high temperature Phase relationships in these alloys, which are creep. These alloys have tungsten and rhenium for basically uncomplicated, prove helpful in under- solid solution strengthening, hafnium or zirconium standing these alloys. Complete solid solubility is for liquid metal corrosion resistance and dispersion demonstrated by all combinations of Nb, Mo, Ta, strengthening, and some have additional carbon for W, and V as employed in these systems. Hence, dispersion strengthening. The Nb-lZr alloy is of these are mutual single-phase solid solution most interest among the niobium alloys because of strengtheners. The predominant element of this good fabricability and corrosion resistance. group in both the niobium and tantalum alloys is tungsten. Two niobium alloys also contain tantalum, and one contains molybdenum and vana- Welding Processes i dium instead of tungsten. The processes which have been selected One cr the other of the reactive correctly for welding the refractory metal alloys elements—zirconium or hafnium—is a necessary can be classified as the nonccmsumable electrode component in all the high strength niobium and processes. These are typified by gas tungsten arc 148 LE88MANN welding and electron beam welding. Laser welding can now be added as a typical method to consider. The weld joint designs are typical of those with an autogenous chemistry. Welds are either fusion welds joining two parts without filler metal, or, if filler wire is used, it is of base metal composition. The exception is that several of these alloy systems are weldable aa bimetals as discussed later. Also, special case joining processes are possible: hot isostatic pressure welding, inertia welding, explosive welding, plasma and resistance welding. Any welding process must provide freedom from interstitial and metallic contamination. The potential for metallic contamination is minimized LONGITUDINAL by proper joint preparation (machining followed by BEND pickling), careful handling, and special inserts in holding fixtures. Interstitial contamination is lim- DUCTILE BENDS ited primarily by control of the welding atmosphere. 100

Welding Atmosphere Control | DUCTILE-BRITTLE eo • BEND TRANSITION The refractory metal alloys are extremely reac- 1 TEMPERATURE -70'C tive and act as getters when heated in contami- 60 nated atmospheres. Contaminants are taken into solution in the weld with a resultant loss in ductil- 40 ity, liquid metal corrosion resistance, usually a loss FRACTURE in strength, and a change in thermal stability. Optimum shielding gas quality control reduces the 20 200 -150 -100 -50 50 100 °C risk of unpredictable behavior of welds in service. nil I I Good data are available to quantify the contamina- 1 -300 -200 -100 6 100 200 "F tion effects. The bend test ductile-to-brittle transi- TEST TEMPERATURE tion temperature provides a sensitive indication of the effect of oxygen and nitrogen contamination. The bend test is described in Fig. 1; the sensitivity NOTE: ARROWS SHOW ROLLING DIRECTION to oxygen pickup in welds and base metal is shown THICKNESS, t = 0.9 mm PUNCH SPEED = 1 IPM in Fig. 2. WIDTH = 12t TEMPERATURE - VARIABLE LENGTH - 24t PUNCH RADIUS - It With careful control of the welding atmosphere, TEST SPAN = 15t very little contamination of welds occurs.2 As an example, an overnight pumpdown cycle on a Fig. 1 Bend test description. vacuum-purged weld chamber complemented by a 90°C chamber bakeout yields a vacuum of make atmosphere control much simpler than was 6 5 <5 X 10~ torr and <3 X 10" torr/min leak possible 20 years ago. The stability of the welding rate. An ultrahigh purity helium backfill produces atmosphere after backfilling of a vacuum-purged an atmosphere with <1 ppm total impurity level weld chamber is summarized for various purge and permits several hours of welding with oxygen cycles in Figs. 3, 4, 5, and 6. The most important and moisture each held at <5 ppm. Niobium and conclusion that can be drawn from these results is tantalum alloys welded in this atmosphere show no that control and monitoring of welding atmo- pickup of nitrogen or carbon within 95% confi- spheres is an entirely realistic requirement in dence limits and actually show a loss in oxygen welding refractory metal alloys. Other conclusions content between 16 and 40 ppm. drawn from this reference are: The use of oxygen and moisture monitors per- 1. By using optimum evaluation and backfill- mitted the development of very definitive guide- ing techniques, a high quality welding lines for management of shielding gas 3 atmosphere having <1.25 ppm total active atmospheres. The basic results are unchanged impurities can be obtained in vacuum- today except that better atmosphere monitors purged chambers. Following backfilling, the WELDING OF RffHACTORY ALLOYS 148

1000 [- 4.-A- MANUAL WATER COOLED, -r TYGON INSULATeD TORCH INSTALLFD 4X ULTIMATE PRESSURE ANU 200 IEAK SATE LISTED 200 100

0 o

-200 -100

-400 -200 .. I -300 200 <00 600 800 1000

2 BllIYI RUBBER 5 * 1C"6 ion 01 ti/MIN) |.oooj- £ ; — 1(538 <

•A200 5 z z 2 2X 100 2 o i e -100 ^ I 4 6 b I? 12 14 16 IB 2! 24 £ -200 -200 ~ TIME, HOURS AFTER BACKFILLING CHAMBER 6UJ.894A J, -400 -300 E Fig. 3 Effect of glove material on oxygen contamination 200 400 600 BOO rcte.

r 24 -

1 n - MANUAL WATER COO1FD TYiJ"'~ IN5IILA7FII Io«CH INSTALLED 20 ULTIMATE PRESSURE AND LEAK RATE LISTED

200 400 600 BOO 1000 l?00 J-300

4000 6000 10.000 >2.00C

ATOMIC,

KEY 1000°f - 538°C . HIGHEST TEST TlMPf"*TU«E -32O"F--195°C = LO*ES. TEST T!MPi»*TU«E O = UN^ILDED • • MELDED LATEX NATURAL RUBBER 7 K 10 lofr (.007p/MIK 1800°F - 98O°C DlP'L'S'ON AN*4:ALING TEMP. CEMENT NEOPRENE 2 « lo"6 lorr (.01 H/MIN) I I I I I I 1 Fig. 2 Bend ductility vs. oxygen content for 3 alloys. 6 8 10 12 14 16 18 20 72 24 TIME, hOURS AFTEK BACKFILLING CHAMBER 60489SA Welding atmosphere gradually deteriorates, permitting 6 or more hours use depending Fig. 4 Effect of glove material on water vapor contamina- on the contamination limit established for tion rate. the particular run. 2. The sources of moisture and oxygen con- chamber atmosphere depends on the gas tamination in weld chamber atmospheres quality, weld box tightness, a moderate differ considerably. Consequently, these evacuation of 10~4 to 10~5 torr, and the containments are not related ard must be backfill techniques employed. The oxygen considered independently. level increases following backfilling mainly 3. The oxygen level in the backfilled weld by diffusion through the weld box gloves. 160 LESSMANN

24 MANUAL WATER COOLED IYGON INSULATEO TORCH INSTALLED 22 ALL GLOVES CEMENT NEOPRENE ULTIMATE PRESSURE, LEAK BATE, AND 20 EVACUATION TIME LISTED •CORRECTED VACUUM GAGE FOR HELIUM 18 SENSITIVITY

16 DOUBLE EVACUATION WITH He PURGE 5 U 2.5 < I0" >"r (.27 H/KAIN) 1ST EVAC. DOUBLE EVACUATION WITH He PURGE 2.5 x I0"5 lotr (.27 (i/VHN) 1ST EVACUATION 6.5 x ID"5 loci (.60| tO MIN. TOTAL 5 8 MANUAL WATEK COOLED, TyGON 2.6. 10 tor. (,26,j/MIN) INSULATED TORCH INSTALLED 52 MIN. ALL GLOVES CEMENT NEOPRENE ULTIMATE PRESSURE. LEAK RATE, AND EVACUATION TIME LISTED •CORRECTED VACUUM GAGE FOR (.01 p/MIN) 18 HOURS He SENSITIVITY I I, 0 12 3 4 5 6 TIME, HOURS AFTER BACKFILLING CHAMBER

i— (01 H/MIN) Fig. 5 Effect of evacuation pressure on oxygen contami- / 18 HC nation rate. 0 1 2 3 4 5 6 TIME, HOUS5 AFTES BACKFILLING 'HAMBER 4. Low moisture levels in the backfilled weld chamber atmosphere are obtained by using Fig. 6 Effect of evacuation pressure on water vapor con- extended pumpdown cycles, conveniently tamination rate. overnight for 16 to 18 h, to the low 10~6 torr range. Atmosphere stability with nitrogen contamination assumed by impli- respect to this impurity is enhanced by cation to be 4 times the oxygen level. The longer and lower pressure pumpdowns source of nitrogen, like oxygen, must be air because outgassing of the chamber interior leaks and diffusion through gioves. and tooling surfaces is the primary source 7. Other active atmospheric contaminants of moisture. that are not generally airborne, such as 5. The leak rate (pressure rise rate of a sealed hydrocarbons, are avoided by judicious chamber) is an excellent measure of the selection of materials, lubricants, and adequacy of a pumpdown cycle because it cleaning techniques for internal chamber represents the sum of leakage and outgas- components. sing rates. Hence, a low leak rate assures 8. Neoprene gloves provided the best overall low moisture and oxygen rise rates in the performance of those tested except in spe- backfilled chamber. A 1 min pressure rise cial applications where sulfur contamina- in the evacuated chamber of 3 X 10~5 torr tion may be detrimental. All the gloves is required for reasonably good stability of tested were permeable to air. Kence, for the the backfilled chamber atmosphere. In this same size and thickness glove, the greater respect, double purge cycles are not benefi- the number of gloves used and the smaller cial. Contrary to a widely held opinion, the weld chamber the more rapid will be welding can be accomplished under a the deterioration of a weld box atmosphere. slightly negative pressure (below 1 atm) without increasing the contamination rate if a good leak rate is obtained and sound Weldability gloves are used. The weldability of refractory metal alloys as 6. Nitrogen as a contaminant appears to be defined by weld quality and variability of weld present in the weld atmosphere in roughly properties has been studied in some detail.2*4 A dis- the same ratio with respect to oxygen as in cussion of the results is handled most easily by air. Hence, oxygen can be monitored and grouping the tantalum and niobium alloys together WB-DWQ OF REFRACTORY ALLOYS 161 as fabricable alloys and by discussing the tungsten transition temperature. Weld process and alloys separately as special purpose high tempera- metallurgical factors combine such that a ture materials which are difficult to fabricate. heat input threshold for ductility impairment Basic weldability is defined primarily as the is observed foi alloys which are dispersion ability to produce defect-free welds over a reason- strengthened. With increased grain stability, able range of parameters, material thicknesses, as realized with the yttrium-modified alloys, and configurations. This was accomplished by this threshold occurs at a higher heat input. welding both sheet and plate materials in both The solid solution alloy did not display this unrestrained and restrained configurations. threshold but rather a continuous ductility An extended but very important definition of loss with increasing size of the heat-affected weldability is the sensitivity of weld properties to zone. welding parameter variation. Alloy systems were • The thermal analysis interpreted in terms of categorized and compared by using the bend a heat input threshold provided a sensitive ductile-to-brittle transition temperature (Fig. 1) as rationale to which the general improved a sensitive indicator. ductility of electron beam welds can be This review avoids presentation of great detail ascribed. on weldability by first presenting the conclusions of 2 The differences between these alloys in terms of Lessmann and then dwelling only on selected sensitivity to welding are summarized in Fig. 7. major aspects. Fabrication using most of these Several important characteristics can be compared. alloys is very practical as can be recognized from the following conclusions. • Maximum transition temperature: lower is better. • Good weldability was exhibited by the • Increase in transition temperature compared niobium and tantalum alloys as demonstrated to base metal: less is better. by restrained weld tests and general accom- • Total spread in transition temperature: less is modation in welding both sheet and plate. better because it implies less sensitivity to Weldability limitations were exceeded for sev- welding variations. eral alloys as described later. Room temperature and elevated temperature weld strength Hence, an excellent measure of weldability in the approached base metal strength for these conventional sense was provided in the weld alloys demonstrating joint efficiencies at all ductility response study. temperatures of nearly 100%. Within the The inherent flexibility in selecting weld respective alloy groups, FS-85 and T-lll dem- parameters which produce sound welds can be onstrated superior combinations of strength gaged from the data summarized in Table 2. The and fabricability. percent of acceptable welding conditions is listed • Welding resulted in a loss of ductility in all for each alloy along with the defect source of unac- alloys as measured by the bend ductile-to- ceptable welds. brittle transition temperature. The compara- An electron beam process limitation for most tive degradation of ductility occurring with alloys at the highest welding speed of 100 ipm was welding provides a convenient measure of evidenced by welds having unacceptable contours. weldability in these systems. Piste weldabil- Some alloys also welded poorly at other electron ity was comparable to sheet weldability for beam parameter combinations. In this respect, B-66 the more fabricable alloys. However, with the welded with difficulty presumably because vana- less weldable alloys, adverse welding charac- dium tends to boil off more readily than other teristics were exaggerated in plate welding. alloying elements. • Tantalum alloys were considerably less sensi- B-66 also welded with greater difficulty in both tive to welding than niobium alloys and, as a manual and automatic arc welding due to a hot result, have superior fabricability. Tungsten tearing tendency. This alloy is a classic example of was not easily welded because of brittleness this effect, which results from an excessive freez- at room temperature: W-25Re is hot tear sen- ing point depression and liquidus-solidus sitive as well as brittle. separation.5 • Niobium alloy weld behavior was rationalized Arc welds in C-129Y welded at the highest with a thermal response analysis. Welding speed had gross shrinkage defects indicating a lim- conditions that tend to stimulate development itation in welding this material. D-43Y tended to of the heat-affected zone and grain size in hot tear along the weld centerline, have porosity, this zone increase the weld ductile-to-brittle and crack (transverse) during welding. Porosity 152 LHSSMANN

1\ 1 —- 500 900 / > 800 HIGHfcbl Ibbl / TEMPERATURE —' — 400 700 \

600 vn - 300 500

400 200

300 Y g 100 I 200 I r*2t 1 (Owtai itn 1 ,—TEMPERATURE —- 0 0 / -320 °F -100 t- 1 I - -100 -200 1 i _L k 1 i • 1 -300 1 / 1 S: -f- = -200 BAS E BAS E E.B . E.B . BAS E E.B . BAS E E.B . BAS E GT A E.B . BAS E E.B . GT A BAS E E.B . E.B . GT A GT A GT A BAS E E.B . BAS E E.B . BAS E E.B . GT A BAS E GT A BAS E BAS E GT A E . B GT A E.B . GT A GT A GT A GT A

C-129Y SCb-29 FS-B5 Cb-752 D-43Y AS-55 B-66 D-43 T-111 To-lOW T-222 W-25R* W

* - All bends H radiui except as noted.

Fig. 7 Summary of bend test results for gas tungsten arc and (ilectron beam butt welds in 0.05-in. sheet Twelve welding conditions for each alloy and process. Weld tested in both longitudinal and transverse directions.

also occurred extensively in the unmodified D-43 Except for hot tearing in B-66, plate welding but was controllable, as explained later in this was accomplished with relative ease with all the report, by special joint preparation. All three available tantalum- and niobium-base alloys. yttrium-modified alloys demonstrated a weld centerline weakness, the source of which is not Tungsten-Base Alloys apparent but is presumed to be an effect of yttrium. Special efforts4 were made to assess the potential of improved weldability of tungsten through Tungsten and W-25Re welded with difficulty alloying and through alternate base metal process- because of high transition temperatures. This ing, namely powder-metallurgy vs. arc-cast prod- resulted in fracturing on weld cooling to ambient uct. Of particular interest were: temperature. Electron beam welding was not a panacea in this respect, performing a little worse • Ductility improvements associated with than gas tungsten arc welding. For both processes rhenium and molybdenum alloying. slower speeds and. hence, higher heat inputs pro- • Advantage of excellent atmosphere control. vided solid welds. In contrast, the thermal shock • Advantage of high preheat, 760°C. associated with faster speed, low heat input, and • Advantage of postweld annealing. the brittleness of these alloys produced defective • Effect of longtime high temperature expo- welds. Cleavage cracks occurred in both the longi- sure. tudinal and transverse directions. W-25Re gas The basic results of this s'udy are shown in tungsten arc welds also displayed axial hot tearing Tables 3, 4, 5, and Fig. 8. In this study all three along the weld centerline and weld interface. alloys were from arc-cast base, but the ternary was WELOWQ OF BB=RACTORY ALLOYS 153

TABLE 2 Weldability Bwillt for Sheet Weld Study

Parameter combinations Caused) producing acceptable welds, % for weld rejection* Alloy Tungsten arc Electron beam Tungsten arc Electron beam

C-129Y 83 '5 7 SCb-291 100 83 1 FS-G5 100 83 1 Cb-752 100 83 1 D-43Y 57 75 3,5,6 1.2 AS-55 100 83 1 B-66 67 50 3,4 2 D-43 83 83 5 1 T-Hl 100 92 1 Ta-lOW 100 100 T-222 100 83 1 W 50 t 8 t W-25Re 62 39 8,3 2,8

•Causea for weld rejection listed below by number (1) Rippled weld appearance occurring in 100 ipm welds, visual reject. (2) Coarse appearance, visual reject. (3) Hot tearing. (4) Microfissures at 60 ipm (destructive test reject). (5) Porosity. (6) Transverse cracks. (7) Shrinkage voids in 60 ipm welds. (8) Transverse and longitudinal cleavage cracks. fFull set of weld paraRteters not developed due to interference of delamina- tions occurring at the joint during welding. also evaluated as a powder-metallurgy product. W-25Re: The weldability of W-25Re is One effect of alloying tungsten for improved ductil- improved over th?t of unalloyed tungsten ity is that the melting point is lowered. Hence, the because of slightly better as-welded and application temperature limits are probably also recrystallized ductility. Improved ductility lowered. coupled with a lower melting point makes this alloy less susceptible to thermal shock Melting point, °C failures. However, the W-Rt phase relationships are such that this alloy exhibits a ten- Unalloyed tungsten 3410 dency toward hot tearing. W-25Re (w/o) 3125 Preheat in welding was not beneficial in W-25Re-30Mo (a/o) 2900 improving as-weided ductility but permitted welding at higher welding speeds and, hence, The basic conclusions of this evaluation are: essentially improved weldability. • Unalloyed tungsten: The weldability of unal- A stress-relief postweld anneal (1405°C) loyed tungsten is marginal because of its high was beneficial for electron beam welds. This dnctile-to-brittle transition temperature in implied high residual stress in electron beam the welded or recrystallized condition. The welded W-25Re tends to correlate with the high melting point and low ductility in combi- thermal shock behavior observed for W elec- nation make tungsten susceptible to failure tron beam welds. Gas tungsten arc welds by thermal shock during welding. Hence, were not improved by stress relief but instead weldability is enhanced by high weld preheat. required a solution anneal (1800° C) implying It is not apparent that use of arc-cast tung- that sigma phase develops at grain bound- sten is advantageous over powder-metallurgy aries during gas tungsten arc welding. In this tungsten except for absence of porosity in espect, electron beam welding was advanta- welds. Postweld annealing was not particu- geous because embrittlement by the sigma larly beneficial in improving ductility. phase and hot tearing were observed only in 164 LESSMMM

TABLES Gu Tungsten Are Weld Parameter Evaluation*

W-2SSe-30Mo (a/o). Unalloyed tungsten, W-25Re (w/o), arc-cast arc-cast Powder-metallurgy Arc-cast Weld speed, No 230-315°C 760°C No 230-31S°C 760°C No 425°C 760°C No «5°C 760°C ipm preheat preheat preheat preheat preheat preheat preheat preheat preheat preheat preheat preheat

Small welds, 3.0 mm wide, nominal (2.0 mm to 3.6 mm)

3.0 L540 L425 • • L,T 7.5 L>540 >540 L425 L310 * • L454 n •DD DD 15 L540 T425 L315 T54i L205 L.T230 L120

25 L290 L230 D D ••L.T •L230 30 L370 L>540 >540 L260 • T175

35 L205 L290

45 L.T425

Large welds, 4.6 mm wide, nominal (3.8 mm to 5.3 mm) • a 7.5 L370 L370 L540 D • D 15 L540 L>540 1230 D 30 L, T425 L>540

, sound weld; D, defective weld; 4t bend transition temperature in °C indicated for longitudinal (L) and transverse (T) bends.

gas tungsten arc welds. Both the development On aging, this alloy tends to behave quite of sigma phase and hot tearing result from simply as a solid solution system. However, constitutional segregation on freezing which the powder-metallurgy material exhibited is apparently more pronounced in gas tung- secondary recrystallization which could well sten arc welds. be a metallurgical instability brought on by W-25Re-30Mo (a/o): The W-25Re-30Mo alloy the dissolution cf highly dispersed impurity displayed generally excellent weldability precipitates. except for an extreme sensitivity to oxygen • In several checks made in this program, welds contamination which causes hot tearing. in powder-metallurgy product always con- Undesirable levels of oxygen contamination tained porosity, whereas arc-cast material occur &u a very low level in the base metal produced porosity-free welds. making detection difficult. Welding atmospheres, however, can be easily controlled if Thermal Stability properly monitored to eliminate welding as a potential source of contamination. The structural effects of longtime exposures at A postweld stress relief was beneficial in elevated temperatures have been determined for improving the ductility of welds in this alloy. promising refractory metal alloys.6-7 This program Otherwise, all thermal treatments to which provided the needed detailed definition of the per- this material was exposed tended to normal- formance of these alloys following longtime ele- ize ductility to that of a large grain size vated temperature exposures. recrystallized structure. This trend persisted The alloys evaluated were primarily those con- even through 1000-h anneals at temperatures sidered to be fabricable by welding. Hence, weld to 1650°C. stability as well as base metal stability was WELOMQ OF REFRACTORY ALLOYS 166

TABLE 4 Electron Beam Weld Parameter Evaluation*

W-26Bc :0Mo (i»/o) Direction of 60 Powder-metallurgy Arc-cart Weld cycle beam de- Unalloyed tungsten speed, flection. 1.27 mm (arc-cart), W-25Be (w/o) (arc-cart), No 42S°C 760°C No 760°C ipm amplitude no preheat no preheat preheat preheat preheat preheat preheat

Transverse • • •L205 L<540 T>540 •L205 15 < Zero T>540 DL<425 DL.T •L315 •L95 •L65 Longitudinal T>540 >540 T>540 L260 T315 T290 T>540

D n • • • • • L<425 L,; L230 L95 L135 L65 L65 L80 25 Longitudinal T>540 >540 T>540 KSO T260 T315 T120 T65

•L.T •L205 f Zero >540 T>540 50-1 D D •L315 •L,T •L105 •L95 •L120 •L65 •L65 I Longitudinal T>540 315 T260 T315 T315 T120 T95

f Zero D D 1001 D • •L480 •L,T I Longitudinal T>540 >540 4.8 12.7 2.4 6.4 9.5 4.8 9.5

*D, defective weld; •, sound weld; 4t bend transition data in °C indicated for transverse (T) and longitudinal (L) test specimens. emphasized. The thermal stability study was pre- 1315° C for each alloy requiring over 600 tensile ceded by a weldability study in which the baseline tests. Final analyses of instabilities were supported parameters were established for use throughout with optical and electron microscopy as required. the thermal stability study. This assured a.consis- The alloys evaluated in this program are listed tent basis for processing and comparing the stabil- in Table 6, which shows their normal composition ity of the various alloys. rnd general metallurgical classification. All these The alloys were screened for thermal stability materials were procured in the recrystallized con- by exposures between 815 and 1315° C and hold dition and to optimum processing schedules where times up to 10,000 h. Sputter ion pumped furnaces they were identified by suppliers. Hence, insofar were used exclusively providing an optimum fur- as possible, all alloys were normalized for long life nace environment of less than j.0~8 torr total pres- testing. sure. This assured that specimens would remain Of the alloys evaluated, T-lll, T-222, and FS-85 uncontaminated even after 10,000-h exposures. eventually received the greatest emphasis because Further, the potential load loss in event of furnace they were identified in the earliest welding phase failure was nil, because ion pumped systems are of this program and in companion creep and corro- totally closed and do not lose vacuum with Ios3 of sion test evaluation as those alloys demonstrating power. optimum combinations of these characteristics. Hence, they demonstrated the greatest potential Bend testing (to determine the ductile-to-brittle for advanced space power system applications. transition temperature) was used extensively for screening because it provides an excellent indication of a wide variety of structural interactions. Aging Parameters Over 3000 bend tests were required. Bend test The aging matrix for this program was as fol- screening was complemented by tensile testing to lows: 166 L£SSMANN

TABLES Postw Results

Change in 4t bend trans, temperature Lowest Weld 1 h anneal DBTT4 AUoy Structure* preheat, °C temp., °C "C Bendtypet •C

W GTA weld None 1405 +40 L 370 W GTA weld 260 1405 -75 L 482 w GTA weld None 1405 Increased L 370 GTA weld 290 1405 -1-95, in- L 425 W-25Re 3-290 creased duc- 425 tility implied W-25Re 4 GTA welds 1-None 1405 Decreased duc- 425 1-290 tility implied W-25Re 4 GTA welds 3-None 1405 W-25Re 3 GTA welds 760 1800 -240 Max,L&T 315 W-25Re GTA weld 760 1800 Questionable 540 W-25Re 11 EB welds None 1405 -295 Max,T 260 W-Re-Mo GTA welds None 14051 -45 to -75 L 175 32L change T205 (PM) (.1760 +4L W-Re-Mo 2-oe metal - 1540 +50 L -100 (PM) +80 T -60 fl315 +10 L +95 T L65 W-Re-Mo EB welds 760 S1540 +10 L +120 T T95 (AC)5 11760 +40 L +120 T

•GTA—gas tungsten arc; EB—electron beam. fBend type: L—longitudinal, T—transverse. tDBTT for annealed or unann«aled, whichever is lower. UPM—Powder-metallurgy product. §AC—Arc-cast product.

Aging times, h Aging temperatures, °C nal and transverse directions. In addition, room and elevated temperature tpsts were conducted. 100 815 Typical results of the thermal stability study 1,000 980 are presented in Figs. 9, 10, and 11. Overall conclu- 5,000 11 .1 sions provide guidelines for future utilization of 10,000 1315 these materials: • The evaluated alloys displayed a wide range All alloys were tested at all temperatures through of responses to the thermal exr'o ires 1000 h of aging. The most promising alloys were employed in this program. In most cases, carried through completely to 10,000 h. T-lll, these responses were most easily understood T-222, and FS-85 were aged at all combinations; in terms of the metallurgy of the respective D-43 and B-66 received modest attention beyond alloy system. 1000 h; and all the other alloys were evaluated only • The alloys generally displayed satisfactory through 1000 h. stability as would typically be required for For each temperature-time combination, each en;trneeri»g applications. Several alloys, alloy was evaluated by determining bend transition however, are temperature limited with temperatures of base metal, tungsten arc welds, respect to thermal stability. D-43 displayed and electron beam welds all in both the longitudi- loss of strength with increasing aging time WELDING OF REFRACTORY ALLOYS 167

Fig. 8 Sammary of bend test results pertinent to thermal stability for W-25Re-30Mo (a/o). and temperature. This would have to be intergranular separations which did not prop- accommodated in setting design stresses for agate from the weld metal into the base long time applications for temperatures above metal. Niobium-base alloys tended to display 1095°C. SCb-291 and B-66 were prone to loss unarrested cracking and a classic, abrupt of ductility with increasing grain size caused transition from ductile-to-brittle cleavage by high temperature aging. Hence, these behavior. Hence, the bend transition tempera- alloys should be used only at the lower tem- ture represents a design limit for niobium peratures for long time applications. Like- alloys but not for the tantalum alloys. wise, Ta-lOW displayed similar grain growth T-lll, T-222, and FS-85 displayed similar related instabilities. Otherwise, the alloys responses to aging. These were detected only investigated were generally acceptable for in measurements of the bend transition tem- high temperature application from the stand- perature; they were not observed in tensile point of structural stability. tests at room temperature or elevated tem- The magnitude of the aging response tended peratures. The fact that no evidence of a to be greatest for gas tungsten arc welds and response was seen in high temperature tensile least for base metal specimens. tests demonstrated that the thermal stability An important difference in bend test fracture is excellent from a design standpoint. Of mode was noted for aged T-lll and T-222 these, FS-85 alone would be limited but only bend specimens compared with niobium-base in an unusual situation requiring periodic alloys. Even though shifts in transition cycling from the aging temperatures to below behavior were noted, fractures were ductile the bend transition temperature. 168 LESSMANN

TABLE 6 Alloys Evaluated for Thermal Stability

Nominal composition, Alloy Classification wt/%

Highest creep strengths Ta-SW-ZHf (lettered alloys* Ta-9.6W-2.4Hf-0.01C Solid solution + dispersion strengthened Nb-27Ta-10W-lZr Solid solution strengthened Ta-lOW Ungettered alloys* Nb-lCTMOTa Nb-5Mo-5V-lZr Solid solution + dispersion strengthened Nb-10W-lZr-0.1C Gettered alloys' Nb-lOW-lOHf+Y Nb-10W-2.5Zr

'Reactive element addition, Zr or Hf, provides corrosion resistance in liquid alkali metals.

50 :oo 1 • na TO AGE 0 • mot TO AGE 0 \ 0 k 100 HI. ACE 0 - >ASE METAL • in HI. AGE O TOO HI. AGE -50 0 1000 HI. AC! & SO0O ML AGE -100 A 5000 MR. AGC 50 Q 10000 ML AGE -100 -100 Q10000 HK. AGE -100 -200 -no - LOWEST TEST TEMFEiATME -150 u /-LOWEST TEST TEMPEIATUtE -150 uS -3O0 -200 5 -xo S i t -f— 2 ' t -200 \ 1 1 1 1 aJ 1

o I | GAS TUNGSTEI " ? I -100 -100 ? -no LOWS! TEST TEM««*TlK-> , , Q -150 3 1-300 _.__ _\ Alt / -200 " 3 t

.in -150 o

0 o ELECTION IEAM WELDS -50 -100 -100 -100 LOWEST nST TEMPEIATUlf - -150 -100 r --^--"^fc -200

tOO TOO 1000 1100 1200 1300 "C «00 900 1000 1100 1200 1300 t AGING TEMPERATURE AGING TEMPEtJklUBE Fig. 9 Bend ductile-to-brittle tranaition temperature of Fig. 10 Bend ductile-to-brittle transition temperature of T-lll • • a function of aging parameters (It bend radius). T-222 as a function of aging parameters (it bend radius). WELDNG OF REFRACTORY ALLOYS 159

411OY; FS-t5 the alloy's usefulness is not limited by the aging or, conversely, to define any application limits due to aging. This latter approach is reviewed under Multipass Waldiag in this report. The particular aging and annealing responses observed in T-lll welds are shown in Pig. 12. Con- clusions based on these results and related supporting data are summarized as follows. • The weld structure of T-lll cannot be stabi- lized with respect to aging by postweld annealing for 1 h at temperatures to 1650°F. • The aging response observed was limited to an observed shift in the bend transition temperature with aging at 115CC (above 815°G and below 1315°C). This was accomplished by weld precipitation associated with cored areas of solute enrichment. • Neither tensile properties nor hardness traverses responded to aging. Weld tensile ductility remained excellent irrespective of thermal history. Hence, the engineering properties of this alloy appear unaffected by aging. • The aging response appears to be peculiar to BOO 900 IOOO 1100 1200 1300 °C the It bend test which, in this case, AGING TEMPISATURE represents a temperature sensitive test for Fig. 11 Bend ductile-to-brittle transition temperature of accommodation of an outer fiber strain of FS-85 as a function of aging parameters (2t bend radius). 3ZVz%. This agrees with the tensile ductility of 25% which, although excellent, was less than that required for It bends. Postweld Annealing Studies of T-lll • All bend test fractures occurred by intergranular duc+:le tearing rather than by brit- This investigation was initiated because T-lll tle cleavage. Hence, the bend ductile-to- welds responded to aging with an increase in the brittle transition temperature is a misnomer bend transition temperature as described above. for the effect observed in this program. The apparent mechanism of this response did not Again, this shows that the observed aging appear to compromise the usefulness of T-lll in response has no detrimental effect on the per- long time application. However, the unquestioned formance of this alloy. ;,nportance of this alloy in space power technology dictated that a more complete understanding of this aging response be developed. Further, the observed behavior was found to be characteristic of Multipass Welding several other gettered refractory metal alloys. Hence, this investigation was important not only in Investigations which fGeused primarily on eval- the application of T-lll but also on a general basis uation of the most promising space power system to the entire field of space power materials tech- alloys (T-lll, T-222, ASTAR-811C) demonstrated nology. Two Dadic approaches lend themselves to that multipass weld fabrication should be closely 8 9 this situation. scrutinized in the future. ' Problems observed also lent themselves to substantial resolution through • First, one can define a thermal treatment compositional control of the base metal. However, such as a postweld anneal which eliminates this is the type of problem requiring diligence in the response entirely. planning future systems, processing, and fabricat- • Second, the aging response can be investi- ing components, as recommended in the final sec- gated in i >ater depth to demonstrate that tion of this report. 160 LESSMANN

T T 200

Open Symbol - Low Heat Input GTA, Solid Symbol - High Heat GTA O Aged without postweld anneal 300 • Aged following 1315°C postweld anneal J50 A Aged following 1370°C postv.ild anneal 1.0 mm Sheet (Program Heat) Q Aged following 1425°C postweld anneal S\ Aged following 1480°C postweld anneal 100 « 200 O Aged following 1650°C postweld anneal CD 0) i etc. Reference 0.9 mm sheet data > < (Reference Heat) 50 V . a > ioo 2 *

I? 1480°C PWA V* is 1425°C PWA 1370° C PWA -50 £ "5 S \ 1650CPWA •r "o> ^ All from referencence hheae t -100 1315°CPWA for 05 mm sheet

Q -100 No Postweld Anneal -200

-15P

-300 A» <-1 -200 I I I 100 1000 5000 10,000 Aging Time, Hours dt 1150'C Fig. 12 Effect of postweld annealing on aging of T-lll GTA welds.

Gross sections of multipass welds in 9.5-mm- Bimetal Joints thick plate subjected to high sensitivity fluorescent penetrant inspection are shown in Fig. 13. Crack- Three combinations of bimetal joints are ing of the root passes occurred as subsequent readily recognized as of potential importance and passes were applied. Detailed evaluation indicated have received attention for space power systems that these were low cycle fatigue cracks that could applications. be eliminated with a small compositional adjust- • Tantalum and niobium alloy combinations.10 ment in rhenium content (Fig. 14). • Tungsten, molybdenum, or rhenium alloys to Sophisticated analytic techniques were applied either tantalum or niobium alloys.11-12 in the solution of this problem. In the process, the • Refractory metal alloy and austenitic com- Varestraint test w,as modified for use with %-in.- binations.13"15 thick specimens and was established as a useful cool for evaluating refractory metal alloys. Reasonable weldability is exhibited in the first Figure 15 shows a Varestraint comparison of the group although each special situation requires spe- refractory metal alloys. The comparison is repeated cific qualification for a particular application. in a narrower band in Fig. 16 wherein the effect of Good results have been achieved in GTA welding lower rhenium content in ASTAR 811-C is quanti- T-lll to Nb-lZr, e.g., using Nb-lZr as the filler fied. wire. WELDING OF REFRACTORY AiJ.OYS

ASTAR-8 llC, WeidT4-A5 ASTAR-811C, Weld T4-A6

7X T-111. WeldTl-A4 T-111, Weld T1-A5

Fig. 13 High sensitivity fluorescent penetrant results for GTA plate welds. Indications of cracking in A8TAH-811C result from high (oot-of-spoe) rhenium content (see Fig. 14 also). Rhenicm overstrengthening of grain volume causes grain boundary cracks. The second group of combinations were a result of brittle zone formation. The best poten- evaluated at great detail for thermionic application tial was demonstrated for tungsten and tantalum with emphasis on interdiffusion behavior. Expo- alloy combinations, which were immunized against sures up to 2000 h and 2000°C were investigated. failure from Kirkendall void formation by special- Interfaces with rhenium performed very poorly as ized methods developed in the referenced project. 162 LESSMANN

7X 7X A5TAR-811C, Weld T4-A3 ASTAR-811C, Wold T4-A7

Plate - Heat Number 650056 Plate - Hc-ut Numbei 650056 Filler - Heat Number 650068 Filler - Hi;at Number 650056

Heat Number Re Analysis , H/Q 650056 1.17 650068 1.34

Fig. 14 High sensitivity fluorescent penetrant results for two ASTAH-8UC plate welds chow no defects when plate only or both plate and filler are within specification for rhenium, 0.8 to 1.2 wt%.

At best, these are difficult combinations from all bimetal tubing (refractory/austenitic). In this case, aspects: design, fabrication, and service. the interfaces were not metallurgies1 lly bonded. The third requirement is the transition to con- This program was quite successful in dealing with ventional alloys, particularly austenitic materials. the mechanics of joining a mechanically clad sys- This requirement can stem from structural design, tem. Such a system would have less use as a heat cost, or environmental considerations. Buckman exchanger than as a structural component for envi- and Goodspeed13 evaluated numerous combinations ronmental protection (in air before launch into bonded by explosive welding. In their tests, combi- space). nations of welding and thermal treatments producing interdiffusion zones in excess of 1,3 X 10~5 m thickness would fail in 23 thermal cycles between WELDING TECHNOLOGY 315 and 730°C. The data are comprehensive in SUMMARY other properties but clearly show the difficulty of producing structures with materials whose inter- Excellent broad-based weldability and fabrica- diffusion zonss are extremely brittle. bility projects were supported by the NASA-Lewis Kass and Stoner15 dealt with the methods of Space Power Systems Division during the decade of producing welded complex thermal loops using the 1960s. This work provides an enormous base for WELDING OF RBFRACTORY ALLOYS 163

1000 AH data for welds made at 15 ipm. E 900 All data for melds made at a wveldinn speed of 15 ipm. b Data for fusion zone cracks only. * I ^ 800 H 200 o - 700 F5-85 High Re ASTAR-811C i 600 T-lll A5TAR-811C a. 500 !00 Ta-lOW < o 3 400 10 r-

300

12 3 200 A5TAH-8I1C 5 AUGMENTED STRAIN. %

IOC 5Cfe-29 Fig. 16 Varestraint test results rank cracking sensitivity Jq-lOW exactly as observed in plate welding. No cracks observed for 2 3 iii-spec (low rhenium) ASTAR-811C or Ta-lOW. AUGMENTED STRAIN. %

Fig. IS Comparison of total crack lengtt vs. augmented able, and reliability constraints will severely strain for all refractory metal alloys evaluated during this program. restrict product design. Molybdenum current plans for space power system development The weldability of molybdenum is like that of utilizing refractory metal alloys. tungsten but without an application advantage for higher temperature operation relative to the fabri- Weldability—Niobium- and cable alloys of tantalum or niobium. Tantalum-Base Alloys The niobium- and tantalum-base alloys most Bimetal Joints likely to be selected for fabricated systems will have excellent weldability. They will be welded • There is a broad range of compatibility autogenously or with matching filler wire in inert •between niobium- and tantalum-base alloys. atmospheres. Hence, common inert gas or vacuum These elements form a continuous solid solu- welding processes are applicable. The most easily tion which exhibits good fabricability utilized processes are gas tungsten arc welding, throughout the composition range. Each electron beam welding, and laser welding. Properly application is specialized and will therefore handled, particularly to avoid metallic and atmo- require process and design qualification. spheric contamination, other processes could be • Tungsten and molybdenum are less compati- qualified but at significant expense to achieve the ble with niobium, tantalum, or each other in same degree of reliability. These processes include welding. Specialized joint designs using tran- plasma welding, resistance welding, explosion weld- sition pieces (or spools) produced by extru- ing, hot pressure welding, and related variants. sion, explosive welding, or hot isostatic pressing would be useful to minimize thickness of Tungsten-Base Alloys brittle diffusion zones. Hence, application of transition joints will be diffusion limited. Tungsten-base alloys will present severe weld- • Bimetal joints between any refractory metal ability problems indefinitely because of high alloy and conventional metal alloys are ductile-to-brittle transition temperatures. Welding extremely specialized in nature. Even with aggravates this problem. The use of tungsten as a specialized methods, joint survival in service reliable leak-tight structural component will will be greatly impaired by thermal cycling severely challenge both designer and fabricator. and diffusion. Specialized opportunities for Problems frequently will yield to expensive special- welding these materials have been demon- ized solutions as typified by very thin materials strated using coextrusion, diffusion bonding, such as tungsten lamp filaments. Broad-based hot isostatic pressure welding, explosion generic welding methods, however, are not avail- welding, and, in thin sections, electron beam 164 LESSMANN

welding. Methods and techniques yielding • Joint design for bimetal joints is a complex solid state joints can produce simple shapes. special situation because of the influence of These frequently have demonstrated good interdiffusion, thermal stresses, and low cycle survivability at modest temperatures. Some fatigue. As an example, a good caste can be confusion could develop in design application made for keeping the interdiffusion zone from since sui ival of bimetallic joints is greatly growing beyond 0.013 mm between austenitic enhanced in very thin sections such as wires alloys and niobium or tantalum alloys. or ribbons. In these applications, leads are usually negligible, whereas in sheet or thicker materials, joint loads are not negligib As a CURRENT NEEDS FOR WELDING result, even residual stresses can cam; frac- REFRACTOHY METAL ALLOYS ture. This section addresses the special requirements Contaminated Control of fabricating highly reliable systems for the space nuclear power program. The first priority require- • Shield quality and accountability are major ments are based on the need to maximize the prob- concerns in welding refractory metal alloys. ability of success in prototype fabrication. The There are no nondestructive methods known most highly leveraged investment in prototype for providing product assurance. Hence, pro- hardware fabrication is in the reliability plan. The cess control and monitoring are essential. reliability plan must address the specialized Conventional welding practice is unaccept- aspects of these systems. In this regard, prototype able. costs are driven hard by the nature of the envisioned project. • Controlled atmosphere welding chambers are most easily monitored. Work funded by • Refractory metals. NASA in the early 1960s provide? a classic • High temperatures. basis for atmosphere control and assessing • Nuclear power. shop practice. A strong case can be made for • System prototyping. monitoring oxygen and moisture for weld • Operation in vacuum required (space). atmosphere control and that both must be • Liquid metal coolants. monitored. • Severe transients. • Metallic contamination (i.e., by Ni, Fe, Cu, • Safety. etc.) can be gross or insidious. Gross metallic • Reliability (design complexity vs. fabricabil- contamination produces a brittle structure ity). which cracks on cooling or causes freezing point depression by constitutional segregation Strategic Plan resulting in hot tearing. Minor contamination can go undetected but can result in severe The welding-needs list provides recommenda- structural impairment and instability, partic- tions supporting advanced fabrication techniques ularly by segregation to grain boundaries. and innovation in the reliability plan with special- Loss in creep strength or low cycle fatigue ized surveillance and diligent accountability. resistance are classic results. A related conse- Recommendations tend to be interrelated and com- quence: multipass weldability cap suffer sig- bine to yield a basic overall advancement in this nificantly from impairment of low cycle technology. Refractory metal alloy welding require- fatigue strength. ments are relatively well understood and are treated accordingly. Joint Designs Material Surveillance • Wide flexibility exists in joint designs for the Much of the material processing will be on a weldable tantalum and niobium alloys. Acces- special order single-lot basis which necessitates sibility and simplicity to maximize quality specialized surveillance. In particular, the reactive are major drivers. All joint preparations nature of refractor/ metals makes them unusually should be machined to final size. Machined susceptible to gaseous and metallic contamination. surfaces for welding must be free of laps and Weldabttity is a severe test for contamination and tears to eliminate contamination. should, therefore, be exploited for in-process WELDING OF REFRACTORY ALLOYS 166

screening; that is, the following tests should be Accessibility (eliminate vacuum-purged weld used to check ingots and intermediate product dur- chambers). An extrapolation of advanced shield ing material processing. gas monitoring and sensors leads to elimination of vacuum-purged weld chambers. Shielding methods • Varestraint and TIGAMAJIG testing can be a can be improved to provide greater accessibility, real asset for surveillance and should be more flexibility in design and manufacture, and a applied at the ingot and intermediate process- greater reliability. Investigators at the Westing- ing stages. house Research and Development Center have • Multipass testing: modification of Varestraint implemented easily constructed displacement and TIGAMAJIG tests can be envisioned chambers or specialized air displacement shields which would be of significant value. which have given excellent results in welding the • Bend testing to check for atmospheric con- reactive alloys of titanium and zirconium. Innova- tamination. tive design and use of monitors to map quality of • Gleeble testing can also be considered in a shielding leads to fully qualified designs. Welding comprehension surveillance plan. can usually be initiated within minutes compared Although envisioned as weldability tests, three with hours in vacuum pumpdown cycles. Two allied of these are, in fact, hot ductility tests. Hence, if advantages of eliminating the vacuum-purged anomalous behavior is identified, then processing, welding chambers are that laser welding utilization extrusion, rolling, forming, and service would prob- will be increased leading to greater automation and ably be compromised as well. Development of the quality, and that Var?straint and TIGAMAJIG multipass test will also detect sensitivity relative tests can be easily run out-of-chamber as low cost to low cycle fatigue performance. Hence, when material screening and production surveillance combined, these test results will address material, tests. fabrication, and system design reliability issues as early as the ingot stage in fabrication. The probability of success is greatly enhanced by this type of Welding Technology Improvements specialized probing of material response charac- Minimize use of vacuum-purged chambers. teristics early in the fabrication stage. Interest- Substitution of local shielding and use of ingly, the recommended tests are relatively inex- displacement chambers will greatly reduce fabrica- pensive, although they require systematic and well tion cost and improve design flexibility. As planned characterization for implementation. described above, an appropriate contamination control plan is required. Contamination Control Advanced methods and technology, (a) Lasers: The use of lasers should be exploited because lasers Three issues need to be addressed in contamina- lend themselves to local shielding methods and tion control: encourage automation and design practice which Accountability (control and monitoring). Classic lends itself to automation, (b) Adaptive Control: ground rules were established in the NASA pro- Adaptive control methods now being developed for gram of the 1960s for shield gas management. factory automation will have a positive impact on These still apply today. Moisture and oxygen moni- quality and productivity. This technology should be tors have improved considerably, so accountability transferred to fabricating space power systems. is significantly simplified. However, accountability Use of adaptive control as a reliability method will must be formalized, because fabricators tend to be cost effective even for prototype hardware fabri- view the knowledge base as optional and not essen- cation. This stems from the extremely high finan- tial. cial and schedular risk associated with refractory Detection (sensor deployment-development). Ac- metal system fabrication. Welding vision and tive monitoring techniques to use during welding image analysis systems now being developed for need to be developed to assure freedom from weld closed-loop control represent an excellent example contamination. Real-time spectrographic analysis of this potential. of the arc or weld pool is an example of this approach. Sensor development is receiving enormous industrial attention for factory automation. Materials Reliability Systems Hence, the probability of applying innovative sen- The fabrication of refractory metal alloys needs sor technology in welding refractory metal alloys is to incorporate modern nondestructive examination very great and should be exploited. (NDE) capabilities. 166 LESSMANN

• Properly done, the system considerations will fabrication standpoint, it represents a reliability lend themselves to automation and synergis- problem which demands evaluation on very practi- tic utilization of on-line sensors also required cal terms. The Varestraint test can be envisioned for adaptive control and shield gas audit. for modification to address this problem, and other • System reliability should also be assessed innovative approaches will undoubtedly be based on material properties and design presented for consideration. requirements of inspection sensitivity. The Laser Synthesis of Improved Structures, use of or development of new techniques for Surfaces, and Materials NDE characterization of materials to certify and monitor material structure should be One can anticipate that laser welding will play evaluated. a significant role in fabricating advanced systems. • The sensitivity of conventional NDE needs to No less recognized at this time are the almost be established as a baseline for this activity. incredible opportunities inherent in using the laser • Methods of assessing joint surface quality for heat treating, surface alloying, and cladding. on-line just prior to welding represents an The laser spans the rarge of power densities which addition to reliability system sensor equip- permit ablative shock hardening (as in rapidly ment. Joint surfaces and filler wire surface solidified structures) and continuous annealing of conditions represent a major risk of contami- alloy surfaces and weld claddings. In addition, nation in welding. Satisfying this sensor need lasers can be applied as a continuous surface treat- will greatly enhance system reliability. ment or locally. Obviously, the possibilities are too • An important overall objective of the NDE great for random cut-and-try approaches, but when reliability systems plan is to achieve on-line well grounded in the basic alloy metallurgy, signif- reliability assessment. icant opportunities will be defined. Weld Repair Working Group Summary of Needs Process design and product design must recog- and Priorities nize the need for weld repair. Hence, the development, reliability, and prototype plan must address Aspects receiving special attention at the ORNL repair needs from the onset. At least 30 percent of seminar workshop and items of perceived high pri- fabrication development costs should address this ority are summarized as follows. problem. Fortunately, most of the needs recommen- dation enhances the ability to make repair deci- • Emphasis needs to be directed to the most sions and to implement and monitor repair pro- fabricable and highest strength alloys: cedures. This synergism is available only if an ASTAR-811C and T-lll. outstanding reliability systems approach is incor- • These require attention in areas affected by porated early in the design phase of the space low cycle fatigue performance: multipass power systems. welding and repair welding. • All systems require use of and definition of Metallurgical Requirements in Weld Structures advanced technology for NDE, in-process surveillance (contamination), and development of Characterization of mechanical properties of metallic contamination detection. refractory metal alloys have not been adequate in • ASTAR-811C needs to be evaluated and evaluating low cycle fatigue (LCF) performance. developed for bimetal fabrication. Weld structure can be expected to adversely affect • Priority use of alloys not evaluated in the LCF; conversely, LCF behavior will be reflected in 1960s requires an enormous data base update problems encountered in multipass welding and for active design consideration, e.g., the Mo- weld repair as already noted in this review. The I3Re and C103 alloys. severe reactivity of these alloys will result in further aggravation of this problem since the potential is great for pickup of tramp elements, particularly in welding. Even minor contamination when REFERENCES segregated to grain boundaries will cause severe 1. D. R. Stoner, Effect vf Contamination Levd on Weldabiiity of degradation. Refractory Metal AUoj/3, NASA CR-1609, September 1970. 2. G. G. Lessmann, The Comparative Weldabiiity of At first glance, this is a classic mechanical Refractory Metal Alloys, Welding Journal Research Supple- property-metallurgical structures problem. From a ment, December 1966. WELDMQ OF REFRACTORY ALLOYS 167

3. D. R. Stoner and G. 6. Lessmann, Measurement and Con- 15. J. N. Kass and D. R. Stoner, Joining Refractory/Austenitic trol of Weld Chamber Atmospheres, Welding Journal Bimetal Tubing, NASA-CR-72353, May 1968. Research Supplement, August 1965. 4. G. G. Lessmann and R. B. Gold, The Weldabllity of Tungsten Base Alloys, Welding Journal Research Supple- ment, December 1969. BIBLIOGRAPHY 5. G. G. Lessmann, Welding Evaluation of Experimental Bond, J. A., Advanced Rankine Cycle Potassium Boiler Develop- Columbium Alloys, Welding Journal Research Supplement, ment Program, Volume 1: Test Rig Design and Fabrication, March 1964. NASA CR-135451, November, 1975. 6. G. G. Lessmann and R. E. Gold, Long-time Temperature Sta- Buckman, R. W., Comprehensive bibliography of Westinghouse bility qf Refractory Metal Alloys, NASA CR-1608, September publications, reports, and contracts completed between 1955 1970. and 1913, personal communication, Westinghouse Advanced 7. G. G. Lessmann, Post Weld Annealing Studies qf T-Ul, Energy Systems Division, Large, Pennsylvania. NASA CR-1610, December 1S70. Ekvall, R. A., R. G. Frank, and W. R. Young, T-lll Alloy Crack- 8. G. G. Lessmann and R. E. Gold, Interaction of High Tem- ing Problems During Processing and Fabrication, Conference perature Strength and Weldability, Welding Journal on Recent Advances in Refractory Alloys for Space Power Research Supplement, July 1973. Systems, NASA-SP-245 (1970). 9. G. G. Lessmann and R. E. Gold, The Varestraint Test for Gahan, J. W., A. H. Powell, P. T. Pilegg, and S. R. Thompson, Refractory Metal Alloys, WeUKrtp Journal Research Supple- Fabrication and Test qf a Space Power Boiler Feed Elec- ment, January 1971. tromagnetic Pump, Volume 1, Design and Manufacture qf 10. R. W. Harrison, E. E. Hoffman, and J. P. Smith, T-lll Pump, NASA CR-1949, April 1972. Rankine System Corrosion Test Loop, VoL 1. Lyon, T. F., Purification and Analysis qf Helium for the Welding NASA-CR-1JM816, June 1975. Chamber, NASA-CR-54168, July 1965. 11. F. G. Arcella, Interdiffusion Behavior of tungsten on Mendelson, I., Design and Fabrication qf Brayton Cycle Solar Rhenium and Group V and VI Elements and Alloys qf the Heat Receiver, Final Report, NASA CR-72872, July 1971. Periodic Table, Part I, NASA-CR-134490, Part 1, September Michelson, D. C, J. M. Fielden, A. C. Jordan, and E. B. Lewis, 1974. Selected Publication Relevant to the Space Reactor Materials 12. F. G. Arcella, Interdiffusion Behavior of Tungsten on Program, ORNL Document, DOE Contract W-7405-ENG-26, Rhenium and Group V and VI Elements and Alloys of the August 1983. Periodic Table, Part II (Appendices), NASA-CR-134526, Part Young, W. R., Fabrication of T-lll Test Loop Systems, II. Conference on Recent Advances in Refractory Alloys for Space 13. R. W. Buckman, Jr., and R. C. Goodspeed, Refractory/ Power Systems, NASA-SP-245 (1970). Austenitic Bimetal Combinations, WANL-PR-(EE)4X>4, Young, W. R., Fabrication of Refractory Alloy Components for August 1969. Advanced Space Power System Studies, in Proceedings of 14. D. R. Stoner, Joining Refractory/Austenitic Bimetal Tubing, Energy 70 Intersodety Energy Conversion Engineering NASA-CR-72275, Contract NAS-3-7621, December 1966. Conference, Volume 1 (1970). Refractory Alloy Component Fabrication

W. R. Young General Electric Company

INTRODUCTION WELDING PROCEDURES

During the years 1962 to 1972, advanced space During the early 1960s, welding specifications power systems which utilized tantalum- and for refractory alloys were generally procedural in niobium-base refractory alloys were being content. It was known that contamination of these developed at General Electric, Nuclear Systems alloys by oxygen or nitrogen was detrimental to Programs (GE-NSP) for NASA. Typically, ductility and corrosion resistance; the limits, how- advanced Rankine systems tests involved boiling ever, were not known. Most welding was conducted and condensing alkali metals to study heat transfer by trial and error, with the adequacy ui the purity or the basic compatibility between alkali metals at the welding environment judged by the and the refractory alloy containment material In appearance of welded test specimens prepared either case, system components such as pumps, before actual welding of components. Chemical boilers, condensers, heaters, valves, and pressure analysis of the weld metal for the interstitial ele- transducers were required. The refractory alloy T- ments, oxygen, nitrogen, and carbon, was used to 111 (Ta-8W-2Hf) was selected for high temperature certify the welding procedure. systems. By 1970, welding specifications reflected princi- During the same time, the fabrication of a solar pally the development of analytical and welding heat receiver for use in a 10 kW(e) Brayton cycle equipment. Welding was conducted in helium hav- power system was completed.1 Because gas flow ing a maximum impurity content of 5 ppm oxygen, was received at 595°C and delivered at 815°C, the 15 ppm nitrogen, and 20 ppm water vapor (10 ppm structural alloy, Nb-lZr, was selected. moisture content for component to contain The purpose of this report is to describe joining lithium). A gas chromatograph system was utilized procedures, primarily welding techniques, which to measure oxygen and nitrogen content, and an were developed to construct reliable refractory electrolytic hydrometer was used to monitor water alloy components and systems. An effort is made vapor contec'.3 These impurity levels resulted in no to use a general nondetailed approach in present- detectable increase in the impurity content of the ing this material in order to give those interested refractory alloys during the welding cycle. Actu- in systems some insight, into the state of the art of ally, tests conducted with up to 200 ppm air or building refractory alloy hardware. However, two 50 ppm moisture content intentionally added to systems, the Nb-lZr Brayton Cycle Heat Receiver the helium atmosphere resulted in no detectable and the T-lll Alloy Potassium Boiler Development impurity element concentration increase. The con- Program, are used to illustrate typical systems and trols used were intentionally conservative to avoid components.2 Particular emphasis i3 given to spe- any alkali metal corrosion effects. Long durat'on cific problems which were eliminated during the testing (up to 10,000 h) of both T-lll and Nb-lZr development efforts. Finally, some thoughts on alloy systems welded under the environmental application of more recent joining technology are requirements of the specification has shown no presented. corrosive effects on weldments. If mechanical

•66 REFRACTORY ALLOY COMPONENT FABRICATION 169 properties of the weldment, and not corrosion re- because the evacuation step, with proper leak sistance, are of primary concern, some relaxation checking, provided a complete check of system of environmental requirements of the specification integrity before the introduction of inert gases. may be tolerated. The specifications for cleaning, The double-walled stainless steel construction per- fusion welding, and postheating tantalum and mitted effective bakeout with hot water while cir- niobium alloys were issued by NASA in 1971.4 culating cold water dissipated heat during the welding cycle. Also, the vacuum environment was EQUIPMENT useful in performing local postheating of welds. The system proved to be very versatile since The inert gas purity requirements can be two 2.4-m dia welding chambers were available for attained by working with a sealed chamber, with extension of the basic chamber. Long lengths of access through glove ports. Continuous gas purg- straight piping could be welded by adding exten- ing, recirculating gas purification system, or initial sion tubes that were attached to flanged ports in vacuum purging may be utilized. A chamber the doors of the basic chamber. employing vacuum purging is illustrated in Fig. 1. The helium purification system and delivery A typical welding cycle involved evacuation to the piping was of welded stainless steel construction. 10~5 torr range, while heating the chamber to 60°C High purity helium was procured in bottles and to remove moistu-e from the chamber walls. The then further purified prior .o introduction into the chamber was then backfilled with helium contain- welding chamber. This was accomplished using sil- ing less than 1 ppm oxygen and water vapor. Dur- ica gel to remove moisture and heated ing welding, the helium environment remained titanium/zirconium alloy chips for oxygen and static except for automatic pressure compensation nitrogen removal. Titanium sponge material was during glove movement. Typically, 5 to 8 h of weld- e?a'-ated during the development of the purifica- ing could be performed before impurity levels (usu tic . ay; tem and found to be unsuitable. Mag- ally water vapor) exceed specification limits. This nesium, a titanium sponge impurity, was type of system is considered superior to others transferred into the welding chamber.

Fig. 1 Vacmun-purge inert gas welding chamber and helium parity control system with welding chamber oxton- sion tank attached to 0.9-m-dia by 1.8-m-long welding chamber. YOUNG 170 The welding machine used for both manual and automatic gas tungsten arc (GTA) welding is shown in Fig. 2. This 400-amp maximum constant current machine provided a complete weld sequence of upslope, weld, and downslope current control with time-delay relays to provide for integration of motorized accessory equipment. One important modification was made to this equipment after a welding torch failure caused excessive melting of a Brayton cycle heat receiver tube-to- header weld. Electrical circuitry was added to provide presetting of maximum welding current. Above the set point, the welding sequence was terminated, thus providing additional protection for the workpiece. A large percentage of the welding was done with the hand-held torches shown in Fig. 3. The molded silicon rubber and aluminum insulator provided complete electrical insulation of the brass torch body. It has been determined that certain refractory alloys, such as T-111, may exhibit brittle behavior when contaminated with minute amounts of nonrefractory metals, particularly copper and nickel. To prevent such contamination, refractory alloy Fig. 2 Automatic tungsten inert gas welding machine chucks ?nd shields were incorporated into the with controlled welding sequence. water-cooled torch design. Joint alignment fix-

iNA INSJLATJK

Fig. 3 Gu tungsten arc -elding torchw modlfled with T-111 alloy shield*. REFRACTORY ALLOY COMPONENT FABRCATON 171

Fig. 4 Refractory alloy fixture for aliening tube joints. tures of refractory alloys were utilized at locations 0.06 WU. where the refractory alloy components and the fixture were in direct contact. It was also speculated (±3 that sagging or movement during welding might Jflilt SMOTE {OH G* BCTP SfDCSJ induce weld cracking.5 The rigid refractory alloy fixture shown in Fig. 4 was designed to prevent HMIZOKTN. PQSITIflR such movement. DOUBLE ICC (BOTH WOES) Because welding was performed in an inert gas chamber, access was limited to specific work sta- tions consisting of a view port and gloves. The visual access to the weld joint was a critical problem requiring rather extensive tooling to position the workpiece. Although automatic welding devices alleviated this problem by movement of the torch SIWH itf ',1& 0« MTU S1DC5) rather than workpiece, this equipment was usually limited to joints of simple geometry. An internal Fig. 5 Topical weld joint designs for niobium and tantalum tube-to-header automatic welding torch was used alloys (butt joints). (Dimensions are in inches.) during Brayton cycle heat receiver fabrication and will be discussed later. by design, alternate joining methods such as electron beam welding should be considered. Single JOINT DESIGN pass electron beam welds of this type have been made in thicknesses up to 12.7 mm without diffi- The dssign of welded joints in refractory alloys culty. must compensate for their relatively high melting point and thermal conductivity. Typical joint designs for niobium and tantalum alloys are WELDING DISSIMILAR presented in Fig. 5. Multipass gas tungsten arc REFRACTORY ALLOYS (GTA) welding of some refractory alloys, such as T-111, has resulted in microcracking in the fusion In the fabrication of several prototype alkali 6 zone. This phenomenon usually occurs in an area metal containment systems, it was found to be approximately two weld passes below that being advantageous to utilize tantalum-base alloys such welded. Thus, thicknesses greater than 3.2 mm as T-111 in the high temperature regions <»nd ths requiring three weld passes may be susceptible to Nb-lZr alloy for lower temperature components. microcracking. If greater thicknesses are required For example, surge tanks, which are not a part of 172 YOUNG) the flowing alkali metal system, were normally loops. The Brayton cycle heat receiver required made of the Nb-lZr alloy. This reduced fabrication Nb-lZr to stainless steel joints for connection to costs but still provided complete refractory alloy the recuperator (inlet) and turbine (exit). containment of the alkali metal. Several joint types are available to affect tubu- Although the welding requirements for the T- lar transitions between refractory to nonrefractory 111 and Nb-lZr alloys are quite similar, the selec- alloys. Coextrusion of Nb-lZr and stainless steel tion of a weld filler wire and postwetd annealing has been utilized to produce a metallurgical bond treatment required a brief welding study. Gas along a tapered joint interface.7 A brazed joint tungsten arc (GTA) weldments were made between design illustrated in Pig. 6 was developed by T-lll and Nb-lZr sheet specimens. The effect of GE-NSP. This tongue-in-groove configuration was weld filler metal and response to postweld thermal applied successfully to the tantalum/stainless steel, treatments were evaluated by bend tests at room Nb-lZr/stainless steel, and Nb-lZr/Haynes Alloy temperature and —70°C supplemented by metallo- No. 25 (Co-base) systems.8 graphic and hardness data. These examinations The basic design provides clearances to accom- indicated that the optimum stability of manual modate differential expansion during the brazing welds was obtained using Nb-lZr filler metal and cycle. After brazing, the composite structure relies postweld anneals in the 1260°C to 1315°C range. It on local plastic deformation to maintain its integ- was found that radiographs of those dissimilar rity. Concern over the effects of thermal cycling refractory alloy joints were difficult to interpret. was evaluated by extensive testing of joints up to The large density difference between the two alloys 6.4-cm in diameter. No deleterious effects were along with some dilution of the base alloys pro- found. Tubular joints 7.6 cm in diameter exhibited duced these effects. local deformation and in one case failed in the braze alloy. This result was anticipated because local joint strain increases as a direct function of JOINING REFRACTORY TO diameter due to differential thermal expansion. NONREFRACTORY ALLOYS Thus, if extensive thermal cycling is present in the service condition, the technology would limit joint The refractory alloys were selected for high size to approximately 6.4 cm. temperature Rankine systems because of strength Another area of concern was the compatibility and their excellent compatibility with alkali between the refractory alloy and the cobalt-base metals. Stainless steels were utilized for alkali braze material (nominal composition: Co-21Cr- metal fill, gas pressurization, and instrumentation 21Ni-8Si-0.8B-0.4C) at elevated temperatures. lines as well as low temperature heat rejection Joints were exposed in vacuum for 1000 h at 870°C,

Fig. 6 Nb-IZr/type 316 stainless steel bimetallic joints; three joints on left shown prior to brazing; braced joint shown on right; slotted nipples are Nb-lZn small foint is 1.0-cm OD. REFRACTORY ALLOY COMPONENT FABRICATION 173 the maximum use temperature of the stainless tored to ensure compliance to the welding purity steel component. Metallographic examination and requirements of the basic welding specification. hardness traverses indicated a minimal and accept- Postweld annealing of entire subassemblies wa3 able reaction zone with no evidence of joint separa- also impractical due to size limitations of qualified tion. vacuum furnaces available in the U. S. at that On the basis of these results, brazed bimetallic time. Small refractory metal furnaces shown in joints between Nb-lZr and stainless steel were Fig. 12 were used for local annealing of tube joints used on most of the GE-NSP refractory alloy sys- which connected subassemblies or components. In tems constructed during this period. some cases, annealing was done in the welding chamber vacuum environment. The seven field welds required to join system subassemblies were REFRACTORY ALLOY LOOP annealed locally using the test rig vacuum cham- FABRICATION ber. Detailed planning for this fabrication sequence The development and application of the above- was completed at the time system design was final- mentioned procedures and techniques transpired ized. /( is eoctremely important in refract

VACUUM CHAMBSR

NoK TERTIARY LOOP

MECHANICAL VACUUM PUMP

II [7~1 Nit I I ARGON

Fig. 7 Potassium Boiler Development Test Rig. REFRACTORY ALLOY COMPONENT FABRICATION 175

RAW MATERIAL NASA APPROVAL RELEASE

ENGINEERING DRAWINGS

STRAIGHT DIMENSIONAL AND TUBE ITEMS MACHINED OR LIQUID PENETRANT ENGINEERING TRIM TO • ORMEOPARTS INSPECTION REVIEW LENGTH REPORTS

ACCEPT, REMAKE OR REWORK QUALITY CONTROL REPORT J

MANUFACTURING INSTRUCTIONS DEFINING SEQUENCE OF SUCCEEDING EVENTS

ACID CLEANING- PROCESS CONTROL RECORD

Fig. 8 Potassium boiler development pro- ANNEAL OF VACUUM FUHNACE MACHINED QUALIFICATION gram genertil fabrication procedures. PARTS

WELDING CHAMBER QUALIFICATION

WELDiNG- PROCESS CONTROL RECORD VISUAL AND ENGINEERING RADIOGRAPHIC REVIEW INSPECTION

ACCEPTOR BEPAIH

IN PROCESS DIMENSIONAL ENGINEERING MACHINING OR INSPECTION REVIEW FORMING REPORT

ACCEPTOR REWORK

POSTWELD ANNEAL OF COMPONENTS ASSEMBLIES OR VACUUM fUBNACE fWIVIDUALWELOS QUALIFICATION PP.DJESS CONTROL RECORD

FINAL DIMENSIONAL INSPECTION 176 YOUNG

THERMOCOUPLE WELL T-'EBMOCOUPLE WELL SPACER, BOILER SPACER, CONDENSER

THERMOCOUPLE WELL, BOILER

NaK EXIT HEADER, CONDENSER

ORIFICES, CONDENSER

TllF.[:'"r>CnilPLE WELL HEADER, CONDENSER pr:TA.ic'';i'', HEADER,

Fig. 9 Typical T-! 11 alloy machined parts.

This weld is illustrated in Fig. 13; a trial electron wrapper inside diameter was machined, honed, and beam weld \a shown in Fig. 13a; the completed inspected dimensionally. The helix was then flange is shown in 13b; and a metaiiographic cross ground to provide a 0.05-mm diametral interfer- section of this weld is shown in Fig. 13c. Typical ence fit with the wrapper. The interference fit was pressure transducers are illustrated in Fig. 14. produced by chilling the helix in liquid nitrogen Radiographic inspection, heat treatment, and mass and inserting it into the wrapper that was at room spectrometer leak tests were performed success- temperature. The inside diameter of the wrapper fully prior to filling of the pressure transmitting was slightly tapered at the helix entry end to pro- capillary with NaK. vide an easier start for the interference fit. Helix As discussed previously, multipass welding of insertion was a very delicate operation that T-111 alloy induced microcracking of the root required a steady hand and a handy rubber mallet pass.5-6 This cracking was not detected at the time for gentle persuasion. After completion of mechan- of transducer fabrication. Although it must be ical assembly and welding, the outside diameter of assumed that cracking of this type did occur, the wrapper was machined to the proper dimension no failures occurred during the extensive test pro- for fit-up with the bore of the EM pump stater. gram. Postweld annealing at 1315°C for 1 h was then performed to complete the fabrication cycle. Lithium and Potassium Electromagnetic (EM) Pump Duct Potassium Liquid and Vapor Valves The typical EM pump fabrication is illustrated in Fig. 15. Machining of the helical flow passage The electron beam welding process was used to was the most critical fabrication step. The outer join the T-111 bellows stem assembly as illustrated REFRACTORY ALLOY COMPONENT FABRICATION 177

["POTASSIUM* CONDENSER* | EM PUMP DUCT

ONE SLACK* DIAPHRAGM TWO DIFFERENTIAL" PRESSURE POTASSIUM' BOILER* PRESSURE HEAD TANK TRANSDUCER TRANSDUCERS

ONE SLACK- FOUR SLACK- LITHIUM EM- CONDENSER POTASS! JM EM I DIAPHRAGM DIAPHRAGM DESUPERHEATER PUMP O'JCT PUMP DUCT SUBASSEMBLY PRESSURE PRESSURE SUBASSEMBLY TRANSDUCER TRANSDUCERS FIELD WELD NO. 1 ' POTASSIUM" POTASSIUM POTASSIUSSIUMM* LITHIUM* POTASSIUM DESUPERKF.ATER" PREHEATER LIQUID VAPORRUNE DUMPTANK VAPOR VALVE SUBASSEMBLY SUBASSEMBLY VALVE

FIELD FIELD BOILER LITHIUM WELD LITHIUM- CONDENSER WELD EM PUMP DUCT POTASSIUM SUBASSEMBLY HEATER OESUPERHEATER SUBASSEMBLY NO. 2 SUBASSEMBLY NO. 3 SUBASSEMBLY

TANTALUM BUS BARS

FIELD WELD NOS. 4. 5. 6, AND 7 TEST RIG ASSFMBLY 'These components or subassemblies were furnace annealed DASHED LINE REPRESENTS ASSEMBLY I prior to welding into next assembly. Assembly welds were IN TEST RIG Vi..- UUM CHAMBER ! annealed locally in the test rig vacuum chamber.

Fig. 10 Fabrication sequence for the T-iii Boiler Development Test Rig.

Fig; 11 Portable weld chamber partially installed during qualification teat*. 178 YOUNG

Fig. 12 Refractory metal furnace for local annealing of weldmenta.

U-2.51 .:in-J

Cut Away Showing Diaphragm and Housing Slack Diaphragm Construction Nb-lZr Alloy

IS REFRACTORY ALLOY COMPONENT FABRICATION 178

2.64 cm

2.54 cm

Fig. 14 Slack diaphragm pressure transducers prior to filling with NaK.

F%. 15 180 YOUNG

BACKUP RING

BELLOWS RING BELLOWS (WALL THICKNESS 0.38 mm)

EB WELD

\ EB WELD

Fig. 16 T-lll alloy bellows stem assembly of boilor development valves illustrating electron beam weldments used to join the various comf

in Fig. 16. Bellows fabrication required careful Lithium Heater control of surface finish to prevent orange peeling and tears.11 Annealing after forming was also criti- The lithium heater includes six heater coils. cal due to the 0.38 mm wall thickness. In one case, Each 27.9-cm-dia coil was formed from 2.2 cm out- T-lll bellows contamination occurred because side diameter by 2.5 mm wall T-lll tubing and unalloyed tantalum foil rather than Nb-lZr foil had a developed length of 3.1 m. Appropriate man- was used as protective wrap. Analysis indicated ifolds and electrodes were required to split the that oxygen had transferred from the unalloyed lithium flow uniformly and to provide attachment tantalum to T-lll. of high current electrical buses. The valve assemblies are shown in Figs. 17 and A fixture was designed to hold the lithium >'8. Both the potassium liquid throttle and vapor heater as shown in Fig. 19. A more detailed view vaives used Mo-TZM plugs because neither valve of the three-tube manifold joints is shown in was required to seat during testing. Later testing Fig. 20. The size and weight of this component indicated that rhenium seats and W-25Re plugs made it difficult to handle manually within the were suitable for metering and isolation valves for welding chamber. To alleviate this problem, the service in 1035°C lithium.11 multipurpose weld positioner shown in Fig. 21 was REFRACTORY ALLOY COMPONENT FABRICATION 181

Mo-TZM ALLOY PLUG

2.54 cm

Fig. 17 T-Ul alloy potassium liquid throttle valve.

2.54 cm

Fig. 18 T-lll alloy potassium vapor valve. 182 YOUNG

Kg. 19 T-lll lithium heater positioned in welding fixture.

utilized. This unit provided motor-driven rotation Condenser and manual translation of the workpiece within the 2.4-m-dia welding chamber. The three-tube potassium condenser was NaK- The heavy vertical tantalum bus bars were cooied on the shell bide and was a basic hockey manually welded to the horizontal bus bars after stick design. From the welding viewpoint, this com- annealing. This feature plus the attachment of ponent, being of multitube design, r luired the flexible copper conductors i i shown in Fig. 22. qualification and production of -epresentative REFRACTORY ALLOY COMPONENT FABRICATION 183

Fig. 20 T-lll lithium heater illustrating butt joints to header. tube-to-header joints. In order to provide a fully tion to the NaK and potassium lines. The installed inspectable butt joint, internal automatic welding condenser is shown in Fig. 28. was selected. In this process, a welding torch is positioned inside the joint and is rotated by a motor drive (Fig. 23), which is controlled by the Boiler sequence control on the previously described auto- The potassium test boiler was C-shaped in matic welder. The three tube-to-header joints design and had a 1.9 cm outside diameter by shown in Fig. 24 were welded using this technique. 1.0 ."nm wall T-lll potassium boiler tube main- After inspection of these welds, the tube bundle tained concentric within a 3.4 cm outside diameter and support cage shown in Fig. 25 r/as slipped into by 2.5 mm wall shell. A helical fin insert shown in the sheli. The potassium outlet tubes were then Fig. 29 extended approximately half the boiler bent and trimmed to proper length, and a reducer length. At the end of the fin a helical wire coil was was welded to each tube, as shown in Fig. 26. j attached and continued throughout the remainder These reducers * ere welded to the exit header, pro- of the boiler tube. The shell and boiler tube were viding the same joint design as that of the inlet bent separately to provide the best assurance of header. The completed condenser ihown in Fig. 27 proper tube concentricity in the formed boiler. had connections for pressure transducers in addi- Several trials with stainless steel mockups pre- 184 YOUNG

Fig. 21 Lithiim heater fixtured for welding in the 2.4-m-dia chamber.

ceded the T-111 boiler tube bending. Initial bending component. The installed boiler is shown in was performed manually using a 360° machined Fig. 33. steel mandrel for the inside diameter contour and a steel roll on the outside diameter. Final sizing to the shell diameter was done on a conventional Test Rig Assembly in three-roll tube bender. Four tube spacers were then the Vacuum Chamber GTA tack welded to the formed boiler tube as shown in Fig. 30. The boiler shell shown in Fig. 31 Seven field welds were required to join the vari- was then slipped over the boiler tube. Mechanical ous subassemblies during installation. As shown in vibration was necessary during this operation to Fig. 10, field welds 1, 2, and 3 were part of the per- overcome the considerable friction between the manent installation. The four final welds installed sliding T-111 components. the boiler and were to be repeated for installation The completed boiler assembly shown in Fig. 32 of boiler No. 2. Some of the various components indicates the complexity of a fully instrumented and subassemblies are shown in Figs. 34 through test boiler. Provision for pressure transducers, 38. A typical setup of local annealing furnaces is insert thermocouples, and bulk fluid thermocouple illustrated in Fig. 39. The portable welding cham- wells necessitated many sequential welding and ber and setup for field welds are illustrated in inspection steps to provide high reliability in the Figs. 40 and 41. REFRACTORY ALLOY COMPONENT FABRICATION 186

i 188 YOUNG

FLEXIBLE COPPER CONDUCTORS

Fig. 22 T-lll lithium heater installed im test rig.

T-lll ALLOY DEFECT STUDIES

INTERNAL lUBE-TO-HEADER A review of T-lll alloy cracking occurrences on WELD the Potassium Boiler Development and other NASA-sponsored programs from 1965 to June 1969 was presented at a NASA-Lewis Research Center conference5 in June 1969. Cracking problems associated with multipass GTA welds and copper contamination were described in separate reports.6'12 Additional cracking problems, which occurred during Boiler No. 2 fabrication, were documented at the end of this program.13 It was concluded that foreign metallic contaminants had been introduced into T-lll tubing products. Inspection of tube hol- lows indicated visual contaminants prior to tube reduction. Analysis indicated the presence of iron, nickel, and chromium, singularly or in combination. To avoid these difficulties, a specification Fig. 24 T-lll three-tube condeiwar jpoturium inlet header. GE-NSP P4AYA21, Chemical Reimmd of Nonr REFRACTORY ALLOY COMPONENT FABRICATION 187

SUPPORT CAGE

Fig. 25 T-lll three-tube roidaas?; tube bundle mud rappwt esge.

-THERMOCOUPLE WELL

Fig. 26 T-lll thres-tubeci^dencjr during fabrication. 188 YOUNG

Fig. 27 Three-tube condenser.

Refractory Metal Contamination from Columbium, was a second tube containing lithium fluoride Tantalum and Their Alloys, was prepared. This which would act as a thermal storage material. specification should be incorporated into future During shade times, as thermal energy continues refractory alloy procurements to preclude cracking to be transferred to the working fluid, the heat problems caused by metallic contamination. storage material freezes. Small cracks were observed on the outside and The lithium fluoride filled heat storage tubes inside diameters of T-lll tubing, particularly after were furnished by NASA-LeRC for incorporation repeated acid cleaning. Correlation of defects with into the assembly. The heat storage tube units ultrasonic inspection results was not included in were fabricated at the Lewis Research Center. Fill- thooe studies. ing of the bellows cavity with lithium fluoride was Leaks in the T-lll lithium heater showed corre- accomplished at the Oak Ridge National lation with nickel-plated clamps used during heater Laboratory.14 The heat storage tubes as assembled fabrication. The iiyiication was that nickel contam- with the gas system can be seen in Fig. 42. ination from these clamps resulted in localized A program was conducted by General Ulectric cracking during the 1315°C heat treating cycle. to determine the compatibility of several promising niobium-base alloys with lithium fluoride under the cyclic thermal conditions which simulate the BRAYTON CYCLE HEAT RECEIVER sun-shade cycle of a heat receiver and to evaluate a design concept for containment of lithium The heat receiver was designed to absorb solar fluoride in a manner which accommodates the 29% radiation from a mirror-collector and function as a expansion upon melting without distortion of the heat exchanger to transfer heat into the Brayton bellows assembly.15 The compatibility tests indi- system inert gas working fluid. The basic heat cated the niobium-base alloys Nb-lZr, FS-85, and exchanger portion of the receiver consisted of a SCb-291 were corrosion resistant to lithium flux cone made up of forty-eight gas tubes sym- fluoride. The results of the compatibility study are metrically arranged around a center axis to form described in a separate report.16 the frustrum of a cone. Surrounding each gas tube (Text continues on page 194.) REFRACTORY ALLOY COMPONENT FABRICATION 169

BRAZED BIMETALLIC JOINT

Fig. 28 T-lll condenser after installation in test rig. 190 YOUNG

AREA A

30 cm

AREA A

2.54 cm

T 111 TAPERED PLUG

J 2.54 cm

Fig. 29 Boiler Nc. 1 insert geometry, final design.

Kg. 30 Boiler No. 1 tube mnd insert awembly after forming into C-fbftpe. REFRACTORY ALLOY COMPONENT FABRICATION 191

30 cm

LITHIUM EXIT LITHIUM INLET END-

Fig. 31 T-Ul potassium boiler shell.

POTASSIUM OUTLET -ACCELEROMETER SUPPORT

30 cm

Fig. 32 Boiler No. 1, completed assembly. 192 YOUNG

REDUNDANT PRESSURE -TRANSDUCER

Fig. 33 Boiler No. 1 installed in tert rig.

ROLLED ANDWELDEDSHELL

SPUN END CAP1

-LIQUID LEVEL PROBE I 2.54 cm

Fig. 214 T-lll alloy potassium head tank. REFRACTORY ALLOY OOWPONENT FABRICATION 193

• UNALLOYED TANTALUM ELECTRODE BARS LIQUID THROTTLE VALVE I f .1

30 cm Fig. 35 T-Ill alloy potassium preheater assembly.

POTASSIUM VAPOR VALVE-

30 cm I FROM BOILER EXIT TO CONDENSER INLET' Fig. 3€ T-lll alloy potassium desuperheater assembly.

T 111 ALLOY LINE EXTENDING TO - BOTTOM OF TANK TO LITHIUM CIRCUIT

T-111 TO Nb-1ZrWELD

FILL AND DUMP LINE

30 cm

Fig. 37 Nb-lZr alloy lithium dump tank. 194 YOUNG

Fig. 38 Lithium EM pump duct subasaembly.

The weight summary of the Nb-lZr gas system for the manual welding requirements and weld assembly is tabulated in Table 1 below. process qualification for each unique joint type used in the fabrication of the heat receiver. Weld TABLE 1 process qualification also included welding Weight Summary of the Brayton equipment, fixtures, and tooling required for each Solar Heat Receiver weld. Weight, Welders were qualified for the manual welding kg requirements by making full fusion butt joints in Gas system assembly (Nb-lZr) sheet and plate. The test piece illustrated in Gas tubes (includes 266 pounds Fig. 43 was fabricated to qualify the Nb-lZr alloy of LiP) 262 weldments required for gas system assembly. As Inlet manifold 118 the design progressed, additional joint types such Outlet manifold 105 Ducts (elbows and bimetallics) 16 as the support brackets-to-manifold and shell sup- Foil insulation and thermocouples 14 port ring-to-manifold were added. Three joints of each type were made and evaluated after prelimi- Total 515 nary welding trials had established welding parameters. Each type weld was made using the welding WELDING DEVELOPMENT equipment, process, and weld position intended for AND QUALIFICATION application to the heat receiver welding. Weld evaluation consisted of radiographic inspection as Weld qualification for the Brayton cycle solar applicable, followed by sectioning and metallo- heat receiver consisted of personnel qualification graphic examination. REFRACTORY ALLOY COMPONENT FABRICATION 186

Fig. 39 Refractory metal furnaces set up for local annealing of weldmenta in the teat rig vacuum chamber.

The automatic weld required for the manifold Boss-to-Manifold Welds circumferential joints was qualified by making two weld specimens from 3.2-mm Nb-lZr sheet using The original boss-to-manifold weld joint design the welding equipment and process intended for is illustrated schematically in Fig. 44. It was antic- application to this weld. The evaluation of these ipated that full penetration would be achieved dur- specimens was the same as for the sheet and plate ing the fusion pass applied to the OD. A second manual weld samples. A full-sized manifold mock- weld pass with filler metal addition would com- up, made from Type 304 stainless steel, was fabri- plete the weld. Initial welding trials indicated that cated to qualify the tooling and fixtures required full penetration could be achieved; however, the for this weld as well as an additional qualification weld heat input caused erratic melting of the ID of the welding equipment and process. edge of the manifold. Although the joint was struc- One of the primary purposes of the weld quali- turally sound, the uneven and rough appearance fication task was to identify problems with weld and possible adverse effects due to potentially joint designs, welding procedures, or fixtures prior excessive geometric distortions were causes for to commitment of hardware. As a result of the rejection of the original welding procedure. welding qualification trials, five of the weld joints The new welding procedure devised for this used in the heat receiver required a change in joint incorporated both automatic and manual either the joint design or welding procedure. welding techniques. Initially bosses were GTA tack 196 YOUNG

POTASSIUM EM PUMP DUCTSUBASSEMBLY

GAS CHROMATOGRAPH INERT GAS ANALYSIS

Fig. 40 Setup for field weld No. 1, potatuium subtMembiy.

welded to the manifold as shown in Fig. 45. The directly to the outlet manifold. This problem was root pass was then made using the internal rotary resolved by enlargement of the manifold hole as welding torch illustrated in Fig. 46. This procedure illustrated in Fig. 47. By maintaining the joint resulted in uniform fusion between the boss and overlap at a maximum of 0.5 mm, it was possible manifold. Weld inspection was simplified because to fully penetrate the joint. This resulted in com- the extent of penetration could be determined visu- plete fusion at the joint ID. A manual filler pass ally at the joint OD. A manual filler pass was then was then applied to form a reinforcing fillet. applied to the OD of the joint. Boss-to-Ferrule Weld Inlet and Exit Tube-to-Manifold Weld This joint was originally planned to be made A problem similar to the boss-to-manifold weld using the orbit arc welding equipment. This equip- was encountered, i.e., uneven and rough melting at ment, illustrated in Fig. 48, consisted of a portable the manifold ID. In this case, however, the auto- welding head with a motor-driven split-ring elec- matic internal welding technique could not be trode holder. Thus, tube-to-tube joints could be applied because a 7.6-cm dia elbow was welded welded automatically without rotation of the tubu- REFRACTORY ALLOY COMPONENT FABRICATION 197

Fig. 41 Field weld No. 2, Uthiom assembly. 198 YOUNG

COEXTRUDED JOINT TO OUTLET ELBOW WELD

• COEXTRUDED JOINT ASSEMBLY TO INLET WELD

Fig. 42 Nb-IZr thermal storage heat receive; for solar Brayton system. lar members. This process was abandoned for the liminary trials indicated lack of weld penetration. heat receiver application when, after repeated Variation in weld parameters to obtain full pene- attempts and redesigns, the equipment failed due tration, such as increased welding current or to breakdown of internal insulation, decreased welding speed, resulted in severe under- During the weld fixture planning stages, design cutting of the tube wall. provisions were made to allow for positioning the The ferrule was redesigned to reduce the wall heat receiver for manual welding of boss-to-ferrule thickness at the socket joint by 0.25 mm. Full pen- joints. This approach was adopted and qualifica- etration welds without undercutting were then tion joints were prepared. During welding, the made without difficulty. joint position was equivalent to that required during final assembly of the heat receiver. That is, Support Pad-to-Tube Weld each joint was welded in 180° increments to simulate the two positions required on the assembly. Metallographic examination of initial joints indicated that weld fusion zones had less cross- Tube-to-Feri-ule Weld sectional area than the base metal. This condition These joints were produced by rotation of the was corrected by adding a filler pass to the weld tubular joint under a stationary GTA torch. Pre- procedure for this joint. REFRACTORY ALLOY COMPONENT FABRICATION 199

Fig. 43 j.traytan cycle heat receiver weld qualification tests. (Dimensions are in inches.)

Boss

Manifold

Uneven Fusion Occurred Fig. 44 Original design typical boss-to-manifold joint. 200 YOUNG

TYPICAL TACK WELDS

30 cm

Fig. 45 Weld qualification teat specimen illustrating GTA tack weldfe used for boss positioning.

WATER AND POWER CONNECTIONS

SUPPORT FRAME DRIVE MOTOR

Nb-1 Zr 15.2-cm DIAMETER MANIFOLD TORUS

Fig. 46 Internal rotary GTA welding torch. REFRACTORY ALLOY COMPONENT FABRICATION 201

developed for welding the refractory metal gas system. Welding was conducted using the procedures described previously. The welding sequences employed were dictated somewhat by the requirement for postweld anneal of all weldments at 1205cC for 1 h in vacuum. Thus, during manifold welding sequences as shown in Figs. 50 to 52, the girth joints were only tacked to provide for later splitting of the manifolds for annealing. This operation was necessary because 0.5 mm MAXIMUM the complete manifolds could not be accommodated OVERLAP OF JOINT in vacuum furnaces which could be qualified to the requirements of the specification. The remainder of the manifold welding followed the natural fabrication sequence. Two circumferential welds produced the basic manifold sections. After machining of boss and nozzle holes, bosses were welded in place by a combination of internal GTA welding of the root pass and a second manual filler pass on the boss OD. The support brackets, shell support ring, and inlet or outlet were then welded to the manifolds. After splitting of the manifold and postweld annealing, the halves were rejoined by a manual girth weld. These two welds were annealed locally using a portable furnace specially constructed for this purpose. The vacuum environment during annealing was provided by the welding chamber. Fig. 47 Design af tube-to -manifold weld joints. The heat storage tubes were supplied by NASA with the lithium fluoride fill of the convolutions FABRICATION AND ASSEMBLY complete. As shown in Fig. 53, each tube was trimmed to the proper length prior to automatic The complete fabrication sequence for the Bray- GTA welding of the inlet ferrule and reducer to the ton cycle heat receiver is depicted in Fig. 49. For 3.18-cm OD gas tube. The 1.9-cm OD outlet tube the gas system assembly, close integration was was then trimmed to length (long leg), and the out- required of the four basic fabrication processes: let ferrule was welded. At this point, the short leg machining or forming, cleaning, welding, and of the outlet tube was trimmed to compensate for postweld annealing. Each component fabrication length variations in the as-received gas tubos. This and assembly was controlled by manufacturing operation produced a matched pair which was instructions which defined the detailed sequences maintained by serial number until the final weld and inspection points. between the extension tube and reducer on the gas The major fabrication effort involved the gas tube. The two welds at the reducsr were then vac- system assembly sequences as shown schematically uum postweld annealed in the voiding chamber in Figs. 50 to 55. Consideration of refractory alloy with the use of a specially designed furnace. Tube- welding and related sequences were of particular to-ferrule welds were annealed later in conjunction significance and are therefore discussed in detail with the final assembly boss-to-ferrule welds. prior to discussing the gas system fabrication The welding of 48 gas tubes to the two mani- effort itself. folds was accomplished with the assembly posi- All gas tungsten arc (GTA) welding of the gas tioned in the large lotatable welding fixture. As system was done using the weld chamber shown in mentioned previously, these 96 welds were origi- Fig. 1. This welding system consisted of the basic nally planned to be made by the orbit arc tech- 0.9-mdia by 1.8-m-long chamber shown on the left nique. As shown in Figs. 54 and 55, the rotating to which two 2.4-m-dia by 1.2-mlong chamber fixture was to be used to position each weld in extensions were attached. Not shown is the univer- front of the welding ^ .ions. Because it was sal welding positioner incorporated in the two necessary to use manual welding techniques, access 2.4-m-dia chamber extensions, which was especially to the ID and OD of each weld was required. This 8to

TUNGSTEN ELECTRODE WATER-COOLED CONTACTOR

SPLIT ROTATING RING

/MSP,' NUCUAR SYSTEMS PROGRAMS MOTOR DRIVE UNIT

Fig. 48 Special water-cooled tube welding head. REFRACTORY ALLOY COMPONENT FABRICATION 203

HEAT RECEIVER OUTIET INLET FERRULE REOULfR TUBE EXTENSION TUBES FERRULE

BASIC TOP ENCLOSURE SUPPORT BASIC SHELL SUPPORT BOSSES SUSSES BRACKETS MANIFOLD BRACKETS RING MANIFOLD RING

OUTLET GAS TUb*S INLET ELBOW (48) OIFFUSER

OUTLET INLET BIMETALLIC BIMETALLIC

GAS SYSTEM OUTLET MANIFOLD INLET MANIIOLO ASSEMBLY

TOP ENCLOSURE FINAL ASSEMBLY APERTURE CONE

FLEX PLATES SUPPORT CONE FLEX PLATES

Fig. 49 Breyton cycle heat receiver fabrication sequence. was accomplished by rotating the entire fixture 90° Manifolds on its trunnions so that the planes of the manifolds were in the vertical position. One-half of each The basic manifold sections were formed by weld joint was made in each of the four welding radial draw forming. Because of the high cost of positions. To equalize weld distortions as much as tooling, this method is generally considered only possible, welds to each manifold were made in six for parts which cannot be formed by more conven- groups of eight, alternating back and forth across tional methods or for production quantities. How- the manifold diameter. These welds were postweld ever, because draw forming required significantly vacuum annealed using two specially designed fur- less raw material than would have been required naces, each capable of accepting eight weld joints. for other methods being considered, the savings in The gas system assembly was completed with material costs more than offset the higher cost of the welding of the bimetallic joint assemblies to tooling. the manifolds. These welds were last in the fabri- As depicted in Figs. 50 and 51, four 180° seg- cation sequence because with the bimetallics in ments were required for each manifold. Tool try- position the gas system assembly could not be outs wert done with stainless steel, and these parts rotated within the welding chamber. To change were later used to fabricate an inlet manifold for welding position, it was necessary to move the tooling, welding, and other checkouts. assembly onto the transfer dolly, rotate, then repo- Forming the Nb-lZr segments wa3 highlj suc- sition in the welding chamber. Postweld anneals cessful in that only nine close-trimmed "pieces of were done with the same furnace previously used material were required to produce the eight mam for the manifold girth weld anneals. fold segments. Radial draw forming was done at As illustrated in Fig. 49, the inlet and outlet the Cyril Bath Co., Solon, Ohio. manifolds and the 48 gas tubes were the major After being accurately machine-trimmed at the components of the gas systems. These hardware 180° center plane, the basic manifold sections fabrications will be discussed in this section. (Text continues on pasre 208.) 204 YOUNG

STEP 1 FORM INNER AND STEP 2 TRIM ENDS TO SIZE OUTER SEGMENTS

STEP 3 TACK WELD ENDS TO FORM HALF-TORUS AND MACHINE CIRCUMFERENTIAL WELD JOINT

OUTER

tNNER

Fig. 50 Manifold forming and machining. REFRACTORY ALLOY COMPONENT FABRICATION 206

STEP 1 t/VELD INNER AND OUTER HALVES TOGETHER

STEP 2 MACHINE HOLES FOR STEP 3 WELD BOSSES TO TORUS BOSSES AND NOZZLE

STEP 4 WtLD INLET OR OUTLET STEP 5 WELD BRACKETS AND TO MANIFOLD SHELL SUPPORT RINGS

Fif. 51 Manifold welding. 206 YOUNG

STEP 2 WELD TIE BARS ACROSS 180° SEGMENT STEP 1 SPLIT MANIFOLD AT TACK WELDS AFTER ESTABLISHING UNSPRUNG TORUS DIMENSION

DIMENSION AS DETERMINED IN STEP 1

STEP 3 POSTWELD ANNEAL AT 1205°C

STEP 4 REMOVE TiE BARS, WELD PREP FIXTURE FOR GIRTH WELDS

STEP 6 FINAL MACHINING OF BOSSES, BRACKETS, AND SHELL SUPPORT RING

STEP 5 LOCAL ANNEAL GIRTH WELD

Fig. 52 Manifold annealing, final welding, and machining. REFRACTORY ALLOY COMPONENT FABRICATION 207

2 WELD INLET FERRULE AND REDUCER 1 TRIM TUBES TO LENGTH WELD WELD

3 TRIM OUTLET TUBE 4 WELD OUTLET FERRULE TO LENGTH WELD

5 TRIM SHORT LEi OF OUTLET 6 WELD HEAT RECEIVER TUBE TUBE TO MA I CS HEAT STORAGE TO OUTLET TUBE TUBE LENGTH i WtLD TRIM LENGTH

7 LOCAL ANNEAL REDUCER WELDS H, I'WW II SPLIT 7l FURNACE LL \

Fig. 53 Heat storage tube weld sequence. 206 YOUNG

MANIFOLD WELDING ' POSITION '

Fig. 54 Welding chamber installation, side view. shown in Fig. 56 were GTA tack welded at the manifold was then made, and the manifolds again girth joints to form 360° half shells. The weld fix- reversed to provide access to the first tack welded turing for a typical inner manifold section is side. The automatic GTA welding of these joints shown in Fig. 57. The inside-the-chamber view of was done in a single pass with filler wire addition. the tack welding operation for a typical outer sec- After completion of the circumferential welds, tion is shown in Figs. 58 and 59. the boss location holes were drilled and reamed on Each of the four 360° half sections was a horizontal boring mill equipped with a rotary machined and weld prepped for the circumferential table. A large Bridgeport head was mounted to the weld joint. An inner section is shown being machine which enabled the machining of the holes machined in Fig. 60. at the proper angle with the manifolds in the hori- The inner and outer manifold sections were zontal position. The inlet nozzle and outlet elbow cleaned, and both manifolds were set up for weld- openings were also machined at this time. The ing of the circumferential joints as shown in inlet manifold after completion of all hole drilling Fig. 61. The joints were aligned and GTA tack is shown in Fig. 62. welded during the first welding operation. The The boss-to-manifold welds required three manifolds were then removed from the welding fix- operations: tack weld, automatic internal root weld, ture, reversed, and tack welded on the opposite and manual filler pass. Alignment of the bosses for side. The. automatic GTA weld of one joint in each tack welding was accomplished by a molybdenum manifold was then made, and the manifolds again plug which extended through each boss and the reversed to provide access to the first tack welded machined hole in the manifold. The alignment fix- side. The automatic GTA weld of one joint in each ture is shown in Fig. 63. The manifold mounting REFRACTORY ALLOY COMPONENT FABRICATION 209

SUPPORT STRUCTURE ANO - WELOINGS POSITIONER

Fig. 55 Welding chamber installation, end view.

of the internal welding head occurred. Welding of 27 bosses had been completed before the failure. The cause of failure was traced to the automatic welding equipment. Although set for 160 amps, a current in excess of 400 amps was recorded, which resulted in failure of the torch end of the internal welding head. Primarily copper, some stainless steel, and Teflon melted off the end of the welding head, and the torch cooling water was released by the failure. The boss-to-manifold joint which ^as being welded was severely embrittled due to contamination, as shown in Fig. 65. Coating-type deposits of Fig. 56 Inner and outer manifold sections, cleaned copper and other materials probably carried by the in preparation for girth tacking. steam were more widespread. Spatter-type deposits were located at the bottom of the manifold directly below the boss which was welded. arrangement during tack welding and subsequent The rework procedure outlined in Table 2 was operations is shown in Fig. 64. implemented. In addition, it was determined that While performing the root pass internal boss- loss of the feedback signal had driven the output of to-manifold welds on the outlet manifold, a failure the welding machine above the 400 amp rating and 210 YOUNG

Fig. 57 Loner manifold section being loaded into weld chamber fot girth tacking. REFRACTORY ALLOY COMPONENT FABRICATION 211

Fig. 58 Outer manifold section set up in weld chamber for tack welding of girth joints.

TABLE 2 caused the welding head failure. To prevent a Rework Procedure for Oitlet Manifold recurrence, a meter relay circuit was incorporated Boss Weld Failure to provide failsafe shutdown in the event that output current exceeds a preset value for any reason. 1. Remove tack welded boss. A new internal welding head was also fabricated 2. Trepan machine 7.6-cm-dia hole at contaminated boss on existing centerline. and checked on trial welds prior to resuming mani- 3. Perform interstitial analysis on disk removed to verify that fold welding. Automatic internal welding of the all contaminated Nb-lZr material has been removed. remaining bosses on the outlet manifold and the 48 4. Perform cleaning trials on internal surface of slug to verify inlet manifold bosses was then completed without adequacy of cleaning techniques. Also check out pickling difficulty. Addition of the manual filler pass on method using foil to ensure hydrogen embrittlement will not occur in manifold. each boss OD completed the boss-to-manifold weld- 5. Mechanically remove local deposits from manifold. ing. 6. Nitric acid pickle and rinse (primarily to remove copper). The attachment of support brackets, shell sup- 7. Nitric HF pickle and rinse. 8. Machine and weld prep insert piece to fit 7.6-cm-dia hole port rings, inlet nozzle, and outlet elbow was with predrilled boss hole. Use piece removed from manifold accomplished with both manifolds in the welding section girth trim so as to match contour. chamber so that at least two welding operations 9. Clean, weld, and X-ray. could be performed during each chamber cycle. The 10. Reclean bosses and retack in preparation for internal root shell support rings and bracket support tubes pass welding. shown in Figs. 66 and 67 represent typical 212 YOUNG

Fig. 59 Tack welding girth joint of outer manifold section.

attachments. Alignment fixtures were used to posi- The manifolds were then set up in the welding tion each component as illustrated in Pigs. 68 and fixture for the manual welds of the girth joints. 69 for the outlet manifold support tube welds and Upon completion of this operation, the girth joints for the outlet manifold-shell support ring. Typical were vacuum annealed using the local furnace weldments for the inlet manifold are shown in setup in the welding chamber. During the vacuum Fig. 70. anneal of the outlet manifold, girth weld No. 1, a Upon completion of welding, each manifold pressure surge aborted the run after 55 min of the diameter was measured at the girth tack weld area 60-rn!n cycle. An air leak occurred in plastic tubing prior to splitting the manifolds. The restraining tie used to connect the sealed motor housing on the bars required for postweld anneal v/ere then weld fixture to a vacuum feedthrough. The results machined to dimensions which maintained the of various tests and analyses performed to deter- actual manifold diameters. The tie Dars, fabricated mine the condition of the manifold indicated that from Nb-lZr alloy, tube and discs, were then contamination was superficial. Polishing with welded to each half manifold as in Fig. 71. abrasive paper was used to remove surface contam- Postweld annealing was performed in two ination. The out3ide of the manifold was then acid 1205°C/l-h cycles of the Brew Model 922 vacuum cleaned and rinsed. furnace, located at Stellite Division, Cabot Cor- The final machining of manifolds required poration, Kokomo, Indiana. The significant stress extensive setup and dimensional checks to ensure relief achieved during the postweld anneal was evi- machining features and dimensions would result. dent by the complete lack of manifold spring-back As anticipated during the design phase, distortions when the tie bars were removed. were present in features and feature locations due I

Fig. 60 Machining for circumferential weld joint on inner manifold section. 214 YOUNG

Fig. 61 Inlet aad outlet manifold being set up for automatic GTA welding. REFRACTORY ALLOY COMPONENT FABRICATION 215

Fig. 62 Inlet manifold after machining of 48 boss holes and nozzle opening.

to general welding distortions and the need to fab- Gas Tube Assembly ricate the manifolds in 180° half-sections (because of annealing furnace limitations). Planning, tool- The fabrication sequence depicted in Fig. 53 ing, and stock allowances proved to be both neces- was followed exactly during gas tube assemblv sary and sufficient. welding. A typical welding operation on the heat storage tubes is shown in Fig. 74. A typical weld All openings were plugged or otherwise sealed joint for inlet ferrule-to-tuLt . shown in Fig. 75. to maintain internal cleanliness. Exterior surfaces A plastic-covered wire rack was constructed tn pro- were protected when they were not involved in the vide handling of eight tube assemblies during each specific work area. The bosses were rough and fin- welding chamber cycle. The typical gas tube ish turned with special hollow mills again usi.ig assemblies are shown in Fig. 76. the Bridgeport head mounted on the Lucas HBM. The horizontal pads on the manifolds were mi!led with a right angle attachment and drilled Gas System Assembly with the Bridgeport head. The facility mounting pads (for attachment of the receiver to the Brayton The assembly fixture consisted basically of system) were milled and drilled with the spindle of stainless steol angle brackets which bolted to the the machine. The shell support rings were rotary manifolds and rotating ring of the welding fixture. milled with a right angle attachment and drilled The manifolds were mounted as shown in Fig. 77, with the machine spindle. aligned, and secured to correct dimensions. The After final machining, cleaning, and inspection, welding fixture was then rotated with the mani- the inlet and outlet manifolds shown in Figs. 72 folds in the horizontal and vertical planes, and and 73 were ready for gas system assembly. dimensions were rechecked tc ensure that no sig- 216 YOUNG

Fig. 63 Boss alignment fixture locating a boss on inlet manifold; other bosses shown have been tack welded.

Fig. 64 Inlet and outlet manifold* mounted on routine weld fixture; fixture shown with outlet manifold routed to lower position after tacking all 48 DOSIMSS. REFRACTORY ALLOY COMPONENT FABRICATION 217

Fig. 65 Outlet manifold illustrating boss location where welding torch failure occurred. nificant shifting would occur during final assembly obtain access for the following welding sequence: welding. (1) outlet manifold, OD side, (2) outlet manifold, ID The 48 gas tubes were then installed as shown side, (3) inlet manifold, ID side, (4) inlet manifold, in Fig. 78. Originally, the plan was to install the OD side. At each position, welds were made in support spider after all pas tubes were in position. groups of eight, rVernating back and forth across However, in an attempt to slowly raise this sup- the manifold diameter. After careful visual exami- port, five gas tube extensions were dented. nation a few joints were reweided to ensure that a Although leak tight, these damaged areas were sufficient overlap was in evidence at adjacent weld reworked by the addition of a small amount of passes. A helium leak test indicated that all joints filler metal followed by postweld anneals. X-ray were leak tight. and helium leak check inspections were also per- Postweld annealing c" these tube welds was formed successfully. accomplished using the same local furnaces that The spider and its support brackets were then had been used previously for gas tube anneals. At modified to eliminate the potential for damaging each joint location, there were two welds actually other tube assemblies. A heavy center disc was annealed: tube-to-ferrule and ferrule-to-bosj. Each removed, and the heavy spider support brackets of the two furnaces covered eight tube joints such were replaced by the smaller and lighter spoke-like that 16 anneals were accomplished during each support shown in Fig. 78. It was not possible to vacuum cycle of the welding chamber Furnace install all the tubes after the spider was posi- locationu were alternated back and forth across the tioned. All the tubes were then reinstalled and manifold diameters for each annealing cycle. shimmed with foil at the spider cutouts. The final gas system assembly welds were those The 96-tube ferrule-to-manifold boss welds were which attached the coextruded joints to the inlet initiated by first tack welding the 48 inlet manifold and outlet manifolds. These were manual welds in welds. The welding proceeded as each joint weld the 7.6-cm-dia duct tubing as shown in Fig. 42. was made in two half-'ircumference passes. The elding fixture was rotated in four positions to (Text continues on page 224.) 218 YOUNG

Fig. 66 Shell support rings for inlet and outlet manifolds; temporary bridge shown tack welded at nozzle cutout location on inlet ring.

2.54 en.

Fig. 67 Typical manifold support tube*. Fig. 68 Welding alignment fixture for outlet manifold support tube.

ro 3 220 YOUNG

Fig. 69 Weld alignment fixture for outlet manifold shell ring.

/BOSS TO MANIFOLD MANIFOLD CIRCUMFERENTIAL WELD

FACILITY MOU'TIMC PAD TO

Fig. 70 Inlet manifold, typical weldmento required during fabrication. BPFRACTORY ALLOY COMPONENT FABRICATION 221

Fig. 71 Inlet manifold, restraining tie bars tack welded to halves for postweld anneal.

Fig. 72 Outlet manifold prepnred for gas system assembly. 222 YOUNG

Fig. 73 Inlet manifold ready for gas system assembly.

Fig. 74 Heat iterate tube* loaded for assembly welding of inlet ferrules and reducers. REFRACTORY ALLOY COMPONENT FABRICATION 223

Fig. 75 Closeup of typical automatic weld inlet ferrule-to-tube, as seen through vacuum chamber viewport.

KFUUCFR \ "7 -»— i \ . in r F R R U L L

1 // -GAR TUBF FXTFKSION

--0UT1 F.T FFRRULF

Fig. 76 Typical heat storage tube assemblies. 224 YOUNG

Fig. 77 Manifold mounted on rotating weld fixture.

The furnace which had been used for manifold alloys resulted in the development of specifications, girth weld anneals also accomplished the postweld equipment, and inspection techniques which pro- anneals of these joints. duced highly reliable, leak free hardware.17 Intermediate leak checks had been performed The development of tubular joints between on each component of the gas system assembly as refractory and nonrefractory alloys provided well as all individual assembly welds. The accep- design flexibility in material selection. Portable tance mass spectrometer leak test of the entire gas field welding procedures and equipment eliminated system was performed by independent evacuation size restrictions on refractory alloy assemblies and of the gas system while it was positioned inside the permitted field repairs and component modifica- welding chamber. Helium was then introduced into tions. thvi welding chamber and thus completely sur- One key to this success was the detailed fabri- rounded the gas system. cation planned which occurred throughout the system design phase. Material procurement, machining or forming, cleaning, welding (sequence, SUMMARY AND fixtures, etc.), and postweld annealing were closely RECOMMENDATIONS integrated with design. The Yielding of T-lll alloy presented more The T-lll Potassium Boiler Development Test problems with quality assurance than previously Rig and the Nb-lZr Brayton cycle heat receiver welded niobium-base alloys. Weldability determi- represent the two largest, most complex systems nation gave needed insight but could not evaluate fabricated during r.-e first space power era. weld cracking encountered in large, complex Experience gamed during the welding of refractory components.18 By 1972, many causes of defects liad REFRACTORY ALLOY COMPONENT FABRICATION 225

Fig. 78 Gat system setup for assembly welding.

been found and corrected; however, more basic 2. J. A Bond, Test Rig Desigv and Fabrication, understanding is needed. In all cases, T-lll defects NASA-CR-135451, Vol. I, January 1W9. were repairable by welding. It should be noted that 3. T. F. Lyon, Purification and AvxUysis of Helium fr • the Welding Chamber, NASA-CR-54168, July 1965. every T-lll component performed successfully 4. T. J. Moore, P. E. Moorhead, and K. J. Bowles, Specification after being qualified and accepted for alkali metal for Cleaning, Fusion Welding, and Postheatvng Tantalum and service. Columbium Alloy. NA.SA TMX-6787° July 1971. Since 1972 significant advancements have 5. R. A. Ekvall, R. G. Frank, and W. R. Young, T-lil Alloy Cracking Problems During Processing and Fabrication, occurred in welding processes, nondestructive Paper VIII, Recent Advances in Refractory Alloy for Space inspection techniques, and control instrumentation. Power Systems, NASP-245, June 1969. • The new welding processes such as laser, GTA 6. G. P. Brandenburg and W. R. Young, Evaluation of Weld Dabber, and pulsed current GTA should be Filler Materials for T-lll Alloy, GESP-3S6, R-71-NSP-1, evaluated for refractory alloys. Since rapid December 1970. 7. H. M. Cameron, Thermal Cycling Test on a S-Inch Diameter advancement in technologj' has occurred during the Columbium-1 Percent Zirconium to 316 Stainless Steel Tran- past decade, fluorescent penetrant, ultrasonic, and sition Joint, NASA TMX-2118, Nov. 1970. radiographic inspection also should be reevaluated. 8. S. R. Thompson, J. D. Marble, and R. A. Ekvall, Development at Optimum Fabrication Techniq^ws for Bnued Ta/Type S16 SS Tubular Transition Joints, GESP-521, Con- tract NAS-3-11846. 9. E. E. Hoffman and J. Holowach, Ch-lZr Rankine System REFERENCES Corrosion Test Loop, NASA Ck-1509, June 1970. 1. I. Mendelson, Design and Fabrication of BrayUm Cycle Solar 10. R. W. Harrison, E. E. Hoffman, and J. P £mith, T-lll Heat Receiver, Final Rvport, GE SP-519, N* SA-CR-72872, Rankine System Corrosion Test Loop, Volumes I an** II, July 1971. NASA-CR-134816, June 1975. 228 YOUNG

11. R. W. Harrison and J. Holowach, Final Report,—High Tem- Bellows Capnde Tests, Topical Report No. Z, NASA Contract perature Alkali Metal Valve Test Program, GESP-508, April NAS 3-8523, GESP-436, January 1970. 1970. 16. R. W. Harrison and W. H. HendrixBon, The Compatibility of 12. S. R. Thompson and W. R. Young, InvestigoMon of Cracking Columbium Base Alloys ivith Lithium Fluoride, Topical in T-lll Alloy Tubing Associated vrith Spot Tack Welding, Report No. 1, NASA Contract NAS-3-8523, GESP-261, Sep- GESP-384, General Electric Company, 1970. tember 1969. 13. W. R. Young, Advanced Rankxne Potassium Boiler Develop- 17. W. R. Young, Fabrication of T-lll Test Loop Components, ment Program, T-lll Alloy Defect Studies, General Electric Paper VII, Recent Advances in Refractory Alloys for Spice Company, 1971, unpublished. Power Systems, NASA SP 245, June 1969. 14. P. A. Gnadt, Fitting Heal Storage Tubes for Solar Brayton 18. G. G. Leasmann, Determination of the WeldabUity and Cycle Heat Receiver with Lithium Fluoride, ORNL-RM-2732, Elevated Temperature Stability of Refractory Metal Alloys, July 1970. I. WeldabUity of Refractory Metal Allays, NASA CR-1607, 15. R. W. Harrison and W. H. Hendrixson, Lithium Fluoride August 1970. Mechanical and Physical Properties of Refractory Metals and Alloys

J. B. CoDway Mar-Test Inc.

INTRODUCTION property evaluations of refractor, metals and alloys. Because the author was directly associated This paper was written to provide some discussion with property evaluations, this effort was one of of the technology of refractory metals and alloys the sources of the information used in this paper to for application to space nuclear power. Specifi- define the status of this subject. Other sources cally, this paper was designed to focus on the included NASA-sponsorea programs at Westing- mechanical and physical properties of this group of house and TRW and AEC/ERDA-sponsored materials with the intent of identifying the 3tatus mechanical property program" at the Oak Ridge of existing information. In addition, another objec- National Laboratory (ORNL). tive of this paper was to make some assessment of In all the subsequent discussions attempting to the existing property information and then to define the available material property data base make use of this assessment in determining those for the refractory metals, it should be recognized future material property measurements that must that the sources mentioned above arf some 12 to be made in order to serve the needs of the space 20 years old. Some consideration was given to the nuclear power programs. development of a more up-to-date description of Any review of the existing material property available mechanical and physical property infor- data for the refractory metals and alloys will mation through the use of a detailed review of reveal that much has been done in this area. Such recent literature. Because certain time limitations a review will also reveal that there have been did exist in preparing for this symposium, such an many fairly extensive programs devoted to the in-depth review was not pursued. It is felt, how- research and development of refractory metal and ever, that the review presented here is of sufficieni alloys. One of these was initiated in the early 1960s depth to enable a reasonable position to be defined at the Nuclear Materials and Propulsion Operation regarding the strength capabilities of refractory (later known as Nuclear Systems Programs) of the metals. Before any subsequent experimental pro- General Electric Company in Evendaie Ohio. This gram to address the future needs of the space program was sponsored by the U ted States nuclear power program is initiated, a supplemental Atomic Energy Commission, Division of Reactor review to update the material property data base Development and Technology, Fuels and Materials defined in this paper would definitely be in order. Branch, and encompassed several facets of refrac- For the purposes of this symposium it was tory metal research including development of decided that the review portion of this paper refractory metal processing procedures, develop- should focus on the following properties: ment of advanced refractory metal alloys, the 1. Thermal expansion physical metallurgy of refractory metal alloys, 2. Thermal conductivity development of refractory metal nuclear fuel clad- 3. Enthalpy and heat capacity ding materials, refractory metal thermocouples, 4. Tensile strength radiation effects on fast reactor cladding and 5. Stress rupture and creep strength structural materials, and physical and mechanical 6. Fatigue strength

227 228 CONWAY

Each property will be discussed individually, ?<\d Hoch17 has tabulated enthalpy data as follows: selected examples of available data will be Nb 1004°C to 2324°C presented. In addition, numerous literature refer- Ta 1042°C to 2349°C ences will be cited in the event that the reader Mo 994°C to 2355°C desires more detailed information. This approach is Wo 935°C to 2976°C necessary because the great volume of available Re 990°C to 2368°C property information is too extensive for inclusion in a single paper. Enthalpy data for Ta, Ta-lOW, T-lll, and T-222 ha"e been reported;20 one ploi of these results is presented in Fig. 9. Data fur Ta-lOW calculated REVIEW OF AVAILABLE from the Ta and W da La using weight, fractions INFORMATION yield excellent agreement with the experimental curve in Fig. 9. Thermal Expansion A fairly detailed study of the thermal expan- Tensile Strength sion characteristics of refractory metals was reported in the GE-NMPO program.1"7 A typical A comprehensive review of the cuength proper- plot of thermal expansion measurements is ties of numerous refractory metals and alloys has represented by the data for tantalum (in helium) been published.22 Ultimate and 0.2% yield strength in Fig. 1. Another interesting plot is shown in data for selected materials have been employed to Fig. 2. It compares expansion curves for tungsten yield the plots shown in Figs. 10 and 11. Trend and rhenium; also included is a curve for W-25Re curves have been drawn to proviue an approximate and some calculated data for W-25Re based on a :> mparison of material strength. Actually, if a simple volume fraction approach. A comparison of more detailed comparison of tensile strength is many of the refractory metal expansion curves desired (and, of course, this would bn the case prior developed in this program is presented in Fig. 3. to the initiation of any further experimental pro- Another comparison of these data is presented in grams in support of the space nuclear power Table 1 where the least squares equation constants effort), special attention should be given to tensile are listed for the curves presented in Fig. 3. strength properties as affected by such factors as: product form (sheet vs. rod, etc.), purity (e.g., vari- Thermal Conductivity ous oxygen contents), heat treatment (annealed, stress-relieved, etc.), heat-to-heiat variations, Experimental studies conducted by A. Feith at degree of cold work, and test environment (e.g., GE-NMPO resulted in high temperature thermal vacuum vs. inert gas or liquid metal). conductivity data for several of the refractory metals.8"12 A comparison plot published by Feith12 A fairly straightforward ranking of the is presented in Fig. 4 and a thermal diffusivity strength of the selected materials can be obtained plot published by Feith12 is presented in Fig. 5. from the comparison presented in Figs. 10 and 11. In terms of ultimate tensile strength at 1400°C the High temperature thermal conductivity data for 26 material ranking (strongest to weakest) is: T-lll, Ta and Ta-lOW were listed in a summary report W-25Re-30Mo (data23 from GE-NMPO program), and have been added to the plot in Fig. 4 to pro- TZM, W (based on some extrapolation), and Ta- vide a comparison of data for many of the lOW (based on some extrapolation); these are fol- refractory metals and alloys. lowed by Nb-lZr, Ta, and Nb at much lower strength levels. The top three or four materials Enthalpy and Heat Capacity appear to exhibit a 1400°C strength that is two to Calorimetric studies in the GE-NMPO program three times that of Nb-lZr. This same type of com- led to enthalpy and heat capacity measurements parison can be made in terms of yield strength for many refractory metals and alloys.13"17 Some (Fig. 11) with the same general results, although it typical enthalpy data14 for rhenium and niobium does appear that at 1400° C the highest yield are presented in Figs. 6 and 'i. Another interesting strength is exhibited by the W-Re-Mo alloy. 61 plot of enthalpy data is presented in Fig. 8 based It has been stated that materials considered on tests of the Mo-Re system. This plot reveals for long term applications above 1100°C will proba- good agreement between measured and calculated bly be used in the fully recrystallized condition. data for Mo-50Re using a weight fraction For this reason, it was suggested that the type of approach. plot shown in Fig. 10 would be more realistic if the PROPERTES OF REFRACTORY METALS AND ALLOYS 229

H 0 ! S N V d X 3 H V 3 N II Fig. 2 Linear thermal expansion cf W, Re, and W-25 Re measured in helium. i':;-ni-U: (Solid curves represent least squares fit of actual data points.)

!•••' +

TEMPERATURE, 1000 1500 2000 2500 , PROPERTIES OF REFRACTORY MF'ALS AND ALLOYS 231

I—I - R. H. FINCEL - A. C. LOSEKAMP-

TANTALUM 10 TUNGSTEN"!

-t -I TUNGSTEN- ,25 RHENIUM -

EXPANSION TECHNIQUE . . MICROMETER TELESCOPE \ fit 12 14 16 20 22 24 26 TEMPERATURE °C x 100

Fig. 3 Linear expansion rf various refractory metab tested in helium.

TABLE I Least-Squares Equation Constants for Refractory Metal Linear- Expansion Data L- - L25-C X 100 = Ao+ AjT + ALjjT^Tin' C; all tests in helium)! L25-C Production Temp, range, Material technique Fjrm A« A, A, °C

Tungsten PM* Rod -8.69 x io"3 3.83 X io- 7.92 X 10"8 25 to 2500 PM Sheet -4.58 X 10-•' 3.65 X 10" 9.81 X IO-8 25 to 2500 ACt Sheet -6.76 X 10"J 3.91 X 10~ 8.98 X io-8 25 to 2500 Rhenium PM Sheet -9.79 X 10 3 5.89 X 10" 8.57 X io-8 25 to 2500 Tantalum AC Sheet -6.69 X 10 "3 5.40 X 10" 1.18 X io-' 25 to 2400 Molybdenum PM Rod -8.88 x io -' 3.94 X 10" 1.85 X 10"' 25 to 2500 PM Sheet -5.80 X 10"4 4.59 X 10"4 1.46 X io-' 25 to 2250 AC Sheet -4.84 X 10"3 4.25 X 10 "4 1.71 X 10"' 25 to 2250 Niobium AC Sheet -4.10 X 10 3 5.96 X 10" 1.34 X 10"' 25 tv, 2100 W-25Re PM Sheet -8.46 X 10 4 3.91 X 10" 1.14 X 10"' 25 to 2500 Mo-50Re PM Sheet -1.08 X 10"2 5.30 X 10" 1.26 X io-' 25 to 2250 Ta-lOW AC Sheet -7.34 X 10 3 5.36 X 1.26 X 10"' 25 to 2350

•Powder metallurgy. t Arc-cast. comparison were confined to fully recrystallized Selecting a temperature of 1200°C refra tory mater5 al. Such a plot is presented in Fig. 12. While metal strength (for recrystallized material) is such this comparison reveals the lower tensile strength that an upper limit near 350 MPa (50 ksi) can De charactei istic of recrystallized material, the same expected on ultimate tensile strength along with an general rankings of Fig. 10 are seen to persist. In upper limit of 210 MPa (30 ksi) on 0.2% yield Fig. 12, however, data for ASTAR-8HC are strength. included and reveal strengths comparable to those In one other study,36 short-term tensile data for of the T-lll alloy. Yield strength plots based on a series of niobium alloys have been reported. recrystallized material provide comparisons similar These results are presented in Figs. 13 and 14 and to those derived from the ultimate strength results. reveal that in the region above 1000°C higher 232 CONWAY

CM-H

RVENI M i

ftUwr / //

ft 4 av' r Vt

/ TOO 100y0 WOO !«• 1600 1100 2000 7200 2400 Fig. 6 Enthalpy temperature data for rhenium.

1000 1200 1400 1600 1800 2000 2200 2400 2600 Ttmpariturt, °C

Fig. 4 Thermal conductivity of various refractory metals as a function of temperature.

0.50 I I T

0.45

0.40

0.J5

« 0.20 I

0.15 W-30Rt-HMo

0.10

304 SS a 05

0 200 400 600 IB 1000 1200 1400 Fig. 7 Enthalpy temperature data for niobium. Timptrthin, °C Fig. 5 Thermal diffusivity as a function of temperature for some refractory metals and stainless steels. PROPERTES OF REFRACTORY METALS AND ALLOYS 233

180 TABLES Comparison" of Ultimate Tensile Strength MOLYBDENUM. 160 Data for Tantalum Alloys

CALCULATED ALLOY DATA Ultimate tensile BASED ON WEIGHT FRACTIl" strength Alloy Temperature, °C MPa ksi 120

ASTAR-811C 1092 390 57 I K' ASTAR-1451C 1092 670 97 ASTAR-1611C 1092 810 118

80 ASTAR^ilC 1262 290 42 ASTAR-1411C 1262 50C 73 ASTAR-1611C 1262 630 Si 60

40 rhenium and selected rhenium alloys, and molybdenum and selected molybdenum alloys. In addition 20 to being a helpful reference source, the 1000 1200 1400 1600 1800 2200 2400 Conway-Flagella24 publication contains certain sec- TEMPERATURE, °C tions that are devoted to meaningful comparisons Fig. 8 Enthalpy temperature data for Mo, Re, and Mo- of material strength. Some of these sections have 50Re (actual alloy composition Mo-48Be). been used in the present paper to define a strengths are obtained for D-43, Cb-75* ind FS-85 summary of what is known about the stress- compared to Nb-lZr (Fig. 13). A similar compari- rupture and creep characteristics of refractory son is obtained in terms of yield strength (Fig. 12). metals and then to use this summary to identify It is to be noted, however, that the ultimate tensile areas of future v^ed. In developing this summary it strength data for these niobium alloys are posi- was decided to focus on the temperature regime tioned just above the Ta-lCW data in Fig. 12 and from 800 to 2000°C and to employ information some 70 MPa (10 ksi) or so below the ASTAR- from supplemental reference material.21"22125"27 811C data. An interesting sum n try plot based on 100-h In another fairly recent study,37 existing data rupture strength is presented in Fig. 15. Selecting for Mo-TZM and several tantalum alloys were a temperature of 1400°C from Fig. 15 and employ- reviewed and evaluated in terms of Brayton-cycle ing some extrapolation of higher temperature data power plants for use in space applications. Tensile to 1400°C reveal that the highest strength is ex- data for ASTAR-811C (Ta-8W-lRe-0.7Hf-0.025C), hibited by Re, Ta-lOW, W-25Re, T-lll, and W. ASTAR-1411C (Ta-14W-3Re-0.7Hf-0.025C), and (Data25 for Nb-lZr are not considered to be repre- ASTAR-1611C (Ta-16W-lRe-0.7Hf-0.025C) were sentative of the material as the unusually high compared in the form shown in Table 2. Viewed in strength is probably the result of contamination terms of Fig. 10, these data define strength values, during test) This plot (Fig. 15) also points out a at both 1092°C and 1262°C, that range from about temperature regime (1200°C to ~1600°C) within the level of tungsten for ASTAR-811C to slightly which the available data are limited. Because this above the T-lll data for ASTAR-1611C. temperature regime is probably an important one for space nuclear power applications, it identifies Stress-Rupture and Creep Strength an area of future work. Another observati >n derived from Fig. 15 relates to the attainable 100-h A detailed discussion of the status of the avail- rupture strength for the refractory metals and able stress-rupture and creep data for the refrac- alloys. For example, at 1400°C the upper limit for tory metals and the refractory metal alloys has 24 100-h rupture strength would appear to be in the already been published. This publication region of 70 to 105 MPa (10 to 15 ksi). presented a comprehensive survey of existing liter- A parameter plot24-60 comparing the relative ature on these properties and also included rather strengths of several niobium alloys is presented in extensive high temperature test results obtained in Fig. 16. Based on 100-h rupture the AS-30 alloy the GE-NMPO program. These data, generated in the period from 1963 to 1969, were derived from (Nb-20W-lZr-0.1C) ranks close to TZM (in Fig. 15) studies of tungsten and selected tungsten alloys, i,Text continues on page 238.) 234 CONWAY

9C 10

/ 35 85 A TA--10W > 80 .._. r 75 1/ 75 Y* o 70 / / < J/ 65 < 65 J o J I 60 60 f 55 /

50 o i{ 50 f i <> / / 15 45 • M Hach and H. \.. Johnston O 1. B.Fieidhcu&e and J. C. Hedge f • _oaat *c uor&s cc mputtid 4ff O E»pe 40 _. J —Leas t squaie i comput ed-GE-NMPO / ..35 1 3;

"lOOO .1200 1400 1600 mo 2C00 :20C 2400 '000 1200 1400 1600 1800 2000 220O 2400 1 J 90 LO /

85 f 85 / T--111 » T- 222 7 80 /I ao 7

{ -J 75 / 75

70 / 70 / tNTri / / / / 65 J 60 60 j 55 55 / / 50 50 / / 45 45 m t xp«rrmeintal • ntal f — L.tatt sqv arci computtd -•0*1 tq jor»» co 40 > 40 /

35 IS i 1000 1200 1400 1600 1800 2000 2200 2400 1000 1200 :400 1600 1800 2000 2200 2400 T, #C T, °C Fig. 9 Enthalpy temperature data for unalloyed tantalum and some alloys of tantalum PROPERTES OF REFRACTORY METALS AND ALLOYS 236

1400 Fig. 10 Ultimate teiuile strength vs. tempera tore for selected refractory metals. (O represents data for W-25Re-30Mo.) 1300 — "t- W-25Re-30Mo 0.12-in. rod and 0.012-ln. sheet Ta-lOW 0.060-in. sheet Ta 0.010-in. PM sheet 1200 — Nfc Comm. purity Nb-lZr Recrystallized TZM Wrought A stress rel. T-lll Rolled sheet; stress rel. W 0.16-in. swaged rod; annealed

"-r-r-- • -*«tr-B

Ll TEMP.y °C

1000 1500 236 CONWAY

1400 _£i Fig. 11 Yield strength vs. temperature for -f- selected refractory metals. (O represents data for W-25Re-30Mo.)

100

TEMP., °C

1000 1500 PROPERTES OF REFRACTORY METALS AND ALLOYS 237

1200 — Fig. 12 Comparison of ultimate tensile strength data" for several refractory metals in the recrystallized condition.

1100 —

1000 —

tttfi-Jrr. if. i i i T n rrrr. 238 CONWAY

ksi 700 -100

- 80

- 60

- 20

- 10

200 400 600 800 1000 1200 1400 TEMPERATURE - °C Fig. 13 Effect of temperature on the tensile • trenfth of niobium alloys.

at 1000°C and appears to be in the strength level Fig. 17 reveais that the maximum attainable as Re, Ta-lOW, etc., at 1400°C. It appears that a 1000-h rupture strength the refractory metals at similar strength ranking could also be assigned to 1400°C is approximately 63 to 70 MPa (9 to the F-48 alloy (Nb-15W-5Mo-lZr-0.05C). 10 ksi). A summary plot similar to Fig. 13 was pub- In a review of the Westinghouse and TRi» liter- lished previously24 based on another refractory ature mentioned above, no additional data were metal survey28 and includes 1000-h rupture provided to extend the considerations of 100-h (and strength. This plot is shown in Fig. 17 and reveals 1000-h) rupture strength just described in connec- some of the same deductions made in connection tion with Figs. 15, 16, and 17. For the mosl part with Figs. 15 and 16. For example, the data are the more recent publications that were reviewed limited between —1200 and 1600°C, and the high were oriented toward a study of creep behavior strength materials at 1400°C are Ta-lOW, T-lll, These results will be treated in subsequent para- W-25Re, and perhaps Cb-TZM. (Sylvania A, W- graphs. 0.5Hf-0.02C, is included in this 1000-h compilation A summary plot based on the accumulation of and appears to be the strongest of all.) Finally, 1% creep strain in 100 h is presented in Fig. 17 PROPERTIES OF REFRACTORY METALS AND ALLOYS 239

ksi

500

- 60

- 40

- 20

- 10

2CO 400 600 800 1000 1200 1400 TEMPERATURE - °C

Fig. 14 Effect of temperature on the yield strength of niobium alloys.

and reveals some of the same general strength TABLE 3 rankings estimated for 1400°C from Fig. 12. Focus- Summary of Recently Reported Survey" of ing on the same materials used in Fig. 12, high Creep Data for Several Niobium AlloyB strengths are exhibited by W-25Re (with some in Vacuum of 10"lo 10 'torr extrapolation), W. T-lll, and Ta-lOW. The T-222 Teat temper- Range of creep alloy was not included in Fig. 12 but is seen here Alloy Comimfition ature*, °C dau,h* to exhibit high creep strength along with the NAS-36 and ASTAR-811C alloys which appear to D-43 Nb-lOW-lZr-O.lC 1204 90 FS-85 Nb-27Ta-10W-0.8Zr 982,1204 350.3000 be the strongest of all. Creep strengths of several B-66 Nb-5Mo-5V-lZr 1204 30 alloys are presented in parametric form in Fig. 15. Nb-753 Nb-5V-1.25Zr 932,1204 35,6 Some additional comments are in order in con- Nb-lZr Nb-lZr 982,1204 2,10 nection with Figs. 18 and 19 based on the more recent publications referred to above. For example, •First value where there are two entries is for 982°C. an evaluation36 of the available creep data for °ev- eral niobium alloys led to plots from which 1% One other comparison of 1% creep strength was creep strengths in 100 h could either be read 38 presented in parametric form and is shown in directly or estimated by extrapolation. (See Fig. 20. It will be noted that this format is the Table 3 for summary of test data.) These values same as that presented in Fig. 19, and it might are not shown in Fig. 18, but it can be noted that seem redundant to include the representation the highest creep strength (1% in 100 h) is ex- shown in Fig. 20. It has been used, however, in this hibited by the D-43 and FS-85 alloys; these values section because it incorporates data not originally at 982°C and 1204°C are close to the creep strength included in the comparison of Fig. 19. For exam- exhibited by the Ta-lOW and T-lll alloys. ple, data for ASTAR-811C are included in Fig. 20 240 CONWAY

100— Fig. 15 Comparison of iOO-h rupture data for selected refractory metals.

500 —

100 •

600 . "C PROPERTIES OF REFRACTORY METALS AND ALLOYS 241

600 H00°C 1200'C 1300°C 14OO°C 500 1111 400

300

200

I00 - 90

60 50

90 34«K>" 10

P = T (15 + log tR); T°R, tR hrs

Fig. 16 Comparison of rupture strength of AS-5S, AS-30, F-48 and Nb-IZr.:

10,000—I TEMPERATURE («F| 10OO 2200 2600 3000 3400 3800

TZM (I) -STRESS RELIEVED ATI23CC TZM (2)- RECRYSTALLIZED AT IS40«C T ZM (3) - 1540'C. WARM WORKED AT IO4O*C, ANO STRESS RELIEVED AT123O*C TZM(4)-ANNEALED 2093*C, RE-EXTRUDEO, AND WARM WORKED AT 13I5*C Cb-TZM (5)-ANNEALED 2O95«C. RE-EXTRUDEO, AND WARM WORKED AT I31S'C Cb-TZM (6) ANNEALED 1927*C AND WARM WORKED AT1370*C SU-l6o-1hr ATI200*C SU-16b-1hr ATt600*C, 3hrATI2OO'C 1000 — T-222-2hr AT1315-C W-25%R«(ARC MELTED)-lhr AT 1700*C W-25% R«(POWDER MET.)-.hr AT 1600°C FS-85o-1hrAT137O*C FS-85b-lhr AT 1595'C D-43 -STRESS RELIEVED 10mm AT1205«C D-43M-V2hr ATI650'C, 25% COLD WO'K, AND Ihr AT1425°C Cb-752o-lhr ATI3I5*C Cb-752b-1hrAT1S95#C B-66-C0LD WORKED 50% Cb-753-ihr ATI2O5«C Nb-l%Zr-COLD WORKED Mo-0 5 %Ti- STRESS RELIEVED To -10% W-COLO WORKED if. 100 T-IH -COLD WORKED SYLVANIA A-UNKNOWN W-2%ThO2-WARM WORKED UNALLOYED W- 5min AT II5O*C UNALLOYED Ta-COLD WORKED

-4— ±i W-2% ThOo

UNALLOYED W I POWDER MET.) 10

T- w-ac%R(i ^•".1 (POWDER MET,)

5»10' 900 1100 1300 1500 1700 1900 i'00 2300 TEMPERATURE

Fig. 1? Stress to produce rupture in 1000 h for selected refractory metals and refractory metal alloys.28

these two materials are stronger than T-222, Ta- and, for some materials, to test durations clos: to 10W, W, and T-lll. Data for W-25Re are limited, 20,000 h. Other NASA-sponsored publications that but it would appear that in this temperature range are pertinent to the creep behavior of refractory of 1200 to 15OO°C this material would be creep- metals are: ASTAR-SilO,46 Mo and Mo-3% to strength-comparable to UAS-36 and ASTAR- 811C. 7% Re,47 C-103,48 FS-85,49 Ta-lOW,50 and T-222.51 Perhaps the most extensive source of long-term, In a recent review of existing data, English37 high-temperature creep data for the refractory cited several factors that were considered impor- metals has resulted from the NASA-sponsored tant in the choice of materials for high- program at TRW Materials Technology performance, long-lived Brayton-cycle power plants Laboratories.39"45 This progr-.m reported creep for use in space. One of these factors was associ- data for W-25Re, Mo-TZC, Mo-T7M, Ta, T-lll, ated with the impressive data base that existed for Ta-lOW, ASTAR-811C, ASTAR-1411C, and two refractory alloys (Mo-TZM and ASTAR-811C; ASTAR-1611C tested in vacuum (10~8 to 10~u some 344,000 test hours of creep data). It was also torr) at temperatures to 1593°C (for ASTAR-81KT noted that each of these alloys is compatible with PROPERTIES OF REFRACTORY METALS AND ALLOYS 243

100— 3 -— Fig. 18 Compar'son of stretiea to give 1% creep in 100 fa far selected refractory metale. '" 1 8 500 —

100 —I

10—

600 800 1000 1200 1400 1630 1G00 2000 244 CONWAY

29 1% Creep in 10,000 !irs the parametric analysis of refractory metal data 1000 —i iooo°c noo°c i2oo°c is presented in Fig. 24 where a comparison is made using data for several materials.

Fatigue Strength

Very little has been accomplished in the evaluation of the fatigue characteristics of the refractory metals and alloys. Some load-control studies have been reported for Ta,25 Ta-lOW,26 and Nb25 but these have involved tests at room temperature. Only two studies of strain-controlled fatigue behavior at temperatures above 800°C were identified in this limited survey. The one study30 involved T-111 and ASTAR-811C (Ta-8W-J Re-0.7Hf-0.025C) at 11.5O°C tested in vacuum in diametral strain 56 39 41 43 45 47 49 5] *J control; the other involved cross-rolled tungsten

P = T (15 + Log t) x 10"3 (T°R, t hrs) tested in diametral strain control at 815°C in a vacuum of 1 X 10~6 torr. Obviously, if fatigue is judged to be an important consideration for nuclear space power applications, much remains to •IAS-36 Ta-a. 7W-1.56Re-O. 7Po-0.25Hf-f>. 13Zr-_015C-0.015N be done in this area. T-I11 Ta-8W-2"f Using room temperature tensile properties in 31 •-222 Ta-10.4W-2.4Hf-0.01C conjunction with the Universal Slopes equations, 21 F:-85 'ltj-28Td-10.5W-0.9Zr Collins, Moteff, and Chandler have made predictions to yield life values for a strain range of 2%. Fig. 19 Comparison of 1 % creep strength of various alloys.32 This evaluation provided the following:

Cycles to failure both inert gases and the alkali metals (lithium and Material at room temperature sodium). Furthermore, the ASTAR-811C alloy was cited as being weldable and ductile in the postweld Nb 1100 condition (Mo-TZM, on the other hand, was found Nb-lZr ctOOO to be brittle in the postweld condition). English Mo 900 reported on a Larson-Miller analysis (see Fig. 21) TZM 1000 of the ASTAR-811C data (time to 1^ creep), iden- Ta 1600 tified 2-sigma :imits, and then used these in T-lll 1300 defining allowable design stresses for \% creep in Ta-lOW 900 40,000 h. A simila • analysis for Mo-TZM data was also reported. Creep behavior cf the T-111 alloy hat, been FUTURE NEEDS noted44 to be affected by preexposures to vacuum or to liquiu lithium particularly in the temperature Before any of the specifics of additional prop- regime below 1100°C. However, no significant erty information that might be required in order to effect was noted in specimens tested at 1316°C fol- achieve the objectives of the nuclear space power lowing a 5,000-h preexposure to lithium. program are discussed, it is important to consider two general aspects of refractory metal behavior. A detailed analysis of refractory meta' data has These have both been alluded to, if not actually been reported24"29 based on parametric correlation. mentioned, in the previous section of this paper but One example from this work is presented in warrant special attention in this section prior to Fig. 22 where the data for tungsten and molybde- any detailed definition of future needs. Actually, num are normalized in terms of the homologous both of these points are all too familiar to those temperature. Another example is presented in associated with refractory metal technology, but it Fig. 23 for tungsten data and involves creep as is felt *hat some reiteration of these response well as rupture information. Another example of characteristics is warranted. PF 0PERT1ES OF REFRACTORY METALS AND ALLOYS 246

1% CREEP IN 1000 HRS 1000°C ' 1100°C 1200°C 1300°C 1400°C 200

100 Spec.Nc Alloy Gravity Ig/en .3 - a. NbiZr 8.6 o." Nb10H» 8.8 FS85 10.6 T 111 167 1222 16.8 ASTAR411C 16.7 V) c/3

10 J I I I L I " 36 38 40 42 44 46 48 50 52 54 56 58 60 62 64 66 63 70 PLM - (T + 460) x TO3 (15 + log t)

Fig. 20 Comparison of the creep strengths of sime refractory metal alloys.

The first observation evolves from the discus- design information for the refractory metals. sion of tensile properties presented in connection Therefore, test procedures and material history with Figs. 10 and 11. It was noted that many fac- must be considered in any interpretations that are tors (product form, method of manufacture, etc.) made or conclusions that are formed based on the can have a pronounced effect on the magnitude of comparison plots presented in the previous sec- the property being n. asured. It is not very mean- tions. ingful to cite a /alue for a tensile property (or any The second important aspect of refractory property for that matter) of a refractory without metal response relates to heat-to-heat variation. an accompanying definition of the material and the This is not unique to refractory metals because it test procedures employed. Numerous investigators is a continuing item of interest in connection with have reported on the type of property variability fhe materials associated with ASME Boiler and that can be expected for refractory metals. For Pressure /essel Code activities. The few studies of example, ten^i'e liata for niobium have been 25 v refractor' metals involving multiple heats have reported that ir. ,lve tensile strengths ranging shown that it is indeed an important consideration. from ~7 to 40 ksi at <70°C; yield strength data for 33 25 For example, data have been reported for five dif- niobium have been found to range from ~140 to ferent lots of sintered tungsten tested in vacuum at 385 MPa (20 to 55 ksi) at 650°C; and tensile 2650°C. Strength variations from heat-to-heat are strength data for different forms of tantalum 26 shown in Fig. 25. tested at room temperature have been reported to In organizing this discussion of specific mate- range from 190 to 1260 MPa (27 to 180 ksi). This rial property needs of the future, the subject was phenomenon must be carefully dealt with in any divided into physical and mechanical properties. future program dealing with the generation of Then, in the case of mechanical properties, some 246 CONWAY

400 TWO - SIGMA VALUES FOR 1 PERCENT 20.000 CREEP IN 400(0 HOURS • Mo 1200° lo 2400°r 1300 K. 72 M 154 MPa 100 0.08 to 133 h, A W 1600° io 3000°C 1400 K. 14 ks 98 MPa 10.0W 300 1500 K. 5ksl 35 MPa y- 0.78 ro 256 h< Iksi 7 MPa K »\ 6,000

o 4.000 • : 200 s

in •"' —2.000 in L±J 1J or or 10 \ or t- j> 1.000' 100 * Mo 2283°K 600 Tu (or W - 3A83°K V [ 400 P "^(11.95 < lo9lR)

r><- 8 IR ,n hou.• 45 50 55 60KIO3 T ,n °K * — U9S0N - MLLER PARAMETER 1 2 4 6 8 10 12 14 T°R (15 + log tR); time in hours Larson-Miller parumeier, P Fig. 21 1% creep of ASTAR-811C including 21 tests Fig. 22 Larson-Miller parameter plot for stress rupture for a total of 149,084 h. data for Mo and W using homologous temperatures.

general needs are listed, followed by a description 2. Establish low-cycle/high-cycle fatigue test of needs within each alloy category (tantalum base, systems for evaluations in ultrahigh vacuum 8 10 etc.). (10" to 1CT torr); a goal should be set to acquire 6 systems. Physical Properties a- Perform a study to e/aluate the need for fracture toughness and crack growth mea- This effort should be initiated by first evaluat- surements, and, if these are needed, plan to ing the need ftr any additional data. It is possible establish appropriate test facilities. that sufficient accurate information is available to 4. Due to the many discrepancies that have satisfy the needs in this area; it is also important been noted (the widely different strength to note that with the exception of thermal conduc- properties reported for Nb-lZr is but one tivity enough is known about these properties that example) in the existing literature on good estimates of property information can be mechanical properties, a detailed reassess- obtained from measured and reported data of ment of all existing data should be per- closely relat-jd materials. (In other words, the vol- formed; emphasis should be given to the two ume fraction and weight fraction approaches noted most likely candidate alloys in the in connection with Figs. 2 and 8 might lead to esti- tantalum-, molybdenum-, tungsten-, and mates of physical property uata that are reliable niobium-base systems. enough to make further testing unnecessary.) 5. As a final general need, the topic of data analysis deserves special attention Since Mechanical Properties* about 1970, an in-depth study of parametric methods for creep and rupture data analysis General has been in progress in connection with what was called the Parameter Task Force. 1. Reestablish creep-rupture test systems for A major objective of this effort was to 8 evaluations in ultrahigh vacuum (10~ to review existing parametric methods, evalu- 10 ~10 torr); a goal should be set to acquire 20 ate their effectiveness, and identify and rec- systems. ommend a standard practice for such analysis. While the latter was not fully achieved, •In thiB discussion it has been assumed that allowable the developments57 of the task force were stresses will be established using creep strain data; on the basis of this assumption the generation of rupture data is not neces- substantial and have contributed to a deeper sary. understanding of data analysis in general. PROPERTIES OF REFRACTORY METALS AND ALLOYS 247

Parameter P fof temperature, °K 25 3E 40 45 50 56

Constant, C

o 1% elongation 20.28 a 3% elongation 15.3* o 5% elongation 13.84 a 10% elongation 12.65 Rupture 1230

40 50 60 70 80 90 100 110 120 Parameter P for temp-mature, °R

Fig. 23 Larson-Miller parameter plot defining the creep-rupture characteristics of arc-cast tungsten.

Ma A-C '200 -3400 0.04-5*0 II .97 1000 W, A-C 1600 - 2100 0.71 - 420 II 10s PM 1600 - 2900 0.4] - 62 9 42 . PM 1600 - 2100 0.34-350 6 _ - _ , A-C 7'00 - 2600 0.37 - 14 9.95 — lo - 10*, A-C 1600 - 2900 1.23 - >M 10 53 T (C • lo W . 2SQ. ->M 1600 - 3100 0.05 -MO 8 89 T = .«..., °K w-. JORr, PM 1600.2000 19-190 0 41 im.a, h, -25S. 30Mo, PM 1600-2300 15-100 5. 70 — t-M cai>> C = •ant w.• 30R>-30Mo. PM I6O0.20O0 3.1 . 330 11 10 **> M 23W 1.1-230 II 47 100 TZ M, A-C 7000 - 2200 1.0 - 6.0 20 70 10' V \\v H— \" 1, PM

\ \ _ — M* - 50R«. Pit s A1 \ 1 Si 10 t. A-C », A-C 10' \

S»a-30Ua —W y Ra « - 23H» PM w*^ Ta, l\\ ' Ha. X.

10,000 20,000 30,000 40,000 50.000 60.000 P

Fig. 24 Summary of Larson-Miller parameter data for atress-rupture behavior of refractory metals and refractory metal alloys tested in hydrogen. 248 CONWAY

30 4000

20 M. i \

^-| — .—— 2000 Lot j 0 X T rs ——• O A V . ———• 10 o O B •• — ^—• o c 8 .— CO —• tf D 1000 H •- 0 E - 800 "5 Ref.34, 2650°< , Cl - _—- 600 I I ffl -3 10" 10" 10 1 Creep rate, es, sec" 30 4000

•** 1 V . 2000 —.

T •^, n ^ •*— •* o~—' -». —•. V O 10 o~—— " . — *^ <= 1000 D -Ot —•—.

1 o A —• — a a •«^ 600 o c V o D E 400 - "

As examples, the minimum commitment mental programs to generate creep and rup- method of Manson and Ensign, the general- ture information. ized regression model approach of Booker, the heat-separation approach of Sjodahl for Tantalum-Base Alloys multiheat analysis, and the preanalysis assessment concept of Conway all form a 1. Supplement existing short-term tensile data part of what might be termed the data anal- for ASTAR-811C and T-lll by performing ysis approach of the 80s. Much of this tech- additional tests to cover the temperature nology has been employed in ASME Boiler range from 1200 to 1500°C. and Pressure Vessel Code activities directed 2. Further amplify the data base for short- toward the determination of allowable term tensile behavior of ASTAR-811C and stresses. It is recommended, therefore, that T-lll by evaluating four heats of each mate- full advantage be taken of this past experi- rial. ence in all analyses of any new data gen- 3. Perform short-term tensile tests to evaluate erated (or any reanalysis of existing data) in strain-rate effects using at least two heats connection with the space nuclear power of ASTAR-811C and T-lll; strain rates of program. An application of these recently 8 X 10~5, 8 X KT4, and 4 X HT3 sec"1 developed concepts is certain to yield a bet- are suggested. Temperature range is to be ter interpretation of material behavior and a from room temperature to 1500°C. more reliable definition of long-term projec- 4. Evaluate existing creep data for ASTAR- tions; furthermore, the suggestions made in 811C and T-lll with the objective of estab- connection with the preanalysis assessment lishing creep strengths for 0.5, 1.0, and 2.0% concept could prove useful in guiding experi- creep strain. This evaluation will suggest PROPERTIES OF REFRACTORY METALS AND ALLOYS 249

any additional new data that must be gen- On the basis of the potential that is judged erated in order to develop reliable defini- to exist for this alloy after this testing is tions of creep strength. Temperature range concluded, tests should be performed to eval- to be covered is 800 to 1500°C, with stress uate strain rate effects as outlined in item 3 levels selected to yield test durations to (tantalum-base alloys); then the creep- 10,000 h. strength studies described in item 4 should 5. Expand on the need defined in (4) by be performed, and, if necessary, the fatigue evaluating four heats of each alloy. evaluations of item 6 should be initiated. 6. Review existing information on the low cycle 2. Evaluate the existing short-term tensile data and high-cycle fatigue characteristic? of for the series cf alloys containing 5 to 15% ASTAR-811C and T-lll and determine the Re; the Mo-50Re alloy should also be additional tests that are needed to provide included in this evaluation. Perform addi- an accurate definition of fatigue life over the tional short-term tensile tests that appear to regime from a few hundred cycles to about be required to establish a reliable appraisal 1 X 106 cycles. Testing should include at of the potential of these materials. least two heats of material for each alloy 3. Based on the results obtained in (2), select and strain rates of 4 X 10~4 and 4 X 10~3 the most promising Mo-Re alloy and oucline sec~'. Test temperatures to be employed are a detailed experimental program to furnish room temperature, 600, 1000, and 1500°C. the necessary design information. Multiple- 7. Data for fracture toughness and cyclic crack heat evaluation, strain-rate effects, creep- growth to be generated for ASTAR-811C and strength determination, and fatigue-strength T-lll as design needs dictate. identification should all be made a part of this experimental effort. The temperature Niobium-Base Alloys range to be covered in all the Mo-base alloy evaluations is from room temperature to 1. A detailed reevaluation of the existing ten- 1500°C. sile and creep data for niobium-base alloys is recommended in order to address the dis- Tungsten-Base Alloys crepancies that have been discussed in previ- 1. This evaluation should follow that discussed ous sections of this report. in items 2 and 3 under molybdenum-base 2. Use the results obtained in (1) and select alloys but with W-25Re and W-25Re-30Mo as three alloys for further testing; Nb-lZi, D- candidate materials. Another difference is 43, and C-103 are recommended. Evaluate that the upper temperature limit should be short-term tensile behavior of two heats 1800°C. each of these alloys and cover the tempera- As a final comment on future needs, the proper- ture range from room temperature to ties of welded material should be mentioned along 1100°C. with the su, ;ect of the effect of radiatic "n mate- 3. On the basis of the results obtained in (2), rial properties. Neither of these topics was covered select the two alloys that appear to be the in this discussion, but both these issues are impor- most promising and perform an in-depth ts at and should be given special consideration in evaluation of these materials. This in-depth ail future program plans. evaluation should involve two additional heats, the strain-rate study described in REFERENCES item 3 under tantalum-base alloys, the creep-strength study proposed in item 4, and 1. J. 3. Conway and A. C. Losekamp, Thermal Expansion the fatigue-strength determination proposed Characteristics of Several Refractory Metals to 2500°C, in item 5. Test temperatures to be ut.ed ii^ TAIME, 236: 702 (1966) this in-depth evaluation are to be in the, 2. J. B. Conway and A. C. Losekamp, Refraction Error in a range from room temperature to 1100°C. ! Comparator Method for Measuring Thermal Expansion, Rev. So. Inst, 36(8): 1245-1246 (1965). 3. J. B. Conway, R. M. Fincel, Jr., and A. C. Losekamp, Effects Molybdenum-Base Alloys of Contaminants on the Thermal Expansion of Tantalum, TAIME, 233: 841-842 (1965). 4. J. A. Edwards, Thermal Expansion of Various Materials, 1. Further evaluation of Mo-TZM should be TM 04-2-9 (1964). pursued through the short-term tensile test- 5. J. B. Conway, Properties of Refractory Metals, III—Thermal ing of three additional heats of the material. Expansion Characteristics of Tungsten, Rhenium, Tantalum,, 250 CONWAY

Molybdenum, Niobium, W-iSRe, Ta-lOW and MoSORe, GE- 29. J. B. Conway, Stress-Rupture Parameters for the Refractory NMPO, GEMP-375, June 29,1965. Metals—Time-Temperature Parameters for Creep-Rupture 6. J. B. Conway, R. M. Fincel, Jr., and A. C. Losekamp, The Analysis, pp. 175-198, ASM No. DS-100, American Society Linear Thermal Expansion of Niobium, TAIME, 233: for Metals, 1968. 844-845 (1965). 30. K. D. Sheffler and G. S. Doble, Thermal Fatigue Behavior 7. High Temperature Materials Program Progress Report No. of T-lll and ASTAR 811C in Ultrahigh Vacuum, Fatigue at 6S, GEMP-63, Dec. 30, 1966. Elevated Temperatures, pp. 491-499, ASTM STP 520, Ameri- 8. A. D. Feith, Thermal Conductivity and Electrical Resistivity can Society for Testing and Materials, 1973. of Molybdenum, GE-NMPO, CE-TM 65-10-1. 31 S. S. Manson, Fatipiie: A Complex Subject—Some Simple 9. A. D. Feith, The Thermal Conductivity and Electrical Approximations, Exp. Meek, 193, Jnly 1965. Resistivity of Tungsten-06 Weight Percent Rhenium, GE- 32. T. A. Moss, Materials Technology Presently Available for NMPO, GEMP-562, November 1967. Advanced Ranldne Systems, Nuc AppL, 3(2): 71 (1967). 10. A. D. Feith, The Thermal Conductivity and Electrical Resis- 33. E. C. Sutherland and W. D. Klopp, Observations of Proper- tivity of Tungsten-26 Weight Percent Rhenium, in ties of Sintered WrougfU Tungsten Sheet at Very High Tem- Proceedings of the Seventh Conference on Tkei-mal Conduc- peratures, NASA TN D-1310, February 1963. tivity, National Bureau of Standards Spec. Tech. Pub. 302, 34. W. V. Green, Short-Time Creep-Rupture Behavior of f.. 371-380, September 1968. "hjngsten at 2250-C to 2800°C, Trans. AIME, 215: 1057-1060, 11. A. D. Feith, Thermal Conductivity and Electrical Resistivity December 1959. of Molybdenum, GE-NSP, GE-TM 65-10-1, October 1965. 35. R. L. Stephenson, Creep-Rupture Properties of Unalloyed 12. Eighth Annual Report—AEC Fuels and Materials Develop- Tantalum, Ta-10%W and T-lll Alloys, ORNL-TM-1994, ment Program, GEMP-1012 Part I, March 31,1969. December 1967. 13. M. Hoch and H. L. Johnston, A High Temperature Drop 36. L. J. Pionke and J. W. Davis, Technical Assessment of Calorimeter; The Heat Capacities of Tantalum and Niobium Alloys Data Bast' far Fusion Reactor Applications, Tungsten Between 1000° and 3000°K, J. Phys. Chem., 65: 855 COO-4247-2, U.S. Dept. of Energy Contract No. (1961). EG-77-02-4247, McDonnell Douglas Astronautics Co., St. 14. J. B. Conway and R. A. Hein, E;ithalpy Measurements of Louis, Missouri, August 1979. Solid Materials to 2400°C by Means of a Drop Technique, in 37. Private communication from NASA-Lewis Research Center Advances in Thennophysica' Properties at Extreme Tem- to E. E. Hoffman; excerpt from unpublished paper entitle^ peratures and Pressures, ASME, 1965. "Refractory Alloys and Advanced Materials for Brayt 15. R. A. Hein and P. N. Flhgella, Enthalpy Measurements of Cytle Application" by R. E. English of NASA-Lewis UO, and Tungsten to S260"K, General Electric Company, Research Center. NSP, GEMP-578, Feb. 16, 1968. 38. SP-100 Conceptual Design Studies First Interim Briefing, 16. J. B. Conway, R. A. Hein, R. M. Fincel, Jr., and A. C. Jet Propulsion Laboratory Contract No. 956474; Westing- Losekamp, Enthalpy and Thermal Expansion of Several house Electric Corporation, Advanced Energy Systems Divi- Refractory Metals to 25WC, General Electric Company, sion, Madison, PA, June 8,1983. NSP, TM 64-?-8. 39. Creep Behavior of T-lll Alloy Under the Influence of Con- 17. M. Hoch, The High Temperature Specific Heat of Body- tivuously Varying Stresses, Topical Report No. 2, TRW-ER- Centered Cubic Refractory Metals, GEMP-696, February 7, 7373, NASA Contract NAS 3-9439; TRW Equipment 1969. Laboratories, Cleveland, Clio. 18. K. K. Kelley, Contributions to the Data on Th rretical Metal- 40 Generation qf Long Time Creep Data on Refractory Alloys at lurgy; XIII High-Temperatvre Heat Content, Heat Capacity Elevated Temiteratu, es. Final Report, TRW-ER-7442, NASA and Enthalpy Data for the Elevienti and Inorganic Com- Contract NAS 3-9439; TRW Material Technology Labora- pounds, Bureau of Mines Bull.. No. 584, 1960. tories, Cleveland, Ohio. 19. I. B. Fieldhouse, J. C. Hedge, and J. I. Lang, Measurements 41. Generation of Long Time Creep Data on Refractory Alloys at of Thermal Properties, WADC TR-58-274, November 1958. Elevated Temperatures, Twenty-Seamd Quarter; Numerical 20. Sixth Annual Report—High Temperature Program, Part A, Creep Data, NAS 3-9439, NASA Contract NAS 3-9439; TRW GEMP-475A, March 31, 1967. Materials Technology Laboratories, Cleveland, Ohio, Sep- 21. C. G. Collins, J. Moteff, and B. A. Chandler, Evaluation of tember 19,1969 to May 7,1970. the Potential of Selected Alloys for Use as a Fuel Cladding 42. Analytical Studies of the Variable-Stress, Variable- Material in an LMFBR, General Electric Company, GEMP- femperaticre Creep Behavior of T-lll Alloy, Topical Report, 573, November 1967. NASA-CR-72771, NASA Contract NAS 3-13469; TRW 22. Aerospace Structural Metals Handbook, Vols. I & II: ASD- Materials Technology Laboratories, Cleveland, Ohio. TDR-64-741, March 1963, and Supplements 1 & 2, December 1963 and March 1965; revisions through 193». 43. Generation of Long Time Creep Data on Refractory Alloys at 23. High-Temperature Materials Program Progress Report No. Elevated Temperatures, Final Report, NAS CR-72997, NASA 5S, Part A, GEMP-53A, November 26,19bC. Contract NAS 3-13469; TRW Materials Technology Labora- 24. J. B. Conway and P. N. Flagella, Creep-Rupture Data for the tories, Cleveland, Ohio. Refractory Metals to High Temperatures, Gordon and Breach V Generation of Long Time Creep Data on Refracton' Alloys at Publishers, 1971. Elevated Temperatures, Final Report, NAS CR-134481, 25. The Engineering Properties of Columbium and Columbium NASA Contract NAS 3-15554; TRW Materials Technology Alloys, DMIC Report 188, September 6,1963. Laboratories, Cleveland, Ohio. 26. The Engineering Properties of Tantalum and Tantalum 45. K. D. Sheffler and R. R. Ebert, Generation of Long Time Alloys, DMIC Report 189, September 13,1963. Creep Data on Refractory Alloys at Elevated Temperatures, 27. The Engineering Properties of Molybdenum and Molybdenur- TRW, Cleveland, OH, TRW-ER-7648, September 1973; also Alloys, DMIC Report 190, September 20,1963. NASA CR-134481,1973. 28. ORNL Metals and Ceramics Division Annual Progress 46. W. D. Klopp, R. H. Titran, and K. D. Sheffler, Long Time Report For Period Ending June SO, 1967, ORNL-4170, Creep Behavior of the Tantalum Alloy ASTAR 8UC, NASA- November 1967. TP-1691 (1980). PROPERTIES OF REFRACTORY METALS AND ALLOYS 261

47. W. D. Klopp and W. R. Witzke, Mechanical Properties of 54. V. C. Davis, Elevated Temperature Tensile Properties of Electron Beam Melted Molybdenum and Dilute Recrystallized Nb-lZr, Fansteel Research Report dated Molybdenum-Rhenium Alloys, NASA TMX 2576 (Mo, Mo- December 12,1966, Far.steel Inc. 3Re, Mo-5Re, Mo-7Re), June 1972. 55. V. T. Bates, TensOe Properties of Fansteel "991" Metalr 48. R. H. Titran and W. D. Klopp, Long Time Creep Behavior of Recrystallized Sheet, Fansteel Researcn Report dated the Niobium Alloy C-1OS, NASA-TP 1727 (C-103: NB- December 12,1966, Fansteel Inc. 10HMT: -0.7Zr), October 1980. 56. R. E. Schmunk and G. E. Korth, Tensile and Low-Cycle 49. Zyde D. White, Apollo Experience Report—Electrical Wiring Fatigue Measurements on Croe- Rolled Tungsten, J. of Nucl Subsystem, NASA-TN-D-7885, March 1975 (Alloy ?S-85). Mater., 103 and 104: 943-948, North Holland Publishing Company, 1981. 50. R. H. Titran and W. D. Klopp, Long Time Creep Behavior of Tantalum-10 Tungsten in High Vacuum, NASA-TN-D-6044, 57. Characterization of Materials for Service at Elevated Tem- October 1970 (Ta-lOW). peratures, pp. 247-536, MPC-7, The American Society of Mechanical Engineers, 1978. 51. R. H. Titran, Creep Behavior of Tantalum Alloy T-2S2 at 58. Fifth Annual Report High-Temperature Materials PrtQram, 1S65 to 1700°K, NASA-TN-D-7673, June 1974 (T-222). GE-NMPO, GEMP-400A, February 28,1966. 52. M. Schussler, Properties of Columbium Alloy Cb-753, in 59. Sixth Annual Report High-Temperature Materials Program, Refractnry Metal", and Alloys IV—Research and Develop- GE-NMPO, GEMP-475A, February 28, 1967. ment, Metallurgical Society Conference, Vol. 41, pp. 387-404, 60. Recent General Electric Developments in Columbium-Base Gordon and Breach Science Publishers, 1967. Alloys, FPLD, Cincinnati, Ohio, February 1962. 53. G. G. Lessmann and R. E. Gold, Determination of the Weld- 61. R. L. Ammon, comments made during review of original ibility and Elevated Temperature Stability oj Refractory manuscript. Metal Alloys. Task Ill—Long-Time Elevated Temperature 62. L. J. Pionke and J. W. Davis, Technical Assessme-.a of Stability of K>fractory Metal Alloys, WANL-PR-P014, Wes- Niobium Alloys Data Base for Fusion Reactor Applications, tinghouse Astronuclear Laboratory, October i960. COO-4247-2, August 1979. Effects of Irradiation on Properties of Refractory Alloys with Emphasis on Space Power Reactor Applications

F. W. Wiffcin Oak Ridge National Laboratory*

INTRODUCTION each alloy system (where such a review is available) and to discuss that review and more recent High service temperatures in a nuclear reactor data judged to be the most useful in establishing core, above — 700°C, will require the use of refrac- likely behavior in high-temperature reactor service. tory metal alleys for containment, support, and fuel-coolant interfaces. Use of these alloys in the TABLE 1 projected neutron flux can lead to changes in many Alloys 1Considered in Evaluating the of the properties that govern successful reactor Effects of Neutron Irradiation on operation. These can include changes in the physi- Refractory Metals cal properties, mechanical properties, dimensional stability, and thermodynamic stabili^. Nominal The refractory metal alloys are based on V, Nb, Alloy composition, Mo, Ta, and W. While the vanadium alloys may Alloy aUtu3* wt.% have too low a temperature capability for use in a f V Model V, unalloyed high-tempera .ure reactor, the similarity of this V-20 Ti Model V-20 Ti a!)oy system to •' » other mentioned alloys make3 V-10 Cr Model V-10 Cr data on v&r.^diun. useful in projecting the behavior VANSTAR 7 Mode! V-9 Cr-3.3 of alloys for which data are not available. Simi- Fe-1.3 larly, data on unalloyed .r.etal must often be used Zr-0.05 C to inf jr the r>robab!e behavior of alloys. The metals Nb Model Nb, unalloyed and alloys t >nsidered in this report are listed in Nb-1 Zr Candidate Nb-1 Zr Table 1, anr! their nominal compositions are given. Mo Model Mo, unalloyed Mo-0.5 Ti Model Mo-0.5 Ti The p.obable effects of irradiation on niobium TZM Model Mo-0.5 Ti- and tungsten alloys . i use as components of 0.08 Zr-0.015 C thermionic converters in a space reactor were Mo-Rp Candidate Mo, with range ^viewed by the author in 1971, based on data of Re to 50% 1 1 available at that time. While cons.:! rnbly more Ta Model Ta, unalloyed data on refractory metals have been generated Ta-10 W Model Ta-10 W since that time, the data have not been reviewed T-lll Candidate Ta-8 W-2 Hf ASTAR-811C Candidate Ta-8 W-l Re- with respect to space reactor applications. This 1 Hf-0.025 C paper attempts such a review. However, no attempt has been made to cite all relevant work. vV Model W, unalloyed W-Re Candidate W, with rpnge of The approach used is to work from the most Re to 25% recently available review of irradiation effects for *Mode' systems are alloys for which radiation effects data are available. Candidate alleys are •Operated for the U. S. Departmeni of Energy under con- those considered for use as space nuclear r-actor tract W-7405-eng-26 with the Union Carbide Corporation. components

252 EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 253

The United States has never mounted a focused, TABLE 2 sustained program to evaluate the effects of reac- Envelope of Irradiation Condition for Space tor irradiation on the properties of refractory Reactor Refractory Met*' Components metals and alloys. There have been short-lived programs to evaluate these candidate alloys for vari- Temperatures: ous space reactor concepts, for thermocouple use, Cold shutdown 20°C Minimum operating 800°C for fast breeder reactor core internals, and for Maximum operating (These will be material fusion reactor structures. There has also been a limited.) small but continuing fundamental radiation effects Near-term goal 1100°C program interest in the bcc refractory metals. By Advanced goal 1400°C piecing together the data generated by these Thermionic converter 1300 to 1700-C programs and evaluating that data within the Operating environment: framework of the much more extensive body of Liquid metals—Li, K Inert gas—He data and understanding on the austenitic stainless Vacuum, with Cs vapor steels and selected other materials, some projec- Fuel—LOa UN, with barriers tions can be made of the probable behavior of the Operating mode: refractory alloys. Evaluation by analogy with more Load following thoroughly studied systems will frequently be used Periodic restart in reviewing the irradiation effects data in these Reactor lifetime: alloys. 2 to 7 years Reactor neutronics: Flux, Fluence, REACTOR PARAMETERS n/(cm!-s) n/cm*

Radiation effects in metals are a strong func- Near term 2 X 1013 4 X 1021 22 tion of temperature and neutron fluence. This Advanced 2 X 10" 4 X 10 dependence requires that the projected range of reactor operating conditions be known to define the parameter ranges over vhich a material's response neutrons/(cm2-s) with lifetime fluences of 0.4 to is to be estimated. A general set of parameters 4 X 1022 neutrons/cm2. Out-of-core components estimated for a space reactor is given in Table 2. will be subjected to lower fluences. In most cases theao values are assumed and are not The reactor will not necessarily be operated provided by designers. Since the purpose of this continuously for its full lifetime. It may be listing is to establish conditions within which to operated for times .it a fraction of full power and examine radiation effects in refractory metal may be subject to periodic shutdown and restart. alloys, precision is not required. This results in some concern for low-cycle fatigue For purposes of this document the shutdown properties and for low temperature properties, as temperature is assumed to be 20°C The actual wfll as for operating temperature properties of the shutdown temperature will be set by the ability of candidate materials for reactor components. the system to dispose of the decay heat of the reac- Other assumed reactor conditions are listed in tor core, shut down after some period of operation, Table 2 for completeness. These parameters are of and may be far from 20°C. Based on the assump- less interest in evaluating the potential effects of tion that stainlass steels or n'ckel-based superal- irradiation on the properties of refractory metal loys world be used in a lower temperature system, alloys. the minimum operating temperature will he near 800°C for a refractory metal system. Maximum temperatures will be set by the particular alloy, by EFFECTS OF REACTOR the stress requirements on components, and by IRRADIATION ON THE compatibility with the rest of the system. The goal PROPERTIES OF METALS of near-term designs is near 1100°C, and more advanced reactors may attempt to use upper limits Neutron Interactions with a of 1400°C. If an in-core thermionic power conver- Crystalline Metal sion system is used, the goal temperatures for tungsten emitters may be 1300 to 1600°C or higher. The basic response of a metal to neutron irradi- The in-core flux for a fast spectrum space reac- ation is either the displacement of lattice atoms, tor will probably be in the range 0.2 to 2 X 1014 through the transfer of the kinetic energy of the WIFFEN neutron to the struck atom, or the capture of the Defect Clustering and Effects on neutron by the target atom and the transmutation Properties of thi.t atom either to a new isotope or a new atomic species. Characterizing the neutron fluence High fluence, high temperature radiation in terms of these effects provides a useful mea- effects that are of concern in evaluating the per sure of the fluence that can be used in comparing formance of reactor structural materials include results produced in reactors with quite different swelling, strengthening, loss of ductility, and neutron spectra. The accumulated experience in changes in phase stability. These phenomena and radiation effects experimentation has established other effects of irradiation are all temperature the validity of this measure of exposure and deter- dependent. Except for strengthening, these effects mined the limits on its usefulness. usually become important at temperatures above The measure of displacement events is the num- some lower limit at v.rhi>h the point defect • begin ber of displacements per atom, dpa, calculated for to have significant mobility (i.e., thermal energy is the flux and energy distribution of neutrons. This adequate for rapid vacancy migration through the measure, which incorporates the experimentally lattice), 30, once created, both vacancies and inter- determined average energy required to displace an stitials move rapidly. An upper temperature limit atom from its lattice sito, accounts* for all elastic for some effects occurs when thermal energy is and nonelastic interactions between the neutrons adequate to overcome binding energies of defects in and host atoms. defect clusters, and thus all vacancies and intersti- The transmutation reactions of greatest general tials migrate until they recombine and annihilate. importance are the gas-producing (n,p) and (n,a) For the range of temperatures where point reactions. These generally h«ve energy-dependent defects are mobile but can be bouina strongly cross sections and are of much greater importance enough to resist further motion, defect clusters can at neutron energies in the MeV range than at nucleate and grow. While most point defects are lower energies. In some particular cases the solid annihilated by defects of opposite sign (vacancies products of the transmutation reactions may also and interrtitials recombine), a few cluster with be of importance. Ono example of technical impor- 1 : their own kind to form dislocation loops, to add to tance in a refractory meta alloy s the transmuta- existing dislocations, or to form voids. tion of tungsten to rhenium and rhenium to osmium, which can lead to decalibration of W-Re For most reactor applications, the use tempera- thermocouples in service in a reactor.2 ture of structural alloys is within the window in which irradiation effects are of importance. Effi- These measures of neutron fluence. calculated 3 ciencies of heat transfer and of the power conver- by Gabriel et al., are given in Table 3 fo,- several sion system favor temperatures as high as possible. The same processes that set an upper temperature TABLE 3 limit on sever, l of the irradiation effects limit the Irradiation Response of Refractory Metals use of metals at higher temperatures. Most often in a Fast Reactor Neutron Spectrum this is a creep strength limit. At temperatures where the thermal energies allow rapid vacancy Response* per fluence unit of 1022 neutrons/cm2 release from defect clusters (such as voids), the At. ppm He At. ppm H creep rate is usually too high for structural use at Element dpa significant stress. V 5.43 0.059 2.07 Nb 3.29 0.169 1.18 Mo 3.43 0.170 0.329 Swelling by Cavity Formation Ta 1.26 -0 0.141 W 1.24 ~0 0.0036 Structural alloys can swell under neutron irradiation This volume increase results when the two •Response calculated for the neutron point defects created by irradiation go to different spectrum at the horizontal midplane of row ? sinks, with a net flow of self-interstitial atoms to of EBR-II. [From Gabriel, Bishop, and Wiffen.3] dislocations or dislocation loops and a net flux of vacancies to cavities or voids. This process may of the refractory metals of interest. The response involve helium from (n,a) reactions or possibly has been calculated for a fast neutron spectrum, impurity gas atoms in the metal. These insoluble using the core center spectrum of the EBR-II reac- gases may be of special significance in the nuclea- tor. A total neutron fluence of 1022 neutrons/cm2 tion stage of cavity formation by stabilizing three was used for the response numbers given. dimensional vacancy clusters. EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 256

fective in controlling swelling, but several data sets DOPPLER BROADENING "| _ have shown that alloying with a more reactive ele-

VOID VOLUME v-Cdpo ment may narrow the range of swelling tempera- 0 7 FROM TEM a lldpa tures, as for zirconium additions to niobium, or * I9dpo 06 may completely suppress swelling, as demonstrated 05 0.5 | for titanium additions to vanadium. o 0 4 > 0.4 j- a Short-Time Strength and Ductility 0.3 The strength properties of the refractory metals 02 0.2 are affected by irradiation at all temperatures up O.I O.I to near 0.5 Tm. While all defect clusters can produce strength changes, the most potent strengthen- 0 0.0 300 600 900 1200 1500 I80C ing results for dislocation loops, with small loops IRRAOIAT'ON TEMPERATURE (°C) providing the greatest strengthening increment. Fig. 1 The void volume -f the irradiation on ranging from 0.4 to 4 X 1022 neutrons/cm2 (>0.1 the mechanical properties is in the reduction of MeV). A maximum swelling of under 1% is shown ductility. Three important classes of ductility for this condition. Swelling values of several per- reduction produced by irradiation are illustrated in cent have been reported in bcc refractory metals, Fig. 2. These stress-strain curves from uniaxial although swelling values near 1% have been com- tensile tests show reductions in both uniform and monly observed for neutron exposures in the low total tensile elongation. If no other change results, 1022 n/cm2 range. The shape of the swelling- ductility and strength usually follow an inverse temperature relationship is representative of all relationship. The strength increase produced by refractory metals. The range of temperatures at irradiation-induced changes in the microsti ucture which swelling is observed is usually approxi- is accompanied by a loss of ductility (Fig. 2a). If the hardening results mainly from a population of mately 0.25 Tn, < T < 0.5 Tm for neutron flux levels typical of fission reactors. While narrower tem- small dislocation loops, the first deformation can perature ranges may be suggested by some data trigger a clearing of deformation paths through sets, this guideline is useful in projecting behavior this microstructure. The i ~sult is a plastic instabil- where experimental data are not available. The ity (Fig. 2b), with lead drop at the yield point and temperature ranges are defined for the unalloyed further deformation to failure at continually metals in Table 4. decreasing load. In some materials, especially Mo, W, and their TABLE 4 alloys, strengthening of the matrix can raise the Melting Point and Eey Temperatures for deformation stress above the cleavage stress of the Refractory Metals alloy, and brittle fracture occurs at very high load (Fig. 2c). This fracture mode is not preceded by 0.5 Tm, 0.25 Tm, measurable deformation. The fracture path usually °C °C °C includes both grain boundary separation and cleavage fracture through grains. A simplified V 1900 814 270 Nb 2468 1098 412 mechanistic understanding of the process is given Mo 2610 1169 448 in Fig. 3, which shows the role of flow, decohesion, Ta 29S. 1362 544 and cleavage stresses in determining the failure W 3410 1569 648 mode. The effect of irradiation is to raise the flow stress, with little or no effect on the two fracture Alloying of the bcc refractory metah may affect stress values. The result is to shift T2, the bound- the swelling, although a range of effects has been ary between ductile flow and brittle fracture, to observed. In some cases the alloying may be inef- higher temperatures. 256 WFFEN

~~IAI DUCTILITY LOSS DUE TO HARDENING~~

~~IRRADIATED~~

~~UNIRHADIATED~~

~~20 ELONGATION 1%)~~

~~800 R~~

~~'( '2 TEST TEMPERATURE —»•~~

Fig. 3 A schematic representation of the temperature dependence of flow stress, grain boundary decohesion stress, and cleavage stress on test temperature. Irradiation can shift the flow stress curve, changing the intercepts 20 40 60 with the other two curves and thus changing the tempera- E1ONGATION l%] ture boundaries of the indicated failure modes. The grain separation region is often seen for Mo and W but not often observed for V, Nb, and Ta alloys. 1 1 800 - (C) DBTT SHIFT observed, is usually decreased as the irradiation fIRHADIATED temperature is increased. 600 - Some effects of alloying on the change in mechanical properties have been reported The 400 - UNIRHADIATED existing data, too sparse to draw general conclusions, will be discussed below when the individual 200 - alloy systems are considered. Hfiium Embrittlement 1 0 20 40 60 A ductility loss mechanism that becomes impor- ELONGATION (%) tant at higher temperatures is helium embrittle- Fig. 2 Comparative stress-strain curves showing' three ment. This embrittlement results from helium pro- major classes of irradiation effects on tensile properties, duced in (n,a) reactions. At temperatures adequate (a) General hardening and an associated loss of ductility for helium mobility in t:ie host metal, usually >0.5 (niobium, irradiated and tested near 400°C). (b) Plastic Tm, the helium accumulates in grain boundaries. instability at the yield stress (Mo-0.5% Ti, irradiated and There is little or no effect on strength or flow tested near 400°C at a low strain rate), (c) Brittle fracture (Mo-0.5% Ti, irradiated and tested near 400°C, at a moder- properties because point defects are mobile and ate strain rate). annihilate, but the presence oi the helium in the boundary leads to low ductility failure by boundary separation. Because the cross section for the (n,a) The ductility loss modes discussed above have reaction is relatively low in the refractory elements been identified experimentally in a number of for fission spectrum neutrons, this ductility loss cases. The plastic instability mechanism, which has process is of little importance in the refractory been seen for alloys of V, Nb, Mo, and Ta, appears alloy systems. An extreme example of this behav- to be associated with irradiation temperatures near ior for Inconel 600 is illustrated in Fig. 4. or below the lower end of the void formation range. Most observations are for temperatures below 0.35 Creep, Irradiation Creep, and Tm. Shift of the DBTT to well above room temperature has been documented for irradiated molybde- Microstructural Stability num and tungsten alloys but has not been observed (Jreep rates of stressed components under irra- for alloys based on V, Nb, or Ta. The effect, when diation can be quite different from those predicted EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 257

interstitials. These atoms then move through the 500 HELIUM EMBRITTLEMENT lattice with the defect, species and concentrate at the sink where che defects have either clustered or 400 been annihilated. The possible results can include formation of phases not predicted from equilibrium UNIRRADIATED phase diagrams or the concentrations of impurity S= 300 atoms. The formation of precipitate phases (films) on grain boundaries or the concentration of surface

H 200 active species at the boundaries can result in embrittlement and susceptibility to grain boundary fracture. There is a possible synergism with helium 100 effects on grain boundaries, but conclusive evidence IRRADIATED is not available.

5 10 15 ELONGATION (%! DATA FIELD ON IRRADIATED Fig. 4 An extreme example oi loss of ductil- REFRACTORY METAL ALLOYS ity due to helium embrittlement. This example is for Inconel 600, irradiated after cold Few data are available for irradiation effects in work. Irradiation and 'tensile tests were at refractory metals, especially at conditions of inter- 700°C. est for space power reactors. However, the data by thermal creep data. Irradiation creep is usually that are available can be used to guide application of importance at temperatures below the range at of the existing knowledge of irradiation effects to which thermal creep is important, <0.5 Tm, and has predictions on refractory metals. This approach some of the characteristics of superplaatic behav- can also serve to identify critic?.! questions that ior. Figure 5, which depicts the results of creep must be answered in future experiments. measurements on an austenttic stainless steel Similarities in the irradiation response of the under light-ion bombardment,5 shows marked devi- refractory metals are greatest for elements in the ation froir thermal creep behavior at the lower temperatures. Three general characteristics of irradiation creep are a weak temperature dependence (shown in Fig. 5), a linear dependence of the creep rate on damage rate (flux), and a slow increase in rate with increasing fluence at a fixed flux. The dependence of the creep rate on composition and fj DAMAGE. »i'[ n .iO" Ipo s ' microstructural variables is not well established. The ductility to rupture in creep under irradiation has been demonstrated to be much greater than measured in postirradiation creep tests. There are very few data available on irradiation creep ir the refractory alloys. Available data on austenitir stainless steel cold i 10"4 worked before irradiation indicate that thtr- may be a small effect of irradiation on recovery and recrystallization. These results suggest that a depression of the recovery temperature may exist but that the effect does not exceed ~50°C. Precipitation and Phase Stability The question of the stability of alloy phases under irradiation can be of concern to the behavior 1O"6 of structural materials during reactor service. The II 12 13 14 15 concern arises from the demonstrated tendency of RECIPROCAL TEMPERATURE IK"'XI04I some solute or impurity species to pref- Fig. 5 Temperature dependence of irradiation creep in erentially couple to the flux of either vacancies or 60%-cold worked 321 stainless steel. (From Hudson et al.') 258 WFFEN same group of the periodic table, rather than for TEM examination showed that there had been no adjacent elements. Therefore, the material will be void formation. The TEM observations showed con- presented in the order V, Nb, Ta, Mo, and W. siderable development of a dislocation microstructure during the irradiation. There may also have Vanadium Alloys been some change in the precipitate structure. The V-20Ti composition has been the most Vanadium alloys we -e evaluated in both the studied of the vanadium alloy family. The resis- United States and West Germany as candidate tance to void formation and swelling is well es- cladding materials for LMFBRs and are currently tablished for a wide range of neutron fluences and under evaluation for use in fusion reactors. As a irradiation temperatures. However, Tanaka et al.9 result of these two interests there are more vana- showed that injection at room temperature of a dium alloy irradiation effects data at elevated tem- uniform distribution of helium to concentrations of peratures and moderate-to-high fluences than for 90 and 200 at. ppm resulted in void formation dur- other refractory metal alloys. Similarities in ing subsequent neutron irradiation to 3 X 1022 behavior of the several refractory metals make the neutrons/cm2 (>0.1 MeV) in EBR-II. Void forma- data useful in predicting the probable response of tion occurred for these conditions over the exam- the other systems. ined range of irradiation temperatures, 400 to The effects of irradiation on vanadium alloys, 700°C. In spite of this helium-promoted cavity for- with emphasis on the V-15Cr-5Ti alloy, were re- mation, the resulting swelling was negligibly small viewed in 1982 for a fusion reactor design project.6 of the order of 0.03%. A more complete review was published in 1978 by 7 More recent ion bombardment experiments to Gold et al. Few data have been generated since simulate neutron irradiation have confirmed and these reviews. extended knowledge of the swelling behavior of Experimental data have been reported for vanadium alloys. Work by Loomis and Ayrault10 swelling in unalloyed vanadium irradiated to fast 22 2 illustrated in Fig. 6 shows high damage level fluences up to about 3 X 10 neutrons/cm . Irra- results at 700°C. This shows that a V-Cr binary diation temperatures have covered the range from alloy is not swelling resistant—in fact swelling at about 385 to 700°C. Maximum swelling values have more than twice the rate of che unalloyed generally been 1.5 to 2% and occur, depending somewhai on purity, for irradiations at 550 to 600°C. The trends are the same as those observed ^ [ T ! i i i r r i I I | in other pure metals, with void concentrations and ION e.-MBflRDMENT AT 7OO°C fs.5%-20% _ sizes dependent on both fluence and irradiation A V,3 2-MeV 58Ni* A v,3.0-MeV 5'v + temperature. Voids are cubic in shape with cube / _• v-15Cf,3.5-MeV 51V perfection increasing with an increase in irradia- • V-! Ti, 3.0- MeV 51V + tion temperature. Void ordering on a superlattine, O V- which has been reported in other bec metals such as Nb, Ta, and Mo, has not been observed in vanadium. Several vanadium-base alloys h /e now been investigated to determine the swelling behavior under neutron irradiation. These studies, which have concentrated on the V-Ti- Cr system, have shown a general resistance to void formation and swelling in the alloys containing titanium. A V-10% Cr alloy, neutron irradiated at about 700 and 800°C, showed8 near 1% swelling for a fluence of 1.5 X 1022 neutrons/cm2. There was insignificant swelling at lower temperatures, although a low concentration of small voids formed. In a more complex alloy, VANSTAR-7, some void formation occurred, but swelling was I I I I I I I I I I I 20 30 40 60 70 very slight. In alloys containing titanium, DOSE (dpol irradiation at 470 to 780°C to fluences as high as 22 2 6 X 10 neutrons/cm produced little or no swell- Fig. 6 Swelling produced by heavy ion bombardment of ing detected by immersion density measurements.8 vanadium alloys at 700°C. (Frc Loomis and Ayrault.") EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 259

~~900 1 ' ' "~l 1 i 1 i r r I i VANf.lAR 7 800 /\o.95 dpa IRRADIATED IN HFIR AT 70°C TO DISPLACEMENT LEVEL SHOWN TENSI_fc TESTS AT 25°C 100 0.095 dpa :~~

~~600 j 0.0095 dpa \ 0.00095 dpa -\ CONTROL bOO - / A—^r~ Y 300~~

~~200 J - ,'IUNCORRECTED FOR 100 /MACHINE MODULUS)~~

~~1 1 i i 1 1 1 1 1 1 2 4 6 10 12 14 16 18 20 22 24 ELONGATION l%!~~

Fig. 7 Load-elongation curves for VANSTAR-7 irradiated at 70°C to neutron fluences up to 1.0 X 1025 neutrons/m*. Plastic instability is observed for fluences producing 0.0095 •'.pa or greater. metal—while two alloys containing titanium were of higher temperatures.12 The results, in Figs. 8 completely resistant to void formation. This and 9, show that the marked hardening produced confirms the swelling resistance of the titanium- by the irradiation and observed in low temperature containing alloys, established by nontron irradia- tests is gradually lost as the test temperature is tion, for the range of ion bombardment tempera- increased. At 750°C the properties of irradiated tures 400 to 720°C, damage levels to 50 dpa, and a range of He : dpa ratios. Confirmation by neutron irradiation to high fluences is still required. \C\J\J 1 The low level of cavity formation noted previ- NEUTRON iRRAOIATiON ously for the neutron-irradiated VANSTAR-7 alloy 1.4 KI D neu*ron/m^ specimens suggests that zirconium alloy additions •000 100°C IRRAOIATION "" •^-1 dpo may convey the same beneficial effects as titanium \ • additions, but verification is required. It is to be • V-2Cr-IOT. _ 800 hoped that analogs of the Ti (in V) swelling sup- a. | A v-5Cr-5Ti pression can be developed in other bcc refractory 5 • V-5Cr-10T, I alloy systems. t- N V-l5Cr-5Ti O Irradiation of vanadium alloys at temperatures 2 600 — a:u_ near room temperature results in an appreciable \- in increase in the yield and ultimate tensile strength O and a loss of ductility, as measured by elongation.11 400 The strengthening probably results from dislocation loop hardening, and the loss of ductility results from the channeling deformation mecha- 200 - nism discussed earlier. The rates of hardening and ductility loss are shown in Pig. 7 as a function of neutron fluence for the VANSTAK-7 alloy irradi- n 1 i i i ated at 70°C and tested at 25°C. A number of 200 400 600 800 1000 V-Cr-Ti ternary alloys have been irradiated to a TEST TEMPERATURE CO 21 2 constant fluence of 1.4 X 10 neutrons/cm (>0.1 Fig. 8 Effect of neutron-irradiation on the yield strength MeV) at 100°C and then tensile tested at a series of vanadium-base alloys. (From Abdou et al.') 260 WFFEN

~~40 1 I ' I NEUTRON IRRADIAT'ON • V 2 Cr -10 Ti 1.4xiO25 NEUTRON/M2 A V-5Cr-5Cr ~100°C IRRADIATION • V-5Cr-10Ti -30 (-1dpa) D V-I5U-5TI~~

~~• n~~

~~200 400 600 800 1000 TEST TEMPERATURE (°C) Fig. 9 Effect of neutron irradiation on the total elongation of vanadium-base alloys. (From Abdou et al.')~~

and unirradiated material were equal. Total elon- of unirradiated material throughout this tempera- gation values of the irradiated alloys were low in ture range. tests at room temperature to 400°C. Recovery The V-10 Cr, V-3 Ti-1 Si, and VANSTAR-7 occurred in tests above 400°C, with elongation results show mucn greater strengthening produced approximately equal to unirradiated material in by the irradiation, even for irradiation and test tests at 650 to 750°C (not shown in these figures). temperatures near 700°C. The ductility was also Even though significant hardening waa pro- rpluced by irradiation, with uniform elongation duced by these irradiations at 70 to 100°C, there values in the range 1.5 to 4%. There was, however, was no indication of cleavage or grain decohesion no indication of a brittle failure mode in any of fractures. All reported fractures were of the ductile these experiments. The few reports of creep- mode, and the ductile-to-brittle transition tempera- rupture testing on irradiated vanadium alloys indi- ture (DBTT) characteristic of bcc metals remained cate little or no effect on either rupture life or below room temperature for tensile tests. Impact elongation at fracture. or high strain rate tests, which could raise the Vanadium alloys are not immune to helium DBTT, have not been conducted on irradiated vana- embrittlcment. It has been amply demonstrated dium alloys. that helium implanted by accelerator bombard- Several experiment? <•;* have reported the results ment or by tritium decay can result in a loss of of irradiation of vanadium alloys at temperatures tensile elongation for test temperatures above 650 9 in the range 400 to 750°C to fluences exceeding to 750°C. Tanaka et al. have also examined the 1 X 1052 neutrons/cm2 (>0.! MeV). The alloys effect of 90 and 200 at. ppm He prt Injected before have included V-20 Ti, V-15 Ti-7.5 Cr, V-10 Cr, neutron irradiation of the V-20 Ti alloy. Material V-3 Ti-1 Si, and VANSTAR-7, with tensile tests at was irradiated at 400, 575, and 700°C and then ten- temperatures from room temperature to about sile tested at the irradiation temperatuie. These 750°C. These results are reviewed in detail by Gold results showed a factor of 2 loss of ductility in and Harrod.7 The trends indicated for these irradi- helium-free V-20 Vi irradiated and tested at 400°C ations are somewhat different from those observed but little effect for irradiation and test at 575 and for low fluence, low temperature irradiations. In 700°C. In the helium-preinjected material a severe the V-Ti or V-Ti-Cr alloy, modest hardening loss of ductility resulted for irradiation and tests results for irradiation in the lower part of this at 700°C. The helium had little effect at the two temperature range, but little strength increase lower irradiation temperatures of 400 and 575°C. results if the irradiation temperature is above Three vanadium-base VANSTAR alloys were —500°C. Ductility values are usually close to those irradiated at 410 to 450°C to fluences in the range EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 261

0.18 to 6.37 X 1021 neutrons/cm2 (>0.1 MeV). Low The magnitude of swelling reported by different cycle fatigue tests at 400° C showed the irradiation investigators, for closely comparable conditions, to have little or no effect on the fatigue life.13 For shows considerable scatter. While much of this these irradiation temperatures, the authors report may be due to differences in material, especially that radiation hardening saturates near a fluence impurity content, and in irradiation conditions, of 1 X 1021 neutrons/cm2. some may also be due to difficulty in examining In summary, the response of vanadium-base the material. Wiffen15-16 and Michel and Moteff17 alloys to low to moderate fluences of fast neutrons used TEM to examine material cut from the same looks encouraging. The effects of radiation are pri- neutron-irradiated coupons and calculated swelling marily to increase the short-time mechanical values that differed by more than an order of mag- strength. This appears to occur at little cost in nitude. The problem was probably due to inaccura- terms of compromising the ductility. The similarity cies in determining foil thickness and in measuring of vanadium-base alloys to Nb and Ta alloys justi- the diameter of very small voids. Comparison fies the hope for the development of radiation- measurements of cavity volume fraction in a single resistant alloys based on either Nb or Ta. microscope sample were made in a round-robin analysis program, where the same sample was cir- 23 Niobium Alloys culated to eight different sites. The result showed a spread of 70% between highest and lowest Niobium and niobium alloys have been irradi- swelling values determined in this sample. These ated and evaluated at conditions of interest to comparisons emphasize the importance of develop- space reactor applications with support from fis- ing adequate techniques for TEM determination of sion reactor, fusion reactor, and basic studies pro- swelling and suggest caution in the use of any sin- grams. However, none of this has been a focused, gle datum, or single data set. mission-oriented program, so, as with the other 20 refractory metals, data are sparse and uncon- Fischer has showr that void concentration is a nected. Much of the recent data comes from the strong function of flux. He reported an increase in fusion reactor program, where niobium alloys have void concentration by a factor of 4 to 7 for a factor been considered for first wall and blanket struc- of 4 ! acreise in flux, at constant fluence. The aver- tures and continue to be candidates for specialized age void size decreased with the same flux increase, resulting in reduced swelling at the components in the high neutron flux zone of the 24 25 reactor. There is some overlap between conditions higher flux. Loomis and Gerber " have shown a examined in these programs and the application in strong dependence on material purity, especially on space power system, although a large portion of oxygen content. Many of the apparent inconsisten- available data was generated for temperatures and cies that arise in comparing earlier work from dif- neutron fluences below the range of current inter- ferent sources are probably related to these unin- est. tentional variables. Relevant radiation effects data on niobium-base The only niobium alloy to be investigated to any alloys were reviewed by Pionke and Davis14 in 1978, extent is the Nb-1 Zr composition. The general for potential fusion reactor applications. Several finding has been that swelling occurs over a mvih researchers have also reported related data sirce narrower range of irradiation temperature in Nb- that time. 1 Zr than in Nb, with a peak swelling temperature near 800°C. One set of alloy samples has been car- The experimental data on neutron irradiated, 23 ried to very high fluences, 1 to 2 X 10 unalloyed niobium indicate that the temperature 2 26 neutrons/cm (>0.1 MeV), in the EBR-II reactor. range over which swelling will result extends from At this high fluence the immersion density indi- at least 400 to 1050°C. While these may not be the cated swelling in the temperature range 400 to actual extremes of the swelling range, data are not 600°C to be between 0.18 and 0.71%. The swelling available outside this range. The scant data avail- dependence on irradiation temperature was slight able from neutron irradiation experiments indicate in this data set. Most other swelling data on this that the temperature of peak swelling is near 15 16 17 18 alloy resulted from TEM examination. The swelling 600°C [Wiffen; - Michel and Moteff; Adda; 19 20 21 data are summarized in Fig. 10. E!en et al.; Fischer ]. Bartlett et al. did not find a peak swelling for irradiation at 450, 550, and 21 22 Bartlett et al. reported on the examination of 650°C. Jang and Moteff found a slight local maxi- Nb-5 Zr and Nb-10 Zr, irradiated to a fluence of mum at 600°C but found the greatest swelling at 22 2 22 2 3.6 X 10 neutrons/cm at 450, 550, and 600°C. 1050°C for a fluenca of 5 X 10 neutrons/cm They found no significant amount of swelling in (>0.1 MeV). any condition, although a few voids were observed 262 WFFEN

3.0 - - • POWELL et.dl D JANG a MOTEFF 5.0 1Mb — 1 Zr A SPRAGUE et.Ql. NEUTRON 2.0 .- O WIFFEN A IRRADIATION O WIFFEN / \ (FLUENCE SHOWN 22 V MICHEL 8 SMITH / / o \V IN UNITS 1O /25 ) n/cm2) g 1.0 \ in 17.6 / / \ 5.0 •^S^ 5.0 0 •-O—oJ-Urj- J rvJ gr D 1 X 1 o_J 400 1l9 500 600 700 800 900 100U IRRADIATION TEMPERATURE (°C)

Fig. 10 Comparison of the available neutron irradiation swelling data on Nb-1 Zr. The numbers alongside tne data points indicate the neutron fluence in units of lO^n/cm2 (E > 0.1 MeV). (From Powell et al.s) for the irradiations at 550 and 600°C. There do not rapid increase in the yield and ultimate tensile appear to be any results of irradiation of the strength. This hardening is produced by dislocation higher strength niobium alloys. loops formed during the irradiation, and the loops Ion bombardment simulation of neutron irradi- can be swept up by dislocation channeling, as dis- ation of niobium and niobium alloys hss been used cussed earlier. The result is easier deformation to examine the effects of several irradiation and after yielding, or plastic instability, and near-zero material variable? on the swelling behav.jr. uniform elongation. The behavior is illustrated in Loomis et al.24"25 have been the most active in this Fig. 2b. Total elongation in niobium deforming in area. Their results have shown swelling over the this mode remains high, as does reduction in area, temperature range 600 to 1150°C, with the upward and the fracture mode remains ductile. Yield shift in temperature, relative t^ neutron -irradia- strengths at room temperature have been raised by tion, produced by the higher damage rate in the irradiation to at 1< &st four times unirradiated ion bombardment. They showed that the void for- values without resulting in brittle (cleavage) frac- mation behavior was particularly dependent on ture. This implies that the DBTT, at least as ma< jiial purity. The presence of free oxygen pro- determined in slow strain rate tensile tests of moted void nucleation (higher cavity concentra- smooth samples, remains below room temperature tions) but often resulted in reduced void volume even after irradiation to moderately hi^h fluences. (swelling). The effect of solute additions seemed to The difference in the room temperature tensile be related to the oxygen activity in the alloy. The properties16127 for irradiation at room temperature swelling was suppressed by additions of Mo, Hf, V, and at 460°C is shown in Fig. 11. (Although the Zr, Ti and Ta, in decreasing onW of effectiveness, fluences are different for these two samples, that but was not affected by Ni or Fe alloying 25 is not considered to have a major effect on the additions. Results of dual beam bombardment, properties.) At the h'gher irradiation temperature, where helium is introduced along with the dis- the uniform elongation is near 10%, as opposed to placement damage, suggest that helium can also be the near-zero value for the sample irradiated at effective in promoting cavity nucleation in 55°C. The difference results because voids formed niobium. The effect of helium is thus a maximum in the higher temperature irradiation; they are not in the absence of oxygen or when no other cavity removed by dislocation motion during deformation. stabilizing species is available. Some work hardening is possible during testing dome data are available on the tensile proper- and nonzero uniform elongation results. ties of irradiated niobium and Nb-1 Zr, but there Void formation is not necessarily a liability in is little coverage of other properties or other terms of the mechanical property response. Tensile alloys. Most of the available data are for irradia- properties for niobium and Nb-1 Zr irradiated to tion temperatures below those of interest for the nearly identical conditions are shown in Fig. 12, space power systems. for tests at 25, 400, and 650°C. While the two Irradiation of niobium and Nb-1 Zr at reactor materials show appreciable hardening, and both ambient temperatures, 50 to 150°C, results in a retain good total elongation, there is a significant EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 263

STRESS difference in the uniform elongation. In the Nb- 3 no Psi) 1 Zr, where voids did not form during irradiation, NIOBIUM channel deformation resulted at all three test tem- 25°C TENSILE peratures and the uniform elongation was less 1 26 2 0.02 min- than 0.2%. Samples of the alloy irradiated at 394°C 3.0 X 10 /?/m to 1.9 X 1022 neutrons/cm2 yielded test results 80 AT 460°C essentially equivalent to the data shown in Pig. 12. \ A previously unpublished set of data28 on Nb- 1,1 Zr is given in Fig. 13. Samples of Nb-1 Zr, doped with 130 at. ppm 10B were irradiated in the HFIR 40 Reactor at temperatures from 450 to 950°C. Expo- 7.5 X 1C124 nlm2 sure for apnroximately 18 months produced neu- 22 AT 55°C tron fluences that ranged from 3 to 6 X 10 nteutrons/cm2, >0.1 MeV, depending on the sample position in the capsule. Except as noted on the fig- 8 12 16 ure, the specimens were tensile tested at the irra- ELONGATION (%) diation temperature. The results show significant strengthening produced by irradiation at 450°O, Fig. 11 The effect of different uiic. ostructurea on the but this effect decreases with increasing irradia- room-temperature tensile properties of irradiated niobium. tion and test temperature. At 800°C the ultimate The irradiation at 55°C produced a micro&tructure dom- inateri by small dislocation loops, and dislocation channel- tensile stress is only a small amount higher than ing during deformation produced a plastic instability and in the unirradiated material, while the yield zero uniform elongation. The sample irrcdiatsd at 460°C strength is still increased appreciably. Perhaps the contained a very high concentration of small voids which most important result is that the uniform and were not subject to removal by deformation dislocation. total elongations stay well above zero throughout The result was some work hardening in the sample ecu* taining voids and nearly 10% uniform elongation. this temperaiure range. In fact, these data are in

~~120 800 NIOB'UM 1OO c o CONIriOL > o CONTROL 22 * • 3.0M0 n/cm2 • 3.7MO22 n/cm2- 600 60 AT 460°C AT 450°C o o o~~

~~200 20~~

~~Z 40 O it i 2p UJ~~

~~200 400 600 0 200 400 600 TEST TEMPERATURE (°C)~~

~~Fig. 12 Tensile properties of niobium and Nb-1 Zr tested as-annealed or after irradiation to the conditions shown. Tensile tests were conducted at a Eitrain rate of 0.02 min~'. 264 WIFFEN~~

~~(IRRADIATED) Nb-IXZr 400 I 0.2% YIELD STRENGTH (~13Ooppm IOB)~~

~~300 "3 s~~

~~200~~

~~0.2% YIELD STRENGTH 100 10~~

~~J 1_ 30 —' 1 r "- ^CONTROL)~~

~~AATOTAL ELONGATION VTUNIFORM ELONGATION~~

~~200 400 6OO 800 TEST TEMPERATURE (*C)~~

Fig. 13 Tensile properties of Nb-1% Zr irradiated in HFIR to a fast neutron fluence of 3 to 6 X 10" neutrons/cm2, >0.i MeV, which produced 14 to 28 dpa. Samples contain —133 at. ppm He from 10B burnup. This is produced r t temperature during the first ~3 to 10 days of irradiation. Irradiation temperature same as test temperature, except for one test noted at 800°C and for the 25°C data which was for a sample irradiated at 550° C. conflict with the data presented for Nb-1 Zr in in niobium, as in all the refractory metals, is quite Fig. 12, in that irradiation at 450°C has not low (see Table 3), so the concern is only for helium reduced the uniform elongation to zero, but about contents below ~2 at. ppm. For the Nb-1 Zr, 1% remains. Whether this result is due to the Fig. 13 shows that helium embrittlement is not a helium produced by burnup of the 10B early in the concern at 800°C. Other results29"30 showed no radiation exposure, or whether some other explana- effect of accelerator-injected helium in tests at tion prevails is not clear at this time. (If the 1000, 1020, and 1200°C, unless helium contents helium is responsible, it may be because bubbles or exceeded it least 200 at. ppm. voids promoted by the insoluble gas prevent chan- There do not appear to be any data on phase neling.) Electror microscopy of the sample set may stability or on 'rradiation creep in niobium-base be completed in the future, to search for hints in alloys. this apparent contradiction in the data. The good ductility retention at 800°C, for sam- Tantalum Alloys ples irradiated at 800 and 950°C, indicates promise Most of the irradiation effects data available on for higher temperature applications of this! alloy. tantalum and tantalum-base alloys resulted from However, little hardening can be expected at these the consideration of using these materials as con- Vrmperatures, and designers will have to work trol elements in fast breeder reactors. Examina- v.-.hin the limits of the strength level of fully tion of irradiated material wa3 limited by the high i.r.r.*2.lwl material. activity levels generated by the irradiation. The A .-. i.oy system is the embrittling effects of the temperatures near 600°C. .".>.. i.r.vint& of helium produced by (n,a) trans- Voids form and swelling results fc- unalloyed iiiv^..,.v-. r

TABLE 5 TANTALUM Swelling of Tantalum and T-lll Produced by FLUENCE Irradiation in EBR-1I 2.5*1O22 n/cr (>0.1 MeV) Swelling,* % Volume Density Fluence, Irradiation increase. decrease. >0.1 MeV temperature, AL Ap Material X UP n/cm' °C L P

~~Ta 1 '. 390 0.37 0.76 Tin 1.9 415 -0.11 -0.024 Ta 1.9 640 0.27 0.36 T-lll 1.9 640 -0.36 -0.34~~

'Uncertainty in volume increase, determined from length measurements, 3-z—, is .rO.08%. Scatter in density decrease is ±0.25%. 600 800 1000 IRRADIATION TEMPERATURE (°C) Ta-10 W specimens is given in Table 6, which Fig. 14 Swelling data and empirically drawn swelling shows that voids weio formed in the alloy, as in curve for neutron-irradiated tantalum as a function of the unalloyed metal, but that swei^ng was proba- temperature of irradiation. bly less in the alloy.35 The available data on the effects of neutron Figure 14 shows an empirically drawn swelling irradiation on the mechanical properties of curve31 based on transmission electron microscopy tantalum alloys are scarce. Irradiation of Ta-10 W at 70°C, t . falum at 70 or 400°C, and T-lll at (TEM) examination of specimens irradiated in 16 EBR-II. A somewhat similar definition of the 400°C res. '^u in increases in room tempera- swelling-irradiation temperature relationship was ture strengtn prope ties, and reduction of uniform produced by Bates and Pitner,32 based on density tensile elongation co 0.1 to 0.2%. Total elongation change results. Their analysis predicts a somewhat values remained higher—i. the range 8 to 10% for broader temperature range for swelling than is the material irradiated at 400°C. The low uniform shown in Fig. '4. Difficulties in determining exact elongation results from the dislocation channeling irradiation temperatures may contribute to this and plastic instability discussed earlier. Similar discrepancy. Murgatroyd et al.33 confirmed the strength increases occurred in Ta and Till irradiated at 640°C, but elongation values were higher void formation during irradiation at 500°C by TEM 16 examination. Their linear dimension measurements for this irradiation temperature 'iecause the of swelling are in approximate agreement with microstructure produced by the higher temperature Fig. 14, but chey found a reduction in irradiations is dominated by voids, not by disloca- swelling in going to higher fluences. They attrib- tion loops, and dislocation channeling does not uted the recovery of swelling as possibly resulting occur. The full data set for the 400 and 640°C irra- from shrinkage due to the transmutation of Ta to diation of Ta and T-lll is given in Fig. 15, which W, with a resulting reduction in lattice shows that tensile tests to 650°C show a decreasing parameter.33 (A similar effect has been examined effect of the irradiation for the samples irradiated in Mo and TZM irradiated at 450°C and will be at 400°C. Significant effects of the irradiation can discussed in the section, Molybdenum Alloys. The still be seen. Samples irradiated at G40°C showed responsible mechanism is tentatively identified as no effect of test temperature between 400 and impurity segregation to void surfaces, leading to 650°C. void shrinkapj.34) There are no mechanical properties data avail- Swelling in tantalum alloys has received little able for material irradiated and tested at tempera- attention. Geometry and density measurements on tures above 650°C. No data exist for irradiation T-lll irradiated nt 450 and 600°C indicated no creep or for phase stability of tantalum alloys swelling'6 but perhaps some densification during under irradiation. irradiation (see Table 5). These specimens were not examined by TEM to establish whether or not Molybdenum Alloys voids had been produced by the irradiation. Quali- Almost every new, high-temperature technology tative comparison by TEM of a set of irradiated starts by specifying a molybdenum alloy as the ref- 266 W1FFEN

~~TABLE 6 Comparison of Main Microstructural Features Observed in TEM of Irradiated Tantalum and Ta-10% W~~

~~Irradiation Fluence, Microstructure of temperature, >0.1 MeV Microstructure Ta-10% W, °C XKPn/cm2 of tantalum compared to tantalum~~

Small black spots (loops), Similar concentration of about 2 X 10"cm"3, diam small loops (black spots) 20 to 100 A No voids identified No voids identified 425 2.5 Few dislocation segments Few dislocation segments Similar channeling Cleared channels resulted from deformation in handling Swelling of 2.4% calculated Fewer voids, but void from 1.9 X 10" voids/cm3, parameters cannot yet average diameter 61 A be determined

585 2.5 Voids ordered on bcc Much higher contents of superlattice, parameter dislocation segments, loops. 205 A May also be a component of very small loops Random dislocation (black spots) segments Swelling of 0.65% calculated Probably fewer voids from 6.1 X 1016 voids/cm3, average diameter 117 A More dislocations, in networks 790 2.5 Void distribution random

~~Random dislocation segments, ..iany pinned by voids~~

Swelling >0.03%, from Fewer voids 3 X 10" voids/cm3, 950 to 4.4 average diameter 130 A Dislocations in networks 1050 Dislocations in networks Higher dislocation concentrations erence structural material. As a result, limited The most thorough review of the effects of irra- irradiation effects programs have been conducted diation on molybdenum and its alloys was done for to test the feasibility of using molybdenum or a the UWMAK-IH fusion reactor study. Thl3 resulted molybdenum alloy ai, fast reactor cladding, space in the 1976 publication of an annotated bibliogra- reactor core components, and fusion reactor first phy to the literature and a discussion and analysis wall and blanket structures. Unfortunately, in each of the reported radiation effects data.37"38 Some case the molybdenum candidate has not survived 300 sources of data were i lentified, but fewer than beyond initial scoping studies, and irradiation pro- 20 covered the effects of neutron irradiation to flu- 22 2 grams have been abandoned. As a result, various ences above 1 X 10 neutrons/cm and tempera- data sets have little overlap in material purity, tures above 600°C. Important data have also been material condition, irradiation condition, test published in the years since this review. methods, or program goals. Assembling a clear and The data field identified in the 1976 review for consistent picture of the irradiation response of microstructural observations and swelling mea- molybdenum and molybdenum alloys from this surements is shown in Fig. 16. Although the data base is a difficult, or perhaps impossible, task. coverage shown is large, only the region with lower EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 267

240 1 1 • 1 - i TANTALUM T-111 — 1600 i i i i o a CONTROL o c CONTROL 200 - • » 1.5»IO2? n/cm2 AT 390°C • * \3M^2i n/cm2 | • » 1.9x1022 n/cm2 AT 640°C AT415°C i--».JJLTIMATE ?2 • » l.9«io n/c m*- YIELD —•.— 4200 "5 '60 ULTIM/ a AT 640°C O Q. o i YIELD 5 § 120 4 • YIELD AND ULTIMATE — in ULTIMATE: IT IE " 1 i- .~^—. •—. YIE'.D"^ I T YIELD v— 400 40 -ULTIMATE .. _4-.- __YIELD h 0 I I * 60 °— TAL 40 __ _. -sr TOTAL a— UNIFORM 1 I • i UNIFORM —-4— 20 ! TOTAL -CT*™-». -4^, TC TAL I •—. UNIFORM -i UNIFC — 1 -y r—^= 200 400 600 0 200 400 600 TEST TEMPERATURE (°C)

Fig. 15 Tensile properties of tantalum and T-lll (Ta-8% V.'-2% Hf). Results for innealed samples and for two irradiation temperatures are shown for each material. Uniform elongations wore less than 0.3% for the lowest irradiation temperature for both materials and these values are not shewn.

bounds 600°C and 1 X 1022 neutrons/cm2 is of at 130U°C and <0.05% at 1500°C. This suggests an interest to space reactors, and few data fall into upper temperature limit on swelling near 1500°C; this zone. however, the tempera' ures quoted are experiment An attempt was made to develop a single design temperatures, for gas-gap controlled tem- sweUing-temperature relationship for unalloyed peratures, and uncertainties are large. molybdenum from ihe available swelling data.37 To Evans34 has reported a decrease in swelling do this, swelling was assumed to be proportional to with inc-easing fluence for Mo and TZM irradiated fluencf to the 0.6 pow«ir, a juggesHon derived from at 450°C. In both materials swelling decreased ion bombardment data. The result of this analysis markedly in the fluence interval from 3.5 X 1022 is shown in Fig. 17, with a spread in swelling of to 8.0 X 1022 neutrons/ cm2 (>0.1 MeV). The segre- greater than one order of magnitude. It should also gation of an unidentified impurity species to void be noted that the higher swelling bound is defined surfaces was tentatively suggested as the control- by upward extrapolation of low-fluence data, which ling mechanism. In any case, the result was not suggests that the simple fluence dependence observed for higher temperature irradiations on assumed in tho analysis is not correct. the same two materials. More recent data on the swelling produced by These more recent results do not change fhe neutron irradiation of molybdenum falls within the general conclusions on swelling of molybdenum. bounds shown in Fig. 17. The highest fluence data Swelling is expected for neutron irradiation at on unalloyed molybdenum is that reported by temperatures between a lower limit that is 300 to Sprague et al.39 for irradiation at 650°C. Fluences 400°C and to continue to temperatures above at of 5.4 and 8.4 X 1022 neutrons/cm2 produced least 1300°C as shown in Fig. 1. The swelling- swelling of 3.0 and 2.3%, respectively. These values temperature relationship has a broad maximum confirm that the upper limit on swelling shown in that extends from about 600 to 900°C. Maximum Fig. 17 is too great. Bentley and Wiffen40 produced values of swelling below 4% are expected for flu- lower fluence data that showed swelling of 0.27% ences to at least 1 X 1023 neutrons/cm2. 268 WIFFEN

~~10" t»20 O E > I MeV TlTiOAL CTR . • E > 0 I MaV CONDITIONS _ 20 4 |° 36 O 2 -2| O 27 O 35 35 17 17 17 17 17 17 O 0 O O o 10" 38 3°e 38 TO 38 38~~

~~LLI o z 2 UJ 21 2 • O * 2 21 io 31 O z o O 3*7 3 e • n 37 37 28 37 37 37 Z O 23 © 14 8 2 -,20 O • • O • •• O 3722 2 37 37 37 37 37 2 o 30 e ! e o o o~~

IO1!. -273 400 800 1200 T (IRRADIATION) FiR. i6 Map of the fluence-irradiation-temperature coverage of data available in 1976 on the microstructure and swelling of molybdenum and its alloys. Identifying numbers refer to tiie UWMAK-HI report.37

Swelling in the alloys Mo-0.5% Ti and TZM has 650°C, and high temperature-high fluer.se swelling been reported by several experimenters. The gen- data are still needed. eral result is that the temperature dependence of Stubbins and Moteff43 have reported swelling swelling for these alloys is nearly identical to that results determined by ion bombardment of of unalloyed molybdenum. However, where direct Mo-0.5% Ti and TZM, for the temperature range comparisons have been made, the swelling in the 750 to 1450°C, probably equivalent to the approxi- alloys is equal to or greater than that in molybde- mate temperature range 450 to 1150°C for neutron num. At 1000°C Bentley and Wiffen41 found irradiation. Their results confirmed the modest swelling near 1.0% in the two alloys, compared to swelling values anu the weak temperature depen- 0.6% in molybdenum. Sprague et al.39 reported d ,:e of swelling in this temperature range. 4.0% swelling in TZM irradiated to 5.4. X 1022 2 Relatively little attention has been given to the neutrons/cm at 650°C, a condition that produced Mo-Re alloys. Wiffen16 reported 0.44% swelling, only 3.0% swelling in molybdenum. measured by immersion density, in Mo-50% Re Gelles et ai.42 and Powell et al.26 have reported irradiated to 5.3 X 1022 neutrons/cm2 (>0.1 MeV) on the temperature and fluence dependence of at a temperature that increased from 855 to 1056°C swelling in TZM, for fluences up to almost during the irradiation. It is likely that some pre- 2 X 1023 neutrons/cm2. Their resu'ts are shown in cipitation had occurred during the irradiation, but Fig. 18, and again a maximum spelling of le&s there was no TEM to determine whether precipita- than 4% resulted. Unfortunately, the range of tion and/or void formation had resulted during temperatures in this experiment extended only to this irradiation. Igata et al.44 have used irradiation EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 269

upward by approximately 100°C and that the peak z r 1 1 swelling for equivalent fluences was reduced by at MOLYBDEN.. M least one order of magnitude. Verification of this SWELLING advantageous effect of rhenium additions using neutron irradiation is needed. A map from the 1976 review37 showing the irradiation conditions for samples on which mechanical properties measurements have been made is given in Fig. 19. As with swelling, this shows sparse coverage in the high fluence-high temperature region. The problem is oven greater than shown, since samples were often tested at temperatures well z _J below the irradiation temperature. _l LLJ 5 Essentially ali available data on irradiated molybdenum and molybde* um alloys show that a 0.1 principal effect of neutron irradiation is to raise the DBTT. This has been demonstrated at least for Mo, Mo-0.5% Ti, and TZM, for irradiation temperatures ranging from room temperature to 45 o FAST RFACTORS 1000°C. A set of results for these three alloys, irradiated at four different temperatures to a flu- A THERMAL RCAC TOR 22 2 D BENTLL'Y (THERMAL REACTORS' ence of 2.5 X 10 neutrons/cm (>0.1 MeV), is shown in Fig. 20. These results were produced by slow three-point bend tests of coupons 2.5 X 3 '^00 40C 600 BOO 1000 1200 1400 mm and define the DBTT as the temperature IRRADIATION TEMPERATURE (°C) boundary between cracking and bending. Tests on unirradiated material showed that, by this mea- Fig. 17 Swelling data for unalloyed molybdenum irradiated under a range of conditions. Data are plotted for a sure, all three alloys were ductile at room tempera- fluence of 2.5 X 10" neutrons/cm' (>0.1 MeV) after nor- ture. However, the DBTT of irradiated material muiazatioD assuming the sv oiling depends on (the O.S power was as high as 650° C for one condition and above of fluence. (From Badger et al.37) room temperature for all irradiation conditions. For temperatures of 585°C and higher, the DBTT decreased as the irradiation temperature increased. Other test results by Wiffen,16 Webster et al.,46 47 TZM NEUTRON IRRADI Steicken, and others confirm the DBTT increase, • 17.8 usually as determined in tensile tests of smooth specimens. Where yield strength values have been measured in these DBTT determinations, the values in irradiated materials are in the range of two to three times the values for unirradiated material at the same test conditions. 5 Tensile tests at test temperatures above the DBTT usually show very low values of uniform elongation in irradiated material. Total elongations in irradiated material are commonly in the range 2 400 450 500 550 600 650 to 10%, for test temperatures above the DBTT. IRRAOWT'ON TEMPERATURE <°C) Data generated in the USSR on TZM48 irradi- 21 51 Fig. 18 The swelling response of TZM. The peak fluences ated to a fluence of 1.2 to 1.6 X 10 neutrons/cm are in licated in the legend. (From Geiles et al.") at 550 and 950°C are shown in Figs. 21 and 22. These show properties are strongly affected by irradiation at the lower temperature, for tests with electrons in a high voltage electron micro- between 20 and 800°C, but little affected by the scope to investigate swelling processes in Mo ai J 950°C irradiation. The results also suggest that for Mo-Re alloys. They found that in a Mo-2.6 at.% Re material irradiated at 550°C, the DBTT may be alloy the peak swelling temperature was shifted near room temperature and, for material irradi- 270 WFFEN

~~X CREEP a° TYPICAL CTR • FATIGUE '* OO CONDITIONS~~

~~10" 46O~~

~~.32 2 <>43 o |0 V4S LU 3 X4S~~

~~817,44 917.21 44 g IOZO ce i- LJ z 30~~

~~10"~~

~~J_ _L _l_ 0^7 200 400 600 600 1000 1200 1400 1600 T (IRRADIATION) °C~~

Fig. 19 Map of the fluence-irradiation-temperatiure coverage of mechanical property data available for molybdenum and its alloys as of 1976. Iden.ifying numbers refer to tbe UWMAK-HI report.3* ated at 950°C, the DBTT remains below room tem- may be due to irradiation temperature or to the perature. higher neutron fluence in Wiffen's experiment. The Mo-50 Re alloy irradiated at 393°C and at The fatigue properties of TZM at 427°C have a temperature that increased from 855 to 1066°C been investigated by Smith and Michel.51 They during irradiation has been tensile tested at tem- showed that irradiation to 5 X 1022 neutrons/cm2 peratures of 400, 650, and 800°C. Brittle failures at at 427°C increased the fatigue life by approxi- very high stresses resulted for all tests.16 mately a factor of 2 and reduced crack growth Postirradiation creep tests have been conducted rates. Furthermore, samples which were injected on few samples, ^fiffen49 tested two irradiated with an additional 10 at. ppm He injection before specimens in creep at 750 and 1200°C after irradia- the irradiation showed a further increase in tion to a fluence of 6.1 X 1022 neutrons/cro2 fatigue life. Quantitative results show that for (E > 0.1 MeV) at a temperature which increased 1 X 1022 neutrons/cm2 at 427°C (without helium from 860 to 1140°C during irradiation. Compared injection) the fatigue life was increased by 60 to to control specimens the creep life was lengthened 100% in the strain per cycle range of 0.12 to 0.16%. considerably with only a small decrease in ductil- Molybdenum and TZM are probably the only 50 ity. Moteff and co-workers irradiated specimens refractory alloys in which the irradiation creep at 70, 700, and 1000°C to a fast neutron fbaence of rate has been determined. Mosedale et al.62 have 20 2 1.4 X 10 neutrons/cm (E > 1 MeV), after which measured the creep of stressed helices in the the specimens were creep-rupture tested at 750°C. Dounreay Fast Reactor and show that the creep They observed increased creep life, particularly for rate, given as strain per unit stress as a function of the higher irradiation temperatures, although not fluence, overlaps the upper bound of data for the to the extent observed by Wiffen. The difference same measurement of stainless steels. EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 271

~~900 SUMMARY OF BCND TESTS IRRADIATED Mo ALLOYS NO SEM SEM BEHAVIOR O • FU.L DUCTILE BEND a • CRACK. DUCTILE .MTIATION A FULLY BRITTLE CRACK —LJ— 0 • COMPLETE BREAK ) -o- -m-~~

~~_ 600~~

~~o o * A 1~~

~~tn 300 o A o -o- I-o- § I f • ™ * 0 Mo 0.5 Ti TZM Mo 0.5 Ti T7M Mo 0.5 Ti TZM Mo 05 Ti TZM~~

~~585 790 IRRADIATION TEMPERATURE ( °C )~~

~~Fig. 20 Bead ductility determinations on 3 mm X 2.5 mm tabs cut from sheet irradiated to 2.5 X 10" neutrons/cm2 (>0.1 MeV) in EBR-II. The shaded zone defines the DBTT. (From Hef. 45.)~~

~~1200 ALLOY~~

~~_ 1000 o a. 2 1 oi 550°C I 800~~

~~95u°c £ 600 ^ -' \ -~~

~~a •^-^ f "00 • —• •CONTRO ; o 200~~

~~] 1 1 .~~

~~0 100 200 300 400 ")0C 600 ?0O 800 100 200 300 40'J 500 oOO 700 800 TEST TEMPERATURE CO TEST TEMPERATURE |°c)~~

Fig. 21 Effect of irradiation and test temperature on 0.2% Fig. 22 Effect of irradiation and test temperature on total yield strength of TZM. (See Ref. 48.) elongation of TZM. (See Ref. 48.) 272 WIFFEN

Data on phase stability during the irradiation 2.0 r of molybdenum are generally not available beyond T • MATOI.ICH, MAHM the qualitative observation that very complex AMD MCTEFF precipitation is observed in TZM. * WIFFEN Tungsten Alloys. Tungsten alloy radiation effects at reactov relevant conditions have received the least attention of the five refractory alloy system*. Applications that have been the focus of these few studies include in-core fuel center-line thermocouples, thermionic diode emitters, and protective components in fast reactors. Because all of these potential applications were of interest for only a short 600 900 1200 1500 time, probably less than half of the material irra- IRRADIATION TEMPERATURE (°C) diated under the programs was examined, and few data were reported. Fig. 24 Swelling in irradiated tungsten, measured by Elec'.ron microscopy examination of neutron- immersion density. The results for tungsten from Fig. 23 are replotted with unpublished result of Wiffen" for mate- irradiated tungsten showed that swelling extended rial irradiated to 4 to 6 X 10" neutrons/cm1 <>0.1 MeV). at least between the limits of 450 to 1500°C, and the swelling value reported at 1300°C suggested that fhis was not nea- the upper temperature limit. The highest fluence data and the most complete temperature coverage were reported by peratures. Swelling at higher temperatures could Matolich et al.,53 and their results are shown in become irnpo?' int in tungsten emitters for in-core thermionic diodes. However, an unpublished set of Fig. 23. 21 2 The shape of the swelling curve given in data a1 lower fluence [4 to 6 X 10 neutrons/cm Fig. 23, as plotted by Matolich ot al., suggests the (>0.1 MeV)], but with overlapping irradiation tem- possibility of a second swelling peak at higher iem- peratures, does not indicate a second swelling peak.54 Immersion density measurements of swelling from this data set are plotted with the Matolich et al.53 data in Fig. 24. The smooth curve 2.5 drawn approximately fits the data, but a curve I with a shoulder on the high-temperature side of EBR -El IRRADIATION the peak could be used with this combined data set. c 5 x 4O22 Geometry measurements and qualitative TEM O TUNGSTEN examination confirmed the general slope of the nw-25 Re swelling curve, although the magnitude of swelling is still not established.54 The immersion density results plotted in Fig. 24 for the Wiffen data set are believed to be an upper limit on the actual magnitude of swelling. Other isolated data points showing void formation18-55 or other measures of swelling56 gen- 0.5 erally confirm the sv.elling of tungsten presented above. Tungsten-rhenium alloys have also been examined after a range of irradiation conditions. 53 300 GOO 900 1200 Matolich et al. found little or no swelling in the 2 IRRADIATION TEMPERATURE (°CI W-25% Re alloy (Fig. 23). Williams et al. examined a range of W-Re alloys, with 5 to 25% Re, Fig. 23 Observed swelling of irradiated tungsten and irradiated to fluences of 4 to 6 X 102' tungsten-25 rhenium, based on immersion density measure- neutrons/cm^ (>0.1 MeV) at tempera' •: 6 from 600 ments. The neutron fluence for EBU-II irradiation was 5.5 X 1022 neutrons/cm2 (>0.1 MeV). (From Matolich to 1500°C. There was no void formaiion for any of et aI.M) these conditions. EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 273

200 The microscopy examination of W-Re alloys did show the precipitation of a rhenium-rich phase not predicted by available phase diagrams,2 This pre- 160 cipitation .occurred even in the 5% Re alloy, which is far from any previously reported two-phase UJ a: 120 regions of the phase diagram. i- cn Tensile res alts on stress-relieved tungsten irra- o 21 ? 90 - diated at 371 or 382°C to fluences of 5 to 9 X 10 neutrons/cm2 have been reported for a range of test temperatures.'17 These results are shown in Fig. 25. Postirradiation tensile tests, conducted in the range 20 to 927°C. indicated thai the tensile yield strength was approximately doubl&d by irradiation for the full test temperature range. The ductile-to-brittle transition temperature, as measured by the values of the tensile ductility (i.e., percent elongation or reduction in area), was raised from an initial value of —65 to -230°C by the irradiation. Postirradiation creep-rupture test results are 800 1000 available on tungsten and W 25% Re irradiated to modest fluences at elevated temperatures.57 Fg- ure 26 shows the effect of the irradiation on creep Fig. 25 Temperature dependence of the strength and duc- ductility, and Fig. 27 shows ereep curves. Creep tility of unirradieted and irradiated tungsten. (From ductility of the alloy is given in Fig. 28. Steichen.'7)

~~I I I I I I /IRRADIATION TEMPERATURE: 1200iMeV)~~

~~TUNGSTEN CREEP RUPTURE DUCTILITY~~

~~1400 °C 5 _ /CONTROL~~

~~/ . IRRADIATED oL^y__J L 0 20 40 60 80 100 120 140 160 TIME (hrl 1 T 1700 °C 1900 °C —~~

~~— .IRRADIATED — i — IRRADIATED roNTROL~~

~~800 1000~~

70 Swelling due to cavity formation will result for I I W-25Re refractory metal alloy components during service CREEP SUPTURE at the proposed operating conditions. At most, — DUCTILITY swelling near 5% can be anticipated (and thus an increase of 1.7% in linear dimensions). However, data on several of the refractory metal alloys indi- £050 cate swelling resistance. In these systems, alloying E has either narrowed the range of irradiation tem- E peratures over which swelling will occur or ha3 decreased the amount of swelling relative to the z unalloyed base metal. In any case, swelling is not expected to be a 2 30 problem in the reactor designs tentatively identi- o fied for space applications. Since maximum dimen- o sions of these reactor designs are of order 0.5 to 20 m, swelling increases in linear dimensions of less O UNIRRADIATED CONTROL SPECHl^N iian 2% do not suggest any operational or reactor • IRRADIATED 42x1019 n/cm2 ( >1 MeV) fe-limiting problems. A IRRADIATED 2.7 x1019 n/cm2 19 Experience with the refractory metals identifies m IRRADIATED \.2 x 10 n/cm' the greatest potential problem to be the loss of I I ductility that results from irradiation. Strength 700 80C 900 1000 1100 1200 values of all candidate alloys will be increased by TEMPERATURE CO irradiation. In alloys based on molybdenum and tungsten, both a shift in DBTT to well above room Fig. 28 Ductility vs. test temperature for unirradiated temperature and a general reduction of elongation and irradiated W-25 Re creep-rupture specimens. Irradia- at temperatures above the DBTT are anticipated. tions conducted at reactor ambient temperature to a fast In the niobium and tantalum alloys, it is expected (>0.1 MeV) neutron fluence range of 1.2 to 4.2 X 10" 1 neutrons/cm2. (See Ref. 57.) • that the DB^"" of irradiated material will remain well below i jm temperature. However, severe reduction in uniform elongation ai;H some reduc- No data on irradiation creep in tmgsten or tion in total elongation will result. tungsten alloys are available, and no TEM results Where postirradiation fatigue tests have been on phase stability exist except for the W-Re sys- conducted, they have shown only small effects of tem. irradiation. There have been no in situ measurements of fatigue properties during irradiation. Littl'1 is known about the effects of irradiation SUMMARY OF THE STATUS on the creep behavior of the refractory metals. OF TECHNOLOGY ON There ?re *nme indications that the hardening pro- IRRADIATION EFFECTS duced by irradiation can reduce the creep rate in postirradiation tests. There is also the expectation, Irraulation effects data on candidate alloys for supported by a single experiment, that irradiation space reactor applications at relevant irradiation creep will result in deformation rates in-reactor and test conditions are almost nonexistent. The few that exceed those measured during out-of-reactor existent data on refractory metals are either for tests. unalloyed metal or simple alloys. The available Phase stability under irradiation can also result data on these alloy systems and the guidance pro- in property changes. This effect has not been vided by experience with stainless steel and other examined in most of the candidate alloy systems. metals are adequate to predict general classes of Of the refractory alloy systems based on V, Nb, irradiation effects that may be of importance. This Ta, Mo, and W, the largest data base for irradia- approach can be used to provide a relative ranking tion effects is for the vanadium-base alloys. In this of the several alloys of interest and can be used to system, emphasis has been on alloys containing estimate the magnitude of property changes. This titanium, and the coverage of irradiation condi- body of information, however, is not adequate for tions includes temperatures to 750°C and fluences 22 2 the detailed prediction of the behavior of com- to 6 X 10 neutrons/cm (>0.1 MeV). The alloys ponents during service in a space reactor. containing titanium have been found to be resis- EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 275

tant to void formation and swelling and to retain demonstrated that the same DBTT shift observed adequate postirradiation ductility in tensile tests for molybdenum occurs during irradiation of tung- conducted over the range room temperature to the sten. Postirradiation creep testing of W and W-Re irradiation temperature. Changes in the amount shows that the creep rate and the elongation to and distribution of precipitate phases occur during fracture are generally reduced by the irradiation. irradiation, but these have not been characterized Irradiation of W-Re alloys has also been shown to in detail. result in the precipitation of a phase not predicted Niobium alloys have been studied extensively by by available phase diagrams. using ion bombardment to simulate neutron irradiation. These results have established the role of oxygen in promoting swelling in this system and TECHNOLOGY NEEDS ON have shown that Nb-Zr alloys are swelling resis- IRRADIATION EFFECTS tant relative to unalloyed niobium. Neutron irradi- 1 ation results show the s elling of Nb-1 Zr to be The few data that do exist for irradiation restricted to a narrow temperature range near effects on candidate alloys for space reactor appli- 800°C. Tensile tests of irradiated Nb and Nb-1 Zr cations can be us«d to predict probable behavior show that large strength increases result for a but are not adequate either to confirm or to reject range of irradiation conditions, includirg irradia- a potential candidate alloy. Steps to improve the tion temperature up to 9JO°C. Total tensile elonga- data base and th? ability to predict reactor compo- tion after irradiation is frequently in the 5 to 10% nent behavior are given below. range, but zero uniform elongation has been Reactor desipuers need to provide information to observed for irradiation conditions that produce allow more precise evaluation of the needed irradi- populations of dislocation loops but few or no voids. This usually results for irradiation tempera ation effects data. These include: tures below the void formation temperature range. • Specificatio I of lifetime fluences on reactor component? in conceptual designs. Little study has been given to tantalum alloys, • Identification of operating temperatures and but strong similarities to niobium alloy systems shutdown temperatures. are observed. The peak swelling temperature for • Identification of critical and desirable proper- unalloyed tantalum is near 600°C, and alloying ties and allowable ranges of uncertainty in with tungsten reduces the amount of swelling. the data I ase. Tensile properties of Ta and T-lll irradiated af. about 400 and 640°C are quite similar to the Near-term experiments must be identified and results for Nb and Nb-1 Zr discussed above. evaluated to provide scoping data by the end of Molybdenum and the alloys Mo-0.5 Ti and TZM FY 85. These will be restricted to irradiation have been studied by a number of different investi- experiments that have completed exposure (or are gators. The results show quite similar behavior for now in-reactor) and can yield critical data on can- the three materials. Swelling peaks for irradiation didate alloys. temperatures between 600 and 800°C but continues • Irradiated but unexamined EBR-II tantalum for irradiation temperature.? at least as high as alloy experiments can provide tensile proper- 1500°C. Swelling in the TZM alloy seems to be ties, including ductility data for T-lll irradi- greater than in unalloyed Mo. The most important ated to ~1 X 1022 neutrons/cm2 in the range effect of irradiation or. mechanical properties is 450 to 1000°C. Lesser coverage of ASTAR- the hardening that produces an increase in the 811C H included, and swelling data will also DETT. Irradiations at temperatures in the range be produced. 400 to 1000°C have all been shown to shift the DBTT to well above room temperature, with the • Similar experiments with other candidate greatest shift produced by irradiation at ~600°C. alloys may be identified in storage at other Molybdenum-rhenium alloys have received too lit- DOE sites. tle attention to allow prediction of their behavior New experiments are required to provide data under irradiation at conditions of interest. that adequately cover the range of reactor The few data available on tungsten alloys show operating parameters and candidate alloys. These swelling of tungsten for irradiation at tempera- experiments cannot yield data by the end of FY 85, tures between 425 and 1500°C, with a prob.ible but the lov>g time involved in reactor irradiation pe^k swelling rate near 800°C. No swelling has experiments mandate an early start. The following been detected for W-Re alloys irradiated under the considerations should be included in planning one same experimental conditions. One experiment has or more major experiments. 276 WIFFEN

• The experiment design-construction-irradia- 12. H. Bohm, W. Dienct, H. Hauck and H. J. Laue, Irradiation tion-evaluation cycle will be a minimum of Effects on the Mechanical Properties of Vanadium-Base Alloys, Effects of Radiation on Structural Metals, pp. three years, probably longer. 95-106, ASTM-STP-426 (1967). A multiparameter experiment can include: 13. G E. Korth and R. E. Schmunk, Low-Cycle Fatigue of Three Irradiated and Unirradiated VANSTAR AHoys, 1. Several candidate alloys. Effects of Radiation on Structural Materials, pp. 466-476, 2. Base metal and weldments. ASTM-STP-683 (1979). 3. A range of irradiation temperatures. 14. L. J. Pionke and J. W. Dads, Technical Assessment of 4. Sample types for several critical property Niobium Alloys Dati, Base for Fusion Reactor Applications, measurements (e.g., tubes for irradiation McDonnell Douglas Astronautics Company-St. Louis, Report COO-4247-2, Decomber 1978. creep and sheet tensile specimens). 15. F. W. Wiffen, The Effect of Alloying and Purity on the Formation and Ordering of Voids in BCC Metals, in End-of-life fluences are easily attained in Proceedings of the 1971 International Conference on Radia- EBR-II or FFTF. tion Induced Voids in Metals, Albany, NY, pp. 386-396, USAEC Symposium Series 26, April 1982. 16. F. W. Wiffen, The Tensile Properties of Fast Reactor Neu- tron Irradiated BCC Metals and Alloys, pp. 176-196, REFERENCES Defects and Defect Clusters in B.C.C. Metals and Their Alloys, National Bureau of Standards, Gaithersburg, 1. F. W. Wiffen, Radiation Damage to Refractory Metals as Maryland, 1973. Related t' Thermionic Applications, Oak Ridge National 17. D. J. Michel and J. Moteff, Rod. Eff., 21: 235 (1974). Laboratory Report ORNL-TM-3629 (February 1972), and 18. Y. Adda, Report on the CEA Program of Investigations of pp. 156-165 in Thermionic Conversion Specialists Confer- Radiation-Induced Cavities in M3tals: Presentation of ence, IEEE, New York, 1971. 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Wiffen, Oak Ridge National Laboratory, unpublished ment for Irradiation Performance Semiann. Progr. Rept results from a niobium alloy irradiation experiment in September SO, 1981, DOE/ER-0045/7, M*.:-cii 1982. HFIR. EFFECTS OF IRRADIATION ON REFRACTORY ALLOYS 277

29. F. W. Wiffen, Creep and Tensile Properties of Helium Molybdenum Alloys Irradiated Between 425 and lOOCC, Injected Nb-1% Zr, CONF-750989, pp. 11-344-361, J. NucL Mater., 85-86: 901-905 (1979). Radiation Effects and Tritium Technology for fusion Reac- 46. T. H. We ..ter, B. L. Eyre, and E. A. Terry, The Effect of tors, March 1976. Fr.dt Neutron Irradiation on the Structure and Tensile 30. J. R. Remark, A. B. Johnson, H. Farrar, and D. G. Properties of Molybdenum and TZM, in Proceedings of the Atteridge, Helium Charging of Metals by Tritium Decay, Omference on Irradiation Embrittlement and Creep in Fuel NucL TecknoL, 29: 369-377 (1976). Cladding and Core Components, organized by the British 31. F. W. Wiffen, The Microstructure and Swelling of Neutron Nuclear Energy Society, London, November 9-10, 1972, Irradiated Tantalum, J. NucL Mater., 67:119-130 (1977). p. 61, The British Nuclear Energy Society, London. 32. J. F. Bates and A. L. Pitner, NucL TecknoL, 16: 406 (1972). 47. J. M. Steichen, Tensile Properties of Neutron Irradiated 33. R. A. Murgatroyd, I. P. Bell, and J. T. Bland, Dimer i >• al TZM and Tungsten, J. NucL Mater., 60: 13-19 (1976). Stability of Tantalum During Fast Neutron Irradiati ., in 48. V. A. Kasakov, A. N. Kolesnikov, V. A. Krassnoselov, Z. E. Properties of Reactor Structure Alloys After Neutron or Ostrowsky, A. S. Pokrowsky, V. I. Prchopov, E. M. Particle Irradiation, ASTM STP 570, pp. 421-432, American Loboda, A. V. Smirnoff, V. K. Shamardin, Effect of Neu- Society for Trsting and Materials, 1975. tron Irradiation on Properties of Potential Structural 34. J. H. Evans, Void Swelling and Irradiation-Induced \oid Materials for Thermonuclear Reactors, paper distributed at shrinkage in Neutron Irradiated Molybdenum and TZM, USSR-US Exchange on CTR Materials, Nov. 1974. J. Nucl Mater., 88: 3L-41 (19S0). 19. F. W. Wiffen, Mechanical Properties of Irradiated Molybde- 3J. F. W. Wiffen, Swelling of Irradiated Tantalum and Tan- num, pp. 27-48, ORNL-TM-4156, Fusion Technology Studies talum Alloys, HEDL-TME 72-144; Quart Prog. Rep. Irradi- P. igress Report for Period Ending Dec. 31, 1972, May ation Effects ~H Reactor Structural Materials, 1973. August-October 1972; also Oak Ridge National Laboratory, 50. J. Moteff, P. C. Rau, and F. D. Kingsbury, Influence of Report ORNL-TM-4055, pp. 30-34, February 1972. Irradiation Temperature on the Creep-Rupture Properties 36. T. T. Claudson and H. J. Pessl, Irradiation Effects on High of Foij crystalline Molybdenum, in Radiation Damage in Temperature Reactor Structural Materials, BNWL-23. Reactor Materials, Vol. II: 269-281,1969. February 1965. 51. H. H. Smith and D. J. Michel, The Effect of Irradiation on 37. B. Badger et al., Radiation Damage and Materials, the Fatigue and Flow Behavior of TZM Alloy, J. NucL UWFDM-150, Chap. IX UWMAK-III, A Nondrcular Mater., 66:125-142 (1977). Tokamak Power Reictor Design, University of Wisconsin, 52. D. Mosedale, G. W. Lewthwaite, and I. Ramsay, Irradiation Madison, WI, July 1976, Creep of Fast Reactor Materials, in Proceedings of an 38. R. E. Schmunk and G. L. Kulciuski, Survey of Irradiation International Conference on the Physical Metallurgy of Data on Molybdenum, University of Wisconsin Report Reader Fuel Elements, held at Berkeley Castle and Berke- UW*'DM-J61, June 1976. ley Nuclear Laboratories, Glouceatershire, Sept. 2-7, 1973, 39. J A. Sprague, F. A. Sniidt, Jr., and J. R. Reed, The Micro- pp. 132-135, J. E. Harris and E. C. Sykes (Eds.), The structures of Some Refractory Metals and Alloys Following Metals Society, London. Neutrr Irradiation at 650°C, J. NucL Mater., 85-86: 53. J. Matolich, N. Nahm, and J. Moteff, Swelling in Neutron 739-743 u979). Irradiated Tungsten and Tungsten-25 Percent Rhenium, 40. J. Bentley and F. W. Wiffen, unpublished data on irradia- Scripta Met, 8: 837-S42 (i974). tion of molybdenum in EBR-II, 1978. 54. F. W. Wiffen, Oak Ridge National Laboratory, unpublished 41. J. Bentley and F. W. Wiffen, Neutron-Irradiation Effects results from a tungsten irradiation experiment in EBR-II. in Molybdenum and Molybdenum Alloys, in Proceedings of 55. R. C. Rau, R. L. Ladd, and J. Moteff, Voids in Irradiated tht Second Topical Meeting on the Technology of Controlled Tungsten and Molybdenum, General Electric Company Nuclear Fusion, pp. 209-218, CONF-760935-P1 (1976). Nuclear Systems Programs, GEMP-692, (1969), General Electric Company, Cincinnati, Ohio; also J. NucL Mater., 42. D. S. Gelles, D. T. Peterson, and J. F. Bates, Void Swelling 33: 324 (1969). in the Molybdenum Alloy TZM Irradiated to High Fluence, 56. V. :j. Bykov, G. A. Birzhevoi, and M. I. Zakharova, Change J. NucL Mater., 103-104: 1141-1146 (1981). in the Density of Single-Crystal Tungsten Curing Neutron iC. •!. F. Stubbins and J Moteff, Swelling Behavior of Mo- Irradiation, translated from At Ehnerg., 32(4): 323-324, O.o Ti and TZM Bombarded with Heavy Ions at Tempera- April 1972. tures Between 750 and 1450°C, J. NucL Mater., 103-104: 1163-1168(1981). 57. F. D. Kingbbury and J. Moteff, Effects of Neutron Irradia- tion on Creep-Rupture Properties of W-25Re Alloy, 44. N. Igata, A. Kohyama, and S. Nomura, In-Situ Observation ASTM-STP-426, The Effects of Radiation on Structural of Void Swelling in Molybdenum Alloys by Means of High Metals, pp. 512-533, (1967), and Progress Report No. 74, Voltage Electron Microscopy, J. NucL Mater., 103-104: GEMP-1006, pp. 37-40, General Electric Company Missile 1157-1162 (1981). and Space Division, AEC Fuels and Materials Development 45. B. L. Cox and F. W. Wiffen, The Ductility in Bending of Program, June 29,1968. Summary of Key Needs for Further Research and Development on Refractory Alloys for Space Nuclear Power Applications

W. O. Harms Oak Ridge National Laboratory*

As stated in the Introduction, the principal PROCESSING AND PRODUCTION objective of this symposium is the identification and documentation of the refractory alloy research • Integrated iacility for producing refractory and development needs for the SP-100 Program. alloy hardware. The purpose of this paper is to summarize these • Designation of a knowledgeable group to take needs, principally as they were developed in the responsibility for procurement and fabrica- workshops held in connection v.'ith the symposium. tion of an inventory of refractory alloys. The key needs are listed according to the technical • Large ultrahigh (ion-pumped) vacuum areas addressed in the symposium. annealing facility, particularly for tubing. • Facilities for surface conditioning of fabri- COMPATIBILITY cated products prior to annealing of welding. • Additional fabrication experience and product • Experimental assessment of the feasibility of evaluation data for ASTAR-811C alloy. using refractory alloys in inert-gas environ- • Determination of the effect of ther- ments for Brayton cycle applications. momechanical processing on the long-term • Reestablishment of systems frr handling, recrystallization behavior of molybdenum purification, and analysis of alkali metals. alloys. • Reestablishment of ultrahigh vacuum equip- • Development of fabrication processes for ment for compatibility testing of refractory electron-beam-melted and arc-cast Mo-Re metals. alloys. • Evaluation of corrosion behavior of ASTAS- 811C in lithium, including effects of oxygen content, heat treatment, and welding parame- WELDING AND COMPONENT ters. FABRICATION • Determinatirn of mass transfer rates in the T-lll/W-2fFe dissimilar metal loop system. • Development of methods for precluding • Determination of mass transfer oehavior ?? cracks in multipass welds of tantalum alloys. Mo-13Re in circulating lithium and under • Weld repair methods for Nb-lZr, C-103, and conditions of evaporative heat transfer. ASTAR-811C. • Assessment of long-term behavior of • In-process screening test for assessing weld- ASTAR-811C as boiler, condenser, and static ability (tramp elements, airborne contam- turbine blade material under prototypic inants, heat-'o-heat variations). potassium Rankine cycle conditions. • Assessment of state of NDE technology for refractory alloy weldments. •Operated for the U. S- Department of Energy under con- • Development of brazinp alloys and braze-joint Tact W-7405-eng-26 with the Union Carbide Corporation. designs for joining of refractory alloys. 278 SUMMARY OF KEY NEEDS 279

~~MECHANICAL AND PHYSICAL PROPERTIES~~

Mechanical Properties Need*

Fatigue Application and/or 5000-b creep Low High Tensile Launch Material component strength cycle cycle strength Mirvivability*

Nh-lZr Secondary and tertiary loops _ X X _ X Postweld and components A X X X C-103 Secondary and tertiary loops - X X - X Postweld and components X X X X JC FS-85 Primary and secondary loops _ V X - X Postv.e'J and components X X X X X Mo-base T?,M Turbines and compressors - X X X Mo-Re Heat pipes X X X X X Postweld X X X X X T-111 Fuel clad and primary loops - X X - X Postweld X X _ X 811-C Fuel ciad and primary loops - X X - X Postweld X X X X X W-Re Heat pipe and emitters X X X X X Postweld X X X X X

~~'Fracture toughness including high cycle fatigue and crack growth, flnrlicates insufficient data to assess feasibility in the 900 to 1250°C range.~~

~~EFFECTS OF IRRADIATION~~

Guidance from concept design teams on life- previously irradiated in EBR-II at 450 to time fluences, operating and shutdown tem- 1000°C (ductility and swelling). peratures, and critical properties. Complete design-construction-irradiation- Evaluation of effects of 1 X 1022 n/cm2 expo- evaluation experiment cycle (estimated to sure on T-111 and ASTAR-811C specimens require a minimum of three years).

~~. GOVERNMENT PRINTING OFFICE 1984-746-013/2436 Contents~~

~~Page~~

Origin and Organization of the SP-100 Program 1 J. Ambrus SP-100 Program Overview 5 V. C. Truscello Potential Refractory Alloy Requirements for Space Nuclear 14 Power Applications H. H. Cooper, Jr. Refractory Alloy Component Accomplishments from 1963 to 1972 18 E. E. Hoffman Compatibility of Refractory Alloys with Space Reactor 34 System Coolants and Working Fluids J. H. DeVan, J. R. DiStefano, and E. E. Hoffman A Review of Tantalum and Niobium Alloy Production 86 R. W. Buckman, Jr. Processing and Production of Molybdenum and Tungsten Alloys 98 W C. Hagel ,!. A. Shields, Jr., and S M Tuominen CVD Refractory Metals and Alloys for Space Nuclear Power Application 114 '' L. Yang, T. D. Gulden, and J. F. Watson Machining Refractory Alloys—An Overview 130 J. D. Christopher Welding of Refractory Alloys 146 G. G. Lessmann Refractory Alloy Component Fabrication 168 W. R. Young Mechanical and Physical Properties of Refractory Metals and Alloys 227 < J. B. Conway Effects of Irradiation on Properties of Refractory Alloys wi h Emphasis 252 v on Space Power Reactor Applications F. W. Wiffen Summary of Key Needs for Further Research and Development 278 ' on Refractory Alloys for Space Nuclear Power Applications W. O. Harms TEMPERATURE CONVERSION TABLES

~~°F K °C °C K °F K °C~~

1000 811 538 600 873 1112 850 577 1071 110r 866 593 650 923 1202 900 627 llfil 1200 922 649 700 973 1292 950 677 1251 1300 977 704 750 1023 1382 . 1000 727 1341 1400 1033 760 SPO 1073 1472 1050 777 1431 1500 1089 816 <'5O 1123 1562 110" 827 1521 1600 1144 871 HC) 1173 1652 1150 877 1611 1700 1200 927 950 12^ 1742 1200 927 1701 1800 1255 982 1000 1273 1832 1250 977 1791 1900 1311 1038 1050 i:i23 1922 1300 1027 1881 2000 1366 1093 1 i 00 1373 2012 1350 1077 1971 2100 1422 1149 1150 1423 2102 1400 1127 2061 2200 1477 12'' 1200 1473 2192 1450 1177 2151 2300 1533 1260 IL'0 1523 2282 1500 1227 2241 2400 158r> 1316 1300 i573 2372 1550 1277 :!331 2500 1614 1371 1350 162" 2462 1^00 1327 2421 2600 1700 1427 1400 1673 2552 1650 1377 2511 2700 1755 1482 1450 1723 2642 1700 1427 2601