ERIAY (j.l 3 445k D5L5450 b DEVELOPMENT AND CONSTRUCTION OF A MOLYBDENUM TEST STAND CENTRAL RESEARCH LIBRARY DOCUMENT COLLECTION LIBRARY LOAN COPY DO NOT TRANSFER TO ANOTHER PERSON If you wish someone else to see this document, send in name with document and the library will arrange a loan OAK RIDGE NATIONAL LABORATORY OPERATED BY UNION CARBIDE CORPORATION e FOR THE U.S. ATOMIC ENERGY COMMISSION Printed in the United States of America. Available from National Technical Information Service U.S. Department of Commerce 5285 Port Royal Road, Springfield, Virginia 22151 Price: Printed Copy $5.45; Microfiche $0.95 This report was prepared as an account of work sponsored by the United States Government. Neither the United States nor the United States Atomic Energy Commission, nor any of their employees, nor any of their contractors, subcontractors, or their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness or usefulness of any information, apparatus, product or process disclosed, or represents that its use would not infringe privately owned rights. (m ORNL-4874 UC-80 — General Reactor Technology Contract No. W-7405-eng-26 METALS AND CERAMICS DIVISION DEVELOPMENT AND CONSTRUCTION OF A MOLYBDENUM TEST STAND Compiled by J. R. DiStefano A.J. Moorhead Principal Contributors N. C. Cole R. E. McDonald J. R. DiStefano A. J. Moorhead December 1972 OAK RIDGE NATIONAL LABORATQORY ¢ Qak Ridge, Tennessee 37830 operated by UNION CARBIDE CORPORATION for the U.S. ATOMIC ENERGY COMMISSION LOCKHEED MARTIN ENERGY RESEARCH LIBRARIES AR 3 Y456 0515450 & —_— —_— CONTENTS A TaCt L L o e e e 1 Introduction . ........ ... .. i i e 1 Design of Test Stand . . ... ... e 3 Fabrication Development . ... ......... e e 6 Primary Fabrication of Molybdenum Components . . . ... ... ... ... ... .. . . . .. ... ... . ..... 6 Tubing Development . ... .. . ... . ... . ... ... e 10 Joining Development .. .. e 14 Welding . ... 14 Brazing . ..o e 27 Mechanical Couplings . . ... ... o 34 CONS U I ON L .. e e e e e 38 Mockup Construction . .. ... ... 38 Fabrication and Prefit of Components . ... ... .. . . . .. 41 Fabrication of Subassemblies . . . ... ... . . 42 Interconnection of Subassemblies .. ... ... .. 43 Acknowledgments . . ... e 43 Appendix A — Specifications for Purchase of Molybdenum Tubing . ... ... ... ... ... ... ... ...... 44 Appendix B — Electron Beam Welding Parameters for Tube-to-Header Joints ... .................. 57 Appendix C — Parameters for Welding of Butt Joints in Molybdenum Tubing Using an Orbiting-Arc Weld Head . ... ... . . 58 iii [ { DEVELOPMENT AND CONSTRUCTION OF A MOLYBDENUM TEST STAND Compiled by J. R. DiStefano A. J. Moorhead Principal Contributors Cole R. E. McDonald N. C. ‘ J. R. DiStefano A.J. Moorhead ABSTRACT The discovery of a process that uses liquid bismuth at 500 to 700°C to remove protactinium and fission products from the molten salt fuel of a breeder reactor led to a search for suitable containment materials. Although several other refractory metals or graphite may be suitable, molybdenum appears most promising. Therefore low-carbon, low-oxygen molybdenum prepared by arc casting was chosen as the structural material for a reductive-extraction test stand that would be representative of typical equipment. We recognized that the use of molybdenum as a structural material would require unorthodox assembly procedures and impose stringent limitations on the system design. However, this material apparently possesses the best combination of properties such as fabricability, and oxidation and corrosion resistance. Final design was determined after the development of appropriate fabrication and joining techniques. ' ' Procedures were developed for the production of closed-end molybdenum half sections by back extrusion. Parts that were generally free from cracks and had high-quality surfaces were produced by the use of ZrOj,-coated plungers and dies and extrusion blank preheat temperatures of 1600 to 1700°C. In cooperation with a commercial vendor, we found that molybdenum tubing with improved ductility could be produced by careful removal of contamination introduced during tubing fabrication. : Complex components were fabricated by welding, using either the gas tungsten-arc or electron-beam process. Welding studies centered on three major types of joint: tube-to-tube, tube-to header, and header-to-header. Two of the most important factors found to minimize weld hot cracking were stress relieving and.preheating of components prior to welding. Mechanical bonding techniques were developed to join small-diameter tubing to back-extruded end sections. Experiments carried out at 250°C in an argon atmosphere produced helium leak-tight joints. Brazing filler metals were developed that are reasonably corrosion resistant to bismuth and molten salt up to 700°C. Techniques involving induction or resistance heating were developed to braze the several types of joints used in the test stand. . The Molten-Salt Reactor Program was terminated before construction of the test stand was completed, but an unjoined mockup using molybdenum components was assembled, and detailed assembly procedures were worked out. All of the techniques required for final assembly were demonstrated. INTRODUCTION A key feature in the conceptual design of the single-fluid molten-salt breeder reactor is the connecting chemical processing plant to continuously remove protactinium and fission products from the fuel salt (Fig. 1). Protactinium, the intermediate element in the breeding chain between thorium and ?°?U, has a significant neutron capture cross section and must be kept out of the core to obtain a good breeding ratio. Rare-earth fission products are also neutron poisons and must likewise be removed for good breeding. In 1968 the chemical feasibility of a process that uses liquid bismuth containing dissolved lithium and thorium as reductants to extract protactinium and rare earths from fuel salt containing both uranium and thorium was demonstrated.! A simplified flowsheet based on this process is shown in Fig, 2. One of the requirements for the development of the reductive-extraction process is identifying materials that are compatible with both molten fluoride salts and bismuth containing reductants. Hastelloy N 1. M. E. Whatley and L. E. McNeese, Molten-Salt Reactor Program Semiannu. Progr. Rep. Feb. 29, 1968, ORNL-4254, pp. 248-51. PRIMARY SALT PUMP I — | — PURIFIED SALT GRAPHITE MODERATOR REACTOR l HEAT EXCHANGER T 566°C = ORNL - DWG 68-{185ER SECONDARY NaBF,—NaF SALT PUMP COOLANT SALT ’ CHEMICAL PROCESSING PLANT FUEL SALT T . LiF ~BeF, - ThF, - UF, ——— STEAM GENERATOR ) /) . TURBO- GENERATOR STEAM Fig. 1. Single-fluid, two-region molten-salt breeder reactor. FUEL SALT ORNL-DWG 71-9004 RECONSTITUTION Ufg URANIUM REMOVAL REACTOR | S— EXCESS U 'FOR SALE UFg PROTACTINIUM RARE EARTH —(-- REMOVAL REMOVAL FISSION PRODUCTS TO WASTE Fig. 2. Simplified flowsheet for processing the fuel in a molten-salt breeder reactor. (Ni—7% Cr—16% Mo—5% Fe) is a likely material of construction for the reactor because it has excellent compatibility with molten salts at 500 to 700°C. However, nickel has appreciable solubility in bismuth at these temperatures and, therefore, is unsuitable for those portions of a processing plant that have to contain liquid bismuth. Other commonly used construction materials such as iron- or cobalt-base alloys have lower solubilities in bismuth but rapidly mass-transfer under conditions involving a temperature gradient. Several refractory metals (tantalum, molybdenum, tungsten, rhenium)Z:3 and graphite4 appear promising, but each has limitations with respect to either fabricability, oxidation resistance, or compatibility with bismuth and salt. When it was proposed to build a reductive-extraction column that would be representative of typical equipment and in which distribution coefficients for various elements could be checked and engineering performance data determined, molybdenum was chosen as the material of construction. We felt that it possessed the best combination of properties required for the processing application, namely, corrosion and oxidation resistance, availability, and fabricability. However, no system this complex had ever been constructed of molybdenum, and we foresaw many difficult fabrication problems that would have to be solved. Molybdenum is a particularly structure-sensitive material; that is, its mechanical properties are known to vary widely, depending upon how it has been metallurgitally processed. The ductile-brittle transition temperature of molybdenum varies from below room temperature to 200—300°C, depending both up.on strain rate and the microstructure of the metal. Maximum ductility is provided in the stress-relieved fine-grained condition, and recrystaltization and grain growth are known to reduce fracture stress and ductility. Interstitial impurities such as oxygen and carbon often segregate at grain boundaries, and this can result in a decrease in grain-boundary mobility which also favors premature fracture and low ductility.> However, recent advances in vacuum-melting practices have led to the production of material with improved and more reproducible metallurgical properties. The arc-melted low-carbon, low-oxygen grade of molybdenum, available commercially, affords relatively good control of grain size and interstitial impurity level. Nevertheless, the use of molybdenum as a structural material requires highly unorthodox assembly procedures and imposes stringent limitations on system design from the standpoint of geometry and rigidity. DESIGN OF TEST STAND The principal molybdenum components of the test stand are the column, disengaging pots, bismuth and salt feed pots, and connectinig piping (Fig. 3). The column consists of a |'-in.-OD Raschig-ring-packed central column for contacting countercurrent streams of bismuth-lithium and molten fluoride salt. Enlarged end sections (37%-in.-OD pots) are provided for deentrainment and separation of the exit stream from the entering stream. The fluids flow countercurrently through the column because of their difference in density, and they are returned to their respective feed pots by an argon gas-lift system. Flow rates of the fluids are controlled by orifices located in the bottom of the 37%-in.-OD feed pots (Fig. 4). The orifices are removable, and different sizes may be used to allow a range of flow rates compatible with the limited head available in the feed pots. The feed pots contain internal baffles and Raschig rings (Fig. 4) to deentrain 2. H. Shimotake, N. R, Stalica, and J. C. Hesson, “Corrosion of Refractory Metals by Liquid Bismuth, Tin, and Lead at 1000°C,” Trans. Amer. Nucl. Soc. 10, 141—42 (June 1967). 3. J. W. Siefert and A. L. Lower, Jr., “Evaluation of Tantalum, Molybdenum, and Beryllium for Liquid Bismuth Service,” Corrosion 17(10) 475t—-478t (October 1961). ‘ 4. Molten Salt Reactor Program Semiagnnu. Progr. Rep, Feb. 29, 1972, ORNL-4782. 5. Molybdenum Metal, Climax Molybdenum Co., 1960, pp. 78—82. ORNL-DWG 72-5404 (3% in. OD x Bin. LONG) FEED POTS SALT FLOW METERING ORIFICE ] 6 | UPPER Bi—SALT T DISENGAGING POT === (3% in. OD x18in. LONG) T TsALT L ‘ ARGON —» fi / .- EXTRACTION COLUMN (1%5in.OD x 5ft LONG) {saLr ( LOWER Bi—-SALT DISENGAGING POT (3%in. OD x 16:in. LONG) SALT Bi ARGON —» = U INTERFACE POSITION INDICATOR Fig. 3. Schematic of molybdenum reduction-extraction test stand. Overall, 11/2 ft in diameter and 17 ft high. liquid from the argon in the gas lift and damp flow surges. The feed pots contain an access port for sampling and for adding thorium or lithium to the system. More cdmplete details on the design of the test stand have been reported elsewhere.6-7 ' One major problem in the fabrication of complex molybdenum equipment is its lack of ductility after recrystallization. For optimum fabricability, it should be subjected to at least a 50% reduction in area after recrystallization and then given a stress-relief treatment. Although we felt that the 37%-in.-OD pots could be produced by machining from bar stock, material in this size range would have poor as-machined properties because it would have received bnly a limited amount of working. The capacity of available metal-working equipment was limited to using starting or blank material only slightly larger than the required finished pots. We chose instead to fabricate these components by back extrusion, a process in which a cylinder with a closed end can be produced by hot working in the range 1200 to 1700°C. -Although molybdenum can be welded, the process generally results in very large grains in both the fusion and heat-affected zones. Molybdenum welds are very brittle at room temperature and have a tendency to hot-crack. We thought that our best chances for successfully welding molybdenum would be 6. E. L. Nicholson, Conceptual Design and Development Program for the Molybdenum Reductive Extraction Equipment Test Stand, ORNL-CF-71-7-2 (July 1971). 7. W. F. Schaffer, Jr., E. L. Nicholson, and L. E. McNeese, Quality Assurance Program Plan for the Molybdenum Reductive Extraction Equipment Test Stand — Job No. 12172, ORNL-CF-73-1-45 (February 1973). ORNL-DWG 7i-3745R2 COUNTERBGRE 2-4mils GREATER THAN TUBE DIAM FOR RAZE CLEARANCE—-| |—~_L = '/gin. 1/4 in H 2 ' \ A T QS % A\ TYPICAL DESIGN FOR l/l/’// P Ll NIY 77777 Ygin. TUBE-TO-HEADER TYPICAL EB WELD DESIGN FOR |.0.020in. ROLL-BONDING (TYP) N 7, TS TN BT TR Z / rA AND BACK BRAZING T LTSS T LI OIS LTSI TTLLLL, &7 XIS DESIGN FOR EB WELD VENT TUBE END DETAIL OUTLET Fig. 4. Feed pot for molybdenum chemical processing loop. with the electron-beam process, which minimizes contamination effects, heat-affected-zone problems, and abnormal grain growth. However, electron-beam welding is not applicable for joining lengths of tubing together or for joining the tubing to a pot where the tube passes completely through the end section. A technique using the gas tungsten-arc welding process with an orbiting electrode was developed to make the tube-to-tube butt welds, and a mechanical joining technique (roll bonding) was developed to make the latter joints. In addition, reinforcement of all welded or roll-bonded joints by brazing served to strengthen the joint as well as to add a barrier to fluid leakage if a crack developed. To utilize brazing, however, required the development of a filler metal that would both be corrosion resistant to bismuth and salt and have a melting temperature below 1200°C, the maximum temperature that we felt could be tolerated before the molybdenum would become overly embrittled as a result of recrystallization and grain growth. The development work undertaken was concomitant with design of the test stand, and there was a definite interrelation between the two functions. The final loop design represented a compromise between engineering requirements and progress made in the development of fabrication and joining procedures for molybdenum. FABRICATION DEVELOPMENT Primary Fabrication of Molybdenum Components R. E. McDonald The choice of molybdenum as the construction material for the test stand presented several fabrication problems. It was originally suggested that the pots and the large-diameter heavy-wall tubing be machined from bar stock. However, our experience was that bar stock more than 1% to 2 in. in diameter had poor mechanical properties. Most available mechanical properties data comes from thin sheet or small-diameter rod which has been heavily worked, and we felt that an end cap machined from 4-in.-diam bar would be weak and crack-prone. Fabrication development started before the test-stand design was completed. One method considered in making the pots was by forging heavy plate to form end caps and then welding them to heavy-wall extruded pipe. We had previously developed techniques to fabricate heavy-wall thngsten pipe, and 4-in.-diam by % in. wall by 50-in.-long sections had been made.® Molybdenum pipe could readily be produced using the same techniques. The forged end caps would then be welded or brazed to the large pipe body, producing the pot, and pots made by this technique would have good mechanical properties because of the mechanical working necessary to shape them. Available equipment aided us in the choice of back extrusion for producing the end caps and forward extrusion for producing heavy-wall tubing. The Metals and Ceramics Division at ORNL has available a horizontal extrusion press with an extensive tooling inventory. Three-, 4-, 5.6-, and 7-in.-diam containers with stems were on hand. The water-nitrogen accumulator system was capable of 1300 tons on the 4-in. stem, and an induction billet heater, 50 kW at 3000 cycles, was capable of “heating 4-in.-diam, 10-in.-long billets to 2200°C in 45 min. ' The first attempt to produce an end cap by back extrusion was made using a 4-in.-ID container with a Zr0,-coated split die having provisions for integral bosses, a solid die backer, and a 3-in.-diam stem with a 3-in.-diam ZrO;-coated tool-steel plunger attached. This is shown schematically in Fig. 5. A 3.950-in.-diam molybdenum blank was heated to 1300°C in the induction heater under an argon atmosphere, transferred to the container, and quickly pushed with the plunger. The blank neatly filled the die and flowed back over the coated plunger, producing a 4-in.-OD product. The tooling was conventionally cleared from the container, the die parted, and the plunger extracted. Several back extrusions were made to determine the optimum temperature, boss configuration, and skirt lengths. Back extrusions were successfully made at 1200, 1300, 1400, and 1500°C, and the maximum skirt length we obtained was 4 in. at 1500°C. During this phase of development, however, cracks were noted in the hemispherical portion and the skirt ends. Although it was first suspected that they occurred during back extrusion, we noted that fine cracks had been generated on the faces of the blank during machining, and we felt they could have been transferred to the product. In an effort to understand and eliminate this cracking problem, a blank was cut lengthwise into two sections, and a '4-in. square grid network of grooves was machined into each half. The grid of one half was filled with small-diameter tantalum wires, and the two halves were pinned together. The blank was heated to 1650°C and back-extruded. The halves of the extrusion were easily parted, and Fig. 6 shows the flow pattern of the molybdenum during extrusion. The flow pattern shows that the cylindrical extrusion blank was first pushed forward, filling the die cavities to form the bosses. After it bottomed out against the solid die backer, it flowed backward over the coated plunger to form the skirt. As the free surface of the skirt flows backward between the plunger and the die body, it is essentially undeformed. Therefore, if fine 8. R. E. McDonald and G. A. Reimann, Floating-Mandrel Extrusion of Tungsten and Tungsten-Alloy Tubing, ORNL-4210. 4 ] ORNL - DWG 70-3877A DIE BACKER PLUNGER DIE HOLDER ¢ BLANK i ‘ 3-in. STEM SLEEVE LINER CONTAINER SLEEVE Fig. 5. Tooling used in capsule fabrication by back extrusion. Y-108581 Fig. 6. Flow pattern in a molybdenum back extrusion revealed by a network of tantalum wires installed in half of the original blank. cracks exist on the back edge of the starting blank, then cracks in the skirt edge will occur. To minimize this type of cracking, the back edge of the starting blank was radiused and highly polished. At this point, our machining procedure was modified to require that all blanks be chemically etched and dye-penetrant- inspected after machining. Molybdenum, like tungsten, has a tendency of smearing over cracks during machining or grinding, and unless etched, it is difficult if not impossible to detect fine cracks. If cracks were detected, very fine grindin'g removed them prior to acceptance. This change in procedure noticeably reduced cracking in the back-extruded product. While demonstrating that end caps could be made reliably, we noted that fairly long lengths of skirt were extruded back over the plunger. We then decided it would be feasible to back-extrude two halves to make a pot. This approach would require one girth weld instead of the two required to attach the end caps to an intermediate section of pipe, and it would also eliminate the need to extrude 3%-in.-OD, %-in.-wall pipe. In order to back-extrude the end cap with a long cylindrical section, we thought that it would be necessary either to do it in several steps or to raise the preheat temperature of the blank. However, this led to further complications, We found that the ZrO, coating on the plunger stayed intact for only one or two pushes before it had to be recoated because the inner surface of the product was adversely affected. Also, the increase in the preheat temperature to 1650°C caused galling and tearing of the outer surface of the extrusion because of a reaction of the molybdenum with the steel container liner (Fig. 6). The lubricant used by ORNL for all extrusions of molybdenum, molybdenum alloys, tungsten, and tungsten alloys is what we term a base-metal oxide lubricant. At the hot working temperature of these metals and alloys a liquid oxide forms which is an excellent lubricant (covered by USAEC patent 3,350,907). However, the residence time of the back extrusion in the container is long when compared with that of a conventional forward extrusion, and the liquid oxide breaks down as a lubricant as the temperature of the steel approaches its melting point. We had observed that the surfaces of the extrusion that were insulated from the steel by the plasma-sprayed ZrO, were not tearing (Fig. 6); therefore the tooling was changed so that molybdenum was in contact only with ZrQ,-coated surfaces during extrusion. To do this the 5.6-in.-ID container was used, in which a 5.6-in.-OD, 4-in.-1D, 9-in.-fong split die was inserted. The entire inner surface of the split die was plasma-sprayed with ZrQ,. At this point, the end design of the pot was also changed from a multiple-bossed hemispherical head to a large single-boss flat head as shown in Fig. 7. Three 8-in.-long back extrusions were made at preheat temperatures of 1600 to 1700°C with a stem load of up to 800 tons. When the length requirement of the half sections for the lower disengaging pot was increased to 9Y,s in., a new [2-in.-long die was made. Three long blanks were then back-extruded at 1600 to 1700°C with stem loads up to 800 tons. These extrusions exhibited good outer and inner surfaces, and skirt cracking did not exceed an inch in length. A typical back extrusion after machining is shown in Fig. 8. We thus showed that by using existing equipment — the horizontal extrusion press — half sections for the bismuth, salt, and the upper and lower disengaging pots could be made by back extrusion. A total of 12 back extrusions were produced for use in constructing the molybdenum test stand, and data concerning these products are summarized in Table 1. The second problem, the production of the large-diameter heavy-wall tubing for the packed column, was solved using techniques developed under the High Temperature Materials and Tungsten Programs at ORNL. About 6 ft of 1.16-in.-OD, 0.080-in.-wall tubing was required. For this, we used a 4-in.-OD, l1-in.-ID, 7-in.-long billet, which was heated to 1600°C and extruded over a ZrQ,-coated mandrel at a reduction ratio of 29:1. A second extrusion produced a tube 11% ft in length that was concentric within 0.007 in. tolerance in radius and with excellent outer and inner surfaces. The process used was the ORNL 1] 107985 ™ .l Fig. 7. Zirconium oxide-coated back-extrusion tooling with molybdenum starting blank and as-extruded pot half section. V111890 Fig. 8. Example of molybdenum back-extruded half section after machining. 10 Table 1. Molybdenum half sections back-extruded for the molybdenum test stand Extrusion Internal Part E::i:éleorn temperature length Results O (in.) Feed Pot 1197 1600 4 No cracks IFeed pot 1250 1600 "4 End cracks, 3.5 in. usable Feed pot 1257 1600 4.5 End cracks, 3.5 in. usable Feed pot 1259 1600 4.5 End cracks, 4 in. usable Spare 1251 1600 4 End cracks, 3.5 in. usable, , surface cracks on top Spare 1256 1600 4 Surface cracks on top Lower disengaging 1258 1600 8 No cracks Lower disengaging 1260 1600 8 No cracks Spare 1261 1600 8.5 Cracksin wall Upper disengaging 1286 [700 9 No cracks Upper disengaging 1290 1700 11 End cracks, 9.5 in. usable Spare 1288 1700 9 Reextruded, crack in wall floating mandrel technique, which is described in another report,” with the lubrication provided by molybdenum oxide. Even though a scale-up of the test stand would be required in construction of an actual chemical processing facility, we feel that we have demonstrated primary fabrication techniques for molybdenum that would be épplicable to containers up to 12 in. in diameter and 36 in. long. Fabrication of larger-sized components would require the adaptation of other fabrication techniques, such as ring rolling or power spinning, or the development of new techniques. Tubing Development J. R. DiStefano Four sizes of molybdenum tubing were required, and these were obtained commercially, as listed in Table 2. All of the material was purchased according to the following specifications: Specification No. (see Appendix A) Title MET-RM-B208 Tentative Specification for Seamless, Arc-Cast Mo Tubing for High Temperature Service MET-NDT-3 : Tentative Specification for Ultrasonic Inspection of Metal Piping and Tubing MET-NDTH4 Tentative Methods for Liquid Penetrant Inspection ' Initial investigation of three heats of this tubing revealed differences in microstructure, hardness, and response to heat treatment, but little difference in interstitial concentration (Table 3). Inspection of the inside of the tubes revealed surfaces which were rough or pitted. Metallographic examination of the as-received Y -in-OD tubing showed it was partially recrystallized. Heating to 925°C for 1 hr resulted in complete recrystallization. However, both the ;-in.- and %;-in.-OD tubing were received in a cold-worked condition. Heat treating to 925°C did not alter the hardness or microstructure of the %;-in. tubing, but the '/,-in. tubing softened and was almost completely recrystallized. Since the interstitial element concentra- tions in the different heats were essentially the same, we can assume that the surprisingly low 9. R.E.McDonald and C. F. Leitten, Jr., *“*Production of Refractory Metal Tube Shells by Extrusion and Flow-Turning Techniques,” pp. 85-92 in Refractory Metals and Alloys IfI; Applied Aspects, vol. 30, ed. by Robert 1. Jaffee (Proceedings of the Third Technical Conference, AIME}, Gordon and Breach Science Publishers, New York, 1966. ) ay 11 Table 2. Molybdenum tubing for chemical processing loop Tubing size No. of e L. Stress relief by Source oD Wall heats Identification manufacturer Vendor Manufacturer (in.) (in.) A 0.020 1 CPM-1 1 hr, 870°C TECO? Superior Tube? % 0.025 3 CPM-2 1 hr, 870°C TECO Superior Tube CPM-6 None TECO TECO CPM-8 1 hr, 870°C TECO TECO A 0.030 2 CPM-3 1 hr, 870°C TECO Superior Tube CPM-8 1 hr, 870°C TECO TECO % 0.080 1 CPM-5 1 hr, 870°C TECO TECO 9Thermo Electron Corp., Woburn, Mass. bSuperior Tube Co., Norristown, Pa. Table 3. Hardness, microstructure, and chemical analysis of molybdenum tubing as a function of heat treatment Size of tubiflg Condition DP};?EC;IE)%SS 2) Microstructure Concentration (ppm) (OD in.) Oxygen Carbon l/4 As received? 237 Cold worked 69 80 A 1 hr at 800°C 258 Cold worked Ya 1 hrat 925°C 250 Cold worked % As received® 200 Partially recrystallized (10—15%) 69 50 % 1 hr at 800°C 243 Partially recrystallized A 1 hrat 925°C 168 ' Completely recrystallized i, _ As received? 242 Cold worked 69 40 YA 1 hr at 700°C 255 Cold worked A 1 hr at 800°C 243 Cold worked. A 1 hr at 900°C 236 Very slightly recrystallized A 1 hrat 925°C 206 Almost completely recrystallized (90%) 4The tubing was stress relieved for 1 hr at 870° C before delivery. recrystallization temperatures for the ¥-in.-'and '%,-in.-OD tubing were the result of working prior to heat treating. " To evaluate the ductility of the tubing, we devised a somewhat qualitative test in which a 0.5-in.-long sample was impact-flattened a predetermined amount at different temperatures using the equipment shown in Fig. 9. The load was applied by a 2300-g weight dropped a distance of 6 cm. To obtain a quantitative measure of deformation, we divided the displacement (original ring diameter minus the minor axis diameter of the tested specimen) by the original diameter of the ring. The results of some of these tests are given in Table 4. The as-received "4-in.-OD tubing was “ductile” (no cracks) at room temperature, while we had to heat samples of the %-in.-OD tubing to 150—250°C and the %, -in.-OD tubing to 300°C before they became ductile. The behavior of the Y;-in.-OD tubing was traced to a brittle layer on its inside surface, as indicated in Table 5. Removal of 0.004 in. from the inside diameter (0.002 in. from the wall thickness) resulted in lowering the temperature at which acceptable ductility was observed from 300 to 125°C; when 0.006 in. 12 Table 4. Ductility of as-received molybdenum tubing as a function of deformation temperature Stress relieved 1 hr at 870°C of . Deformation Temperature . tubing dlsp!acement/ tube ©0) Observation (OD in) diameter (%) 1/4 8 25 Cracked where I} in tension 1/4 8 100 Cracked where ID in tension A 3 200 Cracked where ID in tension A 4 25 Cracked where ID in tension -1/4 4 100 Cracked where ID in tension A 4 150 Cracked where ID in tension A 4 175 Cracked where ID in tension A 4 200 Cracked where ID in tension l/4 4 250 Cracked where ID in tension Y 4 300 No cracks 3/3 6.5 25 Fractured into four pieces % 6.5 100 Cracked in three places % 6.5 150 Cracked in three places 3/8 6.5 250 ‘No cracks 3/8 3.25 25 Fractured into four pieces 3/8 3.25 100 Cracked in four places 3/8 3.25 150 No cracks 3/8 3.25 175 No cracks 3/8 3.25 200 No cracks 1/2 10 25 No cracks Table 5. Mechanical behavior of 1/‘;-in.-OD tubing from the inner surface _as a function of removing incremental layers Material Deformation removed displacement/tube Tem;zerature Observation from diameter (%) co ID (in.) 0.002 4 25 Hairline crack 0.002 4 125 Hairline crack 0.004 4 25 Hairline cracks 0.004 4 125 No cracks 0.006 4 25 No cracks 0.006 4 125 No cracks 0.008 4 25 No cracks 0.008 4 125 No cracks 0.010 4 25 No cracks 0.010 4 125 No cracks @ A 13 DIAL ORNL-DWG 71-8693 INDICATOR / IMPACT LOAD STRAIN 7] INCREMENT . SPUN GLASW?? o Yfifi\/ o é{ INSULATION 77 Ry ), *=ARGON MANUAL VARIAC g ? § CONTROL b \ b CLAMSHELL ? ! FURNACE ZONE CONTROLLED, BY TC-I { \ T g }‘g-spmmsu \\ é / 7= / Z ZONE / CONTRQLLED, / &2 TC-2=2 % § \ N LAVITE 7 "z INSULATION ‘ ELEVATING MECHANISM Fig. 9. Schematic of equipment used to impact-test samples of molybdenum tubing. was removed from the inside diameter, the material was ductile at room temperature. Chemical analysis indicated that material from near the inside surface contained higher oxygen and carbon concentrations compared with the bulk sample analyses (140 and 120 ppm, respectively, compared with 69 and 80 ppm). We suspected that contamination of the tubing occurred during fabrication and that it was not removed during subsequent cleaning or heat treating operations. Removing material from the inside of the %-in.-OD tubing also improved its room-temperature ductility, but its as-received microstructure (partially recrystallized as compared with the cold-worked fine-grained % -in.-OD tubing) led us to believe that the fabrication schedule used in its production might also be responsible for its lack of ductility. Consultation with the manufacturer!® led to the following changes in our specifications: 1. Starting tube shell shall have an average grain size of ASTM No. 6. No grain shall be larger than ASTM No. 3. . ‘ 2. Starting size of tube shell to produce %-in. OD X 0.025-in. wall product shall be 1.125-in. OD by 0.250-in. wall or 1.125-in. OD X 0.187 in. wall. In either case, no more than two intermediate anneals shall be used and the temperature shall be no higher than 815°C. 10. Thermo Electron Corp., Woburn, Mass. 14 3. Prior to the final stress relief, material shall be cleaned in an alkaline solution. The inside surface shall then be mechanically cleaned by wire brushing and then pickled to remove 0.001 to 0.002 in. from the wall. ] 4. Final stress relief shall be 1 hr at 800°C in dry hydrogen or vacuum (<5 X 107% torr). A quantity of %-in.-OD tubing was purchased according to these modified specifications and was found to be ductile at room temperature as measured by our flattening test. The ',-in.-OD tubing that was already on hand was acid etched to remove approximately 0.005 in. from its inside diameter, and selected samples were found to be ductile. It should be noted that, although the as-received “contaminated” tubing was not ductile in the impact flattening test, it could be bent at room temperature without fracturing. Some samples were bent up to 90° without evidence of cracks. The rate of bending was found to be important, but the data were not quantitated. Welding studies indicated that a preweld stress-relief heat treatment for 1 hr at 875 to 900°C was desirable to minimize weld cracking; however, samples of '4-in.-OD material (CPM-3) became embrittled when heated-to 925°C, and one heat (CPM-6) became embrittled when heated to 860°C. Tubing selected for loop construction was heat treated for | hr at 900°C in vacuum after all bends had been made. In this way, we took advantage of the as-received ductility of the material for bending and still satisfied the requirements for a preweld heat treatment. ‘ JOINING DEVELOPMENT - Welding A.J. Moorhead Material selection and preparation. Molybdenum has several characteristics that make it difficult to fabricate into complex structures by welding. It has a tendency to (1) undergo hot-cracking, (2) develop porosity in the weld fusion zone, and (3) undergo abnormal grain growth. In addition, it has a ductile-to-brittle transition temperature in large-grained microstructures (such as occur in the heat-affected or fusion zones of a weld) well above room temperature, which can easily cause fracture at ambient temperature. This latter characteristic makes handling of welded components, such as during subsequent assembly steps, quite difficult. Consideration was given in the selection of the base material and in the design and fabrication phases of the program to ways to overcome or at least minimize these characteristics. The selection of arc-cast molybdenum with low impurity element levels was significant to the welding development portion of the program. Low impurity levels are very desirable for this metal (especially O, N, and C) since small amounts of these elements have been shown to have adverse effects on the ductile-brittle tran- sition temperature and weldability of arc-cast molybdenum.?1—1!3 Oxygen is especially detrimental, as it forms low-melting eutectic films with molybdenmfi, and the presence of these films at grain boundaries can cause hot-cracking in welds. Molybdenum produced by powder metallurgy techniques was not considered for the test stand because welds in commercially available materials of this type have repeatedly been found to contain large amounts of porosity.!? Therefore, the selection of low-carbon, low-oxygen arc-cast base material helped to minimize two of the detrimental factors in welding molybdenum, namely, hot-cracking - and porosity formation. Great care was also taken in all subsequent operations to ensure that these 11. T. Perry, H. S. Spacil, and . Wulff, “Effect of Oxygen on Welding and Brazing Molybdenum,” Welding J. 33(9), 442-5—448-5 (1954), 12. W. N. Platte, “Influence of Oxygen on Soundness and Ductility of Molybdenum Welds,” Welding J. 35(8), 369-s—381-s (1956). 13. W. N. Platte, “Lffects of Nitrogen on the Soundness and Ductility of Welds in Molybdenum,” Welding J. 36(6), 301-5—306-s (1957). 14. N.E. Weare and R. E. Monroe, Welding and Brazing of Molybdenum, DMIC Report 108, p. 3 (March 1959). L )] 15 elements were not introduced into weldments by surface contamination or by an impure welding atmosphere. _ In order to ensure that the welds were not adversely affected by surface contamination, all components were cleaned using the portion shown below of a complex cleaning procedure attributed to. Ryan by [0 Thompson: 13 1. Degrease with acetone. 2. Immerse for 5 min in a 65—80°C solution of 10 wt % sodium hydroxide, 5 wt % potassium permanganate, and 85 wt % distilled water. 3. Rinse in flowing tap water, brushing to remove smut. 4. Immerse for 5—10 min in a room-temperature solution of 15 vol % sulfuric acid (95—97 % H,S0,), 15 vol % hydrochloric acid (37—38% HCI), 70 vol % distilled water, plus 6—10 wt % chromium trioxide (CI’Og). 5. Rinse in tap water, 6. Rinse in distilled water. 7. Dry with hot air. Although we did not find any correlation (as far as leak-tight welds were concerned) with the time between cleaning and welding, when parts had been excessively handled after the initial cleaning, we did take the precaution of repeating steps | and 4—7 of the procedure just prior to welding. After chemical cleaning, all parts used were given a vacuum stress-relief treatment for 60 min at 875 to 900°C at a pressure of 5 X 107° mm Hg or less. In our preliminary work, we found that a stress-relief heat treatment had a major effect on the weldability of the various forms of molybdenum. For example, one half of a length of ¥%-in.-OD, 0.025-in.-wall tubing was cleaned as previously described and then stress relieved in vacuum for 60 min at 875°C. The other half of the tube was chemically cleaned only. In a series of bead-on-tube “field” welds on these two pieces by the orbiting gas tungsten-arc process, we found that the welds on the cleaned and stress-relieved tube were crack-free, while those on the “cleaned only” tube had extensive center-line cracking. Similar results occurred on other sizes of tubing and on back-extruded half sections. We tried several other stressrelief temperatures to determine the lowest temperature that would produce the desired improvement in weldability and at the same time have the least effect on base-metal ductility due to recrystallization. A temperature of 875 to 900°C was apparently optimum from a weldability standpoint for these components, and it had little effect on the base-metal microstructure, as shown by Fig. 10. The combination of the chemical and thermal treatments left the components with surfaces that were lighter gray in color than in the as-received condition, and there was no noticeable grain-boundary attack by the etchant.. Welding procedure development. Procedures were developed for welding three major joint types for the test stand, two of which are illustrated in Fig. 4. A “tube-to-header” type of joint was required in seven locations in which a line terminated in the end of a back-extruded half section. A joint which we refer to as a “cylindrical girth” joint was welded to attach pairs of half sections together to form the four-fvessels or pots. Many welds of the “tube-to-tube’ type were required to attach various lines to stubs projeéilting from the pots or to one of the six tees. Each of these joint types will be discussed separately below. 15. E. G. Thompson, “Welding of Reactive and Refractory Metals,” Welding Research Council Bulletin 85, p. 7 (February 1963). : 16 V107759 ¥107763 Y-107758 (a)) (b)) (c)) 10. Effect of heat treatment (in vacuum) on the microstructure of -in.-OD, 0.030-in.-wall molybdenum tubing. Etchant: 50 vol % H30,-50 vol % NH4OH. (@) As received, DPH 241; (b) 60 min at 900°C, DPH 236 (¢) 60 min at 925°C, DPH 206 Tube-to-header joint. There are seven major tube-to-header joints in the test stand. Four join 0.875-in.-OD, 0.080-in.-wall tubing (which is machined to 0.050-in. wall in the joint area) to the feed pots; two attach 1.125-in.-OD, 0.060-in.-wall, 7.25-in.long stubs of the packed column to the disengaging pots: and the seventh attaches a 5.25-in.-long machined tube (with a %-in.-diam section at the weld) to the upper disengaging pot. There are also other miscellaneous joints of this type in the stand; for example, it was used in attaching the weirs to the horizontal baffle plates shown in Fig. 11. To facilitate making this type of weld between the massive bosses on the extruded half sections and the relatively thin-walled tube, a groove or trepan was machined inside the pot around the hole to produce a corner-flange joint, as previously illustrated in Fig. 4. This design provides a good heat balance between the two components and excellent mechanical support for the weld. Its major drawbacks are that it is relatively inacce: ible for welding or nondestructive examination. The use of the electron-beam process overcomes the accessibility problem for welding, and has the added benefit that it minimizes abnormal grain growth in the fusion zone because it is a process with a high energy density and a low total energy input. Because this type of joint is difficult to inspect by radiography or ultrasound, we relied on close parameter control, fluorescent dye penetrant inspection, and/or helium leak detection as our quality assurance techniques for these welds. 17 ¥-117087 Fig. 11. Weirs electron-beam welded (at arrowheads) (o one of the baffle plates shown installed in the upper half of the bismuth feed pot. Procedures were developed for electron-beam welding tube-to-header joints in tube sizes of 'y, %, Vs Uy, and 1% in. outside diameter. The parameters for welding all of these sizes of tubing are given in Appendix B, and an example of these two welds s shown in Fig. 12. The electron-beam welder used is of the high-voltage, low-current type (150 kV, 0.040 A max) with a chamber 36 in. wide, 23 in. deep, and 24 in. high. In our preliminary welds, in which flat plates simulated the vessel ends, we made welds both by the conventional method of rotating the work under the beam and by using u simple system for manually % in. in diameter) on the workpiece. Although we made acceptable rotating the beam in a circle (up to welds using both techniques, we had greater difficulties in making reproducible welds with the rotating-beam technique: so all subsequent prototy pe welds were made by rotating the joint rather than the beam. This latter technique was somewhat complicated for the two vent tube joints, which were located away from the center line of the pot, as seen in Fig. 13. However, this was compensated for by a simple diam fixture with cross slides for centering the weld under the beam. A close-up of a 0.375-ir tube-to-header weld is shown in Fig. 14, and a photomicrograph of a helium leak-tight weld joining a 1.125-in.-0D, 0.060-in.-wall tube to a back-extruded molybdenum header is shown in Fig. 18 Pio70 103077 ¥-111689 Fig. 12. Back-extruded half section with two tube-to-header welds made by the electron-beam process. Fig. 13. Off-centered vent tube weld (arrow) made by welding a “washer” to the tube and then welding the washer to alip in the feed pot bottom. 19 A Fig. 15. Cross section through a weld joining a 1.125-in.-OD tube to a half section. Etchant: 20 vol % Hy05-10 vol % H,504 (96% H2504)~70 vol % H,0. 29X ~~Yszin. | Y%in. Fig. 16. Joint design used for connecting back-extruded half sections by the gas tungsten-arc proce: of % in. was typically used A root opening PHOTO 194271 ¥-103241 Fig. 17. Manual gas tungsten-arc weld joining back-extruded molybdenum half sections. (a) Completed pot. (5) Cross section through the weld: etchant: 20 vol % H305 10 val % H,804 (96% H,504)-70 vol % H,0:6x 21 Cylindrical girth joint. Both the gas tungsten-arc and electron-beam welding processes were investigated for the welds required to join two half sections to form each of the four vessels. The joint design used for the gas tungsten-arc welds is shown in Fig. 16. The tungsten-arc welds were made manually in a vacuum-purged, argon-backfilled glove box. This chamber has the basic dimensions of 36 in. in inside diameter and 56 in. in length, but with available extensions the length can be increased to over 12 ft. The chamber was evacuated to a pressure of 5 X 10 mm Hg or less prior to backfilling with argon of 99.997% minimum purity by volume. Filler metal was 0.060-in.-diam low-carbon, low-oxygen molybdenum wire prepared commercially from cast stock. The electrode was 2% thoriated tungsten, % in. in diameter. No auxiliary preheating was required for the gas tungsten-arc welds. Evidently the high conductivity of the molybdenum promotes a rapid distribution of heat throughout the parts (i.e., preheat) before the fusion temperature is reached at the joint. A photograph of one of the gas tungsten-arc cylindrical girth welds (which had a helium leak rate less than 1 X 107 atm cm?® sec ™) and a cross section through it are shown in Fig. 17. Note that although this weldment has undergone abnormal grain growth, radiographically there was no evidence of porosity, and only two small cracks (on the inner surface in the heat-affected zone) were detected by fluorescent penetrant inspection. Electron beam girth welds were made using the rotary fixture shown in Fig. 18. A general requirement for welding with the process, if filler metal is not added, is that the joint must be held tightly together so that coalescence will occur and so that the resulting weld bead will not be underfilled. The harmful effects of an underfilled, concave weld bead are especially pronounced in crack-sensitive materials such as PHOTO 1217-72 Fig. 18. Molybdenum pot formed by an electron-beam girth weld joining two back-extruded half sections. 22 molybdenum. Unfortunately, the complex geometry of the vessel half sections, with their protruding tubes and internal baffle plates, made the application of end loading by conventional means (such as external pots or internal rods) quite difficult. To overcome this problem, a technique was developed in which three 0.032-in.-diam molybdenum pins through the step joint were used to hold the halves in intimate contact during welding. These pins were subsequent'ly fused into the weld. Parameters for joining molybdenum vessel half sections using the electron-beam process are listed below. The welds were made using a self-accelerating triode design electron gun (S-32). Manual beam current downslope was used at the end of the weld to prevent crater formation. 1. Joint type — step (illustrated in Fig. 4). . Joint thickness — 0.094 in. 2 3. Accelerating potential — 150kV. 4. Beam current — 12 mA preheat and 25 mA weld. 5 . Number of revolutions — 6 for preheat (2 preheat passes on either side of and 1 in. away from the joint, and then 2 passes on the joint itself) and 1%, for weld. 6. Travel speed — 32 in./min. 7. Focus — 0.015 A defocus. 8. Work distance — 4 in. Note that in this procedure, extensive preheating was carried out with the defocused beam prior to welding. A phdtomicrograph of a cross section through a preliminary electron-beam girth weld in which penetration through the step was not achieved is shown in Fig. 19. Tube-to-Tube Butt Joint. Tube-to-tube welds joining five sizes of molybdenum tubes were successfully made in an argon-filled glove box by the gas tungsten-arc process both manually and using an automatic orbiting-arc technique. Using either technique, we were readily able to produce welds that were helium leak-tight and that had little or no porosity as shown by radiographic and metallographic examination. However, the orbiting-arc technique was of particular interest after we found that our welders had considerable diffiéulty in making manual butt welds in thin-walled molybdenum tubing without excessive root reinforcement and without misalignment. These problems were compounded by the necessity of making the welds in a glove box, which reduces the senses of sight and touch. Several commercial orbiting-arc weld heads were evaluated, including the head shown in Fig. 20. In this unit the tungsten electrode is carried around the tube (which remains stationary) by a copper ring, which also holds a set of boron nitride gas cups. Inert gas is fed through the handle to provide an atmosphere suitable for welding in the cavity formed by the cups and body housing. The clectrode carrier is rotated by a small dc motor (with tachometer feedback) in the handle. The tubes are held together and in alignment by a pair of clamshell clamps, which have removable inserts to adapt to the various tubing sizes. In order to further reduce any stresses applied to the weld (such as by tube misalignment), we used several sizes of auxiliary yoke-type fixtures outside the weld head clamps, such as the one shown in Fig. 21. Devices of this type are invaluable, as molybdenum weldments are highly susceptible to cracking under the influence of even small stresses. The Ya-, Y-, and Yp-in-diam tube-to-tube welds were made using a magnetic amplifier power supply with capabilities for programming both variations in the welding current and in the rotational speed of the orbiting electrode carrier. The Y- and 1Y%-in-diam tube-to-tube welds were made with another programmable power supply, but we did not have a power supply available for varying the speed of the stepping motor in the orbiting-arc head used for the larger sizes of tubing. L 23 Fig. 19. Cross section through electron-beam cylindrical girth weld shown in Fig. 18. Etchant: 20 vol % H;0, - 10 vol % H3804 (96% H3504)~70 vol % H,0. 6x Y.117806 Fig. 20. Orbiting-arc weld head and %-in.-diam tube-to-tube butt weld. Figure 21 also shows a feature which we added to the orbiting-arc head to permit us to make molybdenum welds outside the confinement of a glove box: molded rubber sleeves which seal tightly around the head and the tubes. Without these sleeves we were unable to make crack-free welds outside the inert gas chamber (i.e., in the “field”), whereas with the sleeves we were able to make helium leak-tight welds in the Y-, %-, and %-in.-diam tubes. However, the results were not as reproducible with the field 24 — YI13975 Fig. 21. Orbiting-arc weld head with attached rubber sleeves for welding molybdenum tubing outside a glove box with yoke-type fixture. welding technique as with the orbiting-arc unit inside the glove box. Therefore we controlled our assembly procedure and sequence so that a minimum of field welds would be required in construction of the test stand, as explained later in this report Weld inserts (such as the one shown in Fig. 22 between two tees) were used in the orbiting-arc welds to aid in controlling root reinforcement and joint misalignment. Because it protrudes slightly above the tube surface, it has the additional function of providing a small amount of filler metal for the joint, thus producing convex weld beads, which we have found to be less prone to cracking than concave beads. The parameters for butt welding the tubing are given in Appendix C. The important features of the welding cycles are the relatively long portion at a lower “initial current,” which provides preheating without fusion of the joint; and the gradual increase in rotational speed (in the cycles for the %-, %-, and Yy-in-diam tubes), which minimizes *“weld-bead-widening” that occurs in tube welds as the overlap portion is approached. A cross section through a typical tube-to-tube butt weld made with the orbiting-arc process is shown in Fig. 23. 25 : e 2t TUBE WALL OF ; 2 jrmcmsssm r— W Tt L 2 (b) Yyin: Fig. 22. Orbitingarc welds in molybdenum tubing. () Molybdenum tees with attached lines and insert in position prior to assembly for welding. (b) Design of typical insert for %-, %-, and %-in.-OD tube-to-tube welds. Fig. 23. Cross section through an orbiting-arc field weld joining two 0.250-in.-OD, 0.020-in.-wall tubes. Etchant: 50 vol % NH4OH-50 vol % H,0;. S0X. » 27 ‘ " Brazing. N, C.‘Cole Filler metal development. In building complex structures from molybdenum brazing is attractive from two standpoints: (1) as a replacement for brlttle welds and (2) as a reinforcement for welds. But to avoid severe loss of ductility to the base metal, the filler metal must flow below 1200° C. Several commercial brazing filler metals will braze molybdenum, but they are rich in such metals as nickel, copper, silver, or gold, which are .not corrosion resistant to molten bismuth at 700°C. In addition, silver is not resistant to corrosion by molten fluoride salts. Unfortunately, few materials adequately resist attack by bismuth. The refractory metals, such as molybdenum, tantalum, rhenium, and tungsten, possess the best resistance; however, it is difficult to depress the melting point of a refractory-metal-based alloy below 1200°C without completely sacrificing 1ts corrosion resistance by the addltlon of large amounts of alloying elements. Under certain conditions, iron and some of its alloys are reasonebly'compatible with bismuth. By alloying iron with small amounts of selected elements, we were able ‘to depress the melting point of iron-based filler metals to betow 1200°C. Figere 24a shows portions of the binary diagrams!6:17 of Fe-C, Fe-B, and B-C. The diagrams show that. 1 to 4% B or C added to Fe depresses its melting point significantly. By adding small percentages of both boron and carbon to iron, we were able to depress the melting point below 1150°C. We also felt that the addition of molybdenum to the filler metal would enhance its corrosion resistance to bismuth and minimize dissolution of the molybdenum base metal during brazing. Even with the addition of molybdenum, we found we could still keep the melting point of the alloy below 1200°C. Figure 245 shows the binary phase diagrams!®¢ of Fe-Mo, Fe-B, and Mo-B. Since the temperatures of the liquidus and solidus-of the Fe-Mo diagram decrease as the molybdenum content increases from 0 to 35%, we were able to add as much as 25% Mo to the Fe-C-B ternary alloy without raising the melting temperature above 1175°C. From various wettability and flowability tests we found that additions of 15% Mo, 4% C, and 1% B gave the optimum ‘compositioh In an attempt to further depress the melting temperature, we added small amounts of germanium. Adding 5% Ge depressed the melting point to 1050°C and, as a bonus, also improved flowablllty To determine mechanical properties of brazed joints made with these experimental filler metals, we shear-tested two of the most promising.compositions. Miller-Peaslee shear-test specimens!8:'9 were brazed with alloys Fe—15% Mo—5% Ge—4% C—1% B and Fe—15% Mo—4% C—1%B. They were pulled with a tensile machine at a strain rate of 0.002 in./min at both room temperature and 650°C. Average test results are shown in Table 6. Room-temperature shear strengths of these alloys were greater than 30,000 psi, and these compare favorably with base-metal shear strengths of about 45,000 psi. At 650°C the shear strengths of the joint brazed with Fe-Mo-Ge-C-B averaged 29,000 psi, whereas those with Fe-Mo-C-B were 18,000 psi. This lower strength is, however, more than a’fde'quate‘ for most applications. The ductilities of joints brazed with both brazing alloys were outstanding at 650°C (elongations of 42 and 50%); and even at room temperature, where welded molybdenum is brittle, the.-'el,oh'gations of the brazed joints were at least 10%. 16. M. Hansen, Constitution of Binary Al?oys_, 2d ed., McGraw-Hill, New York, 1958.. . 17. R. P. Eliot, Constitution of Binary Alloys, First Supplement, McGraw- Hill, New York, 1965. 18. F. M. Miller and R. L. Peaslee, “Proposed Procedure for Testing Shear Strength of Brazed Joints,” Welding J. (N.Y.) 37(4), 144-s—150-s (1958). 19. R. G. Gilliland and G. M. Slaughter, “The Development of Brazing Filler Metals for High Temperature Service,” Welding J. (N.Y.) 48(10), 463 -s—468-s (1969) 28 ORNL - DWG 70-4215 c * % & 3, X é‘/ \", o v & % X7 %P7 2000 v A v v 1800 1600 N 600 ol e, 1400 K, 1400 ?@\’5 4}0 ,‘g" Re- . 1200 y ’ 1200 ¢, Fe BORON (wt %) B 5 1200 £ w A 3 1400 x ] w n \ 5’ \\_I = 1600 ) (a) ORNL-DWG 70-4216 / B * 3 q.‘? %Lf T AAAVAVAVAVAY o L] ° . . o & ‘ 2400 2 - NSy 2a00 . 2000 § ,00'0 000 ) é’hps Ry “\)QS, g ., 1600 1600 ?g@“ % W Srop, 1200 BAVAAVARY VARN 1200 © Fe MOLYBDENUM (wi %) , | Mo 1 ; o ‘eoo e N R ] < 1450°C,(37.5%) / o S— — 7 g, et T~o| A 5 1600 < . 14 Y [y W A Y v a N = z . = 2000 (b) Fig. 24. Portions of binary diagiams used to select alloy compositions for brazing studies. (¢) Fe-C, Fe-B, B-C; (b) Fe-Mo, Fe-B, B-Mo. : ' ' . b I 29 Table 6. Mechanical properties of brazed molybdenum joints, o ] Shear strength (psi) Elongation (%) Composition of brazing alloy R (Wt %) oom 650°C Room 650°C temperature : temperature Fe—15 Mo—5 Ge—4 C—1 B (42M) 30,000 29,000 10 50 Fe—15Mo—-4 C-1B (35M) 31,000 18,000 11 .42 The experimental brazing filler metals listed in Table 7 were tested in bismuth and fluoride salts at 700°C. All brazes survived the fluoride salt thermal convection loop tests with no metallographically detectable corrosion after 1032 hr. A typical example of the after-test appearance of these alloys is shown in Fig. 25. For chemical processing applications, compatibility of the braze speéimens with bismuth was considered the most stringent requirement, since the solubility of pure iron in bismuth is 50 ppm at 600°C and it is known to undergo temperature gradient mass transfer in liquid bismuth.29:2! Molybdenum tee-joint samples brazed with four of the iron-based filler metals (Table 7) were exposed to static bismuth for 671 hr at 600°C. The most extensive dissolutive attack occurred in the fillet area of the sample brazed with the Fe—4% C—1% B alloy. The area between the vertical and horizontal areas of the tee was more resistant to attack, and electron beam microprobe analysis showed that this area contained 5% Mo, no doubt as a result of alloying during brazing. Furthermore, the specimens brazed with filler metals containing molybdenum were considerably more resistant to dissolution in the fillet areas. Seven lap-joint specimens brazed with the four brazing filler metals were also exposed to flowing bismuth in a quartz thermal convection loop at 600 to 700°C for 2000 hr. All seven specimens were intact after testing. Three had been placed in the hot leg and four in the cold leg of the loop. (The Fe-C-B braze was placed only in the cold leg.) Weight-change data were not obtained, since some bismuth adhered to the specimens after test. Chemical analyses of the bismuth drained from the loop indicated that the amounts of iron and molybdenum, if present, were below the limit of detection, 3 ppm. Metallographic examination indicated the presence of one or more layers on the surface of each of the braze fillets. Microprobe analyses showed that the outer layer was rich in iron and the layer immediately under it was rich in molybdenum. 20. A.J. Romaro, C. J. Klamut, and D. H. Gurinsky, The Investigation of Container Materials for Bi and Pb Alloys. Part I Thermal Convection Loops, BNL-811, Brookhaven National Laboratory (July 1963). 21. B. R. T. Frost, C. C. Addison, A. Chitty, G. A. Geach, P. Gross, J. A. James, G. J. Metcalfe, T. Raine, and H. A. Soloman, “Liquid Metal Fuel Technology,” Proceedings of the Second United Nations International Conference on the Peaceful Uses of Atomic Energy, Geneva, 1958, 7, 13965, United Nations, New York, 1958. Table 7. Characteristics of ekperimental filler metals for brazing molybdenum . : Brazing Wettability Composition (wt %) temperature and Joint integrity €O flowability Fe—-4(C-1B 1150 Excellent Good Fe—15 Mo—4 C-1B 1150 Excellent Good Fe—25 Mo—-4 C-1 B 1175 Excellent (a) Fe—15 Mo—-5 Ge—-4 C—-1B 1050 Outstanding Good 2 Some tendency for cracking along base-metal—braze-metal interface. 30 Y-106672 1% B). (a) As brazed, (b) 25. Cross section of a molybdenum lap joint brazed with 35M (Fe-15% Mo-4% after testing in fluoride salts at 700°C for 1032 hr. Phases within the braze and along the interface between braze and base metal were also rich in molybdenum. Also, certain areas were enriched in bismuth, and it is interesting to note that those areas were associated with the regions of high molybdenum concentration. It appears that a complex intermetallic compound formed containing molybdenum, bismuth, and other unidentified eclements. Because of limitations of the microprobe equipment, we were unable to analyze for the presence of boron and carbon, the other elements alloyed in the brazing filler metal The results for each specimen were much the same except that a larger amount of bismuth was found in the fillet of the Fe—4% ( obtained in the static capsule test reported above, where dissolution occurred. We feel that the layers rich in 1% B brazement. However, the fillet was still inta , contrary to the results iron and molybdenum near the surface may have passivated the underlying material in this test After reviewing all filler metal requirements and the data generated on each of the alloys, we selected the composition Fe—15% Mo-5% Ge—4% C—1% B as the filler metal for brazing those sections of the chemical processing test stand that would contain bismuth. Development of techniques for brazing. Brazing techniques were developed for two purposes: to provide reinforcement of the weld zone, which is brittle at room temperature, and to serve as a backup seal should a leak develop through a cracked joint. In building a complex system of material that has the impact sensitivity and reactivity (with oxygen) of molybdenum, many unique techniques had to be developed to ensure reliable brazed joints. Since specific brazing techniques were needed for the variety of joints required, much of our work was necessarily devoted to basic joint design and process development and improvement. We developed several methods of brazing using both resistance furnaces and high-frequency induction as the heat sources. Furnace brazing — large pots. The four vessels with attached tubes were brazed in a resistance-heated vacuum furnace with a hot zone of 7 X 7 X 30 in. In each end of the pots, two or more short lengths of tubing which were joined to the pots by roll bond or welding were back-brazed. At the same time, a split ring was brazed around the girth welds joining the two half sections. These subassemblies were brazed at 1075 torr by heating at a rate of 5°C/min until flow of the brazing alloy was achieved. The temperature was monitored by several thermocouples placed at various locations throughout the furnace. The parts were 31 positioned in the furnace so the flow of the filler metal could be observed. Figure 26 shows an example of a 3%-in.-diam molybdenum pot in which tubing Jomed by roll bondmg or welding was back-brazed, and a split ring covering the girth weld was brazed into place In our preliminary work, we learned that a joint gap of at least 0.001 in. was necessary to ensure proper flow of the brazing filler metal into the joint. However, the joint preparation for roll bonding and welding required a tighter fit. Therefore, the portion of the joint to be back-brazed was counterbored an extra 0.001 to 0.002 in. on the radius to a depth of % in. Filler metal feeder holes were drilled into the bosses of the vessels to serve as reservoirs for the brazing alloy powder and to ensure that the filler metal did not prematurely flow away from the joint until the thick-walled boss reached the melting temperature. We were concerned that if the filler metal were placed in the fillet area and the thin-walled tubing reached the brazing temperature faster than the heavy pot, the filler metal might flow along the hotter tube (and away from the joint) rather than into the joint. Although the procedure generally was satisfactory, cracking of the braze metal in the feeder holes proved to be a significant problem. Cracking was detected by a helium leak-check on the brazed joint and probably was the result of shrinkage stresses as the braze alloy solidified: To avoid this problem a wire was inserted into the small-diameter portion of the feeder hole before the filler metal was added, and a molybdenum cap was brazed over the outer end of the reservoir. Furnace brazing of nozzle bodies. Several nozzle bodies extended through the stainless steel flange of the test stand. The bodies were made of nickel or stainless steel and were brazed to at least one molybdenum line. The joint design included a combination of trepans and feeder holes to obtain proper flow of brazing filler metal into the joint. We successfully brazed several mockups of these dissimilar metal joints (Fig. 27). , Since a small amount of bismuth vapor could reach the stainless steel nozzle body in the area of the ~ body-to-molybdenum-tube braze, it was to be brazed with the iron-based filler metal.-None of the nickel nozzle bodies were to be exposed to bismuth, so they could be brazed with a commercial brazing filler metal. Several filler metals were evaluated for the latter use, and Au—18% Ni was selected because it flowed adéquately. Each of the nozzle brazes was to be made vertically in a vacuum furnace by resistance heating in the area of the nozzle only. To avoid several additional welded and brazed joints in the molybdenum | tubing, an extension was added to the furnace to accommodate tubing up to 6 ft long. Glove box brazing. For brazing of sections of the 17-ft-long test stand that would not fit into any available furnace, it was necessary to design and build portable furnaces for localized heating in a glove box. The controlled atmosphere of the glove box was utilized to prevent oxidation of all portions of the assembly that were heated during the brazing operation. We investigated two types of localized heating sources: resistance and induction. Each has merit for different braze joints, depending on the size and location. . : Two types of resistance-heated furnaces were built, both utilizing tantalum heating elements. One had a helical heating element for brazing those joints over which it could be slipped on and off. The other had a split tantalum-sheet heater which could be opened and placed over the welded'joint in situ and then removed after brazing. This feature was necessary on sections where the furnace could not be slipped over a large or complex section to reach the braze region. We investigated induction heating for use both inside and outside an atmosphere chamber. We experienced problems with arcing in the glove box under argon or helium atmospheres as well as in vacuum when the brazing alloy binder (used for preplacing the filler metal) volatilized. We overcame this difficulty by installing an auxiliary transformer to reduce the high voltage from the power supply to low voltage and high amperage at the coil. With this new attachment, we were able to braze inside the glove box in argon or helium without arcing problems. By changing the geometry of the copper leads from thin-walled tubing to 32 114502 Fig. 26. Molybdenum pot after back-brazing of tube-to-header joint and brazing of ring around cylindrical girth weld. 33 Y.111576 Fig. 27. Mockup of stainless steel nozzle body brazed to molybdenum tubes. Body diameter, 2 in. closely spaced bus bars, we were able to easily achieve brazing temperature on even the largest size of tubing with attached heat sinks. Figure 28 shows a split sleeve brazed around a weld joining a length of 1Y-in.-diam tubing to a stub attached to a prototype vessel. A helical induction coil was slipped over the sleeve, and brazing was completed in 35 min by heating only the immediate area of the part. The heating rate was closely controlled, and the brazing temperature was monitored by a thermocouple placed under the edge of the split sleeve. Alternatively, we brazed this same 1%-in. diam configuration by heating locally with a portable resistance furnace. The portable resistance furnace is also capable of brazing all of the smaller tube-to-tube joints, provided that there is enough space around the tubing for the 3-in.-diam, S-in.-long insulated heater and that the assembly will fit into a vacuum chamber. Field brazing. Seven tubes have to be joined in the field to connect the subassemblies. Tubing of the size Y-, %-, Yh-, and possibly 7-in. outside diameter will first be welded and the split sleeves back-brazed on site. In this regard, we field-brazed (by induction) sleeves on mockups of all of these sizes of tubing. We surrounded the part to be brazed with a quartz tube placed inside the induction coil. Helium gas flowed over the heated area to protect the braze from air. However, the induction coil was still operated in air. When the induction unit was activated, the 2-in.-long section of tubing with the split sleeve was quickly brought to brazing temperature. Because of the small mass of material being heated, adjacent areas remained relatively cool, and the entire brazing cycle took less than 5 min. v.16502 Fig. 28. Mockup of column joined o upper disengaging section. A split sleeve was brazed with Fe 15% Mo- 5% Ge 4% C - 1% B over u tungsten-arc weld. We have also investigated the use of split induction coils. Split coils have an advantage in removability, as described for the split resistance heater. With a I-in. split coil, we reached brazing temperature on a "y-in.-diam tube and matching split-sleeve assembly. Unfortunately, the I-in. split coil did not couple well enough electrically with other sizes. a result, we obtained additional split coils for the other sizes of tubing, both smaller and larger, but this phase of our investigation was not completed. Mechanical Couplings J.R. DiStefano Although litde was known relative to mechanical couplings of molybdenum, we felt that such joints would be desirable because they would (1) allow relatively easy replacement of components in case of a failure and (2) allow us (o avoid making u difficult tube-to-header weld in locations where a tube passed through the end section of a feed pot To investigate resealable joints, experimental metal seal couplings for %-in.-OD tubing were obtained from Gamah Corporation?2 and Aeroquip Corporation.23 Both of these couplings used molybdenum seal rings, but a threaded nut applied the force that sealed the Gamah joint, while an external compressive force was required to seal the Aeroquip joint and then a gate or pin used to maintain the compressive force on the metal seal ring. This latter design is aimed at remote applications and is attractive because molybdenum components easily gall and a threaded nut is particularly susceptible to this problem. However, problems encountered in sealing the Acroquip joint resulted in our cr; king one of the molybdenum components, and further evaluation was discontinued. An additional experimental coupling was built at ORNL with the design shown in Fig. 29. The components of this coupling were all made of molybdenum, but the seal gasket was made of a laminated carbon product called Grafoil.2* Grafoil is a low-density graphite foil (70 I1b/ft*) that is useful as a gasket material: i compressive modulus increases from 7000 to 50,000 psi as load is applied. Both this coupling and the Gamah coupling were tested for tightness by measuring their helium leak rates before and after thermally cycling from room temperature to 650°C. Each was leak-tight (<5 X 107* atm cm?® sec™) prior to being thermally cycled ten times to 650°C. After thermal cycling, the ORNL-designed joint remained leak-tight, but the Gamah joint had a helium leak rate of 5 X 107 atm em? see ™. It was disassembled and returned to the company. and no further evaluation of this joint was made. We disassembled the Grafoil 22. Ganuh Corp., Santa Monica, Calif, 23. Aeroquip Corp. Marman Division, Los Angeles, Calif 24. Registered trademark of Union Carbide Corp. 35 ORNL-DWG T71-3181 MOLYBDENUM NN SONNNNNAN VIS IIO T IO IO IIIN 0.25in. | Fig. 29. Molybdenum mechanical coupling. 7 //\/U00»E—-Z» ORNL-4874 H. Inouye P. R. Kasten J. W. Koger A. L. Lotts R.N. Lyon R. E. MacPherson . R. Martin . McClung McCoy McDonald Moorhead: Nicholson TATDO>PPFPCOORIITRE . W .E. . E. T .E. .S.Meyer L .C. . L. . B. .Pa triarca . M. Perry . W. Rosenthal . L. Scott . F. Schaffer . C. Schaffhouser . M. Slaughter . B. Trauger .M. Weinberg . R. Weir . F. Wiffen E. L. Youngblood Leo Brewer (consultant) Walter Kohn (consultant) G. V. Smith (consultant) W. S. Williams (consultant) BABCOCK & WILCOX COMPANY, P.O. Box 1260, Lynchburg, VA 24505 B. Mong BLACK AND VEATCH, P.O. Box 8405, Kansas City, MO 64114 ‘C. B. Deering 59 60 "BRYON JACKSON PUMP, P.O. Box 2017, Los Angeles, CA 90054 G. C. Clasby CABOT CORPORATION, STELLITE DIVISION, 1020 Park Ave., Kokomo, IN 46901 T. K. Roche CONTINENTAL OIL COMPANY, Ponca City, OK 74601 J. A. Acciarri EBASCO SERVICES, INC,, 2 Rector Street, New York, NY 10006 D. R. deBoisblanc T. A. Flynn THE INTERNATIONAL NICKEL COMPANY, Huntington, WV 25720 J. M. Martin THERMO ELECTRON CORP., 9 Crane Court, Woburn, Mass. 01801 L. W. Shaheen UNION CARBIDE CORPORATION, CARBON PRODUCTS DIVISION, 12900 Snow Road, Parma, OH 44130 R. M., Bushong : USAEC, DIVISION OF REACTOR DEVELOPMENT AND TECHNOLOGY, Washington, DC 20545 David Elias ' J.E. Fox Norton Haberman C. E. Johnson T. C. Reuther S. Rosen J. M. Simmons USAEC, DIVISION OF REGULATIONS, Washington, DC 20545 A. Giambusso USAEC, RDT SITE REPRESENTATIVES, Oak Ridge National Laboratory, P.O. Box X, Oak Ridge, TN 37830 D. F.Cope : Kermit Laughon C. L. Matthews USAEC, OAK RIDGE OPERATIONS, P.O. Box E, Oak Ridge, TN 37830 Research and Technical Support Division Patent Office USAEC, Technical Information Center, P.O. Box 62, Oak Ridge, TN 37830 For distribution as shown in TID-4500 under General Reactor Technology category (25 copies — NTIS) (92) |