& e f P rsartcey—of OAK RIDGE NATIONAL LABORATORY operated by UNION CARBIDE CORPORATION % NUCLEAR DIVISION for the _ U.S. ATOMIC ENERGY COMMISSION (}4 c ORNL- TM- 2483 i6 PE men BT T MAR 2 e et MMM i s TR - i L e s P PREIRRADIATION AND POSTIRRADIATION MECHANICAL PROPERTIES OF HASTELLOY N WELDS H. E. McCoy D. A. Canonico —_———————— LEGAL NOTICE ——————— — This report was prepored os an gccount of Government sponsored work. Majther the United States, nor the Commission, nor any person acting on behalf of the Commission: A. Makes any warranty or representation, expressed or implisd, with respect to the occuracy, completeness, or usefulness of the information ceontoined in this report, or that the use of any information, opparctus, method, or process disclosed in this report may not infringe privotely owned rights; or B. Assumas any liabilities with respect 1o the use of, or for domages resulting from the use of any information, apporatus, methed, or process disclosed in this report. As used in the ohove, "'person acting on behalf of the Commission” includes any employe= or contractor of the Commission, ot employee of such controctor, to the extent that such employes or contractor of the Commission, or employee of such contractor prepores, disseminates, or pravides occess to, any information pursuont to his employment or contract with the Commissien, or his employment with sueh contractor, ORNL-TM-2483 Contract No. W-7405-eng-26 METALS AND CERAMICS DIVISION PREIRRADIATION AND POSTIRRADIATION MECHANICAL PROPERTIES OF HASTELLOY N WELDS H. E. McCoy and D. A. Canonico Submitted to the Welding Journal without the Appendix. MARCH 1969 OAK RIDGE NATIONAL LABORATORY Oak Ridge, Tennessee operated by UNION CARBIDE CORPORATTION for the U.S. ATOMIC ENERGY COMMISSION A 0 BN A e et [P e ¥Ry R R 1 iii CONTENTS Abstract . Introduction . Experimental Details . Experimental Results . Tensile Properties Creep-Rupture Properties Postweld Heat Treatments Metallography . . « « « « + « « & SUMMAYY « « o« e e e e e e e e e Acknowledgments References . Appendix . PREIRRADIATION AND POSTIRRADIATION MECHANICAL PROPERTIES OF HASTELLOY N WELDS H. E. McCoy and D. A. Canonico ABSTRACT Welds were made by the TIG process in several heats of Hastelloy N. The mechanical properties of transverse weld samples and the base metal were compared in tensile tests over the range of 75 to 1600°F and in creep tests at 1200°F. The as-fabricated welds exhibited lower fracture strains than the base metal under all test conditions, but the properties of the welds were improved markedly by post- weld heat treatments. The postirradiation tensile and creep properties of the welds and base metal at elevated temperatures were about the same, although the properties were widely different before irradiation. INTRODUCTION Hastelloy N is a trade name given the solid-solution-strengthened nickel-base alloy developed at the Oak Ridge National Laboratory specif- ically for use in molten fluoride salts up to 1500°F (ref. 1). The alloy was originally designated INOR-8 and has a nominal composition of Ni—17% Mo—7% Cr—5% Fe. This material is the primary metallic structural material in the Molten Salt Reactor Experiment (MSRE) which achieved criticality on June 1, 1965, at Oak Ridge, Tennesse.? This alloy has been used for numerous other applications, and we anticipate that a slightly modified composition will be used in a future molten salt breeder reactor experiment. The weldability of this material has received considerable atten- 3,4 tion in our program. We found it necessary to establish a weldability test to determine whether specific heats of material could be joined satisfactorily. Through this screening process, it was possible to select heats for use in fabricating the MSRE components. During the past few years we have found that Hastelloy N as well as other nickel- and iron-base alloys are subject to a type of neutron irradiation damage that decreases the creep-rupture strength and fracture 556 Thisg damage 1is rather general for strain at elevated temperatures. all austenitic iron- and nickel-base alloys and is attributed to the production of helium within the material due to the interaction of a neutron of thermal energy with '“B to produce “Li and “He (refs. 7-14). The property changes due to the helium must be evaluated for base metal and welds. In the present paper, we shall show how the mechanical properties of transverse welds differ from those of the base metal and what types of annealing treatments can be used to improve the properties of welds. We shall also show how neutron irradiation alters the mechanical proper- ties of the welds and the base metal. The mechanical properties measured in this study were tensile properties over the temperature range of 75 to 1600°F and creep properties at 1200°F. ok b e R A R PR B R R e vt SR et ke R e ok e G e 1 e b R AR 20, Rttt L e L EXPERIMENTAL DETAILS Several heats of Hastelloy N were used in this study and their chemical compositions and other pertinent details are given in Table 1. Heats 65-552 and 2477 were wvacuum melted; all others were air melted. All base metal was annealed for 1 hr at 2150°F prior to testing unless otherwise specified. Seven weldments were involved in this study and the details are given in Table 2. The welds were all highly restrained and made by the manual tungsten-arc welding process. A standard welding procedure was used on all the welds, although some deviation was necessary in welds 7 and & because of the type of filler metal being used. The joint configu- ration and pass sequence are shown in Fig. 1. The primary working direction in the plates was perpendicular to the weld axis. We used small mechanical property samples throughout the study, since this geometry was required for the irradiation experiments (Fig. 1). The samples were cut perpendicular to the weld axis and parallel to the stringers. Three layers of specimens were cut from all the welds except No. 3; we could find no systematic variation in properties from top to bottom. However, the gage portions of the transverse weld specimens did contain different amounts of weld metal since the joint tapered (Fig. 1). The tensile tests were run in an Instron Universal testing machine 1 (10,000-1b capacity) at strain rates of 0.05 or 0.002 min~ The strain measurements were taken from the crosshead travel. A similar machine was located in a hot cell for testing the irradiated samples. The labo- ratory creep-rupture tests were run in standard lever-arm creep machines. The strain measurements were taken from the pull-rod displacement after Table 1. Chemical Analysis of Test Materials Heat . Composition Designa- Meltl?g etiin Fractice Cr Fe Mo c si Co W (w;n%j v P S AL Ti Cu B (pgm) N 5065 Air 7.2 3.9 16.5 0.065 0.e0 0.08 0.04 0.55 0,22 0.004 0.007 0.01 0.01 0.01 10-37 16 110 5067 Air 7.4 4.0 17.3 0.060 0.43 0.08 0.06 0.50 0.30 0.013 0.007 0.01L 0.01 0.015 10-24 12 100 5101 Air 6.9 3.9 16.4 0.05 0.63 0.05 0.44 0.34 0,001 0.009 - 0.022 — 3.5 5055 Air 7.9 3.8 16.2 0.06 0.6l 0.03 0.69 0.21 0.006 0.008 0.06 0.02 50 2477 Vacuum 7.05 4.25 16.3 0.057 0.015 0.14 0.47 0.04 0.008 0.003 0.055 0.10 0.10 8 5 7 65-552 Vacuum 6.89 4.06 16.2 0.045 0.16 0.050 0.006 0.45 < 0.0005 0.002 0.006 0.25 0.5 1.3 9 aCom.bined amount of eluminum and titanium. Table 2. Identification of Welds Made in This Study Weld Thickness Nurber Base Metal Filler Metal of.Weld {in.) 1 5065, 5067 5101 1 3 2477 2477 1/2 4 5065, 5067 5101 1 5 5065, 5067 5055 1 6 5065 65-552 11/8 7 5065 5055 + Al,052 11/8 5065 5055 + WC? 11/8 8piller material was 1/8 in. wire of heat 5055 plasma sprayed with 0.002 in. of indicated material. the load was applied. The postirradiation creep-rupture tests were run in 1ever-arm creep machines located in the ORNL hot cells. The strain was measured by an extensometer with rods attached to the upper and lower specimen grips. The relative movement of these two rods was measured by a linear differential transformer. All tests were run in an air environ- ment. We used a standard equilibrating time of 1/2 hr for each test prior to loading. The irradiations were conducted in two facilities: the ETR, Idaho Falls, Idaho, and the ORR, Oak Ridge, Tennessee. A core facility was used in the ETR where the thermal and fast (> 1 Mev) fluxes were each 3.2 X 10%* neutrons em™? sec™!. The fluence obtained there was 5 x 10°° neutrons/cmz. The ETR experiments were uninstrumented and the design temperatures were either less than 300°F or 1112 + 180°F. A core facility was also used in the ORR where the thermal flux was ORNL-DWG 68-12878R DIMENSIONS IN INCHES < FILLET WELDED TO STRONG BACK TO PRODUCE HIGH RESTRAINT WELD 29 1, +O L - Ya 16 MAX JOINT DESIGN AND WELDING SEQUENCE +0.0000 in. Q.0008 in. Q.0000 in. 0.0005 in. +Q.0000C in, Q.Q00% in. ! e > @® @® ! nE 3 @ N o 63 o s 53 | i £ ! oX - — — Q= _ t ~(E S \ | | 00008 \— 01875 R(TYP) ~—-3in. 1425 in. *'77/5 in. MECHANICAL PROPERTIES SPECIMEN. Fig. 1. The INOR-8 High-Restraint Weldability Test Specimen Used to Provide Samples for the Mechanical Properties Study. 2.5 x 10'% neutrons cm™? sec™l. The fluences obtained were 8.5 x 102° neutrons/em® thermal and 7.0 x 102° neutrons/cm? fast (> 1 Mev). The temperature in this facility was 110°F. We could not observe any effect of fluence over the small range involved here, so the data are presented without reference to a particular experiment. Irradiation temperature was important in some cases and will be indicated. EXPERIMENTAL RESULTS Tensile Properties¥ The variation of the fracture strain in a tensile test with test temperature is shown in Fig. 2 for both transverse welds and base metal. The behavior of this particular heat of base metal is typical for Hastelloy N. The decrease in fracture strain above 1100°F is due to ORNL-DWG 6812875 % l ] T T T T HEAT 5065 WELD1 WELD 4 AS PROCESSED ° A & AFTER IRRADIATION . . 80 POST WELD ANNEAL OF 8hr AT 1600°F UNIRRADIATED 0 IRRADIATED - -~ o UNIRRADIATED BASE METALX v ¥ o [=] - -~ I o I L o “ e - - n o - ,/ Vo L |1 “NIRRADIATED BASE METAL \ \\/! \ o \ \ FRACTURE STRAIN (%) o O o o . \ A‘—-—- 9 \ 1 20 ——] UNIRRADIATED WEL[;?\\\ / 0 "'\\ / o S o 200 400 600 800 1000 1200 1400 1600 1800 TEST TEMPERATURE (°F) Fig. 2. Ductility of Hastelloy N Welds and Base Metal in Tensile Tests at a Strain Rate of 0.05 min™?!. *See Tables A-1 and A-2, Appendix, for the tabulated tensile data. the transition from transgranular to intergranular fracture. The recovery above 1400°F is associated with increasing grain boundary mobility. The fracture strain of the welds is much lower. Most of the welds fractured in the weld metal, but several samples tested at 75 and 392°F did fail in the base metal. The fracture strain after irradiation is also shown in Fig. 2. Since these samples were irradiated at less than 300°F, there is some displacement damage in the base metal and the fracture strain is lower at test temperatures of 75 and 392°F. At test temperatures of 1200 and 1600°F, where intergranular fracture predominates, the fracture strain of the base metal is reduced dramatically. The postirradiation fracture ductility of the irradiated welds is also shown in Fig. 2. The strain at fracture of the welds is reduced slightly by irradiation, but the changes are much less than those observed for the base metal. Thus, although the ductility of the welds is much lower than that of the base metal before irradiation, the ductilities of welds and base metal at elevated temperatures are quite similar after irradiation. The yield strengths of typical welds and base metal are compared in Fig. 3. The yieldlstrengths of the transverse weld samples are con- sistently higher than those of the base metal. This same observation was also made by Gilliland and Venard.’ After a postweld anneal of & hr at 1600°F, the welded and base metal samples have very similar yield strengths. Irradiation increases the yield strength of the base metal and welded samples at low temperatures. This damage anneals out as the test temperature is increased, and the welds and base metal show similar ORNL - DWG 6B-12876 120 l . | ! ‘ HEAT 5065 WELD1 WELD 4 s L,R,fsf‘ED'GE%EL AS PROCESSED o 0 a 0o | ¥ AFTER IRRADIATION o s R N A POST WELD ANNEAL 7 \\\\\\ OF 8hr AT 1600°F o 80 —— % UNIRRADIATED o e 8 a ‘-~::::§\ IRRADIATED | | oo = . —— —~— |, UNIRRADIATED WELDS ; — o \ T 3 40 O S —O u UNIRRADIATED BASE METAL — & N 20 0 0 200 400 600 BOO {000 {200 1400 1600 1800 TEST TEMPERATURE (°F) Fig. 3. Yield Strength of Hastelloy N Welds and Base Metal in Ten- sile Tests at a Strain Rate of 0.05 min~1. recovery. Thus, above 1200°F the yield strengths of base metal and welds annealed for & hr at 1600°F are equivalent and are unaffected by irradiation. Figure 4 shows that the ultimate tensile strength is about the same for welded samples and for base metal. This property is not affected significantly by postweld heat treatment at 1600°F. Irradiation causes a slight (approx 10%) increase in the ultimate tensile strength at the iower test temperatures and a decrease at test temperatures of 1200°F and greater. This reduction at higher temperatures is due to the reduced ability of the material to deform plastically after irradiation (i.e., fracture occurs before the stress increases to the higher wvalues noted for unirradiated materials). Thus, the effects of welding on the strength parameters measured by standard tensile tests are relatively small and the most significant factor is the reduction in the fracture strain. The fracture strains observed for the welds involved in this study are summarized in Fig. 5. FRACTURE STRAIN (%) 140 120 100 80 60 40 ULTIMATE TENSILE STRESS (1000 psi) 20 Fig. 4. 10 ORNL-DWG 68-12877 o g 0 -._--'-'-1 e ; b\o\q\ ‘\ g \ 5 & - HEAT 5065 WELD { WELD 4 A AS PROCESSED o o A AFTER IRRADIATION . a [ POST WELD ANNEAL OF 8 hr AT 1600°F UNIRRADIATED o I IRRADIATED . 0 200 400 600 800 1000 1200 1400 1600 1800 TEST TEMPERATURE {°F) Ultimate Tensile Strength of Hastelloy N Welds and Base Metal in Tensile Tests at a Strain Rate of 0.05 min~?!. ORNL -DWG 68- 12888 80 T B As wELDED - — = . 72 7?,, ANNEALED 8hr AT 1600°F o ] o 8 | AS WELDED - IRRADIATED R e - T TTTT 7T T ANNEALED 8hr AT 1600°F - IRRADIATED 13 T T 64 Fu—O—F — e = —0 - E T ®© I . ek e A G S e R Tk IS L e o e 11 Welds 1, 4, and 5 involve air-melted materials with relatively high boron levels. Heat 2477, a vacuum-melted heat containing & ppm B, was used as the base and filler metal in weld 3. Weld 6 utilized filler metal from heat ©65-552, a vacuum-melted heat containing 1 ppm B, and base metal from heat 5065, an air-melted heat. Welds 7 and 8 utilized air-melted base and filler materials. The filler rod for weld 7 was plasma-spray-coated with Al,03 in an effort to reduce the interdendritic spacing in the cast weld metal and to provide additional sites for helium collection. However, most of the Al,03 floated on the weld and had to be removed mechanically before making the next pass; the aluminum content of the deposited weld metal was only 0.06% compared with 0.84% for a typical cross section from the filler metal before deposition. Weld 8 involved a filler material coated with WC. The deposited weld metal contained 2% W and 0.1% C, so the WC coating dissolved in the melt altered the properties of the weld appreciably. The fracture strains shown in Fig. 5 are compared for 1. base metal given a standard solution anneal of 1 hr at 2150°F, 2. base metal given the 1 hr at 2150°F treatment followed by an anneal of 8 hr at 1600°F, 3. transverse weld samples in the as-welded condition, and 4. transverse weld samples annealed for & hr at 1600°F. These properties are considered for unirradiated and irradiated condi- tions where data are available. At a test temperature of 75°F the following points are illustrated in Fig. 5. 12 1. The fracture strains of all welds are lower than those of the base metal. 2. Welds 1, 4, and 5 all involved air-melted alloys and exhibited the lowest fracture strains. 3. Welds 3 and 6 involved vacuum-melted weld metal and have the best properties. 4. Welds 7 and 8, which involved air-melted alloys with the filler metal modified with Al,05; and WC, exhibited intermediate properties. 5. Postweld heat treatments generally improved the properties. 6. Irradiation decreased the fracture strain of all welds and base metal. The welds had lower postirradiation ductilities than the base metals, but the welds were affected less by irradiation. 7. The postirradiation ductilities of welds did not depend appreciably on whether they had received a postweld heat treatment prior to irradiation. 8. The postirradiation properties of all welds except No. 8 were equivalent. Figure 5 also illustrates several important points at a tensile test temperature of 1200°F. 1. The fracture strains of all welds were lower than those of the base metal. 2. The as-welded ductilities of all welds except No. 8 were about equivalent, ranging from 12 to 18%. Weld & had 25% strain at fracture under these conditions. i3 3. A postweld heat treatment of 8 hr at 1600°F markedly improved the fracture strain, with values of 22 to 30% being observed. Weld & again was better. 4. After irradiation, all welds were less ductile than the base metal. There were some differences in the ductilities of the welds, but no consistent trends are apparent. 5. Postweld annealing for 8 hr at 1600°F had little effect on the postirradiation properties. At 1600°F, Fig. 5 shows the following trends. 1. The fracture strains were drastically lower for welds than for base metal. Welds 3 and 8 had properties superior to those of the other welds. DPostweld heat treating had a beneficial effect on the ductility. 2. Irradiation reduced the fracture strain to about 2% independent of whether the test sample was base metal or a transverse weld. All of the postirradiation results presented thus far have been for materials irradiated at less than 300°F. We have irradiated several samples of these heats and others at 1200 to 1400°F and find that there are at least two significant differences in the results obtained. The displacement damage anneals at these irradiation temperatures and the tensile properties up to about 1000°F are the same for irradiated and unirradiated materials. At higher test temperatures the fracture strain is reduced even Turther if the irradiation temperature is in the range of 1200 to 1400°F. For example, heat 5065 was found to have a fracture strain of 11.3% (irradiated at 1200°F) at 1200°F compared with the value of 22.2% (irradiated at less than 300°F) shown in Fig. 5. Weld 7 was observed to have fracture strains at 1200°F of 7.2 and 10.8% after 14 irradiation at 1200°F and less than 300°F, respectively. Thus, the trends shown in Fig. 5 for samples irradiated at 300°F seem to hold for irradiation temperatures in the 1200 to 1400°F range, although the actual fracture strains may be lower for the higher irradiation temperature. Creep-Rupture Properties* The stress-rupture properties at 1200°F of several heats of base metal are shown in Fig. 6. The data are described reasonably well by a single line, although both air- and vacuum-melted materials are involved. The line for the base materials is shown in Fig. 7 where a comparison is made between the stress-rupture properties of the base metal and the various welds involved in this study. All of the welds had lower creep-rupture strength in the as-welded condition than the base metal. Weld 3, which involved only vacuum-melted materials, had significantly higher strength than the other welds. Welds 7 and 8, which involved A1,0; and WC additions, also showed notably better performance than welds 1, 4, and 6. The rupture lives of the welds approached those of the base metal as the stress level was reduced; this is likely due to thermal recovery of the weld metal during the long time at 1200°F. The strength was improved greatly by a postweld anneal of 8 hr at 1600°F. The fracture strains of the welds and several heats of base metal are compared in Fig. 8. The fracture strain of the base metal was about 25% for short rupture lives of a few hours and decreased to values of 10 to 15% when the rupture life was 1000 hr. There is no appreciable difference between air- and vacuum-melted alloys. The welds had fracture *See Tables A-3 and A-4, Appendix, for the tabulated creep-rupture data. 15 ORNL-CWG 68-12879 70 TR Il ° 5065 \ a 5067 60 o\} o 2477 b\\“ 50 N 0 \ N o 40 Q o N l&‘ 30 D\\ o @ N \\ 20 N 10 0 o 2 5 10 2 5 10% 2 5 100 2 5 10° RUPTURE TIME (hr) Fig. 6. Stress-Rupture Properties of Hastelloy N Base Metal at 1200°F. ORNL-DWG 68-12880 70 N n BASE METAL WELD 3 60 ~ 1 \ N N \\\ N\ 0\\2 Y’ \\\ i 50 — ] /,)N\ \\ ™ \\ 1: 0N NG RR | = N 2 \ NG TN \\ ) 8 a0 &{ Sh, ; , Q N AN ‘ z \\ ™ \\A\ 1 h A\ \\ \\\N ! ) WELDS 1,4,6 \3 NG -\\\ | ~ 30 O ;\\ 2 * NG TN NLWELD 8 NG TN NG TN O WELD1 N N AWELD3 \fiL______ BN 20 O WELD 4 vVWELD®SB ) 0 WELD 7 j 10 O 8 A Fa 6 v A A . T G 4 T F & o O O ’:]JP VC‘ S |G <> 0 0 2 %* v J 0 107! 10° 101 102 103 104 RUPTURE TIME (hr) Fig. 8. Variation of Fracture Strain with Rupture Life for Welds and Base Metal Tested at 1200°F. 17 strains of only about 3% for rupture lives up to a few hundred hours. Samples having rupture lives over 1000 hr showed some improvement with strains in the range of 3 to 5%. Weld 3 had much superior fracture straing of about 6% for all test conditions. All welds and even the base metal exhibited much higher fracture strains after an anneal of 8 hr at 16C0°F. As shown in Fig. 9, irradiation produces some rather dramatic changes in the stress-rupture properties. The lines shown in Fig. 9 were obtained from Fig. 6 for the unirradiated base metal, Fig. & for the minimum strength of unirradiated welds, and from Ref. 6 for the strength of irradiated base metal. Several points are included for base metals irradiated under exactly the same conditions as the welds. The properties of heat 2477 (low boron, vacuum-melted) are superior to ORNL —DWG 68 — {2882 70 N L "N\ UNIRRADIATED BASE METAL ” \l\ NWMINIMUM STRENGTH | NNIRRADIATED N 50 WELDS \\F\ \\ \ - ™ a \H"N\ \ O 40 T~ r S o '\ \\ Q@ \\. \\ — N - 'On.\ho *fl\ a NooI~g 00 | (@ R w30 < — \““4\1 » O HEAT 5065 ] N\ & HEAT 2477 Y N A M~ \\\ 20 WELD 3 _ | L] Iy AVERAGE IRRADIATED T ® WELD 4 BASE METAL v WELD 6 l 10 ¢ WELD 7 < WELD 8 ’ * ANNEALED 8hr AT 1600 °F o L1 1L oLl 0% 2 5 0! 2 5 t0° 2 5 10> 2 5 0% RUPTURE TIME (hr) Fig. 9. Influence of Irradiation on the Stress-Rupture Properties of Hastelloy N Welds and Base Metal at 1200°F. All welds in the as- welded condition unless otherwise indicated. 18 those of heat 5065 for the low irradiation temperature of less than 300°F, but further work has shown that these materials have comparable properties when irradiated at 1200°F or higher.® The rupture lives of the welds seem quite comparable with those of the base metals and all welds seem to have about the same properties. A single specimen was given a postweld heat treatment of 8 hr at 1600°F and had superior post- irradiation strength, but this observation should be repeated before con- cluding that this anneal is effective. Moteff and Smith!® also found that the postirradiation creep properties of Hastelloy N base metal and welds were equivalent at 1200°F. They also found that a postweld anneal of 4 hr at 1600°F did not improve the postirradiation properties of welds. The fracture strain is shown in Fig. 10 as a function of rupture life. The fracture strain of the base metal had a low value of about 0.5% for rupture lives of only a few hours and increased to aboutVB% for rupture lives of 1000 hr. The welds fall into this same pattern, but the strains were consistently slightly lower. Again the single test on a sample having a postweld anneal exhibited better properties. We also compared the minimum creep rates for these same materials to obtain a measure of strength. The creep properties of several heats of base metal in the irradiated and unirradiated conditions are shown in Fig. 11. Most of the data are contained within the band in Fig. 11, with heat 5065 favoring the right side and heat 2477 the left side. This same band is transferred to Fig. 12, where the data for the welds are shown for comparison. The data do not deviate significantly from the band obtained for the base metal. Thus, the creep rate of Hastelloy N is not affected appreciably by welding or by irradiation. 19 ORNL-DWG 68—12883 ] HEAT 5065 HEAT 2477 WELD 3 WELD 4 WELD 6 WELD 7 + WELD 8 * ANNEALED 8nr AT {600°F &« 4 B P DO STRAIN AT FRACTURE (%) n o a® X % o 2 s w0 2 5 102 2 5 03 2 s 0% RUPTURE TIME {hr) Fig. 10. Influence of Irradiation on the Fracture Strain of Hastelloy N Welds and Base Metal in Creep at 1200°F. All welds in the as-welded condition unless otherwise indicated. ORNL—-DWG 68-12884 70 AT / /1 60 AL Y | A a / A | /] Y = / //fi 8 40 { "... A _9; ,/ . / fi 30 .:‘_I’-.Qn . o HEAT 5065 & .,/ F/ 4 HEAT 5067 " ° / o HEAT 2477 20 3 OPEN SYMBOLS — UNIRRADIATED 1 A CLOSED SYMBOLS - IRRADIATED 10 0 10°% 2 5 073 2 5 02 2 5 ' 2 5 10° 2 5 10 MINIMUM CREEP RATE (%/hr} Fig. 11. Influence of Irradiation on the Creep Rate of Hastelloy N Base Metal at 1200°F. 20 ORNL-DWG 68-12885 ST T ’ / g O WELD 1 s/ 4 A WELD 3 s . 60} o WELD 4 // . v WELD 6 p 4 0 WELD 7 v, A 1400°F _ 2 s o 1200°F & M E 12 : > /] / o // / & A / 8 A /? /J / LA L/ 4 = ,.—/ /1 et i 41 T ? T 5 109 2 s w0 2 5 102 2 5 103 TIME AT TEMPERATURE (hr) Fig. 15. Influence of Postweld Heat Treatments on the Reduction in Area of Weld 1 Tested at 1200°F and 650°C. 23 obtained by postweld anneals at 1400 and 16CO°F. The unusual behavior due to annealing for 50 hr at 1600°F again appears for two of the samples. The reduction in area is not affected appreciably by annealing st 1200°F, indicating that the improved elongation at fracture (Fig. 14) is due primarily to the formation of cracks during testing. We examined the influence of whether the postweld annealing environ- ment was argon or hydrogen at 16C0°F, since Gilliland and Venard® had reported better properties for samples annealed in hydrogen than for those annealed in argon. Comparative points are shown in Figs. 13, 14, and 15 for these annealing environments, and no appreciable differences are apparent. Metallography Many of the samples were examined metallographically and a few typical microstructures will be presented. Figure 16 shows the fusion zone of weld 4. The structure of the base metal is characterized by stringers of Mg(C~type precipitates and a relatively fine grain size. The weld metal contains a fine cellular structure within coarse grains. The fusion line is quite sharp and the transition of some of the large M¢C precipitates to a lamellar phase is apparent. The exact identifica- tion of the lamellar phase has eluded us for some time, but our evidence indicates that it is a Mo-Ni-Si compound.!® Microprobe studies have shown that the precipitates are enriched in silicon and that the dark etching portions of the dendritic cast structure are also higher in silicon. 24 A o =R - - - - 3 3 s 5 L Re - : A= R TR R R - FONr. ] 2 . : 2 - = el ¥, 1 5 SR > L SRR TR - -~ N oy - > j T ;‘\}‘-:1—: - \ "-'} " : < S St . oy e R AN v .fi\_- - ¢ T N L 2 R it B fflqf:’ :‘.P'& Ok \l.:-t \ - ~ ) ; - : S SRt A SEvte et N —] Rt \.““in & ¥ ARG A A O R B, 8 o = % Ny SR NN ‘ s = = a e e e, s : DR R S S, y A :'.: - = Ji : I".\ """'-_._;’:1 l'Lh ““T‘.-vh\ H'::‘ Il\"hq. e g S e R USSR \ \'-,_\‘ ¢ 'F'I ‘i\ N < -_\“‘- NG T s A ) - l\r ,‘).--\_"‘\‘:‘ 'l---J L’; F : A % A . . £ L “‘xfk"\_l ‘iw 1 . 3 E‘_\ - . F‘?.*\ .'5_-\'. ; A \___L - Sy . = c T A LTk e Jitt\ Wi ™ _ TN e R S W BSRN TR T . 4 « sl v R NI B K oy ; HEETES = N . g TR AT N ' . < iy . _4_".-%.‘__‘ ; v ; _'.. *_"' j? 7 L o 7 - -4 2 Sl war PRI TR £ X% ¢ w\ ol ol ; ; =t - " L I o '\;J‘t—fif,‘.']'.- :“"-‘. f - "l \ . \' £ ’ . %, % ‘ (a) » f b 13 00 Jr\ 3 \:J,' X \; ",-'Ir N ¥ t e g < - - Wty Y-63005 - B i y oy 1 » & v o - \ e - ~—g™ o b a .l "‘. - h - . \_h - H - ' é. 'w \l"“ - - . - b - . . S L - ,l'"\‘__‘ - - - 3 § o " W i g vy . - e i3 - . p /J # I‘1.. — - “* 2 ?‘h 3 & - h-'k_ - r-fl?-'— o A “ ;"' - n *a ‘*"-14 ‘ Y Y - - - g Sk Y < " 3 - ¥ YR e .- (b) b — Fig. 16. Photomicrographs of the Fusion Zone of Weld 4. (a) Base metal on left and weld metal on right. 100x. (b) Transformation of stringer precipitates to lamellar product at the fusion line. 500x. Etchant: glyceria regia. This same general description applies to the fusion zones of the other welds except 3 and 8. The base metals used in weld 3 were vacuum- melted and the stringers are minimal. The bulk silicon content was low (0.015%) and the precipitates were primarily of the MosC type and were found to dissolve easily. Thus, not much of the lamellar transformation 25 product was present at the fusion line. Weld 8 involved the filler mate- rial that was coated with WC, and the structure of the deposited weld metal is much finer (Fig. 17). A typical microstructure of a creep-rupture fracture in Hastelloy N base metal is shown in Fig. 18. This sample was tested at 40,000 psi and 1200°F and failed after 312 hr with 16.6% strain. There is considerable grain boundary cracking and the fracture is predominantly intergranular. Typical microstructures of a tested sample from weld 1 are shown in Fig. 19. This sample was tested at 40,000 psi and 1200°F in the as- welded condition; failure occurred in 18.7 hr with 3.1% strain and only 0.77% reduction in area. The fracture occurred in the weld metal and followed the grain boundaries; however, little grain boundary cracking adjacent to the fracture is evident. Postweld heat treatments generally improved the tensile and creep properties. Figure 20 shows the fracture of a sample from weld 1 that was annealed for 2 hr at 1600°F and tested at 40,000 psi and 1200°F. This sample failed in 200 hr with 9.4% strain and 6.4% reduction in area. The failure still occurred in the weld metal; however, intergranular cracks did form throughout the weld zone, indicating that deformation occurred generally throughout this zone. SUMMARY This study has shown that the mechanical properties of welds in Hastelloy N are generally inferior to those of the base metal. The strain at fracture is the property affected most severely, although appreciable reductions in the high-temperature creep-rupture strength e -~ : o 20s - AT N N s i Ntk y b} Y-662 g- X _\::r:::‘ - -‘--:_‘fl\n\-'-l\g F‘ 'Qtfi*- fiI"“% fik}r“:’\r‘fi | ". 1'11 Rk o . '-\"h. gt gt e { ki n‘._. o T 2 e e e TR ? . -1\,-.-“-11. ”_‘1.\“ 3 Pt .\ ] SN SCTEANGR AR T b L & T:';;"' "\7\1{" Ffi Y .}(\F"fl‘:'“x.-\\# Py """ 0 N ~ :fi.dfl?_ E:E o - ;\‘ o \ h o 'li aj ‘u;\:',lu‘ ANy [\‘H-'Air" \. I. : N = Vol e ol YW L | AL Ry N < C 1 INEEL o :fi"“‘n' At E.':_‘;'ffi‘ A .: ¥Lli'-l.1;{t I’l“. .!1‘_ in' ;.l;‘ wfi :‘ ‘:,‘" y -'. .\ W b [ I -JL “”“”; "'\}1\ \\ 4 s 1_ & b ‘l i F " l\1I|'ul - \:i: Y‘L_‘_?‘!\Q!""fi ?- v X : Ly \ 1‘ I;l. J ‘;‘xl"-l'.": )‘;l._ .I.-'?.'I'-‘l ‘I t‘.‘ .-{ ‘.\lh P v . \ e NS g _'.. = % A . % e AN MEIRL B, A - r.¥ .‘.i‘fi..‘}i;\%‘:“ I g \ g 3 1IN ety &, g W A -‘.'].'il o fi o W { j-" % "‘,J' ™, iy A ¥ e N i ¥t he ‘ol : Vg i 0 - ' o fol 1 ::I b 3 | - - % 'I * = c f i .n. i { oy iy W iy . ' L | 6 . I w1 g s " : I v 3 . = L i I - ¢ I','_ . y . wis . i AT A Y o ¥ § ‘l., i A . LT 2 ¥ ¢ ' | R Nt TN % = ¥ fx. N * o el 1 A ' #E I I—| |'i.; 1 ol I,{ “I'l B A 'y \ g . LT Al i | Vi 4 g Fig. 17. Photomicrograph of the Fusion Line of Weld 8. Weld metal in the upper left and base metal in the lower right. 100X. Etchant: glyceria regia. ‘ : e ,.f | : o T ’ A b g Y-6L729 Y~ ¥ J,li 1 \l . j / E ?\ E (.\. o { X QYT R S ¥ Py Y, PSRk e 7 CEF Ly el e A - ( ) ; I { P NS :4 = K \ y Eal? L i M- } ] e 5 : { -,.1/ "I,_f‘ : ;}:w e ) - ot Py il J: 1s Yoy r L E‘r=J" 4 > 3 l / Y ol ‘-( i - " A e i / >’ 1 : . ‘ \ \ f s = ~.r = LS A |_. , s i \ i j . ' o .r : BN =l T Y G 7 - i I i 3 " - e )( - Fig. 18. Photomicrograph of the Fracture of a Hastelloy N (Heat 5065) Sample Tested in Creep at 40,000 psi and 1200°F Following a Pretest Anneal of 1 hr at 2150°F. Failed in 312 hr with 16.6% strain. 100x. Etchant: glyceria regia. - - - - t . - - - - = § i & 44 N L ¥ v -l v - ' - A 1 L Pl ? . i - 4 i 8 g o 1 g f ol 5 = \ L - N e . i # . ’ ‘ 2 } . iy I i \ 5 b ' & - - . Ly -~ - £ - J . S _II ¥ . g - 1 ¥ & 2 1 - - ; ; > Bl A f F gk ¥ = L - i . 1 ry & o r r ™ e i ¢ = = P . - Y-66857 Fig. 19. Photomicrographs of Sample from Weld 1 Tested at 40,000 psi and 1200°F in the As-Welded Condition. Failed in 18.7 hr with 3.1% strain. 100x. (a) Fracture in the weld metal. Etchant: glyceria regia. (b) An unetched photomicrograph of a cracked area away from the fracture. (o) L Fig. 20. Photomicrographs of Tested Creep Sample from Weld 1, Sample given a postweld anneal of 2 hr at 1600°F in argon and tested at 40,000 psi and 1200°F. Failed in 200 hr with 9.4% strain. 100x. (a) Fracture. Etchant: glyceria regia. (b) An unetched photomicro- graph of deformed area in the weld metal. 29 also occurred. The properties of welds can be improved by suitable post- weld heat treatments to almost equal those of the base metal. Several welds were studied which involved air- and vacuum-melted materials. The weld involving vacuum-melted base and filler metal (weld 3) had superior creep-rupture properties in the as-welded condition. Weld 8 involved air-melted materials and the filler metal was coated with WC. The properties of this weld were generally superior to those of the other welds, but the fracture strain under creep conditions was not improved. Neutron irradiation caused a decrease in the fracture strain of both welds and base metal. The postirradiation fracture strains in tensile tests of the welds and the base metal became closer as the test temperature was increased; they were identical at 1600°F. Although post- weld heat treatments improved the fracture strain in preirradiation tensile tests, the postirradiation fracture strains of welds seemed independent of postweld anneal. The postirradiation creep-rupture proper- ties at 1200°F were about the same for welds and base metal, even though the preirradiation properties differed greatly. Although some of the irradiated welds showed improved performance in tensile tests, they all had about the same creep-rupture properties. ACKNOWLEDGMENTS The welds were made under the supervision of T. R. Housley and the mechanical property tests were run by B. C. Williams, E. Bolling, N. O. Pleasant, J. T. Feltner, and V. G. Lane. H. R. Tinch was respon- sible for the metallographic work. We are also indebted to 30 G. M. Slaughter and J. R. Weir for their interest in this work and for their careful review of the manuscript. The drawings were prepared by the Graphic Arts Department and the manuscript was prepared by the Metals and Ceramics Division Reports Office. REFERENCES 1. W. D. Manly et al., "Metallurgical Problems in Molten Fluoride Systems," Progr. Nucl. Energy Ser. IV 2, 164~179 (1960). 2. H. G. MacPherson, "Molten Salt Reactor Shows Most Promise to Congerve Nuclear Fuels, Part 2," Power Eng. Zé(Z), 56-58 (1967). 3. R. G. Gilliland and J. T. Venard, "Elevated Temperature Mechanical Properties of Welds in Ni-Mo—Cr—Fe Alloy," Welding J. (N.Y.) 45(3), 103-5-110-s (1966). 4. R. G. Gilliland and G. M. Slaughter, private communication. °. W. R. Martin and J. R. Weir, "Postirradiation Creep and Stress Rupture of Hastelloy N," Nucl. Appl. 3, 167 (1967). 6. H. E. McCoy, "Variation of the Mechanical Properties of Irradiated Hastelloy N with Strain Rate," submitted to Journal of Nuclear Materials. 7. D. R. Harries, "Neutron Irradiation Embrittlement of Austenitic Stainless Steels and Nickel Base Alloys," J. Brit. Nucl. Energy Soc. 5, 74 (1966). 8. G. H. Broomfield, D. R. Harries, and A. C. Roberts, "Neutron Irradia- tion Effects in Austenitic Stainless Steels and a Nimonic Alloy," J. Iron Steel Inst. (London) 203, 502 (1965). 1 e g~ gD A B A R e S N ast L, L L 31 9. F. C. Robertshaw et al., "Neutron Irradiation Effects in A-286 Hastelloy Y and René 41 Alloys," Spec. Tech. Publ. No. 341, p. 372, American Society for Testing and Materials, Philadelphia, Pa., 1963. 10. N. E. Hinkle, "Effect of Neutron Bombardment on Stress-Rupture Properties of Some Structural Alloys," Spec. Tech. Publ. No. 341, p. 344, American Society for Testing and Materials, Philadelphia, Pa., 1963. 11. P.C.L. Pfeil afid D. R. Harries, "Effects of Irradiation in Austenitic Steels and Other High-Temperature Alloys," p. 202 in Flow and Fracture of Metals and Alloys in Nuclear Environments Spec. Tech. Publ. No. 380, American Society for Testing and Materials, Philadelphia, Pa., 1965. 12. J. T. Venard and J. R. Weir, "In-Reactor Stress-Rupture Properties of a 20 Cr—25 Ni Columbium-Stabilized Stainless Steel,” p. 269 in Flow and Fracture of Metals and Alloys in Nuclear Environments Spec. Tech. Publ. No. 380, American Society for Testing and Materials, Philadelphia, Pa., 1965. 13. P.C.L. Pfeil, P. J. Barton, and D. R. Arkell, "Effects of Irradiation on the Elevated Temperature Mechanical Propertles of Austenitic Steels,” Trans. Am. Nucl. Soc. 8, 120 (1965). 14. P.R.B. Higgins and A. C. Roberts, "Reduction in Ductility of Austenitic Stainless Steel after Irradiation,” Nature 206, 1249 (1965). 15. J. Moteff and J. P. Smith, Sixth Annual Report of High Temperature Materials Program, Part A, March 31, 1967, GEMP-475A, p. 185 (1967). 32 16. R. E. Gehlbach and H. E. McCoy, "Phase Instability in Hastelloy N," paper presented at the International Symposium on Structural Stability in Superalloys, Seven Springs, Pa., Sept. 4—6, 1968; to be published in the proceedings. APPENDIX 35 Table A-1l. Tensile Properties of Welds in the Unirradiated Condition - Test . - Reduction Specimen Weld Postweld S;:::n Tenpera- Stress, psi Elongation, % in Area Number Number Anneal . =1 ture Yield Ultimate Uniform Total (%) ( min ) ( o F) 13 1 None 0.05 75 g1,100 113,100 22.0 22.3 37.7 14 1 None 0.05 390 59,100 81,200 11.3 12.3 24 4 15 1 None 0.05 1200 50,000 64,800 6.0 7.8 13.1 16 1 None 0.05 1200 40,800 42,000 3.0 5.8 6.9 17 1 None 0.05 1600 23,200 23,400 1.8 21.8 25.5 18 1 a 0.05 1600 41,600 42,100 3.9 19.2 18.6 3000 1 None 0.05 75 76,300 111,900 27.1 28.3 29.5 3001 1 None 0.05 1200 59,300 79,800 10.2 11.7 26.5 3005 3 None 0.05 75 71,700 122,800 35.2 37.2 42,4 3006 3 None 0.05 1200 53,200 77,300 12.6 14.4 23.6 10285 3 None 0.05 1600 41,300 44,900 T 13.8 14.8 10287 3 a 0.05 75 61,000 123,700 44,0 46.1 41.02 10286 3 a 0.05 1200 43,700 82,400 21.4 22.5 26.93 10300 3 a 0.05 1600 40,900 44,200 5.3 21.7 18.15 3007 4 None 0.05 75 74,300 108,900 24.9 25.9 47.8 3008 4 None 0.05 1200 55,300 72,300 13.5 15.3 31.7 10288 4 None 0.05 1600 40,900 44,700 3.1 5.9 7.66 10289 4 a 0.05 75 57,700 107,500 31.4 31.9 38.4 10290 A a 0.05 1200 39,300 76,900 29.6 30.4 35.4 5 None 0.05 75 73,300 119,800 33.0 34.2 45.1 5 None 0.05 1200 51,100 70,100 14.2 15.9 42.8 10299 5 None 0.05 1600 43,300 45,400 1.7 11.7 18.0 10296 5 a 0.05 75 63,800 119,500 32.7 33.1 41.1 10297 5 a 0.05 1200 49,200 91,300 25.6 25.9 38.8 10295 5 a 0.05 1600 39,000 39,400 1.4 36.1 62.6 1144 6 None 0.05 75 68,200 120,900 38.7 40,7 38.3 1145 6 None 0.05 1200 49,800 81,300 16.6 18.6 20.9 1147 6 None 0.05 1600 39,000 44,400 2.2 2.9 3.04 1140 6 a 0.05 75 63,100 119,100 37.0 38.5 47.3 1141 6 a 0.05 1200 44,400 85,400 29.0 29.9 36.8 1142 6 a 0.05 1600 39,900 44,500 4.5 13.9 16.7 1164 7 None 0.05 75 71,900 125,700 32.3 33.8 45.0 1165 7 None 0.05 1200 52,000 79,100 13.7 14 .4 11.9 777 7 None 0.002 46,900 65,700 7.3 8.6 18.7 1166 7 None 0.05 1600 41,500 45,100 2.0 3.9 7.8 10284 7 a 0.05 75 64,100 119,400 33.8 34.5 34.9 10283 7 a 0.05 1200 45,100 86,100 24.0 24.6 37.6 10295 7 a 0.05 1600 39,000 39,400 1.4 36.1 62.6 3011 8 None 0.05 75 72,300 124,900 35.8 38.8 49.4 g None 0.05 1200 54,000 91,300 22.9 24.9 29.1 778 8 None 0.002 1200 39,000 66,600 11.8 16.7 28.2 10280 8 None 0.05 1600 38,400 44,600 4.5 13.0 8.1 10282 8 a 0.05 75 58,900 120,100 33.9 34.0 28.4 10281 8 a 0.05 1200 41,700 91,200 35.7 37.0 37.7 10294 8 a 0.05 1600 39,800 43,800 4.3 53.5 6.4 29 hr at 1600°F. Table A-2. Tensile Properties of Welds in the Irradiated Condition Test Temper- ) ] . Specimen Weld Postweld Experiment S;::;n ature Stress, psi Elongation, % R?iuZEQZn Number Number Anneal Number (min-1) (°F) Yield Ultimate Uniform Total (%) 469 3 None ETR-41-30 0.05 1200 57,600 70,800 5.9 6.8 18.6 470 3 None ETR-41-30 0.002 1200 56,400 62,200 3.0 3.6 12.6 699 4 None ORR-149 0.05 75 109,900 120,500 16.8 19.2 48.7 700 4 None ORR-149 0.05 392 92,900 113,300 18.9 19.1 25.3 701 4 None ORR-~149 0.05 1200 53,300 69, 300 8.6 8.7 12.0 703 4 None ORR-149 0.002 1200 51,900 63,400 4.7 5.2 12.8 702 4 None ORR-149 0.05 1600 41,200 41,600 1.4 2.0 1.3 704 A None ORR=-149 0.002 1600 26,000 26,000 0.9 2.0 9.2 708 4 a ORR-149 0.05 75 104,500 122,800 17.2 18.7 14.5 709 4 a ORR-149 0.05 392 82,700 107,600 21.0 22 .2 27.3 710 4 a ORR~149 0.05 1200 42,200 63,300 9.8 10.3 21.9 712 4 a ORR-149 0.002 1200 45,400 58,300 5.6 6.6 28.1 711 4 a ORR-149 0.05 1600 38,700 39,600 1.4 2.1 1.8 713 4 a ORR-149 0.002 1600 25,300 25,300 1.0 2.0 5.9 457 4 None ETR-41-30 0.05 1200 57,200 69,800 5.9 7.3 20.1 458 4 None ETR-41-30 0.002 1200 57,900 64,900 2.5 3.3 11.3 461 4 a ETR-41-30 0.05 1200 44,600 70,400 16.2 16.6 32.6 462 4 a ETR-41-30 0.002 1200 43,200 59,700 8.9 11.0 15.6 714 6 None ORR-149 0.05 75 111,000 130,700 22.6 25.4 47.0 715 6 None ORR-149 0.05 392 88,800 112,100 20.4 22.4 27.4 716 6 None ORR-149 0.05 1200 45,100 72,000 14.2 14.6 13.2 718 6 None ORR-149 0.002 1200 51,700 63,300 6.1 7.0 30.4 717 6 None ORR-149 0.05 1600 38,900 40,500 1.4 2.5 3.2 719 6 None ORR-149 0.002 1600 24,900 24,900 0.8 1.4 6.0 723 6 a ORR-149 0.05 75 106,500 127,100 21.1 22."7 33.4 724 6 a ORR-149 0.05 392 83,900 111,200 25.0 25.9 34.9 725 6 a ORR- 149 0.05 1200 44,700 60,300 8.8 9.5 6.1 9¢ Table A-2. (continued) Test . . . Stress, psi Elongation, % uctio Specimen Weld Postweld Experiment S;:i:n Temper- Yield U;_:') . — g ,t ] Rii A;c':a.n Number Number Anneal Number (min- atlgre 1€ imate niform Tota (%) (°F) 727 6 a ORR-149 0.002 1200 43,500 53,600 4.6 5.4 9.2 726 6 a ORR-149 0.05 1600 36,400 40,100 2.0 5.3 11.9 728 6 a ORR-149 0.002 1600 18,600 18,900 0.9 1.1 5.1 529 7 None ORR-149 0.05 75 112,400 135,100 18.4 21.4 25.7 530 7 None ORR-149 0.05 392 89,700 114,000 18.8 20.1 34.0 531 7 None ORR-149 0.05 1200 45,900 75,000 10.6 10.8 5.6 533 7 None ORR-149 0.002 1200 53,400 65,200 5.0 6.4 5.6 532 7 None ORR-149 0.05 1600 40,300 41,300 1.5 1.5 5.6 534 7 None ORR-149 0.002 1600 25,100 25,100 1.0 1.6 5,9 538 7 a ORR-149 0.05 75 107,700 122,900 14.6 16.7 54.2 539 7 a ORR-149 0.05 392 88,100 110,500 19.1 20.3 31.8 540 7 a ORR-149 0.05 1200 44,300 66,400 10.2 15,2 7.4 542 7 a ORR-149 0.002 1200 45,100 56,600 4.2 4.6 10.4 541 7 a ORR-149 0.05 1600 39,500 40,200 1.3 2.4 10.5 543 7 a ORR-149 0.002 1600 24,800 24,800 1.0 1.4 2.4 1162 7 None ETR-41-31 0.05 1200 50,000 65,600 7.2 7.2 1163 7 None ETR-41-31 0.002 1200 48,300 56,300 3.4 3.5 729 8 None ORR-149 0.05 75 110,300 137,100 25.0 29.0 48.91 730 8 None ORR-149 0.05 392 86,700 121,200 30.0 33.8 47.7 731 8 None ORR~149 0.05 1200 45,900 75,000 14.6 14.9 16.7 733 8 None ORR-149 0.002 1200 48,200 64,100 6.6 6.8 11.3 732 8 None ORR=149 0.05 1562 38,700 41,400 1.7 2.0 2.7 734 8 None ORR- 149 0.002 1562 27,500 27,500 1.0 1.3 8.9 738 8 a ORR~149 0.05 75 108,900 134,800 27.5 30.5 11.7 739 8 8 ORR-149 0.05 392 88,400 117,900 25.4 26 .4 31.9 740 8 a ORR-149 0.05 1200 41,800 70,500 13.5 13.8 14.5 742 8 a ORR-149 0.002 1200 42,700 61,400 9.0 9.2 9.2 741 8 a ORR~149 0.05 1562 2.8 3.0 20.4 743 8 a ORR- 149 0.002 1562 27,900 28,000 1.1 1.6 6.0 &2 nhr at 1600°F. LE 38 Table A-3. Creep-Rupture Properties of Welds at 1200°F in the Unirradiated Condition Minimum Sample Test Weld Postweld Stress Rupgure Creep F;act?re R?ductlon Number Number Number Anneal (psi) Life Rate Strain in Area (nr) (%) (%) (%/br) 35 399¢ 1 None 55,000 4.6 0.15 3.10 2.71 1 3780 1 None 40,000 18.7 0.013 13.1 0.77 33 5476 1 None 20,000 3864.6 0.0013 4.7 5.1 3 4080 1 None 20,000 1215.6 0.001 9.4 0.32 4éy 5583 1 None 20,000 3184.3 0.00080 3.6 1.9 2 3779 1 None 30,000 80.9 0.005 4.7 0.79 31 3996 1 a 40,000 308.8 0.029 19.4 13.2 30 5202 1 a 40,000 551.9 0.028 28.0 9.4 3345 5889 3 None 55,000 26.5 0.11 6.0 6.4 3345 5890 3 None 47,000 77.9 0.046 6.2 1.8 3346 5891 3 None 35,000 468.9 0.0072 7.4 1.8 4016 3 None 40,000 273.7 0.0163 7.8 g.2 5110 3 None 30,000 1228.0 0.0028 5.2 3.6 3395 5893 4 None 47,000 6.2 0.024 3.0 1.5 2 3997 4 None 40,000 25.2 0.0265 3.1 1.6 3 4013 4 None 30,000 80.2 0.0098 4.7 0 5894 4 None 21,500 495.9 0.0012 1.7 0.7 7060 5140 6 None 55,000 2.8 0.584 6.3 6.4 5881 6 None 47,000 6.3 0.0388 2.0 0.8 7061 5141 6 None 40,000 14.5 0.0326 3.2 3.2 1158 5882 6 None 30,000 158.2 0.0046 1.9 0.8 7127 5462 6 a 40,000 423.7 0.028 17.0 4.5 7071 5195 7 None 55,000 7.3 0.11 3.1 7.1 1168 5883 7 None 47,000 27.9 0.0406 3.6 1.0 7069 5192 7 None 40,000 47.8 0.016 3.1 0.8 7129 5464 7 None 32,400 179.0 0.070 2.5 0.6 7164 5312 7 None 30,000 194.8 0.0050 2.9 2.4 7135 5472 7 None 21,500 2173.2 0.0030 3.4 2.1 7126 5464 7 a 40,000 372.0 0.030 18.3 3.2 7073 5221 g None 55,000 129.1 0.080 4.4 4.8 3340 5884 8 None 47,000 26.5 0.048 3.0 0.8 7070 5194 g None 40,000 82.4 0.014 3.1 1.6 7125 5460 8 None 30,000 520.9 0.0035 3.5 4.0 7190 5576 g8 None 21,500 3172.2 0.0012 5.2 4.0 7128 5463 8 a 40,000 692.9 0.026 29.9 23.1 %2 nr at 1600°F. Table A-4. Creep-Rupture Properties of Welds at 1200°F in the Irradiated Condition Rupture Minlmum Fract Reducti Sample Test Weld Postweld Experiment Stress L?f Creep gtc gre ? uirlon Number Number Number Anneal Number (psi) (1 S Rate raln in Area hr) ey () (%) 471 R-90 3 None ETR-41-30 32,400 175.5 0.0064 1.59 706 R-36"7 4 None ORR-149 39,800 7.6 0,037 0.38 3.5 705 R-363 4 None ORR- 149 32,400 47.8 0.0037 0.20 4,3 459 R-83 4 None ETR-41-30 32,400 111.3 0.0031 0.59 463 R-71 4 a ETR-41-30 32,400 585.2 0.0056 3.52 720 R~ 107 6 None ORR-149 35,000 9.6 0.020 0.37 721 R-370 6 None ORR-149 30,000 51.4 0.0045 0.39 12.7 535 R-37"7 7 None ORR-149 39,800 17.7 0.0085 0.29 536 R-116 7 None ORR-149 32,400 45,1 0.0084 0.38 735 R-108 8 None ORR- 149 35,000 3.8 0.0605 0.37 736 R-365 g None ORR-149 30,000 210.8 0.0032 0.87 4.5 6¢ %% hr at 1600°F. Table A-5. 1Influence of Postweld Heat Treatment on the Creep Properties of Weld 1 at 1200°F and 40,000 psi Postweld Treatment Rupture Minimum Fracture Reduction Sample Test : . Creep . . Numb er Number Tempfrature Time Environment Life Rate Strain in Area (°F) (hr) (hr) (%) (%) (%/hr) 1 3780 18.7 0.0125 3.1 0.77 22 3979 1200 100 Ar 20.3 0.0265 3.1 0.78 24 3958 1200 300 Ar 25.3 0.0557 9.4 2.4 27 5028 1200 1000 Ar 79.7 0.027 6.3 1.6 26 5112 1200 1000 Ar 39.8 0.027 9.4 1.6 42 5111 1400 2 Ar 46.6 0.014 4.7 2.4 21 3984 1400 8 Ar 97.3 0.020 4.7 3.6 19 3976 1400 20 Ar 77.6 0.032 6.3 5.1 32 3972 1400 50 Ar 320.9 0.023 12.5 7.9 34 4024 1400 200 Ar 296.6 0.030 15.0 17.2 45 5315 1400 1000 Ar 434.3 0.024 15.6 14.6 3996 1600 0.5 Ar 54,5 0.023 3.1 4.0 3989 1600 2 Ho 200.0 0.023 9.4 6.4 4 3793 1600 2 Ar 193.9 0.029 9.4 8.6 7 3817 1600 8 Ar 416.8 0.033 21.9 11.0 31 3966 16C0 8 Ho 308.8 0.029 19.4 13.2 30 5202 1600 8 Hp 551.9 0.028 28.0 9.5 37 3995 1600 50 Ar 151.7 0.032 6.3 7.3 38 4065 1600 50 Ar 193.6 0.046 12.5 9.4 43 5249 1600 50 Ar 266.3 0.033 15.6 21.3 0% 1-3. 4=5. 6—15. 16. 17. 18. 19. 20. 21. 22. 23. 2. 25. 26. 27 28. 29. 30. 31. 32. 33. 34. 35. 36. 37. 38. 39. 40. 41, 42. 43. by, 45, 46. 47. 48. 49-53, 54. 55. 56. 57. 58. 59. 60. 6l.. 41 INTERNAL DISTRIBUTION Central Research Library ORNL — Y-12 Technical Library Document Reference Section Laboratory Records Department Laboratory Records, ORNL RC ORNL G. 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