Figure G.7c Temperature Dependent True Stress-strain Curves Of A508 .
100 90 72 F 550 F 80 300 F True Stress (ksi) 70 700 F 60 932 F 50 1112 F 40 30 1292 F 20 1472 F 10 2732 F 0 0 0.025 0.05 0.075 0.1 0.125 0.15 0.175 0.2 0.225 0.25 Plastic Strain Figure G.7b Temperature dependent true stress-strain curves at A516 Grade 70 100 72F 90 600F 80 True Stress (ksi) 70 60 1000F 50 40 30 1400F 20 10 1800F 0 0 2 4 6 8 10 12 14 16 18 20 True Strain (%) Figure G.7c Temperature dependent true stress-strain curves of A508 Class 3 tested by ORNL G-15
90 80 75F 300F 70 True Stress (ksi) 550F 60 700F 900F 50 1100F 40 1300F 30 1652F 1500F 1832F 20 2012F 2550F 10 0 0 0.025 0.05 0.075 0.1 0.125 0.15 0.175 0.2 0.225 0.25 Plastic Strain Figure G.7d Temperature dependent true stress-strain curves of Type 316 and Type 309 100 90 75F 80 300F 550F True Stress (ksi) 70 700F 60 900F 1100F 50 1300F 40 1500F 1652F 30 1832F 2012F 20 2550F 10 0 0 0.025 0.05 0.075 0.1 0.125 0.15 0.175 0.2 0.225 0.25 Plastic Strain Figure G.7e Temperature dependent true stress-strain curves of Type 304 G-16
Table G.6 Temperature dependent creep constants for all the materials As ε& s A sσ ns ns T M ATERIAL: A508 Clas s 2 1.0000E-26 4.0000 70 2.2910E-12 6.0451 1000 3.2670E-07 4.8865 1200 3.2670E-07 4.8865 2500 M ate rial: A516-70 1.0000E-26 4.0000 70 2.5060E-13 6.3261 900 1.9920E-09 4.4071 1000 6.9010E-08 4.5039 1100 6.9010E-08 4.5039 2500 M ATERIAL: S309, S304, S316 1.0000E-26 4.0000 70 9.2650E-25 9.7800 887 4.6900E-24 9.9700 932 1.6410E-21 9.0600 977 3.9710E-19 8.2000 1022 2.7540E-18 8.2000 1067 1.7060E-17 8.2000 1112 1.1700E-16 8.1800 1157 7.2180E-16 8.1600 1202 3.4110E-14 7.4200 1247 1.3300E-12 6.7200 1292 2.0930E-11 6.2500 1337 3.2310E-10 5.7700 1382 M ATERIAL: INCO182 1.0000E-26 4.0000 70 1.0000E-26 4.0000 990 2.1478E-16 6.1709 1000 4.6025E-15 6.6426 1100 4.6025E-15 6.6426 2500 G-17
End of Butter (70F) 50. 30. 10. End of Heat-Up (1100 F) -10. -30. -50. 20. 12. End of PWHT (1100-70F) 4. -4. -12. -20. End of PWHT (4-Hour) (1100F) Figure G.8 Axial stresses during heat treat process End of Butter (70F) 50. 30. 10. End of Heat-Up (1100 F) -10. -30. -50. 50. 30. 10. -10. -30. End of PWHT (4-Hour) (1100F) End of PWHT (70F) Figure G.9 Hoop stresses during heat treat process G-18 -50.
End of Butter (70F) End of PWHT (70F) 0.05 0.05 0.04 0.04 0.03 0.03 0.02 0.01 0.01 0.005 0. Figure G.10 Equivalent plastic strains End of PWHT (70F) Consistent With Murakawa Results 0.003 0.003 0.002 0.002 0.002 0.001 0.001 0.0009 0.0006 0.0003 0. Figure G.11 Equivalent creep strains G-19
50. 40. 30. 20. 10. 0. -10. -20. A516-70 SS304 16 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 17 -30. -40. SS316 -50. Figure G.12 Residual stresses final (axial) at room temperature 22C (70 F) 50. 40. 30. 20. 10. 0. -10. -20. A516-70 SS304 16 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 17 -30. -40. SS316 -50. Figure G.13 Residual stresses final (axial) at operating temperature 291 C (556 F) G-20
50. 40. 30. 20. 10. 0. -10. -20. A516-70 SS304 16 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 17 -30. -40. SS316 -50. Figure G.14 Residual stresses final (hoop) at room temperature 22 C (70 F) 50. 40. 30. 20. 10. 0. -10. -20. A516-70 SS304 16 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 17 -30. -40. SS316 -50. Figure G.15 Residual stresses final (hoop) at operating temperature 291 C (556 F) G-21
for Inconel 182/82 weld metal and for A508 steel. It is important to recognize that the material properties of the weld material must be obtained on annealed weld samples because the weld modeling itself models the heating and cooling strain hardening explicitly. Hence, while the experimental residual stress measurements did not provide direct validation of the weld modeling, the insight that was obtained by considering all of the above processes was very important. Indeed, after all of these effects were considered, and re-analysis of the cold leg completed, the residual stresses predicted were lower than those originally predicted. However, they were still higher than the measurements. The fact that the hoop residual stresses measured at both the inside and outside surfaces are so low clearly indicates that the measurements were not accurate. Because the constraint in the weld direction (hoop direction) is high, as the weld bead cools, it shrinks and is constrained by the already cool material, producing high tensile residual stresses in all cases the present authors have seen in over twenty five years. through thickness stresses favor axial crack growth via stress corrosion cracking mechanisms if entirely driven by residual stresses. Figures G.16 (a) and (b) show comparisons of axial weld residual stresses to measurements, while Figures G.16 (c) and (d) show the corresponding hoop stress comparisons. The measurements were made at Battelle at our West Jefferson, Ohio site, where the cold leg pipe has been stored since 1988. The ‘chip removal’ or trepanning technique of Reference G.6 (and many references sited therein) was used for the measurements. The trends for the axial residual stresses comparisons (Figure G.16 (a) and (b)) are similar, but the measurements are lower than the predictions. The hoop residual stresses (Figures G.16 (c) and (d)) measurements are quite low compared with predictions. Hoop residual stress measurements in bimetallic welds have not been reported in the literature as far as can be determined. However, from prior measurements and predictions of pipe (Refs. G.6 and G.9) for same material welded pipe and many reference sited therein), hoop residual stresses are nearly always tensile and approaching yield, especially in the regions of the weld for both thick and thin pipe. The measured stresses here (Figure G.16 (c) and (d)) are actually compressive in this region. This is considered unrealistic. Despite efforts to resolve this quandary, no errors in the measurement technique could be found. Measurement of residual stresses in bimetallic welds should be pursued in the future, perhaps using the new deep hole drilling procedures developed by Professor Smith of Bristol University (Ref. G.10). Regarding the trepanning method of measuring residual stresses, it has served very well in past studies at Battelle in the late 1970’s and should be regarded as a viable method for measuring residual stresses. However, it requires a skilled and experienced technician to carefully remove the pyramid shaped chips from the pipe. Therefore, the main purpose of this analysis effort for the cold leg, to validate the VFT weld modeling procedure for bimetallic welds, was not successful. However, the results are useful and provide insight for the hot leg analysis discussed next. When the residual stress measurements were obtained, and the low values were measured, the weld modeling procedure was completely re-evaluated. The post weld heat treatment was then considered in the analysis process. The weld processes and procedures for both the cold leg and hot leg were carefully reevaluated. The material properties used for the weld analysis were carefully evaluated. In fact, a separate test program was initiated at Oak Ridge National Laboratory (ORNL) to obtain better temperature dependent material properties Figure G.17 illustrates equivalent plastic strains. Figure G.18 shows the corresponding axial, hoop, and shear plastic strains after welding. It is interesting to note that the axial plastic strains are compressive for the most part in the buttering region while the hoop plastic strains are tensile in the butter and weld. Moreover, from Figure G.18 (c), rather large values of shear strain develop in the region of the butter. While PWSCC growth is considered to be driven by tensile stress, or stress intensity factors, it may be useful to consider the role of tensile plastic strains in SCC growth in future studies. G-22
Axial Residual Stresses Along OD 60. 51. IN182 (Weld.) 316 SS IN182 (But.) 43. 34. Prediction 25. A516-70 16. Data 7.5 -1.3 -10. -8.5 -7.9 -7.2 -6.6 -5.9 -5.3 -4.6 -4. Y Coordinate System 0 Figure G.16(a) Residual stresses final (axial) at operating temperature 291 C (556 F) Axial Residual Stresses Along ID IN182 IN182 (But.) (Weld.) 75. 64. 316 SS 54. 43. 33. 22. A516-70 11. 0.6 -10. -8.5 -7.9 -7.2 -6.6 -5.9 -5.3 -4.6 -4. Y Coordinate System 0 Figure G.16(b) Residual stresses final (axial) at operating temperature 291 C (556 F) G-23
Hoop Residual Stresses Along OD IN182 (Weld.) 80. IN182 (But.) 316 SS 69. 58. Prediction 46. A516-70 35. Data 24. 13. 1.3 -10. -8.5 -7.9 -7.2 -6.6 -5.9 -5.3 -4.6 -4. Y Coordinate System 0 Figure G.16(c) Residual stresses final (hoop) at operating temperature 291 C (556 F) Hoop Residual Stresses Along ID IN-182 (But.) 80. 67. 316 SS 54. IN-182 (Weld.) 41. 28. 14. 1.3 A516-70 -12. -25. -8.5 -7.9 -7.2 -6.6 -5.9 -5.3 -4.6 -4. Y Coordinate System 0 Figure G.16(d) Residual stresses final (hoop) at operating temperature291 C (556 F) G-24
0.1 0.08 0.06 0.04 0.02 Figure G.17 Residual equivalent plastic strains in cold leg at room temperature G-25 0.
(a) Axial 0.02 0.012 0.02 0.015 0.004 -0.004 -0.012 -0.02 (b) Hoop 0.01 0.005 Figure G.18 Residual axial (a), hoop (b), and shear (c), plastic strains in cold leg at room temperature G-26 0. -0.005
(c) Shear 0.0382 0.0227 0.00733 -0.00808 -0.0235 -0.0389 Figure G.18 Residual axial (a), hoop (b), and shear (c), plastic strains in cold leg at room temperature As with the cold leg, the PWHT was modeled via creep analysis applied to the buttered weld residual stress state. G.6 RESULTS HOT LEG ANALYSIS This section presents the axis-symmetric results for the hot leg analysis. The results of this analysis were used to calculate stress intensity factors so that PWSCC predictions could be made (Section G.7). The buttered nozzle along with the stainless steel pipe was then pre-heated again and weld metal was deposited from the inside of the pipe to a depth of 18 mm (0.7 inch). After this amount of weld metal was deposited, the weld was rejected. In preparation for weld grind out (of the original 18 mm (0.7 inch) of weld metal), a weld bridge was deposited. The weld was then ground out from the pipe inside. There were then two weld sequences that were considered in the analysis since it was uncertain whether the weld repair was deposited from the bridge first on the outside of the pipe, followed by the inner weld or vice-versa. Both were modeled to examine the effect of the repair sequencing on the final residual stress state. G.6.1 Hot Leg Computational Weld Model The geometry of the hot leg bimetallic weld joint is illustrated in Figures G.19 and G.20. Note that the hot leg analysis for the V. C. Summer plant is similar to the cold leg analysis discussed in Section G.5 except that the geometry is different (smaller diameter and thickness), and the materials are different for the nozzle (compare with Figure G.3). Please follow Figure G.21 for the description of the weld modeling process. The modeling sequence is quite complicated since the V. C. Summer hot leg in question had several repairs made to it. The sequence of the repairs was not entirely known, so two repair sequences were considered. Figure G.21 illustrates the welding sequence modeled. The nozzle was first preheated and a buttering layer deposited. The nozzle was then post weld heat treated (PWHT). All of the processes listed in Figure G.21 were considered in the model. Figure G.22 further illustrates the modeling process pictorially. Figure G.22 (a) shows the original buttering model results. Figure G.22 (b) shows the PWHT modeling process. Figure G.22 (c) shows the completion of the weld prior to weld rejection, building of a weld bridge, and then G-27
Figure G.19 Geometry of V.C. Summer bi-metallic weld joint 2.33’’ INCO182 A508 class 2 Lump-Pass Weld SS304 SS309 cladding Figure G.20 Axis-symmetric model of V.C. Summer bimetallic weld joint G-28
Preheat Butter PWHT Preheat Weld to 0.7 inch from inside Build a weld bridge Ground out the weld Weld Inside to the bridge Weld outside from the bridge Weld outside from the bridge Weld inside to the bridge Figure G.21 Welding process simulated on hot leg G-29
Cladding model (c) Rejected weld (a) A508 class 2 SS304 A508 class 2 0.7’’ Welding to 0.7 inch 20” Bridge 10” A508 class 2 (b) 1100. 688. SS304 Welding bridge and grounding out rejected weld 276. (d) Figure G.22 Cladding (butter) and rejected weld model state in the pipe as material is removed and the precise modeling of the chipping process is not necessary. grinding out the original weld so that only the bridge remains. Finally, the weld was either deposited from the inside surface first, then the outside surface, or vice-versa. Both were modeled since the precise repair weld deposition sequence was not known. There is another source of grinding that occurs after the entire weld repair is completed. Reference G.12 provides summary of the metallurgical investigation of the cracking in the V. C. Summer plant. As discussed on page 9 of Reference G.12, ‘The surface appeared highly irregular with evidence of significant surface grinding and machining distress marks’. Photographs and micrographs clearly show small ‘scratch marks’ along the inner pipe surface at the weld location (Figures 10 and 19 from Reference G.12). This grinding was presumably performed in order to remove the weld repair ‘bulging’ at the pipe inside surface in order to permit more uniform flow through the pipe. The grinding will redistribute residual stresses (as discussed above regarding the grinding before weld repair). However, because the material ground out is a small volume, it is not included in the analysis (i.e., the final geometry, already ground, is modeled). An important point regarding the analysis steps is in order at this point. Referring to Figures G.21 to G.23, grinding of weld material prior to deposition of the final weld passes was included in the modeling process. For instance, from Figures G.22 (c) to G.22 (d), material was ground out to make a bridge of weld metal prior to deposition of the weld repair layers. This grinding process simply consisted of removing material ‘computationally’. By this we mean that the material was removed mathematically by eliminating the stiffness of these elements and therefore redistributing the residual stresses. The actual grinding process, whereby a rigid (or nearly rigid) sharp tool impacted the weld region and material was ‘chipped away’ was not considered. This is a complex modeling problem, but it can be done. However, the main effect of the grinding is to redistribute the residual stress G-30
Buttering Thermal analysis Buttering: 1. Preheat to 150F 2. Butter Buttering Stress analysis PWHT 1. Heat to 1100F 2. Hold for 3 hours 3. Cool Output stress Welding Creep analysis cooling Combine plastic strain plastic strain component Welding thermal analysis Welding stress analysis Loading plastic strain component effective plastic strain Output stress Apply 1. Operation temperature 2. inside pressure 3. Axial force 4. Moment Output stress K K calculation Map results to 2D coarse model Revolve 2D coarse model to 3D model with results transfer Figure G.23 Finite element analysis process flow and should be avoided. It may be a useful exercise to include the actual modeling of grinding in such a model as an additional step in order to further prove this hypothesis. Moreover, since grinding is common practice, and is apparently not specifically considered by the code bodies, such a series of ‘grinding’ model studies may be of use in setting standards in future construction and aging repair. In essence, if additional material was added to the inside surface weld, and then removed, the final residual stress state should be very similar to that from ignoring it except for very localized grinding stresses. The very local residual stresses from the grinding process are ignored in this case. Typically, additional residual stresses from grinding are considered to be important for only a very short depth into the thickness of the pipe. Certainly, after the crack grows a very short distance into the pipe thickness, these local residual stresses are eliminated and the weld induced residual stresses dominate for most of the PWSCC growth life. G.6.2 Hot Leg Computational Weld Model – Buttering and PWHT Results Figure G.23 illustrates the entire analysis procedure for the hot leg. As seen, after the weld modeling is completed, results were mapped to a coarser two-dimensional model. The coarser two-dimensional model was then revolved to a three-dimensional model in preparation of the three dimensional PWSCC crack growth analysis. Service loads were then applied and However, the geometric effects of the scratches are expected to be very important. These scratches should be considered as crack initiation sites for PWSCC, fatigue, or any possible cracking mode. Such grinding, which produces scratches, may serve as PWSCC initiation sites G-31
Figure G.7b Temperature dependent true stress-strain curves at A516 Grade 70 Figure G.7c Temperature dependent true stress-strain curves of A508 Class 3 tested by ORNL 0 10 20 30 40 50 60 70 80 90 100 . steel. It is important to recognize that the mater-ial properties of the weld material must be
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