ML20133A609
| ML20133A609 | |
| Person / Time | |
|---|---|
| Issue date: | 12/23/1996 |
| From: | Kuo T, Samson Lee, Wichman K NRC (Affiliation Not Assigned) |
| To: | |
| Shared Package | |
| ML20133A607 | List: |
| References | |
| NUDOCS 9612310181 | |
| Download: ML20133A609 (18) | |
Text
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FLAW EVALUATION OF THERMALLY AGED CAST STAINLESS STEEL IN LIGHT-WATER REACTOR APPLICATIONS 1
S. Lee, P. T. Kuo, and K. Wichman Office of Nuclear Reactor Regulation, U. S. Nuclear Regulatory Commission, l
Washington, District of Columbia 20555, USA j
\\
- 0. Chopra i
Argonne National Laboratory, Argonne, Illinois 60439, USA ABSTRACT Cast stainless steel may be used in the fabrication of the primary loop piping, fittings, valve bodies, and pump casings in light-water reactors. However, this material is subject to embrittlement due to thermal aging at the reactor temperature, that is, 290 C (550 F).
The Argonne National Laboratory (Ahl) recently completed a
)
research program and the results indicate that the lower-bound fracture toughness of thermally aged cast stainless steel is similar to that of submerged arc welds (SAWS). Thus, the U. S.
Nuclear Regulatory Commi:,1on (NRC) staff has accepted the use of SAW flaw evaluation procedures in IWB-3640 of Section XI of the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code to evaluate flaws in thermally aged cast stainless steel for a license renewal evaluation. Alternatively, utilities may estimate component-specific fracture toughness of thermally aged cast stainless steel using procedures developed at ANL for a case-by-case flaw evaluation.
THERMAL AGING 0F CAST STAINLESS STEEL Cast stainless steel materials may be used in the fabrication of reactor coolant pressure boundary components, such as piping, fittings, valve bodies, i
and pump casings, in light-water reactors. However, cast stainless steel is l
susceptible to thermal aging at the reactor operating temperature, that is, l
about 290*C (550 F). Thermai aging of cast stainless steel results in embrittlement, that is, a decrease in the ductility, impact strength, and fracture toughness, of the material. Depending on the material composition, the Charpy impact energy of a cast stainless steel component could decrease to 1
9612310181 961223 PDR ORG NRRA PDR
l j
a fraction of its original value after exposure to reactor temperatures during service.
The cast stainless steel components in use in light-water reactors in the
~
United States are generally fabricated to the American Society of Mechanical Engineers (ASME) Specification SA-351 Grade CF-3, CF-3A, CF-8, CF-8A, and CF-8M.
The cast stainless steel component is either statically cast or centrifugally cast.
The susceptibility of the material to thermal aging increases with increasing ferrite contents.
Grade CF-8M contains a larger amount of molybdenum and shows increased susceptibility to thermal aging, compared with the other cast stainless steel grades listed above.
Because of the potential for embrittlement, an evaluation of the flaw tolerance of cast stainless steel components in reactor applications should consider the effects of thermal aging to ensure the structural integrity of the reactor coolant pressure boundary. However, there is no generally accepted bounding level of fracture toughness for thermally aged cast stainless steel which may be used in flaw evaluations, i
i Section XI of ASME Boiler and Pressure Vessel Code,' which has been endorsed by the USNRC, provides little guidance as to how flaws in cast stainless steel components should be evaluated to determine acceptability for continued service.
Section XI provides flaw evaluation procedures for austenitic piping 3
in IWB-3640 if the provisions in IWB-3641 are satisfied.
Subparagraph IWB-3641(c) states:
"For cast stainless steel materials, adequate toughness for the pipe to reach limit load after aging shall be demonstrated." However,Section XI does not contain specific procedures to demonstrate reaching limit load.
l RECENT DEVELOPMENT j
In 1994, the Argonne National Laboratory (ANL) completed an extensive research program in assessing the extent of thermal aging of cast stainless steel materials.
The ANL research program measured mechanical properties of cast stainless steel materials after they had been heated in controlled ovens for long periods of time. ANL compiled a database, both from data within ANL and i
from international sources, of about 85 compositions of cast stainless steel
+
exposed to a temperature range of 290-400 C (550-750 F) for up to 58,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br /> (6.5 years).
From this database, ANL developed corr ations for estimating the extent of thermal aging of cast stainless steel.g*
Recently, a U. S. utility owners group submitted a topical report for review by the U. S. Nuclear Regulatory Commission (NRC) addressing the aging managemept of their reactor coolant system piping components for license renewal.
Power reactors are licensed to operate in the United States for a 40-year duration and there are specific license renewal requirements for extending the operating license for a 20-year period. The focus of license renewal review is to ensure that the effects of aging will be adequately managed during the period of extended operation.
Although the main scope of the owners group topical report addresses stainless steel clad carbon steel 2
4
primary piping, it also includes cast stainless steel valve bodies.
The potential loss of fracture toughness of cast stainless steel valve bodies ~due to thermal aging was identified as an aging effect to be managed for the period of extended operation.
Based on the ANL results, the USNRC staff has accepted flaw evaluation procedures for thermally aged cast stainless steel components as part of this topical report review.
The purpose of this paper is to discuss the approved flaw evaluation procedures.
FRACTURE TOUGHNESS ESTIMATION PROCEDURES ANL developed fracture toughness estimation procedures g* r,' thermally aged cast stainless steel based on correlating experimental data.
The estimation procedures are described in a flow chart with three distinct options, called
" lower bound," " saturation," and " service time." The selection of different options would depend on the information available for the specific material and the special interest of the utility.
The first and simplest option provides a lower-bound estimate.2,3 No information from the certified material test report (CMTR) of the cast stainless steel component is necessary for estimating the lower-bound fracture toughness. However, if the ferrite content of the component is measured in the field using a ferrite gauge or is calculated based on the chemistry composition, this option provides a lower-bound estimate specifically for materials containing a similar range of ferrite.
The other two oQ' ig' s are more elaborate and require the use of information from the CMTRs.
The second option provides an estimate for the
" saturation" toughness, that is, the minimum fracture toughness remaining after thermal exposure for an infinitely long period of time.
However, in some instances, " saturation" may occur at a time after the cast stainless steel component is no longer in service.
Then, it may be beneficial to estimate the toughness of the component at some time prior to " saturation,"
and this is addressed by the third option fcr a specified time in service.
ANL compiled an extensive database of material properties for thermally aged cast stainless steel samples. ANL developed the fracture toughness estimation procedures by correlating data in the database conservatively as discussed below. After developing the correlations, ANL validated the estin>4 tion procedures by comparing the estimated fracture toughness with the measured value for several cast stainless steel components removed from actual plant service.
It is observed that the ANL procedures produced conservative estimatp's that were about 30 to 50 percent less than actual measured values. ' Such conservative estimate-are expected based en the discussion in the next section.
CONSERVATISM AND LIMITATIONS OF ESTIMATION PROCEDURES ANL built the following conservatisms into the procedures to ensure that the 3
l
l estimation procedures would produce conservative results:
1.
Two best-fit correlations of the room temperature Charpy impact energy at saturation with different combinations of chemistry factors and ferrite content were developed from the ANL database.
Then, for a i
specific material composition, the lower value from these two i
correlations is used in the estimation procedures.
Figure I shows that theestimatedroomtemperatureCharpyimpactenergyatsaturption,which is part of the estimation procedures is within 5 to 10 J/cm (3 to 6 ft-lb) of the measured impact energy value for most of the samples and is usually conservative, that is, under-estimates the actual value.
Figure I shows on more than 20 J/cm]y three estimates out of 101 being over-estimated by (12 ft-lb).
I 2.
The coefficient "C" in the "J-R curve" estimate is the best-fit correlation of "C" with the room temperature Charpy impact energy at j
saturation from the ANL database, lowered by one standard deviation.
(The form of the "J-R curve" is shown in the next section.) This results in a 84 percent probability that the estimated value of "C" will be conservative, that is, the measured value of "C" will be larger than the estimated value.
3.
The coefficient "n" in the "J-R curve" estimate is the lower-bound correlation of "n" with the room temperature Charpy impact energy at saturation from the ANL database.
4.
The " lower bound" estimate is developed assuming the material has the lowest room temperature Charpy impact energy at saturation in the ANL 2
database. Specifically, these values are 30, 25, and 20 J/cm (18, 15, and 12 ft-lb) for CF-3 and CF-3A, CF-8 and CF-8A, and CF-8M, respectively.
The ANL estimation procedures have the following limitations / uncertainties:
1.
Niobium increases the susceptibility of the cast stainless steel to thermal aging. The ANL estimation procedures do not consider niobium.
However, in most cases, cast stainless steel components in reactors in the United States do not contain niobium, the evaluation of niobium-containing cast stainless steel material is outside the scope of this paper.
2.
Because the maximum ferrite content of statically cast CF-8M in the ANL database is 25 percent, there may be uncertainties in the application of the ANL procedures to estimate fracture toughness for statically cast CF-8M material with a ferrite content exceeding 25 percent.
However, it is unusual for cast stainless steel components in reactor applications in the United States to have such a high ferrite content.
3.
There are uncertainties in estimating the fracture toughness of a material because the fracture toughness depends on the actual microstructure of the material. The ANL database may not encompass all metallurgical factors that can arise from differences in production heat 4
treatment or casting processes, such as ferrite morphology and casting grain structure. Although microstructural studies indicate that the mechanism of thermal embrittlement for samples aged in a laboratory is the same as that observed in material aged inservice, there may still be uncertainties with the effects of service conditions on thermal aging.
Recent data' show that some thermally aged statically cast CF-8M samples may have fracture toughness properties which would be over-estimated by the ANL procedures. However, the samples from these recent data (1) may contain niobium, (2) have a ferrite content exceeding 25 percent, and (3) may have grain structures not included in the ANL database. Thus, the ANL procedures should provide a conservative estimate within the above limitations.
LOWER-BOUND ESTIMATES The " lower bound" option of the ANL procedures provides a lower-bound fracture toughness estimate thefollowingform:g-r thermally aged cast stainless steel as a "J-R curve" of Jd = C ( Aa )"
where J is the." Deformation J" per the American Society for o
Testing and Ma})erials (ASTM) Specifications E813-85 and Ell 52-87 (kJ/m Aa is the crack extension (mm)
C and n are parameters listed in Table 1 The parameters C and n are listed in Table 1 according to the ferrite content of the cast stainless steel.
If the ferrite content information is not available, the C and n parameters corresponding to a ferrite content exceeding 15 percent should be used.
Lower-bound "J-R curves" have been plotted in the ANL reports.2'3 From these plots, the following general observations can be made:
1.
Grade CF-8M has the smallest "J " value for a given crack extension d
among the grades considered due to thermal aging.
The value of "J " at d
a given crack extension for grade CF-8M is generally about half of that of the other grades.
2.
Grades CF-3, CF-3A, CF-8, and CF-8A show a similar extent of thermal aging. The value of "J " at a given crack extension for grades CF-8 and d
CF-BA is generally less than 10 percent lower than that for grades CF-3 and CF-3A.
3.
Statically cast stainless steel is more susceptible to thermal aging than centrifugally cast material. The value of "J " at a given crack extension for statically cast stainless steel is g,enerally about 20 5
-- - -. _ ~
i percent lower than that for the corresponding centrifugally cast material with a similar ferrite content.
4.
Although grade CF-8M shows a similar extent of thermal aging when the "J-R curve" is measured at both room temperature and at reactor operating temperature, the other grades considered show a decreased value of "J ' at a given crack extension when the "J-R curve" is g
measured at reactor operating temperature.
The value of "J " at a ghen g
crack extension for grades CF-3, CF-3A, CF-8, and 0-8A is generally about 20 percent lower when measured at reactor operating temperature than when measured at room temperature.
i ASSESSMENT OF FRACTURE TOUGHNESS i
i Thermal aging reduces the fracture toughness of cast stainless steel components in the reactor coolant pressure boundary. The potential loss of fracture toughness of cast stainless steel valve bodies due to thermal aging was an identified aging effect in the owners group topical report.
The USNRC staff reviewed the ANL results and ASME Section XI evaluation procedures in order to assess the effects of thermal aging and determine the appropriate evaluation procedures.
Although Section XI contains little guidance for evaluating cast stainless steel,Section XI does contain specific flaw evaluation procedures for flux welds which have reduced fracture toughness compared with piping of wrought i
base metal.
Specifically,Section XI contains criteria in IWB-3640 for the i
evaluation gf flaws in flux welds accounting for the reduced fracture toughness.s, The USNRC staff compared the lower-bound fracture toughness estimate developed by ANL for thermally aged cast stainless steel with the fracture toughness used in IWB-3640 of Section XI for evtluating flux welds.
The comparison is shown in Figures 2 and 3.
j Figures 2 and 3 show the lower-bound fracture toughness at 290 C (550*F) of thermally aged statically and centrifugally cast CF-8M, respectively, and the fracture toughness of submerged arc welds (SAWS).
Specifically, Figures 2 and 3 show the following information:
1.
The ANL lower-bound fracture toughness estimates for thermally aged CF-8M with various ferrite contents.
Data for Type 316 and 304 SAWS}'used in developing the fracture 2.
toughness for SAWS in IWB-3640.
(Note that the fracture tou data for the SAW were measured according to the " Modified J"" ghness procedures and the " Modified J" values are shown in Figures 2 and 3.)
3.
Reconstructed "J-R" curve used in IWB-3640 for evaluating flaws in SAWS.
Because References 8 and 9 do not explicitly provide a "J-R" curve for SAWS, the "J-R" curve was reconstructed from the " Tearing Modulus" curve and material properties for SAWS documented in References 8 and 9.
The crack extension for "J " at initiation was calculated using the ASTM g
6 i
f F813-85 procedures, that is, "J," at initiation was defined on the 0.2-mm (0.008-in) offset line. The "J-R" curve was constructed numerically from the." Tearing Modulus" curve using crack extension increments.
Then, a power law was fit to the resulting "J-R" curve.
The power law equation is shown in Figures 2 and 3.
From Figures 2 and 3, it is observed that the power law fit matches the SAW data.
(Note that the
" Tearing Modulus" curve in References 8 and 9 shows " Modified J" values and thus, the reconstructed "J-R" curve also shows " Modified J" values even though it is labelled as "J " in Figures 2 and 3.)
o Figures 2 and 3 show comparisons of the fracture toughness at 290'C (550 F) used in IWB ~c.90 for evaluating SAWS with the ANL lower-bound fracture toughness curves for thermally aged statically and centrifugally cast CF-8M, respectively, with various ferrite contents.
Reference 12 also shows similar figures.
The SAW fracture toug evaluationofSAWsinIWB-3640.grgsswr;usedbySectionXIindeveloping Figures 2 and 3 are for CF-8M material because it is conservative to consider CF-8M which is more susceptible to thermal aging than the other grades of stainless steel, that is, CF-3, CF-3A, CF-8, and CF-8A.
From Figures 2 and 3, it can be observed that the lower-bound fracture toughness of thermally aged cast stainless steel is similar to the fracture toughness used in IWB-3640 to evaluate SAWS.
Specifically, Figures 2 and 3 show that the fracture toughness of SAWS is less than the lower-bound for thermally aged statically and centrifugally cast stainless steel with a ferrite content exceeding 15 percent at small crack extensions, but it becomes larger than the lower-bound at larger crack extensions. However, the ANL-lower-bound curves have been developed conservatively and actual material properties are expected to have fracture toughness exceeding these lower-bound estimates. Another conservatism is that the SAW data and associated reconstructed "J-R" curve show " Modified J" values while the ANL lower-bound estimates show " Deformation J" values.
When measuring the fracture toughness of a material, the values of " Modified J".are larger than the corresponding values of " Deformation J."
Thus, it can reasonably be concluded that the lower-bound fracture toughness of thernCly aged cast stainless steel is similar to the fracture toughness used in M -
3640 of Section XI to evaluate SAWS.
FLAW EVALUATION BASED ON LOWER-BOUND TOUGHNESS As discussed above, Figures 2 and 3 show that the lower-bound fracture toughness of thermally aged cast stainless steel is similar to the fracture toughness used in IWB-3640 to evaluate SAWS. The orocedures in IWB-3640 of ASME Section XI reduce the load bearing capacity of the stainless steel componenttoaccountforthereducedfracturgAtoughness of the SAWS based on elastic-plastic fracture mechanics analyses.
Because the lower-bound fracture toughness of thermally aged cast stainless steel is similar to the fracture toughness of SAWS used in the elastic-plastic fracture mechanics analyses of IWB-3640, the procedures in IWB-3640 for SAWS are directly applicable to cast stainless steel.
Further, the flaw evaluation is based on the unaged material ultimate stress.
This is conservative because 7
i experimental data indicate that the flow stress, that is, half of the sum of theultimateandyieldstresses,isincreasedbyabout10,14,.and24 percent, for CF-3 and CF-3A, CF-8 and CF-8A, and CF-8M materials, respectively.
A higher ultimate stress would increase the load bearing capability of a component.
Therefore, the USNRC staff has accepted the procedures developed in IWB-3640 of Section XI for SAWS for evaluating flaws in thermally aged cast stainless steel for a topical report addressing license renewal, j
To summarize, the use of the IWB-3640 SAW procedures for evaluating flaws in thermally aged cast stainless steel components are conservative because:
1.
The lower-bound fracture toughness of thermally aged cast stainless steel is similar to the fracture toughness used in IWB-3640 to evaluate SAWS.
2.
The actual fracture toughness of a thermally aged cast stainless steel component in a U. S. utility plant would likely be higher than the ANL 1
lower-bound fracture toughness.
3.
The thermally aged components would be able to bear mora load due to the increased ultimate stress resulting from thermal aging.
This is not credited in this flaw evaluation procedure.
Thus, the use of the IWB-3640 SAW procedures for evaluating flaws in thermally aged cast stainless steel components may be considered as a " screening" step to determine if further detailed flaw evaluation accounting for actual plant-specific material properties should be performed.
However, it is expected that the flaw evaluation procedures based on the lower-bound fracture toughness of thermally aged cast stainless steel would be sufficient for the vast majority of cases. This is because the procedures in IWB-3640 for SAWS have been available since the Winter 1985 Addenda of Section XI and the IWB-l 3640 procedures have been applied by utilities successfully without resulting in unnecessary component repairs or replacements.
Administratively, the USNRC staff acceptance is currently restricted to the approval of a specific topical report submitted by a certain U. S. utility owners group addressing the aging management of their reactor coolant system piping components for license renewal. A draft code case outlining this procedure will be submitted to the ASME Code Section XI for its review and approval.
1 FLAW EVALUATION BASED ON COMPONENT-SPECIFIC TOUGHNESS Although the flaw evaluation procedures based on the lower-bound toughness are sufficient for the vast majority of cases, there may still be a few instances where a more detailed case-by-case review may be beneficial to the utility to remove excessive conservatism due to the underlying lower-bound assumptions.
In such instances, as accepted by the USNRC staff, a utility may use the procedures developed at ANL to estimate the fracture toughness of the plant-specific cast stainless steel components using specific material properties 8
i d at a. 2'I'5
The estimated fracture toughness would be applied in subsequent fracture mechanics analyses to demonstrate structural integrity to determine its acceptability for continued service on a case-by-case basis.
As discussed earlier, ANL provides fracture toughness estimation procedures with three options:
" lower bound," " saturation," and " service time." The following example il thesethreeoptionsjustratestheestimatedfracturetoughnessresultingfrom j
4 ANL has measured and estimated the fracture toughness after 53,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br /> (6 years) of thermal exposure at 320*C (610*F) and at saturation for a statically cast CF-8M sample with a ferrite content of 16 percent (designated Heat 74).
Figure 4 shows the susceptibility of this sample to thermal aging by plotting the estimated Charpy impact energy of the samp*le over time due to thermal aging at 290*C (550*F) and at 320*C (610 F). The estimated values show good agreement with the measured Charpy impact energies during service at 320 C (610 F).
Also, for this 2
sample,anaveragemeasuredsaturationimpactenergyofp3J/cm (37 ft-lb) agrees very well with the estimated value of 62 J/cm (36 ft-lb).
From Figure 4, it is observed that the Charpy impact energy decreases over time, and would reach its minimum or " saturation" value during service at 320 C (610 F), but not at 290 C (550 F), within the current license term, that is, 40 years or 32 effective full power years (EFPYs).
Figure 5 shows the experimental fracture toughness data for this sample and the estimated fracture toughness "J-R curves" at 290 C (550 F) corresponding to the " lower bound" for CF-8M with a ferrite content exceeding 15 percent, " saturation," and after a 32-EFPY " service time."
The unaged estimate of the "J-R curve" based on the measured initial Charpy impact energy is also shown. Measured data show that the estimates are conservative by about 40 percent.
A conservative estimate is as expected because that is the intent of the ANL procedures as discussed previously.
From Figure 5, it is observed that the fracture toughness estimates at
" saturation" and at 32 EFPY are about 80 and 100 percent, respectively, higher than the lower-bound estimate.
Because the lower-bound estimate is the technical basis for accepting the IWB-3640 SAW evaluation procedures, there may be certain advantages in performing a more detailed case-by-case flaw evaluation using either the fracture toughness estimate at " saturation" or 32 EFPY.
CONCLUSIONS Cast stainless steel may be used in the fabrication of the reactor coolant pressure boundary components in light-water reactors.
However, this material is subject to thermal aging at the reactor temperature, that is, about 290 C (550 F). Thermal aging results in the gradual loss of fracture toughness, that is, embrittlement, of the cast stainless steel.
Because of thermal 9
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aging, there are no generally accepted procedures in evaluating flaws in cast stainless steel components to determine acceptability for continued service.
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ANL recently completed an extensive research program on thermal aging of cast
~,
stainless steel. The results of the program indicate that the lower-bound fracture toughness of thermally aged cast stainless steel is similar to that of SAWS. On this basis, the USNRC staff has accepted the use of SAW flaw evaluation procedures in IWB-3640 of Section XI of the ASME Code to evaluate flaws in thermally aged cast stainless steel for a topical report addressing license renewal. The USNRC staff recognizes that this conclusion is based on 1
lower-bound fracture toughness of thermally aged cast stainless steel, and l
thus, in some instances, utilities may estimate component-specific fracture i
toughness using procedures developed at ANL for a case-by-case fracture mechanics flaw evaluation.
i REFERENCES 1.
" Rules for Inservice Inspection of Nuclear Power Plant Components,"
Section XI, ASME Boiler and Pressure Vessel Code, the American Society of Mechanical Engineers, New York, NY, July 1, 1989.
2.
O. K. Chopra and W. J. Shack, " Assessment of Thermal Embrittlement of Cast Stainless Steels," NUREG/CR-6177, U. S. Nuclear Regulatory Commission, Washington, DC, May 1994.
3.
O. K. Chopra, " Estimation of Fracture Toughness of Cast Stainless Steels During Thermal Aging in LWR Systems," NUREG/CR-4513, Revision 1, U. S.
Nuclear Regulatory Commission, Washington, DC, August 1994.
4.
S. Lee and P. T. Kuo, " Aging Management of Reactor Coolant System Piping fcr License Renewal," submitted to the 14th International Conference on I
Structural Mechanics in Reactor Technology (SMirt 14), Lyon, France, August 17-22, 1997.
5.
O. K. Chopra, " Estimation of Mechanical Properties of Cast Stainless Steels During Thermal Aging in LWR Systems," Transactions of the 13th International Conference on Structural Mechanics in Reactor Technology (SMiRT 13), Vol. 2, M. M. Rocha and J. D. Riera, eds., Escola de Engenharia - Universidade Federal do Rio Grande do Sul, Porto Alegre, Brazil, August 13-18, 1995, pp. 349-354.
6.
O. K. Chopra, "Effect of Thermal Aging on Mechanical Properties of Cast Stainless Steels," Proceedings of the 2nd International Conference on Heat-Resistant Materials, K. Natesan, P. Ganesan, and G. Lai, eds., ASH International, Materials Park, OH, August 1995, pp. 479-485.
7.
S. Jayet-Gendrot, P. Ould, and T. Meylogan, " Fracture Toughness Assessment of In-Service Aged Primary Circuit Elbows Using Mini C(T)
Specimens Taken from Outer Skin," Fatigue and Fracture Mechanics in Pressure Vessels and Piping, PVP-Vol. 304, H. S. Mehta, ed., American 10
f Society of Mechanical Engineers, New York, NY,1996, pp.163-169.
8.
"Evaluatit ' of Flaws in Austenitic Steel Piping," ASME Code Section XI Task Group for Piping Flaw Evaluation, NP-4690-SR, Electric Power Research Institute, Palo Alto, CA, July 1986.
9.
" Evaluation of Flaws in Austenitic Steel Piping," Journal of Pressure Vessel Technology, Vol. 108, No. 3, August 1986, pp. 352-366.
]
10.
J. D. Landes and D. E. McCabe, " Toughness of Austenitic Stainless Steel Pipe Welds," NP-4768, Electric Power Research Institute, Palo Alto, CA, October 1986.
11.
H. A. Ernst, " Material Resistance and Instability Beyond J-Control Crack Growth," in Elastic-Plastic Fracture: Second Symposium, Vol. 1:
Inelastic Crack Analysis, ASTM STv 803, American Society for Testing and Materials, Philadelphia, PA.,1983.
12.
R. Nickell and M. A. Rinckel, " Evaluation of Thermal Aging Embrittlement for Cast Austenitic Stainless Steel Components," TR-106092, Electric Power Research Institute, Palo Alto, CA, March 1996.
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2 Table 1 Parameters C and n for lower-bound "J-R curve" measured at room temperature and at reactor operating temperature for thermally aged '
statically and centrifuga11y cast stainless steel i
Statically Cast Centrifuaally Cast i
21*C (70 F) 290*C (550*F) 21*C (70 F) 290*C (550 F) l l
Material Grade C
n C
n C
n C
n i
l l
Ferrite Content > 15%
l CF-3 and CF-3A 287 0.39 264 0.35 334 0.39 347 0.35 CF-8 and CF-8A 261 0.37 251 0.34 304 0.37 330 0.34 CF-BM 119 0.33 167 0.31 149 0.33 195 0.31 i
)
Ferrite Content 10-15%
CF-3 and CF-3A 342 0.40 290 0.36 398 0.40 382 0.36 i
CF-8 and CF-8A 307 0.38 274 0.35 357 0.38 360
.0.35 CF-BM 149 0.35 192 0.32 186 0.35 223 0.32 j
Ferrite Content < 10%
l CF-3 and CF-3A 400 0.40 331 0.39 507 0.43
-435 0.39 CF-8 and CF-8A 394 0.40 313 0.17 458 0.41 412 0.37 CF-8M 211 0.36 238 0.33 264 0.36 276 0.33
)
12 l
.._._..__m i
Figure 1 Experimental and estimated values of room temperature Charpy impact energy at saturation for various compositions of thermally aged cast stainless steel.
Figure 2 Estimated lower-bound fracture toughness at 290 C (550 F) of statically cast CF-8M thermally aged at the same temperature for various ferrite contents and fracture toughness of submerged arc welds used in IW8-3640 evaluations of Section XI.
Figure 3 Estimated lower-bound fracture toughness at 290"C (550 F) of I
centrifugally cast CF-8M thermally aged at the same temperature for various ferrite contents and fracture toughness of submerged arc l
welds used in IW8-3640 evaluations of Section XI.
i Figure 4 Experimental and estimated room temperature Charpy impact energy j
over time due to thermal aging at 290 C (550 F) and at 320 C (610 F) for a sample of CF-8M with 16 percent ferrite.
i i
Figure 5 Experimental and estimated fracture toughness at 290 C (550'F) for a sample of CF-8M with 16 percent ferrite for the following conditions:
unaged, after service for 32 EFPYs, " saturation," and lower bound.
13 I
1 I
Measured Impact Energy at Saturation (ft Ib) 0 25 50 75 100 n^ 200 r''I8
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E O
p e
3 O
100 e
e c
o o
2 m
a E 150 E
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2 E
o 75 3
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6 A
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g A
O 9
5 100
/G 0 42 E
6 s
o3 50 W
A o
S A
O E
a o
o E
E A
u 50 11 o
CF-3 u
25 e
2 o
CF-8 E
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A a
CF-8M
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E y
Total 101 Compositions -
w i
....i....i....i..
0
._?O 0
50 100 150 200 i
2 Measured Impact Energy at Saturation (J/cm )
Figure 1.
Emerirnental and estirnatal t alues of voorn ternperature Charpy impact energy at saturation for various corapositions of thermally aged cast stainless steel l
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Crack Extension (in) 0.0 0.05 0.10 0.15 500 i
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Je=178.2Aa 513 Reconstructed J-R Curve for p
400 ASME IWB-3640 SAW Evaluation p
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c 2000 x
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Q 300
<10% Ferrite g
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p' 'g W A' C
v M
2
'a A
.9 200
~
5 10-15% Ferrite-1000 E
S
~
^
~
_~
ci y
.A
>15% Ferrite g
h.
'p 100 u /
O o
Type 316 SAW -
a a
Type 304 SAW -
0 0
0 1
2 3
4 Crack Extension (mm)
Figure 2.
Estimated louer-boundfracture toughness at 290D (550'F) of statically cast CF-8M thermally aged at the same temperaturefor variousferrite contents andfracture toughness of submerged arc welds used in IWB-3640 evaluations ofSection X7 i
l 1
I i
15
)
q j
Crack Extension (in) l O.0 0.05 0.10-0.15 l
-500 i
i i
i i
i i
iiiii_
- J =178.2Aa W M
~
0 d
Reconstructed J-R Curve for k
^
p 400 ASME IWB-3640 SAW Evaluation I
c 6
E
~
3 pq 2000 y 300 10-15% Ferrite A
E i
o
- c p.'
c.
o c
s
.o
>15% Ferrite -
g g
200 u
a_
1000 E
o-A g
S
~
/A y
3
,, g f 'a O
100
[.
o Type 316 SAW -
1 A A
Type 304 SAW -
' l '
0 0
0 1
2 3
4 Crack Extension (mm)-
F1gure 3.
Estimated lower-boundfracture toughness at 290*C (550cF) of centrifugally cast CF-BM thennally aged at the same temperaturefor variousferrite contents andfracture toughness ofsubmerged arc welds used in 1%73-3640 cualuations ofSection XI
\\
i i
k 16 4
~..
Aging Time Duration (hours) 103 4
10 105 250 3 i
e i i i ii q i
i i i i ii q i
, i,,i o
o' 3
o s
g 100 x 150 a
s O
C N
gx h Estimated o\\
h 100 s
O 5
320 C 50 W
-f
~
- 290cC
~
T5
~
v-50
- Experimental Measured /
o 320oC Saturation = 63 J/cm2 E
0 -
''I
'I O
1 0'1 100 101 102 Aging Time Duration (years)
Figure 4.
Experimental and estimated room temperature Charpy impact energy over time due to thermal aging at 290'C (550=F) and at 320'C (610'F)for a sample of CF-8h1 with 16 percentferrite i
17 s
I J
~
1 i
Crack Extension (in) l 0.0
' O.1 0.2 0.3 1500
,i 1
i i i i
i i i
- i i i i
- i i i i-3
- o - - - - - - - Unaged
- A
' Saturation o
- 7500 o-
- - 6 EFPY at 320 C or o
^
m W
32 EFPY at 290 C 0
m o
- 9 c
- Lower Bound o
2-1000 o o 4
o 5
o o
5000 C-o
.. - ;' o oa* ^
^ ^
^:
7 c
o o
o a
.9 8
- g4 C
- &*,,o la
._9 g
go
~
E' U. '&
m B
500 -
fo9 E
3
- g a
o, 2500 o
,o o
g
- ~'
0 0
0 2
4 6
8 10 Crack Extension (mm)
Figure 5. Experimental and estimatedfracture toughness at 290*C (550*F)for a sample of CF-8M with 16 percentferritefor thefollowing conditions. unaged, aper servicefor 32 EFPYs, " saturation," uad lower bound 18 i
l
.