ML20094N758
| ML20094N758 | |
| Person / Time | |
|---|---|
| Site: | Seabrook  |
| Issue date: | 06/30/1984 |
| From: | Clark H, Lee Y, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19277H799 | List: |
| References | |
| WCAP-10566, NUDOCS 8408160378 | |
| Download: ML20094N758 (40) | |
Text
.
WCAP-10566 TECHNICAL BASES FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUC-TURAL DESIGN BASIS FOR SEABROOK UNITS 1 & 2 June 1984 S. A. Swamy Y. S. Lee H. F. Clark, Jr. R. A. Holmes j !
f APPROVED: \. L . /~) APPROVED:
J. 'N. Chirigos , ger E. R. Johnson, Manager Structural Materia s Structural and Seismic Engineering Development APPROVED: I W
. J. McInerney, Manager Mechanical Equipment and Systems Licensing Work perfomed under NAH-950 WESTINGHOUSE ELECTRIC CORPORATION NUCLEAR ENERGY SYSTEMS P.O. Box 355 PITTSBURGH, PENNSYLVANIA 15230 8408160378 840009 l PDR ADOCK 05000443 A PDR i )
l FOREWORD This document contains Westinghouse Electric Corporation proprietary i information and data which has been identified by brackets. Coding associated i with the brackets set forth the basis on which the information is considered proprietary. These codes are listed with their meanings in WCAP-7211.
The proprietary information and data contained in this report were obtained at considerable Westinghouse expense and its release could seriously affect our competitive position. This information is to be withheld from public disclosure in accordance with the Rules of Practice,10 CFR 2.790 and the information presented herein be safeguarded in accordance with 10 CFR 2.903.
Withholding of this information does not adversely affect the public interest.
This information has been provided for your internal use only and should not be released to persons or organizations outside the Directorate of Regulation and the ACRS without the express written approval of Westinghouse Electric Corporation. Should it become necessary to release this infomation to such persons as part of the review procedure, please contact Westinghouse Electric Corporation, which will make the necessary arrangements required to protect the Corporation's proprietary interests. .
The proprietary information is deleted in the unclassified version of this report.
iii
TABLE OF CONTENTS Section Title Page FOREWARD
1.0 INTRODUCTION
1-1 1.1 Purpose 1-1 1.2 Scope 1-1
, 1.3 Objectives 1-1 1.4 Background Infomation 1-2 2.0 OPERATION AND STABILITY OF THE PRIMARY SYSTEM 2-1 4
2.1 Stress Corrosion Cracking 2-1 2.2 Water Hammer 2-2 2.3 Low Cycle and High Cycle Fatigue 2-3 3.0 PIPE GE0E TRY AND LOADING 3-1 1
f 4.0 FRACTURE ECHANICS EVALUATION 4-1 4.1 Global Failure Mechanism 4-1 4.2 Local Failure Mechanism 4-2
! 4.3 Material Properties 4-3 4.4 Results of Crack Stability Evaluation 4-5 ;
5.0 LEAK RATE PREDICTIONS 5-1 1
6.0 FATIGUE CRACK GROWTH ANALYSIS 6-1 7.0 ASSESSENT OF MARGINS 7-1
8.0 CONCLUSION
S 8-1
9.0 REFERENCES
9-1 APPENDIX A- [ Ja.c.e A,l l
l v
- - ~ ~
,e-LIST OF TABLES P
i Table Title g 1 Seabrook Primary Loop Data T-1 2 Fatigue Crack Growth at [ ]a,c,e T-2 f
i 4
r i
- i h
! E i
l r I
1 s
I f
vit
LIST OF FIGURES Figure, Title g 1 Reactor Coclant Pipe F-1 2 Schematic Diagram of Seabrook RCL Showing Pipe to Equip- F-2
. ment Weld Identifications 3 [ Ja.c.e Stress Distribution F-3 ,
4 J-aa Curves at Different Temperatures, Aged Material F-4
[ ]a c.e (7500 Hours at 400'C) 5 .
Critical Flaw Size Prediction F-5 6 Typical Cross-Section of [ ]a,c.e F-6
-7 Reference Fatigue Crack Growth Curves for F-7
[ 3a c.e 8 Reference Fatigue Crack Growt5 Law for [ Ja.c.e F-8 :
in a Water Environment at 600'F A-1 Pipe with a Through-Wall Crack in Bending A-2 i
I W
ix j
T 1.0 INTR 000CTION 4
1.1 Puryose
- This report applies to the Seabrook plants reactor coolant system primary loop
' piping. It is intended to demonstrate that specific parameters for the Seabrook plants are enveloped by the generic analysis performed by Westinghouse in WCAP-9558. Revision 2 (Reference 1) and accepted by the NRC (Reference 2).
l 1.2 Scope
- The' current structural design basis for the Reactor Coolant System (RCS) primary loop requires that pipe breaks be postulated as defined in the .
3 approved Westinghouse Topical Report WCAP-8082 (Reference 3). In addition, protective measures for the dynamic effects associated with RCS primary loop pipe breaks have been incorporated in the Seabrook plants design. However, I
Westinghouse has demonstrated on a generic basis that RCS primary loop pipe i breaks are highly unlikely and should not be included in the structural design '
basis of Westinghouse plants (see Reference 4). In order to demonstrate the appifcability of the generic evaluations to the Seabrook plants, Westinghouse l has perfomed: a comparison of the loads and go.utry for the Seabrook plants with envelope parameters used in the generic analyses (Reference 1), a i
j fracture mechanics evaluation, a determination of leak rates from a j through-wall crack, fatigue crack growth evaluation, and an assessment of l margins.
1.3 Objectives j i l The conclusions of WCAP-9558 Revision 2 (Reference 1) support the elimination f
, of,RCS primary loop pipe breaks for the Seabrook plants. In order to validate
- this conclusion the following objectives must be achieved. ,
I s
1-1
- a. Demonstrate that Seabrook plants parameters are ' enveloped by generic Westinghouse studies. l
'l
- b. Demonstrate that margin exis'ts between the critical crack size and a l
. postulated crack which yields a detectable leak rate. l 1
- c. Demonstrate that there is sufficient margin between the leakage through a L
postulated crack and the leak detection capability of the Seabrook plants.
d .- Demonstrate that fatigue crack growth is negligible.
1.4 Background Information Westinghouse has performed considerable testing and analysis to demonstrate that RCS primary loop pipe breaks,can be eliminated from the structural design basis of all Westinghouse plants. The concept of eliminating pipe breaks in the RCS primary loop was first presented to the NRC in 1978 in WCAP-9283 (Reference 5). This Topical Report employed a deterministic fracture mechanics evaluation and a probabilistic analysis to support the elimination of RCS primary loop pipe breaks. The approach was then used as a means of addressing Generic Issue A-2 and Asymmetric LOCA Loads.
Westinghouse performed additional testing and analysis to justify the elimination of RCS primary loop pipe breaks. As a result of this effort, WCAP-9558, Revision 2, WCAP-9787, and Letter Report NS-EPR-2519 (References 1,
! 6, and 7) were submitted to the NRC.
The NRC funded research through Lawrence Livermore National Laboratory (LLNL) l to address this same issue using a probabilistic approach. As part of the LLNL research effort, Westinghouse performed extensive evaluations of specific plant loads, material properties, transients, and system geometries to demonstrate that the analysis and testing previously performed by Westinghouse and the research performed by LLNL applied to all Westinghouse plants, including Seabrook (References 8 and 9). The results from the LLNL stu(y were released at a March 28, 1983 ACRS Subcommittee meeting. These studies which l
1-2 l
I' are applicable to all Westinghouse plants east of the Rocky Mountains,
' determined the mean probability of a direct LOCA (RCS primary loop pipe break) to be 10-10 per reactor year and the mean probability of an indirect LOCA to
' be 10-7 per reactor year. Thus, the results previously obtained hy
. Westinghouse (Reference 5) were confirmed by an independent NRC research study.
Based on the studies by Westinghouse, hy'LLNL, the ACRS, and the AIF, the NRL completed a safety review of the Westinghouse reports submitted to address 4
asymmetric blowdown loads that result from a number of discrete break locations on the PWR primary systems. The NRC Staff evaluation (Reference 2) concludes that an acceptable technical basis has been provided so that asymmetric blowdown loads need n'ot be considered for those plants that can 4
demonstrate the applicability of the modeling and conclusions contained in the' Westinghouse response or can provide an equivalent fracture mechanics demonstration of the primary coolant loop integrity.
This report will demonstrate the applicability of the Westinghouse generic evaluations to the Seabroot plants.
8 J
f n
i f
1-3 i
2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM The Westinghouse reactor coolant system primary loop has an operating history which demonstrates the inherent stability characteristics of the design. This includes 'a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress corrosion cracking), water hammer, or fatigue (low and high cycle). This operating history totals over 400 reactor-years, including five plants each having 15 years of operation and 15 other plants each with over 10 years of operation.
2.1 3ressCorrosionCracking For the_ Westinghouse plants, there is no history of cracking failure in the reactor coolant system loop piping. For stress corrosion cracking (SCC) to occur in piping, the following three conditions must exist simultaneously:
high tensile stresses, a susceptible material, and a corrosive environment (Reference 10). Since some residual stresses and some degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimized by proper material selection immune to SCC as well as preventing the occurrence of a corrosive environment. The material specifications consider compatibility with the system's operating environment (both internal and external) as well as other materials in the system, applicable ASE Code rules, fracture toughness, welding, fabrication, and processing.
The environments known to increase the susceptibility of austenitic stainless steel to stress corrosion are (Reference 10): oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g., sulfides, j sulfites, and thionates). Strict pipe cleaning standards prior to operation and careful control of water chemistry during plant operation are used to prevent the occurrence of a corrosive environment. Prior to being put into service, the piping is cleaned internally and externally. During flushes and preoperational testing, water chemistry is controlled in accordance with written specifications. External cleaning for Class 1 stainless steel piping f includes patch tests to monitor and control chloride and fluoride levels. For o
. 2-1
preoperational flushes, influent water chemistry is controlled. Requirements on chlorides, fluorides, conductivity, and pH are included in the acceptance criteria for the piping.
During plant operation, the reactor coolant water chemistry is monitored and maintained within very specific limits. Contaminant concentrations are kept below the thresholds known to be conducive to stress corrosion cracking with the major water chemistry control standards being included in the plant operating procedures as a condition for plant operation. For example, during nomal power operation, oxygen concentration in the RCS is expected to be less than 0.005 ppm by controlling charging flow chemistry and maintsining hydrogen in the reactor coolant at specified concentrations. Halogen concentrations are also stringently controlled by maintaining concentrations of chlorides and'
- fluorides within the specified limits. This is assured by controlling charging flow chemistry and specifying p. roper wetted surface materials.
2.2 Water Hammer Overall, there is a low potential for water hammer in the RCS since it is designed and operated to preclude the voiding condition in normally filled f lines. The reactor coolant system, including piping and primary components, is designed for normal, upset, emergency, and faulted condition transients.
The design requirements are conservative relative to both the number of transients and their severity. Relief valve actuation and the associated l
j hydraulic transients following valve opening are considered in the system l
design. Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlled. Temperature during normal operation is maintained within a narrow range by control rod position; pressure is cont 911ed by pressurizer heaters and pressurizer spray also within a narrow range for steady-state conditions. The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant Additionally, pump characteristics are controlled in the design process.
L Westinghouse has instrumented typical reactor coolant systems to verify the l
i 2-2
flow 'and vibration characteristics of the system. Preoperational testing and operating experience have verified the Westinghouse approach. The operating transients of the RCS primary piping are such that no significant water hammer Can occur.
2.3 Low Cycle and High Cycle Fatigue Low cycle fatigue considerations are accounted for in the design of the piping system through the fatigue usage factor evaluation to show compliance with the rules of Section III of the ASE Code. A further evaluation of the low cycle fatigue loadings was carried cut as part of this study in the form of a
. fatigue crack growth analysis, as discussed in Section 6.
High cycle fatigue loads in the system would result primarily from pump vibrations. These are minimized by restrictions placed on shaft vibrations
- during hot functional testing and operation. During operation, an alann signals the exceedance of the vibration limits. Field measurements have been made on a numb'er of plants during hot functional testing, including plants similar to Seabrook plants. Stresses in the elbow below the RC pump have been found to be very small, between 2 and 3 ksi at the highest. These stresses are well below the fatigue endurance limit for the material, and would also result in an applied stress intensity factor below the threshold for fatigue crack growth.
4 1
l l
1 l
l 2-3
r 3.0 PIPE GE0 METRY AND LOADING l
A segment of the primary coolant hot leg pipe is shown in Figure 1. This segment is postulated to contain a circumferential through-wall flaw. The inside diameter and wall thickness of the pipe are 29.2 and 2.37 inches, respectively. The pipe is subjected to a normal operating pressure of 1
[ ]a,c.e psi. Figure 2 identifies the loop weld locations these six pipe to nozzle weld locations were found to envelope all the intermediate welds. The material properties and the loads at these locations resulting from deadweight, thennal expansion and Safe Shutdown Earthquake are indi-cated in Table 1. As seen from this Table, the junction of hot leg and the reactor vessel outlet nozzle is the worst location for crack stability analy- '
sis based on the highest stress due to combined pressure, dead weight, thermal expansion, and SSE (Safe Shutdown Earthquake) loading. At this lo-cation, the axial load (F) and the bending moment (M) are [ 3a,c.e (including axial force due to pressure) and [ ]a,c,e, respectively.
The loads of Table 1 are calculated as follows:
The axial force F and transverse bending moments, My and Mz, are chosen for each static load (pressure, deadweight and thennal) based on elastic-static analyses for each of these load cases. These pipe load components are c mbined algebraically to define the equivalent pipe static loads Fs' Mys, and M zs. Based on non-linear time history SSE analyses, amplifiedpipeseismicloads,F'hd,Mzd d are obtained. The maximum pipe loads are obtained by combining the static and dynamic load components as follows:
F= F s
+ F d
M= M y
+M z
where:
i h*lNys
- M yd "z " zs zd l
r i
l 3-1
. - , - _s- . ~ ,
The corresponding geometry and loads used in the reference report (Reference I
'1) are as follows: . inside diameter and wall thickness are 29.0 and 2.5 l inches; axial load and bending mcment are [ ]C inch-kips.- The outer fiber stress for Seabrook is [ ]a,c.e ksi, while for the reference report it is [ 3a,c.e ksi. This demonstrates conservatism in
.the reference report which makes it more severe than the Seabrook analyses.
The normal operating loads (i.e., algebraic um s of pressure, deadweight, and
- 100 percent power thermal expansion loading) at the critical location, i.e.,
the junction of hot leg and the reactor pressure vessel outlet nozzle are as follows:
F=[ ]a,c.e (including internal pressure) g,[ 3a,c.e The calculated and allowable stresses for ASME Code equation 9 (faulted, i.e.,
pressure, deadweight, and SSE) and equation 12 (thermal) at the critical location are as follows:
Calculated Allowable Ratio of Equation Stress Stress Calculated /
Number (ksi) (ksi) Allowable a,c.e m
3-2
l 4.0 FRACTURE MECHANICS EVALUATION 4.1 Global Failure Mechanism l
Determination of the conditions which lead to failure in stainless steel must be done with plastic fracture methodology because of the large amount of deformation accompanying fracture. A conservative method for predicting the failure of ductile material is the [
]"' C d This methodology has been shown to be applicable to ductile piping through a large number of experiments, and will be used here to predict the critical flaw size in the primary coolant piping. The failure criterion has been obtained by requiring [
la.c.e (Figure- 3) when loads are applied. The detailed development is provided in Appendix A, for a through-wall circumferential flaw in a pipe with internal pressure, axial force, and imposed bending moments. The [
]a,c.e for such a pipe is given by:
~
a,c.e where:
a,c.e m
4-1
s a,c.e The analytical model described above accurately accounts for the piping internal pressure as well as imposed axial force as they affect [
7 3a.c.e Good agreement was found between the analytical-predictions and the experimental results (Reference 11).
4.2 Local Failure Mechanism The local mechanism of failure is primarily dominated by the crack tip behavior in terms of crack-tip blunting, initiation, extension and finally crack instability. Depending on the material properties and geometry of the pipe, flaw size, shape and loading, the local failure mechanisms may or may not govern the ultimate failure.
l The stability will be assumed if the crack does not initiate at all. It has been accepted that the initiation toughness, measured in terms of Jgy O .e. , J Ic from a J-integral resistance curve is a material parameter defining the crack initiation. If, for a given load, the calculated J-integral value is shown to If the be less than J yy of the material, then the crack will not initiate.
L initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:
dJ E T,pp = da 2 a The notation JIN instead of J cI was used in Reference 1 to designate the j value of the J-integral at crack initiation; the J NI notation will be used L in this ' report in keeping with Peference 1.
4-2
where:-
T,p = applied tearing modulus ,
= modulus of elasticity of = [ ]a c.e (flow stress) a = crack length
[ 3a,c.e In summary, the local crack stability will be ' established by the two-step criteria:
l J<J IN app < T,,, if J > J gy 4.3 Material Properties The materials in the Seabrook Units 1 and 2 primary loops are wrought stainless steel pipe (SA 376 304N), cast stainless steel fittings (SA351 CF8A) and associated welds. The tensile and flow properties of the limiting location, the hot leg and reactor vessel outlet nozzle junction, are given in Figure 5, wh'ich will be discussed further in the next section. For this location, the material of interest is the wrought seamless pipe material (SA 376 304N). The fracture properties of this material are equivalent to the data reported in Reference 1. Cast pipe (SA 351-CF8A) fittings, such as elbows, are subject to themal aging. For this reason, an additional location with cast material properties was evaluated. This location is the junction between the cast elbow fitting and wrought pipe section at the inlet of the reactor pressure vessel.
The fracture properties of CF8A cast stainless steel have been determined
- through fracture tests carried out at 600"F and reported in Reference 12.
This reference shows that J gy for the base metal ranges from [
l ] # # for the multiple tests carried out.
l 4-3
l Cast stainless steels are subject to thermal aging during service. This f thermal aging causes an elevation in the yield strength of the material and a
-degradation of the fracture toughness, tne degree of degradation being proportional to the level of ferrite in the material. To determind the effects of thermal aging on piping integrity a detailed study was carried out in Reference 13. In this report, fracture toughness results were presented for a material representative of [
- Ja,c.e Toughness results were provided for the material in the fully aged condition and these properties are also presented in Figure 4 of this report for information. The J value for this material at operating IN 8
temperature was approximately [ 3 'C d and the maximum value of J obtained in the tests was in excess of [ ]a c.e The tests of this material were conducted on small specimens and therefore rather short crack extensions, (maximum extension 4.3 mm) so it is expected that higher J values would be sustained for larger specimens. [
l
].a.ce Therefore, it may be concluded that the degree of ,
thermal aging, expected by end-of-life for these units is much less than that which was produced in [ 3a,c.e of Reference 13, and therefore the l
J values for'the Seabrook plants 1 and 2 after end-of-life would be l IN expected to be higher than those reported for [ ]a,c .e in Figure 4 (also see l
! (Reference 14). In addition, the tearing modulus for the Seabrook Units 1 and 2 materials would be greater than [ ].a,c.e values for the Available data on stainless steel welds indicate the JIN worst case welds are of the same order as the aged material, but the slope of the J-R curve is steeper, and higher J-values have been obtained from fracture 2
tests (in excess of 3000 in-lb/in ). The applied value of J integral for a flaw in the weld region will be lower than that in the base metal because the yield stress for the weld material is much higher at temperature. Therefore, weld regions are less limiting than the cast material.
4-4
4.4 Results of Crack Stability Evaluation Figure 5 shows a plot of the [- ]a,c,e as a function of through-wall circumferential flaw length in the [ ]a,c.e of the main coolant piping. This [ 3a,c.e was calculated for Seabrook data of a pressurized pipe at [
]a,c.e properties. The maximum applied bending moment of [ 3a,c.e in-kips can be plotted on this figure, and used to determine the critical flaw length, which is shown to be
[ 3a,c.e inches. This is considerably larger than the [ 3a,c.e inch reference flaw used in Reference 1.
J'
[
3a.c.e Therefore, it can be concluded that a postulated [ 3a,c.e inch through-wall flaw in the Seabrook plants loop piping will remain stable from both a local and global stability standpoint.
The J integral value was estimated for a [ ]a,c.e long through-wall flaw. The estimation was based on extrapolation of the available [
3a,c.e long flaw. The j amplification factor for estimating J corresponding to the larger crack was obtained by detemining the J integrals using simplified handbook methods l l
l 4-5
.' (Reference 19) for .the two cracks when Reference 1 maximum loads were applied. !
This amplification factor was determined to be [ J.a.c.e Since the J value f for a [. ]a,c.e. crack corresponding to the maximum Seabrook loads is
} ^[. ]a.c.e the J,pp for a 9. inch flaw is estimated to be about
!. -[. ].a,c.e This is within the range of the maximum J (Jmax)
, _obtained in tests oh wrought materials (Reference 1). The applied tearing modulus [ ]a c.e fora [ 3a,c.e crack is 'significantly lower than
]a.c.e (Reference 1). Therefore, a postulated [ 3a.c.e theT,,g[
l flaw will remain stable.
The loads at the elbow (cast stainless steel SA351-CF8A) are significantly lower thantheloadofthecritical[ 3a.c.e (Figure 2). Because of the difference in material, a stability calculation was performed for the elbow carrying the j maximum load which is identified as [ ]a c.e (Figure 2). At [
3a.c.e(seeFigure2),themaximumbendingmoment is [
]a,c.e For this loading condition and the [ 3a,c.e ,
f ,
reference flaw size, a J value of less than [ '
]a c.e is estimated.
Thus, crack initiation is not expected. In addition, the applied tearing modulus
).a,c.e As stated in the previous sec-
- T applied as taken from Reference 13 is [
tion, the estimated tearing modulus for Seabrook (cast stainless steel elbows in !
j the fully aged condition) is at least [60 (80 best estimate)la.c.e Therefore, a postulated [ 3a.c.e flawat[ 3a.c.e will also remain stable, Fora [ ]a,c.e flaw in the elbow, J is estimated to be less than [
f
! ]a.c.e using Reference 19. This value is significantly less than the Ja.c.e J max value obtained from testing the [
]' '# ' '
l~
[
]a,c.e ,
]
material of Reference 13. T,pp wascalculatedforthe[ -
3a ,c .e using the appropriate fonnula of 3eference 1 and was found to be [
]a c.e will remain stable. l i, Thusa[
l .
l At [ 3a.c.e (see Figure 2), the thickness of the weld region is near that f of the pipe. However, the yield stress for the weld material is much higher than the base metal [ 3a.c.e(Reference 6). The yield stress is I
! higher than the appliedstressof[ ].a.c.e Therefore a linear elastic for-mulation for K will provide a reasonable estimate for J. Using Reference 19, J was 4-6
- :1a,:_ z_ . - l found.to'be[. ]a,c.e for both a [ '!
]a,c.e This value is well below the J max E 3**
for aged welds. T,pp was calculated to be [ .]a.c.e for both flaw sizes using the appropriate fomula of Reference 1. Thus flaw sizes of [ ]a,c.e till remain stable in the welds.
j b
i i
l 1
I i
j 4-7
5.0 LEAX RATE PREDICTIONS Leak rate estimates were performed by applying the normal operating bending moment of [ 3a,c.e in addition to the normal operating axial force of [ ]***** These loads were applied to the hot leg pipe containing a postulated [ 3a ,c.e through-wall flaw and the crack opening area was estimated using the method of Reference 15. The leak rate l was calculated using the two-phase flow formulation described in Reference 1.
l The computed leak rate was [ ]."'C In order to determine the sensitivity of leak rate to flaw size, a through-wall flaw [ Ja,c.e in length was postulated. The calculated leak rate was [
j,a,c e The Seabrook plants have an RCS pressure boundary leak detection system which is consistent with the guidelines of Regulatory Guide 1.45 of detecting leakage of 1 gpm in one hour. Thus, for the [ ]a,c.e inch flaw, a factor of approximately [ 3a,c.e exists between the calculated leak rate and the criteria of Regulatory Guide 1.45. Relative to the [
3a,c.e I
5-1 l ._ _
6.0 FATIGUE CRACK GROWTH ANALYSIS To determine the sensitivity of the primary coolant system to the pre.sence of small cracks, a fatigue crack growth analysis was carried out for the [
3a,c,e region of a typical system. This region was selected because it is typically one of the highest stressed cross sections, and crack growth calculated here will be conservative for application to the entire primary coolant system.
A[
Ja,c.e of a plant typical in geometry and operational characteristics to any Westinghouse PWR System. [
3a ,c.e All normal, upset, and test conditions were considered and circumferential1y oriented surface flaws were postulated in the region, assuming the flaw was located in three different locations, as shown in Figure 6. Specifically, these were:
Cross Section A:
a,c.e Cross Section B:
Cross Section C:
Fatigue crack growth rate laws were used [
3a,c.e The law for stainless steel was derived from Reference 16, with a very conservative correction for R ratio, which is the ratio of minimum to maximum stress during a transient.
6-1
r The law is:
h=(5.4x10-I2)K,ff 4.48 inches / cycle where K,f f = K,,x (1-R )0.5 R=K min5max
[
3a,c.e
~
a,c.e where:
The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is summarized in Table 2, and shows that the crack growth is very small, regardless [
3a,c.e 6-2
7.0 ASSESSMENT OF MARGINS l
i In Seabrook Units 1 and 2 the primary piping is wrought whereas the fittings are cast. The most highly stressed locations are considered for both svtuations and also for welds.
In Reference 1, the maximum design load was [ ]a,c.e, whereas, the maximum load as noted in Section 3.0 of this report is significantly less,
[ ]a c.e For the current application, the maximum value of J
[ ]a,c.e as deter-minedby[ ]a.c.e analyses compared with the value of [
] a,c.e in Reference 1. Thusinwroughtpipe,the[ ]a c.e flaw will be stable. In addition, at the maximum load location of the wrought -
pipe the Seabrook plants Unit 1 and 2 applied J-value is enveloped by the J,,x of Reference 1. For the cast material Section 4.3 shows that the testing of fully aged material of chemistry worse than that existing in Seabrook plants cast piping extended to J values of [ ],a,c.e this is greater than the maximum value of applied J of [ 3a,c.e for a[ ]a,c.e in cast stainless steel elbows. Crack initiation would not occur in the cast material. For a [ ]ac.e flaw in a weld, J is less than [ ]a c.e which is well less than the J max obtained from thermally aged welds. T,pp islessthan[ ].a c.e Thus, welds will remain stable for the reference flaw. Insummary,a[ ]a.c.e will remain stable in the primary piping system regardless of where it is.
In Section 4.4, it is seen that a [ 3a c.e flaw in wrought pipe has a J value at maximum load of [ ]a,c.e which is also enveloped by the J of Reference 1. T app is[ ] less than T mat' I" Section 4.4, the " critical" flaw size using [ ]a,c.e methods is calculated tobe[ ]a.c.e Based on the above, the " critical" flaw size will, of course, exceed [ ]a,c.e in the wrought pipe.
In Section 4.4, the applied J values for a [ 3a,c.e in cast elbows are well less than J max obtained from the [ 3a,c.e T,pp was calculated to be [ ]a c.e which is well less than that for [ ]a,c.e Thus a [ 3a,c.e flaw in the elbow will remain stable. A similar conclusion holds for welds based on an evaluation accounting for the elevated yield strength.
Specifically, both cast materials and welds in the loop are stable for flaws for
[ 3a.c.e andamarginof[ 3a,c.e exists against tearing instability.
7-1
In Section 5.0, it is shown that a flaw [ 3a.c.e would yield a leak rate of almost [ ]C Thus, there is a factor of at least [ 3a,c.e between the flaw size that gives a leak rate of [ ]a,c.e that of Regulatory Guide 1.45 and the " critical flaw size of [ 3.a.c.e In summary, relative to
- 1. Loads
- a. Seabrook Unit 1 and 2 are enveloped both by the maximum loads and J values in Reference 1 and the J values employed in testing of fully aged material and the forged material properties at the critical locations.
- b. At the maximum load critical location, a margin of [ ]a,c.e on faulted condition stresses and a margin of [ 3a.c.e on themal stresses exists relative to ASME code allowable values.
- 2. Flaw Size
- a. A margin of at least [ 3a c.e exists between the " critical" flaw and the flaw yielding a leak rate of [ 3a,c.e
- b. A margin exists of at least [ 3a,c.e relative to tearing,
- c. A margin exists of at least [ ]a,c.e relative to global stability.
If [ 3a,c.e is used as the basis for " critical" flaw size, the margin for global stability would be at least [ 3a,c.e with respect to a flaw size [
3a,c.e,
- 3. Leak Rate A margin of almost 1. ]a.c.c exists between calculated leak rates for the reference flaw [ 3a.c.e and the criteria of Regulatory Guide 1.45 at the critical location.
7-2
% 3. Y
8.0 CONCLUSION
S ,
This report has established the applicability of the generic Westinghouse evaluations which justify the elimination of RCS primary loop pipe breaks for the Seabrook plants as follows:
- a. The loads, material properties, transients, and geometry relative to the Seabrook Units 1 and 2 RCS primary loop are enveloped by the parameters of WCAP-9558, Revision 2 (Reference 1) and WCAP-10456 (Reference 16).
- b. Stress corrosion cracking is precluded by use of fracture resistant materials in the piping system and controls on reactor coolant chemistry, temperature, pressure, and flow during normal operation.
- c. Water hammer should not occur in the RCS piping because of system design, testing, and operational considerations.
- d. The effects of low and high cycle fatigue on the integrity of the primary piping are negligible,
- c. Anple margin exists between the leak rate of the reference flaw and the criteria of Reg. Guide 1.45.
- f. Ample margin exists between the reference flaw chosen for leak detectability and the " critical" flaw.
- g. Ample margin exists in the material properties used to demonstrate end-of-life (relative to aging) stability of the reference flaw.
The reference flaw will be stable throughout reactor life because of the ample margins in e, f, and g, above, and will leak at a detectable rate which will assure a safe plant shutdown.
Based on the above, it is concluded that RCS primary loop pipe breaks should not be considered in the structural design basis of the Seabrook plants.
8-1
9.0 REFERENCES
- 1. WCAP-9558, Rev. 2, " Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall Crack," [ -
Westinghouse Proprietary Class 2, June 1981.
_a
- 2. USNRC Generic letter 84-04,
Subject:
" Safety Evaluation of Westinghouse =-
7 Topical Reports Dealing with Elimination of Postulated Pipe Breaks in PWR ,_
Primary Main Loops", February 1,1984.
Primary Coolant Loop," Class 2. January 1975. . , ,
-1
- 4. Letter from Westinghouse (E. P. Rahe) to NRC (R. H. Vollmer),
NS-EPR-2768, dated May 11, 1983.
- 5. WCAP-9283, "The Integrity of Primary Piping Systems of Westinghouse --
Nuclear Power Plants During Postulated Seismic Events," Westinghouse -g Proprietary Class 2, March,1978.
- 6. WCAP-9787, " Tensile and Toughness Properties of Primary Piping Weld Metal for Use in Mechanistic Fracture Evaluation", Westinghouse Proprietary Class 2, May 1981.
4
- 7. Letter Report NS-EPR-2519, Westinghouse (E. P. Rahe) to NRC (D. G.
Eisenhut), Westinghouse Proprietary Class 2, November 10, 1981. l-a
- 8. Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated April ._
25, 1983.
- 9. Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated July 25, 1983.
=
- 10. NUREG-0691, " Investigation and Evaluation of Cracking Incidents in Piping !
in Pressurized Water Reactors", USNRC, September 1980.
9-1 _
- 11. Kanninen, M. F., et. al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks", EPRI NP-192, September 1976,
- 12. Bush, A. J., Stouffer, R. B., " Fracture Toughness of Cast 316 SS Piping Material tieat No. 156576, at 600'F", W R D Memo No. 83-5P6EVMTL-M1, Westinghouse Proprietary Class 2 March 7,1983.
l 'i . WCAP-10456, "The Effects of Themal Aging on the Structural Integrity of Cast Stainless Steel Piping For W NSSS," W Proprietary Class 2, November 1983.
- 14. Slama, G. , Petrequin, P. , Masson, S. H. , and Mager, T. R., "Ef fect of Aging on Mechanical Properties of Austenitic Stainless Steel Casting and Welds", presented at SMiRT 7 Post Conference Seminar 6 - Assuring Structural Integrity of Steel Reactor Pressure Boundary Components, August 29/30, 1983, Monterey, CA.
- 15. NUREG/CR.3464,1983, "The Application of Fracture Proof Design Methods using Tearing Instability Theory to Nuclear Piping Postulating Circumferential Through Wall Cracks"
- 16. Bamford, W. H., " Fatigue Crack Growth of Stainless Steel Piping in a Pressurized Water Reactor Environment", Trans. ASME Journal of Pressure Yessel Technology, Vol.101, Feb.1979.
a,c.e 17, 1
- 1 a,c.e 18.
~
- 19. Kumar, V., Geman, M. D., and Shih, C. F., "An Engineering Approach for Elastic-Plastic Fracture Analysis", EPRI, NP-1931.
9-2
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TABLE 2 FATIGUE CRACX GROWTH AT [ Ja,c.e (40 YEARS)
FINAL FLAW (in)
( )a,c.e .
INITIAL FLAW (IN) STEEL [ Ja.c.e [ ja.c.e 0.292 0.31097 0.30107 0.30698 0.300 0.31949 0.30953 0.31626 0.375 0.39940 0.38948 0.40763 0.425 0.45271 0.4435 0.47421 T-2
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in a Water Environment at 6000F F-8
I APPENDIX A 4.
LIMIT MOENT a,c.e I
l 4
=
l i
4 W
r A-1
M 4.C.A l
t 4
. 3 3
z .
z M
W 4
3 W
I 3
x s
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l 1
A-2
.