ML20112H524
| ML20112H524 | |
| Person / Time | |
|---|---|
| Site: | Harris  |
| Issue date: | 09/30/1984 |
| From: | Clark H, Lee Y, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML18018B847 | List: |
| References | |
| WCAP-10700, NUDOCS 8501170177 | |
| Download: ML20112H524 (42) | |
Text
.
WESTINGHOUSE PROPRIETARY CLASS 3 WCAP-10700 TECHNICAL BASES FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS A STRUCTURAL DESIGN BASIS FOR SHEARON HARRIS UNIT 1 September 1984 S. A. Swamy J. C. Schmertz Y. S. Lee R. A. Holmes H. F. Clark, Jr.
n APPROVED:
APPROVED: J. N.'Chirigos, Manag W E. R. Johnson, Manager Structural Materials Engineering Structural and Seismic l
Development 5bk1 JJML2 APPROVED:
l J. J. M(Iner'ney, Managef/
l Mechanical Equipment fd Systems l
Licensing l
WESTINGHOUSE ELECTRIC CORPORATION NUCLEAR ENERGY SYSTEMS P.O. Box 355 Pittsburgh, Pennsylvania 15230 8501170177 850114 PDR ADOCK 05000400 A
FOREWORD This document contains Westinghouse Electric Corporation proprietary information and data which has been identified by brackets. Coding associated with the brackets set forth the basis on which the information is considered proprietary. These codes are listed with their meanings in WCAP-721i.
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 1
information presented herein be safeguarded in accordance with 10 CFR 2.903.
Withholding of this information does not adversely affect the public interest.
l 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 information to such persons as part of the review procedure, clease 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.
O e
iii
TABLE OF CONTENTS j -
Section Title Pggt
1.0 INTRODUCTION
1-1 1.1 Purpose 1 -1 1.2 Scope 1-1 3
1.3 Objectives 1-1 1.4 Background Information 1-2 2.0 OPERATION AND STABILITY OF THE PRIMARY SYSTEM 2-1 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 GEOMETRY AND LOADING 3-1 i
4.0 FRACTURE MECHANICS 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 1
5.0 LEAK RATE PREDICTIONS S-1 6.0 FATIGUE CRACK GROWTH ANALYSIS 6-1
7.0 ASSESSMENT
OF MARGINS 7-1
~
8.0 CONCLUSION
S 8-1
~
9.0 REFERENCES
9-1 APPENDIX A- [
]**'
A-1 V
LIST OF TABLES Table Title Page 3-1 Shearon Harris Unit 1 Primary Loop Data 3-3 6-1 Fatigue Crack Growth at [
]C
6-3 6-2 RCL Design Transients 6-4 6-3 Inside & Outside Surface Stresses for Nonnal, Upset, and 6-5 Test Transients with Applied Pipe Loads Included 1
O 1
e 4
9 O
vii
LIST OF FIGURES Figure Title Page 3-1 Schematic Diagram of Primary Loop Showing Weld Locations 3-4 Shearon Harris Unit 1 3-2 Reactor Coolant Pipe 3-5 4-1
[
j*** Stress Distribution 4-7 4-2 J-aa Curves at Different Temperatures, Aged Material 4-8
[
]**' (7500 Hours at 400*C) 4-3 Critical Flaw Size Prediction 4-9 6-1 Cross-Section of [
3
6-6 6-2 Reference Fatigue Crack Growth Curves for 6-7 j.c.e a
6-3 Reference Fatigue Crack Growth Law for [
6-8
]b *** in a Water Environment at 600*F A-1 Pipe with a Through-Wall Crack in Bending A-2 ix
O l
1.0 INTRODUCTION
1.1 Purpose This report applies to the Shearon Harris Unit 1 plant reactor coolant system primary loop piping.
It is intended to demonstrate that specific parameters for the Shearon Harris plant are enveloped by the generic analysis perforned by Westinghouse in WCAP-9558, Revision 2 (Reference 1) (i.e., the reference report) and accepted by the NRC (Reference 2).
1.2 112E' The current structural design basis for the Reactor Coolant System (RCS) primary loop requires that pipe breaks be postulated as defined in the I
approved Westinghouse Topical Report WCAP-8082 (Reference 3).
In addition, 1
protective measures for the dynamic effects associated with RCS primary loop pipe breaks have been incorporated in the Shearon Harris plant design.
However, Westinghouse has demonstrated on a generic basis that RCS primary loop pipe 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 applicability of the generic evaluations to the Shearon Harris plant, Westinghouse has performed a comparison of the Shearon Harris plant loads and geometry with the envelope parameters used in the generic analyses (Reference 1), a fracture mechanics evaluation, a determination of leak rates f rom through wall cracks, a fatigue crack growth evaluation, and an assessment of margins.
1.3 Obiectives The conclusions of WCAP-9558, Revision 2 (Reference 1) support the elimination of RCS primary loop pipe breaks for Shearon Harris Unit 1.
In order to validate this conclusion the following objectives must be achieved.
~
a.
Demonstrate that the Shearon Harris plant parameters are enveloped by generic Westinghouse studies.
1-1
b.
Demonstrate that margin exists between the critical crack size and a postulated crack which yields a detectable leak rate.
l c.
Demonstrate that there is sufficient margin between the leakage through a postulated crack and the leak detection capability of the Shearon Harris plant.
d.
Demonstrate that fatigue crack growth is negligible.
1.4 Backaround 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). That Topical Report employed a deterministic fracture mechanics evaluation and a probabilistic analysis to support the elimination of RCS primary loop pipe breaks. This approach was then used as a means of addressing Generic Issue A-2 and Asymmetric LOCA Loads.
l Westinghouse performed additional tests and analyses 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.
l The NRC funded research through Lawrence Livermore National Laboratory (LLNL) to address this same issue using a probabilistic approach. As part of the j
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 Shearon Harris (References 8 and 9). The results from the LLNL study were released at a March 28, 1983 ACRS Subcommittee meeting. These studies, which are applicable to all Westinghouse plants east of-the Rocky 1-2
Mountains, determined the mean probability of a direct LOCA (RCS primary loop
-10 pipe break) to be 10 per reactor year ar.d the mean probability of an indirect LOCA to be 10- per retctor year. Thus, the results previously obtained by Westinghouse (Reference 5) were confirmed by an independent NRC research study.
Based on the studies by Westinghouse, LLNL, the ACRS, and the AIF, the NRC completed a safety review of the Westinghouse reports submitted to address asynnetric 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 not be considered for those plants that can 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 Shearon Harris Unit 1.
O O
l-3 s
I 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM The Westinghouse reactor coolant system primary loop has an operating history
' ~
that 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 Stress Corrosion Crackina 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 properly selecting a material 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 ASME Code rules, fracture toughness, welding, fabrication, and processing.
The environments known to increase the susceptibility of austenitic stainless steel to stress cerrosion are (Reference 10): oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g., sulfides, 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. External cleaning for Class 1 stainless steel piping includes patch tests to monitor and control chloride and fluoride levels.
During flushes and preoperational testing, water chemistry is controlled in accordance with written specifications.
2-1
?
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 normal power operation, oxygen concentration in the RCS is expected to be less than 0.005 ppm by controlling charging flow chemistry and maintaining 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 proper wetted surface materials.
2.2 Water Hammer Overall, there is a low potential for water hammer in the RCS since its design and operation precludes the voiding condition in normally filled lines.
Tne 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 hydraulic transients following valve opening are considered in the system design.
Other valve and pump actuations result in relatively slow transients with no sfgnificant effect on the system's dynamic loads. To ensure dynamic system stability, reactor coolant parameters are str'..c,r.ly controlled. Temperature l
during normal operation is maintained withir narrow range by control rod l
position; pressure is controlled by pre" e heaters and pressurizer spray also within a narrow range for s'teady-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 pump characteristics are controlled in the design process. Additionally, Westinghouse has instrumented typical reactor coolant systeam to verify the flow and vibration characteristics of the system.
Preoperational testing and
\\
2-2 i
~ - -
m.,._._,__._
.._y L.
m
s 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 Hiah Cycle Faticue Low cycle fatigue considerations are taken into account in the design of the piping system through the fatigue usage factor evaluation to show compliance with the rules of Section III of the ASME Code. A further evaluation of the low cycle fatigue loadings was carried out 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 shaf t vibrations during hot functional testing and operation. During operation, an alarm signals the exceedance of the vibration limits.
Field measurements have been made on a number of plants during hot functional testing, including plants similar to Shearon Harris. Stresses in the elbow region below the reactor coolant 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 appliec stress intensity factor below the threshold for fatigue crack growth.
9 2-3
3.0 PIPE GEOMETRY AND LOADING The loop weld locations for Shearon Harris are identified in Figure 3-1.
The material properties and the loads at these locations resulting from deadweight, thermal expansion, and Safe Shutdown Earthquake (SSE) are provided in Table 3-1.
The primary loop material is SA-376-TP304N. Fittings are made of SA-351-CF8A. As seen from this table, the junction of the hot leg and the reactor vessel outlet nozzle (Location 1) is the most limiting location for crack stability analysis based on the highest stress due to the combined pressure, deadweight, thermal expansion, and SSE loadings. A segment of the primary coolant hot leg pipe is sketched in Figure 3-2.
This segment is postulated to contain a circumferential through-wall flaw. This location will be referred to as the critical location. The inside diameter and the wall thickness of the pipe are 29.2 and 2.37 inches, respectively. At this location, the axial force (F,) and the bending moment (M ) are [
b J c.e (including the axial force due to pressure) and [
a
]C respectively. The pipe is subjected to a nor1aal operating pressure of 2235 psig. The method for calculating the loads found in Table 3-1 is described below.
The axial force F and transverse bending moments, M and M, are chosen y
7 for each static load (pressure, deadweight and thermal) based on
~
elastic-static analyses for each of these load cases. These pipe load components are combined algebraically to define the equivalent pipe static loads F,, My3, and M,. Based on elastic SSE response spectra analyses, g
amplified pipe seismic loads, F '
are btained. The maximum d
yd' zd pipe loads are obtained by combining the static and dynamic load components as fallows:
+ l F
I l
F, l F
=
d
[M 2
,g 2
M, =
y whore:
lM,l lM,I M
+
=
y y
y 3-1
z"
!M I
- IM I
M zs zd The corresponding geometry and loads used in the reference report (Reference
- 1) are as follows:
inside diameter and wall thickness are 29.0 and 2.5 inches; axial load and bending moment are [
]" # The outer fiber stress for Shearon Harris is [
]' '" #
while in the reference report it is [
] #
This demonstrates
. conservatism in the reference report which makes it more severe than the Shearon Harris analyses.
The normal operating loads (i.e., algebraic sum of pressure, deadweight, and 100 percent power thermal expansion loading) at the critical location, i.e.,
the junction of the hot leg and the reactor vessel outlet nozzle, are as follows:
F=[
]C (including internal pressure)
M-t J
The calculated and allowable stresses for ASME Code Section III, NB-3600 equation 9 (faulted, i.e., pressure, deadweight, and SSE) and equation 12 l
(thermal) at the critical location-are as follows:
i Calculated Allowable Ratio of Equation Stress Stress Calculated /
Number (ksi)
(ksi)
Allowable
~
~
~
arcre a,c.e 97
$g,y 12 52.8 O
L 3-2 i -
l e.
c, a+
ATAD POO L
YRAM I
1 R
3 P
E 1
L B
T I
AT NU S
I RRAH NO RAEHS erusserp l
n a
o n
i r
t e
a t
c n
o i
L l
e a
d c
u i
l t
c i
n r
I C
A I
"w
anactor Pressure
~ Vessel a)
=
h i
j
'- W
(
Anactor coolant Pop steso emnerator
.j===
f I
HOT LEG l
Temperature: 619'F; Pressure: 2235 psig CROSS 0VER LEG Temperature: 555'F; Pressure: 2202 psig COLD LEG Temperature: 555'F; Pressure: 2299 psig FIGURE 3-1 Schematic Diagram of Primary Loop Showing Weld Locations - Shearon Harris Unit 1 3-4 s
t 1
Crect M
2.37"
~
-_.1._
(*
F o r
1 1
a
.h Q
.a.--__---
p u
h29.2"M
~
2,235 psig P
=
a,c.e FIGURE 3-2 Reactor Coolant Pipe 3-5
1 4.0 FRACTURE MECHANICS EVALUATION 4.1 Global Failure Mechanism Determination of the conditions that lead to failure in stainless steel must be done with plastic fracture mechanics methods because of the large amount of deformation accompanying fracture. A conservative method for predicting the failure of ductile material is the [
] C # This methodology has been shown through a large number of experiments to be applicable to ductile piping and will be used here to predict the critical flaw size in the primary coolant piping. The failure criterion has been obtained by requiring [
j"'C# (Figure 4-1) 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 [
]"' C # for such a pipe is given by:
a,c.e g
3 where:
a,c.e usuinee MD 4-1
a,c.e The analytical model described above accurately accounts for the piping internal pressure as well as the imposed axial force as they affect [
aJ.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 terns of crack-tip blunting, initiatien, extension, and finally crack instability. Depending on the material properties and geometry of the pipe, flaw size, shape and loading, the local failure mechanisms say or may not govern the ultimate failure.
The ttability will be assumed if the crack does not initiate at all.
It has been accepted that the initiation toughness measured in terms of Jgy (i.e.,
gg)" from a J-integral resistance curve is a material parameter defining J
the crack initiation.
If, for a given load, the calculated J-integral value is shown to be less than J, of the material, then the crack will not g
initiate.
If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:
a The notation JIN instead of Jrc was used in Reference 1 to designate the value of the J-integral at crack initiation; the JIN notation will be used in this report in keeping with Reference 1.
1132E:10/092584 4-2
J 12 L_
7 app da 2
,f where:
T,,, = applied tearing modulus E = modulus of elasticity f=[
] # (flow stress) a a = crack length
[
3,c.e a
In swunary, the local crack stability will be established by the two-step criteria:
J<J IN 1
app
- mat, IN t
4.3 Material Pronerties 4
The materials in the Shearon Harris Unit 1 primary loops are wrought stainless steel pipe (SA-376-TP304N), cast stainless steel fittings (SA-351-CF8A) and associated welds. The tensile and flow properties at the critical location, the hot leg and the reactor vessel outlet nozzle junction, are given in Fig. 4-3, which will be discussed further in the next section. For this location, the material of interest is the wrought seamless pipe material (SA-376-TP304N). The fracture properties of this material are equivalent to the data reported in Reference 1.
The fracture properties of CF8A cast stainless steel have been determined through fracture tests carried out at 600*F and reported in Reference 12.
1 4-3 y
--e.
.4
_..n-.,mw,,,..,, -.,
.,y%,
..mw,-.,.,,,,,,_a
,.,._,._rm_mmm
._,.re.%..-.,-. ~.,.-,.c~e.
This reference shows that J for the base metal ranges from (
gy J.c.e for the multiple tests carried out.
a Cast stainless steels are subject to thermal aging during service. This thermal aging causes an elevation in the yield strength of the material and a degradation of the fracture toughness, the degree of degradation being proportional to the level of ferrite in the material. To determine the effects of thermal aging on piping integrity, a detailed study was carried out in Reference 13.
In that report, fracture toughness results were presented for a material representative of [
] C # Toughness results were provided for the material in the fully aged condition and these properties are also presented in Figure 4-2 of this report. The J, value for this material at operating temperature was g
approximately [
]C and the maximum value of J obtained in the tests was approximately [
]' ' ' d The tests for this material were conducted on small specimens and therefore rather short crack extensions resulted.
(Maximum extension was 4.3 mm.) Therefore it is expected that higher J values would be sustained for larger specimens.
[
),a,c.e l.
Therefore, it may be concluded that the degree of thermal aging expected by end-of-life for this plant is much less than that which was produced in
[
]# of Reference 13, and therefore the end-of-life J values IN for the Shearon Harris plant would be expected to be higher than those reported for [-
]C # in Figure 4-2 (also see Reference 14)'.
In addition, the tearing modulus for the Shearon Harris Unit 1 would be greater than [
].****'
Available data on stainless steel welds indicate that the J, values for the g
worst case welds are on 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
l tests (in excess of 3000 in-lbs/in ).
The applied value of the J-integral 4-4 4
-rm--
.-y
..-e
---,,---..,.-o-.w-.---,-
for a flaw in the weld region will be lower than'th'at in the base metal because the yield stress for the weld material is much higher at the operating temperature. Therefore, weld regions are less limiting than the cast material.
4.4 Results of Crack Stability Evaluation i
Figure,4-3 shows a plot of the [
]
as a function of through-wall circumferential flaw length in the [
] # of the main coolant piping. This [
l# was calculated for Shearon Harris using data for a pressurized pipe at 2235 psig with an axial force of [
]****', operating at 619'F with ASME Code minimum tensile properties.
The maximum applied bending moment of (
] # can be plotted on this figure. and used to determine a critical flaw length, which is shown i
to be [
). # This is considerably larger than the [
J# reference flaw used in Reference 1.
4 J-integral calculations were performed in Reference 1.
Based on the-calculationsitwasshownthata[
j c.e a
The pipe under present investigation is 2.37 inches thick with a 29.2 inch inside diameter. These[
]' d The axial load used in the present case is lower than that l
used in Reference 1.
The (
j.c.e a
percent of the moment load used in Reference 1.
The maximum outer fiber j
stress for Shearon Harris is only 82 percent of that of Reference 1.
[
]' # On this basis it is 2
1 judged that the conclusions of Reference 1 are applicable to the Shearon Harris primary loops. Actually, for the Shearon Harris loads and geometry, 2
the applied J was calculated by
]Cin-lb/in
~
which is significantly lower than the J value (Reference 1) for wrought IN value is nearly. equal to the J Of material.
In addition, this J
[
Jac.e forthe[pplied
],a,c.e max a
~
4-5 t
d
. _.... _. ~ _.... _. _... _ _ _.,. _ _ _ _.. _. _. _ _. _, _. _,. _ _
1 and is substantially less than J f[
]
for the weld material.
max T
frthe[
]a,c.e flaw as taken from Ref.13 is [
].a,c.e applied Further T f r the cast material would be greater than[ la.c.e As stated mat earlier, weld regions are less limiting than the cast material.
In addi-tion the applied value of J for a flaw in the weld will be lower than that in the base metal because the yield stress for the weld is much higher (esti-mated to be less than[ ]$li-lb/in ).
Therefore, it can be concluded that a postulated [
]a,c.e through-wall flaw in the Shearon Harris loop piping will remain stable from both local and global stab"ity standpoints.
A similar estimate was obtained for a [
Ja.c.e through-wall flaw. The purpose of the evaluation was to investigate the crack stability for a postu-lated flaw larger in size than the [
]a,c.e reference flaw.
For the Shearon Harris Unit 1 maximum moment of [
3a,c.e the maximum ap-plied J was estimated from available ja.c.e for the base metal. This is lower than even the 2
J valueof[
]" 'C # in-lb/in for
]a.c.e-inch flaw obtained applied in Reference 1.
For a postulated [
]baw in the weld material, the value is estimated from available[ 2 Japplied Ja,c.ein-lb/in and is of course lower than 2
the J-value of about [
3a,c.e in-lb/in obtained from fracture test on welds. Therefore, it is further concluded that a postulated [
ja.c.e through-wall flaw in the Shearon Harris Unit 1 primary loop piping will re-main stable from both local and global stability standpoint.
9 4-6
a,c,e III/H i
2.
/ww a.
Y)
Ly i
FIGURE 4-1
[
]Cd Stress Distrhtien e
4-7
FIGURE 4-2 J-aa Curves at Different Temperatures. Aged Material [
3ac.e (7500 Hours at 400*C) 4-8
a,c.e t I o
i FLAW GEOMETRY OD
= 33.94 in.
2.37 in.
t
=
p
= 2,235 psil l
l I
.J Figure 4-3 Critical Flaw Size Prediction l
4-9 l
l l
L
5.0 LEAK RATE PREDICTIONS Leak rate estimates were performed by applying the normal operating bending moment of [
]*** in addition to the normal operating axial force of (
]****' to the hot leg pipe containing a postulated
[
]' ' * # through-wall flaw.
The crack opening area was estimated using the method described in Reference 15. The leak rate was calculated-using the two-phase flow formulation described in Reference 1.
The computed a
leak rate was (
J c.e In order to determine the sensitivity of leak rate to flaw size, a through-wall flaw (
]
in length was postulated. The calculated leak rate was [
j.c.e a
The Shearon Harris Unit 1 plant has 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 [
]
inch flaw, a factor in excess of 160 exists between the calculated leak rate and the criteria of Regulatory Guide 1.45.
Relative to the [
]
flaw, a factor of over 75 exists.
e 0
)
5-1
6.0 FATIGUE CRACK GROWTH ANALYSIS To detennine the sensitivity of the primary coolant system to the' presence of small cracks, a fatigue crack growth analysis was carried out for the [
]a.c.e region of Shearon Harris Unit 1 [
]C
This region was selected because crack growth calculated here will be typical of that in the entire primary loop. Crack growths cal-culated at other locations can be expected to show less than 10% variation.
A[
]C of a plant similar in geometry and operational characteris-tics to the Shearon Harris plant. The transients which yielded the available stress results were in most instances more severe than the Shearon Harris Unit 1 transients.
For the few cases which were less severe, the stress results were ratioed upward to assure conservative results. [
3a.c.e i
All nonnal, upset, and test conditions were considered and circumferentially oriented surface flaws were postulated in the region, assuming the flaw was located in three different locations, as shown in Figure 6-1.
Specifically, these were:
Cross Section A:
a,c.e Cross Section B:
Cross Section C:
Fatigue crack growth rate laws were used [
]a.c.e The law for stainless steel was derived from Reference 16, with a very conservative correction for the R ratio, which is the ratio of minimum to maximum stress during a transient.
For stainless steel, the fatigue crack growth fonnula is:
6-1
h=(5.'4x10-12) g,,4.48 inches / cycle where K,ff = K,,, (1-R) U
" " min # max
[
a3,c.e a,c.e
[
)
where: [
] a,c.e The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is summarized in Table 6-1.
I Th'e results show that the crack growth is very small, regardless [
j j,c.e a
Table 6-2 lists the normal, upset, and test transients used to obtain the fatigue crack growth results presented in Table 6-1.
Table 6-3 lists the stresses from these transients at the inside and outside walls of the inlet nozzle safe-end, with the stresses due to the normal condition piping loads also included. At the inlet nozzle safe-end, these normal condition piping loads are [
]' # in the axial direction, combined with a bending
~
l moment of (
].
The Shearon Harris specific analysis described in this section was necessary due to more severe transients specified for Shearon Harris than for other Westinghouse pressurized water reactors.
7 6-2 I
TABLE 6-1 FATIGUE CRACK GROWTH AT [
]a,c.e(40 YEARS)
FINAL FLAW (in.)
INITIAL FLAW (in.)
a,c.e
{
3a,c.e
{
3a,c.e 0.120 0.12262 0.12135 0.12134 0.241 0.25161 0.24490 0.24657 0.361 0.37702 0.36969 0.37521 0.481 0.50329 0.49577 0.50737 O
6-3
_v 4-a
.._A-u a
- =_
rA_
m m
.A-4_
4 a__
a.
TABLE 6-2 RCL DESIGN TRANSIENTS NORMAL CONDITION NO. OF CYCLES a c.e -
I e
e s
0 e
e 6-4
TABLE 6-3 INSIDE AND OUTSIDE SURFACE STRESSES FOR NORMAL, UPSET. AND TEST TRANSIENTS WITH APPLIED PIPE LOADS INCLUDED TRANSIENT NAME Max.
Corres.
Min.
Corres.
Inside Outside Inside Outside Stress Stress Stress Stress (ksi)
(ksi)
(ksi)
(ksi) a,c.e m
O e
e 6-5
G.
U.
I I
h
=
U I
5 w
O 5
Y eft eft Eu
'Te b
4 i
e e
i 6-6
E d
Ee i
1 4
i 4a 1i 5
FIGURE 6-2 Reference Fatigue Crack Growth Curves For [
ga.c.e 6-7
-..,-.--,_---,---...e.e----,-,,_,-,,._....,,-,-,,---mm.
e-,
e,--,e, n-
,.me-,-
e-
i e,c, e 1
2 4
I
~
d i
i Figure 6-3 Reference Fatigue Crack Growth Law for [
},c.e in te Water Environment at 6000F i
6-0 8
0 m-
-,.--~.n-,,.
7.0 ASSESSMENT
OF MARGINS In Reference 1, the maximum design moment was [
]' #
whereas, the maximum moment as noted in Section 3.0 of this report is [
].a c.e The maximum 4
calculated value of J [
] # as compared with the value of [
J" # in Reference 1.
Furthermore, Section 4.3 shows the testing of fully aged material with chemistry more limiting than that existing in Shearon Harris Unit 1 cast piping extended to J values of [
]
which envelope the applied J values for cast elbows.
l As shown in Section 3.0, margins of factors of greater than 4.6 and 2.6 exist between calculated and ASME Code allowable faulted condition and thermal s
i stresses, respectively.
s Referring to Section 4.3, the estimated tearing modulus for Shearon Harris Unit 1 cast stainless steel elbows in the fully aged condition is at least [
].' # T,,, for the reference flaw as taken from,
Reference 13 is [
].
Consequently, a margin on local stability of at least 3 exists relative to tearing.
In Section 4.4, it is seen that a [
] d flaw in the base metal has a J value at maximum load of [
l
which is also enveloped by the J, of Reference 1.
Apostulated[
]Ilaw in the weld metal will also be stable.
In Section 4.4, the critical flaw size using[
3*'C methods,is calculated to be [
]a.c.e Based on the above, the critical flaw size will, of course, exceed [
]a,c.e In Section 5.0, it is shown that a flaw of [
]
would yield a leak rate in excess of [
]
while for a [
]' d flaw, the leak rate is [
3. # Thus, there is a margin of at least 4 between the flaw size that gives a leak rate satisfying the criterion of Regulatory Guide 1.45 and the " critical" flaw size of [
J.c.e a
7-1
In sumary, relative to 1.
Logds a.
Shearon Harris Unit 1 is enveloped both by the maximum loads and J values in Reference 1 and the J values employed in testing of fully aged material.
b.
At the critical locations, margins of 4.6 and 2.6 on faulted conditions and thermal stresses, respectively, exist relative to ASME Code allowable values.
2.
Flaw Size a.
A margin of at least 4 exists between the critical flaw and the flaw yielding a leak rate satisfying the criterion of Regulatory Guide 1.45.
b.
A margin exists of at least 3 relative to tearing.
~
c.
If [
]' # is used as the basis for critical flaw size, the margin for global stability would exceed 4 when compared to the reference flaw.
3.
Leak Rate A margin in excess of 160 exists for the reference flaw ([
] #) between calculated leak rates and the criteria of Regulatory Guide 1.45.
7-2
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 Shearon Harris Unit 1 plant as follows:
a.
The loads, material properties, transients, and geometry relative to the Shearon Harris Unit 1 RCS primary loop are enveloped by the parameters of WCAP-9558. Revision 2 (Reference 1) and WCA'P-10456 (Reference 13).
b.
Stress corrosion cracking is precluded by the 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.
e.
A large 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.
e 8-1
l l
The reference flaw will be stable throughout reactor life because of the ample l
margins in e, f, and g above and will leak at a detectable rate which will hssure a safe plant shutdown.
Based on the above, it is concluded that RCS primary loop pipe brWaks should not be considered in the structural design basis for Shearon Harris Unit 1.
J i
- 'w g
1 9
l l
l s-2 i
l L
i
9.0 REFERENCES
l 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.
I 2.
USNRC Generic letter 84-04,
Subject:
" Safety Evaluation of Westinghouse Topical Reports Dealing with Elimination of Postulated Pipe Breaks in PWR Primary Main Loops", February 1,1984.
1 3.
WCAP-8082 P-A, " Pipe Breaks for the LOCA Analysis of the Westinghouse Primary Coolant Loop," Class 2, January 1975.
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," 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.
7.
Letter Report NS-EPR-2519 Westinghouse (E. P. Rahe) to NRC (O. 6.
Eisenhut), Westinghouse Proprietary Class 2, November 10, 1981.
8.
Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated April l
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 I
i
- 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., Stouf fer, R.
B., "Fractere Toughness of Cast 316 SS Piping Material Heat No. 156576, at 600*F", Westinghouse R D Memo No.
83-5P6EVMTL-M1, Westinghouse Proprietary Class 2, March 7,1983.
- 13. WCAP-10456, "The Ef fects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping For Westinghouse NSSS," Westinghouse 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, 'The Application of Fracture Proof Design Methods using Tearing Instability Theory to Nuclear Piping Postulating Circumferential Through Wall Cracks' 1983.
- 16. Bamford, W. H., " Fatigue Crack Growth of Stainless Steel Piping in a Pressurized Water Reactor Environment", Trans. ASME Journal of Pressure Vessel Technology, Vol.101. Feb.1979.
17.
(
3,C,e 18.
[
a,c,e
)
9-2 L
A L_.*_
- 4,,
a-e I
l 1
APPENDIX A s
M m i4,Cet O
e 1
e j
G O
e W
6 i
M W
e O
A-1
-, -. - - -. + -, - <
m,--e..ew
- m --,.
-,--+.-%-
-,w--
,m,
---e
.p- - - -
-4
+a,c.e i
t I
t i
FIGURE A-1 Pipe with a through-wall crack in bending e
~
i e
l f
l i
A-2
- - ~ - - - -..
. _. _ _ _. _ _ _ _ _ _., _