ML20090K673
| ML20090K673 | |
| Person / Time | |
|---|---|
| Site: | Vogtle  |
| Issue date: | 05/31/1984 |
| From: | Clark H, Lee Y, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19269A161 | List: |
| References | |
| WCAP-10551, NUDOCS 8405240315 | |
| Download: ML20090K673 (35) | |
Text
_ _
WCAP 10552 e
)-
TECHNICAL BASES FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUC-TURAL DESIGN BASIS FOR ALVIN W. V0GTLE UNITS 1 AND 2 May 1984 S. A. Swamy Y. S. Lee H. F. Clark, Jr.
R. A. Holmes APPROVED:
C APPROVED:
n J.N.Chirigos,Manageh E.RfJohnson, Manager Structural Materials Engineering Structural and Seismic-Development h[,,. h 6 APPROVED:
JbJ.McInerney,'Madger Mechanical Equipment and Systems Licensing Work perfonned under Shop Order EEAJ 950 WESTINGHOUSE ELECTRIC CORPORATION NUCLEAR ENERGY SYSTEMS P.O. Box 355 Pittsburgh, Pennsylvania 15230 A
m 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 proprieta ry. 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 information to such persons as part of the review procedure, please contact Westinghouse Electric Corporation, which will make the necessary arrangements required to protec.t, the Corporation's proprietary interests.
The proprietary information is deleted in the unclassified version of this i
report.
9 9
d 111
TABLE OF CONTENTS Section Title Page
~
1.0 INTRODUCTION
1-1 1.1~ Purpose 1-1 1.2 Scope 1-1 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 GE0 METRY AND LOADING 3-1 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 4-4 5.0 LEAK RATE PREDICTIONS 5-1 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- [.
Ja,c.e A-1 y
LIST OF TABLES
' Table Title Page
~
4 3-1 Vogtle Primary Loop Data 3-3 6-1 Fatigue Crack Growth at [
.]a,c.e 6-3 f
J t
1 e
6 I
J e
. S:
vii i
__..,m..-.-.._
LIST OF FIGURES Figure Title Page 3-1 Reactor Coolant Pipe 3-4 3-2 Schematic Diagram of Primary Loop Showing Weld Locations 3-5 4-1
[
-]a,c.e Stress Distribution 4-7 4-2 J-aa Curves at Of fferent Temperatures, Aged Material 4-8
[.
Ja,c e (7500 Hours at 400*C) 4-3 Critical Flaw Size Prediction 4-9 6-1 Typical Cross-Section of [
]a,c e 6-4 6.2 Reference Fatigue Crack Growth Curves _ for 6-5
[
3a,c.e 6-3 Reference Fatigue Crack Growth Law for [
6-6
]a,c.e in a Water Environment at 600*F 1
A-1 Pipe with a Through-Wall Crack in Bending A-2 I
I e
l ix
1.0 INTRODUCTION
1.1 Purposa This report applies to the Alvin W. Vogtle (Vogtle) plant reactor coolant system primary loop piping.
It is intended to demonstrate that specific
~
parameters for the Vagtle plants are enveloped by the generic analysis perfonned by Westinghouse in WCAP-9558, Revision 2 (Reference 1) (i.e., the reference report) and accepted by the NRC (Reference 2).
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 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 Vogtle plant design.
- However, Westinghouse has demonstrated on a generic basis that RCS primary loop pipe breaks are highly unlikely and s.hould not be included 1n the structural design
~
In order to demonstrate this basis of Westing. house plants (see Reference 4).
applicability of the generic evaluations to the Vogtle plants, Westinghouse has perfonned a comparison of the loads and geometry for the Vogtle plants with envelope parameters used in the generic analyses (Reference 1), a fracture mechanics evaluation, a determination of leak rates from a through-wall crack, a fatigue crack growth evaluation, and an assessment of
. margins.
1.3 Objectives The conclusions of WCAP-9558, Revision 2 (Reference 1) support the elimination of RCS primary loop pipe breaks for the Vogtle plants.
In order to validate this conclusion the following objectives must be achieved.
a.
Demonstrate that Vogtle plant parameters are enveloped by generic Westinghouse studies.
1-1
b.
Demonstrate that margin ex.... between the critical crack size and a postulated crack which yields a detectable leak rate.
Demonstrate that there is sufficient margin between the leakage through a c.
postulated crack and the leak detection capability of the Vogtle plant.
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). That Topical Report employed a deterministic fracture mechanics evalua' tion and a probabilistic analysis to support the elimination of RCS primary loop pipe breaks. That 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) to address this same issue using a probabilistic approach. As part of the LLNL research effort, Westinghouse performed axtensive evaluations of specific plant loads, material properties, transients, and system geometries to demonstrate that the analysis and testing previously performed hy Westinghouse and the research performed by LLNL applied to all Westinghouse plants l
including Vogtle (References 8 and 9). The results from the LLNL study were released at a March 28, 1983 ACRS Subcommittee meeting. These studies which l
are applicable to all Westinghouse plants east of the Rocky Mountains, determined the mean probability df 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 by Westinghouse (Reference 5) were confirmed by an independent NRC research study.
1-2
Based on the studios by Westinghouse, LLNL, the ACRS, and the AIF, the NRC completed a safety review of the Westinghouse reports submitted to address 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 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 the Vogtie plants.
o e
(
l l
l G
i 0
l l-3
2.0 OPERATION AND STABILITY OF THE REACT 0.R 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. Stress Corrosion Cracking 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 l
(Reference 10). Since some residual stresses and some degree of material l-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 4
- (both internal and external) as well as other materials in the system, applicable ASE Code rules, fracture toughness, welding, fabrication, and processing.
l The environments known to increase the susceptibilty of austenitic stainless
~
steel to stress corrosion 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. During flushes and preoperational testing, water chemistry is controlled in accordance with written specifications.
External cleaning for Class 1 stainless steel piping includes patch tests to monitor and control chloride and fluoride levels. For 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 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 j
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 hy controlling charging. flow chemistry and specifying proper wetted surface materials.
2.2 Water Hammer Overal.1, there is a low potential for water hammer in the RCS since it is designed,and operated to' preclude the voiding condit,fon in normally filled 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 hydraulic transients following valve opening are considered in the systen I
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 coolent parameters are stringently controlled. Temperature during normal operation is maintained within a narrow range by control rod position; pressure is controlled 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 l
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 systems to verify the c
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 l
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 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 shaft 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 Vogtle.- Stresses in the elbow below the reactor coolant pump have j
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 l
also result in an applied stress intensity factor below the threshold for l
fatigue crack growth.
l l
l l-1 2-3
3.0 PIPE GEOMETRY AND LOADING j
A segment of the primary coolant hot leg pipe shown below to be limiting is i
sketched in Figure 3-1.
This segment is postulated to contain a circumferential through-wall flaw. The inside diameter and wall thickness of the pipe are 29.0 and 2.45 inches, respectively. The pipe is subjected to a normal operating pressure of
[.
]a,c.e psi. Ficure 3-2 identifies the loop weld locations. The material properties and the loads at these locations resulting from deadweight, themal expansion and Safe Shutdown Earthquake are indicated in Table.3-1.
As seen from this table, the junction of the hot leg and the reactor vessel outlet nozzle is the worst location for crack stability analysis based on the highest stress due to combined pressure, dead weight, themal expansion, and SSE (Safe Shutdown Earthquake) loading. At this location, the axial load (F ) and the bending moment (M ) are [
x b
la.c,e (including axial force due to pressure) and [
3a,c.e, respectively. This location will be referred to as the critical location. The loads of Table 3-1 are calculated as follows:
The axial force F and transverse bending moments, M and M, are chosen 1
y z
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 4
loads F ' Mys, and Mzs.
Based on elastic SSE response spectra analyses, s
amplified pipe seismic loads, F ' Myd' Mzd are obtained. The maximum d
pipe loads are obtained by combining the static and dynamic load components as follows:
x=
lF l F
+
F s
d M
/M 2g 2 b=
y z
where:
y" INys!
lMyd!
N z"
!Mzsj lMzd !
M
+
'f 3-1
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 [-
Ja,c.e inch-kips. The outer fiber stress for Vogtle is [
]a,c.e ksi, while for the reference report it is [
]a,c.e ksi. This demonstrates conservatism in the reference report which makes it more severe than the Yogtle analyses.
The normal operating loads (i.e., algebric 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=L ja,c.e The calculated and allowable stresses for ASME Code equation 9 (faulted i.e.,
pressure, dead weight and SSE) and equation 12 (thermal) at the critical location i
are as follows:
Calculated Allowable Ratio of l
Equation Stress Stress Calculated /
4 Number (ksi)
(ksi)
Allowabl e a,c.e l
1
~
l l
i I
i 3-2.
TkBLE3-1 l
V0GTLE PRIMARY LOOP DATA a,c.e I
w I
i i
a The Highest Stressed Location b
Includes the axial load due to internal pressure
I l
l l+
CRACK I
g4 ig 2.45 180.I g
-~ ~
R T'
< ' = >
I r=h k
r i//
i% f; a
M i
l.
4-29 i n. -*l a,c.e FIGURE 3-1 REACTOR COOLANT PIPE 9
5
Resc:cy.
Pressure Vessel CCLD L g HOT Lgg 2
f Q
'f 1(
i, di n Pump Steam Generator CROSS-0VER LEG d
Pa I
I.
D (b
e a,C,e b
FIGURE 3-2 SCHEMATIC DIAGRAM 0F PRIMARY LOOP SHOWING WELD LOCATIONS 3-5
4.0 FRACTURE MECHANICS EVALUATION 4.1 Global Failure Mechanism 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 [
3a,c.e This methodology has been shown to be applicable to ductile piping through a large number of experimen',s and will be used here to predict the critical flaw size in the primary coolant piping. The failure criterion has been obtained by requiring C
]a,c e (Figure 4-1) when loads are applf'ed. 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 tur:
g 3
a,c.e where:
a,c.e 4,
d W
G 4-1
-c a, c.e The analytical model described above accurately accounts for the piping internal pressure as well as imposed axial force as they affect [
]a,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 tenms 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 gpvern the ultimate failure.
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 J II****
IN JIc) from a J-integral resistance curve is a material parameter defining the crack initiation.
If, for a given load, the calculated J-integral value i
is shown to be less than J f the material, then the crack will not IN i nitiate.
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 J c was used in Reference 1 to designate I
the value of the J-integral at crack initiation; the JIN notation will be used in this report in keeping with Reference 1.
4-2
dJ E
T,pp = 77, q
'f where:
c.
T,pp = applied tearing modulus E = modulus of elasticity of = [
.]a,c.e (flow stress) a = crack length I
[
ja,c.e In summary, the local crack stability will be established by the two-step criteria:
J<J IN T,pp < T I
mat IN 4.3 Material Prope'rties The materials in the Vogtle Units 1 and 2 primary loops are cast stainless steel (SA 351 CF8A) and associated welds. The tensile and flow properties of the critical location, the hot leg and the reactor vessel nozzle junction, are given in Table 3-1.
The fracture properties of CF8A cast stainless steel have been determined through fracture tests carried out at 600*F and reported in Reference 12.
Th'is reference shows that J for the base metal ranges from [
yy 3a,c.e for the multiple tests carried out.
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 4-3
in Reference 13.
In this report, fracture toughness results were presented for a material representative of [
]a,c.e 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 for information. The J value for this material at operating IN temperature was approximately [
3a,c.e and the maximum value of J obtained in the tests was in excess of [
3a,c.e The tests of this material were conducted on small specimens and therefore rather short crack extensions, (maximum extension 4.3 m) so it is expected that higher J values would be sustained for larger specimens.
The'effect of the aging process on loop piping integrity for Vogtle was addressed in Reference 13, where the plant specific material chemistry for all the loop materials was considered [
].a,c.e This reference shows that the degree of thermal aging expected by end-of-life for these units is less than that which was produced in [
].a.c.e Therefore the JIN**I"*8 I I the Vogtle Units 1 and 2 at end-of-life would be expected to be considerably higher than those reported for [
3a,c.e in Figure 4-2 (also see
~
Reference 14).
In addition, the tearing modulus for the Vogtle Units 1 and 2 materials would be greater than [
]a,c,e Available data on stainless steel welds indicate the J values for the IN worst case welds are of the same order as the aged material, but the slope of l
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 the J-integral for a flaw in the weld regions will be lower than that in the base metal because the yield stress for the weld materials is much higher at L
temperature. Therefore, weld regions are less limiting than the cast material.
l 4.4 Results of Crack Stability Evaluation Figure 4-3 shows a plot of the [
3,c.e as a function of a
through-wall circumferential flaw length in the [
3a,c.e of the main
~
coolant piping. This [
]a,c.e was calculated for Vogtle from data for a pressurized pipe at [
]a,c,e properties. The maximum applied bending moment of [
3a,c.e in-kips can be plotted on 4-4
sc
\\
\\
,4 this figare, and used ito d5termina a critical flaw length, which is shown to be [
la,c.e inches.NTMs is considerably la'rger than the [
Ja,c.e inch reference flaw used in Refarence 1.
[.
'~
s y
'N i
\\:,
]bCJTheaxialloadus(e'd in the present case is 9 percent higher
(
than that used'in Reference 1.
Hcwever, the ['
]a,c.e percent of'the moment load used in Refereace 1.
The maximum outer fiber stress for Vogtle is only 75 percent of that of Reference 1-E s.h _%
3a,c.e On this basis it is judged that the conclusion's of Reference 1.are applicable to the s
Vogtle primary. loops. Specifically, it can b'e concluded that a postulated
[
]a,c.e inch through-walt flaw in the Vogtle loop piping will remain stable fr'm both a local and global. stability standpoint.
o
.+
Actually for,the Yogtle loads, the applied J was estimated to be 2
[
}"'C in-lbs/in which exceeds the J for the [
IN
]a,c.e by [
]. a,c.e For the actual Vogtle piping mate'rf al, initiation would not be expected.
T A[t
]a,c e analysis was performed for a q
[
[
- 3a, c, e through-wall flaw using the same approach and material I
propeNties; described in detail 'in Rpference 1.
The' purpose of this N'
calculation $s to investigate the' crack sta!5flity for a postulated flaw larger'in. size than the' [
]C reference flaw.
For the Vogtle Units 1 and 2. maximum moment of [
]a,c.e the maximum applied J was estimated to be [.
]a,c.e The applied tearing 4
\\
s
's 4-5 s
3 i
(
i e 1
modulus, T,pp, 'was calculated using the methodology of Reference 13 and was less than [ <] a,c.e which is significantly less than T for even the worst mat case [
-] a,c.e material of Reference 13.
Therefore it is further con-cluded that a postulated [
] a,c.e through-wall flaw in the Vogtle Units 1 and 2 primahy loop piping will remain stable from both a local and global stability standpoint.
In order to investigate the additional sensitivity of J to flaw size, a [
] a,c e analysis was performed for a [
] a,c.e through-wall flaw. Here, the maximum applied J was found to be less than 2
[
.]a,c.e in-lb/in,
S 6
5 4-6
f
,, a
,s,
/
i t
t F
y 1
i j'
+
.8 t
+
i I
Y p
/
\\
5 s.
,/
t f
7 a,c.e
////////////
I s
3 2a
)
NeutralAxis
--r-k 1
.s 5
o
+
2 6
I r
. i f [
~
s 5
i
?
t.
i
+
ll l
\\
1 p
,r
' t
'. i ',
'4 g,
('
. 4 x
j j
i e
a 4
- i_.
i
['
] STRESS DISTRIBUTION u,c.e
'i.
8' FIGURE 4-l 5
?
s_,
A
/
g i
t-
. [ ',
/
f 9
if f
a e
f e
te 4-i k
/
3 e,
i 1
(
I
.T
MJ i..
i L.
l-a,c,a 5
t L
N-q 2
E,
[-
!^
l-l.
FIGURE 4-2 J-da CURVES AT DIFFERENT TEMPERATURE. AGED MATERIAL [
3a,c,e (7500HOURSAT400C) 48
a,c.e FLAW j
k) D MD a
1 l
l i
l 1
1 i
I e
i i
4 O
O M
FIGURE 4-3 CRITICAL FLAW SIZE PREDICTION 4.9
5.0 LEAK 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 [
]a,c.e These loads were applied to the hot leg pipe containing a postulated [
]a,c e through-wall flaw and the crack opening area was estimated using the method of Reference 15. The leak rate was calculated using the two-phase flow formulation described in Reference 1.
The computed leak rate was [
]a,c.e In order to determine the sensitivity of leak rate to flaw size, a through-wall flaw [
]a,c.e in length was postulated. The calculdted leak rate was [
3a,c.e The Vogtle 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 [
3a,c.e inch flaw, a factor in excess of [ ]a,c.e exists between the calculated leak rate and the criteria of Regulatory Guide 1.45.
Relative to the [
3a,c.e 9
4 9
5-1 i
1
6.0 FATIGUE CRACK GROWTH ANALYSIS To determine the sensitivity of the primary coolant system to the presence of siaall cracks, a fatigue crack growth analysis was carried out for the [
l
]a,c e region of a typical system (see Location 12 of Figure 3-2).
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 [ finite element stress analysis was carried out for the inlet nozzle safe 4
end region]a,c.e of a plant typical in geometry and operational characteristics to any Westinghouse PWR' System. [
]a,c.e All normal, 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:
4 a,c.e Cross Section A:
Cross Section B:
Cross Section C:
1-Fatigue crack growth rate laws were used [
3,c.e The law for stainless steel a
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 formula is:
0 3
6-1
4.48 h = (5.4 x 10-12) g inches / cycle eff where K,ff = K,,x (1-Rb R = K,9,/K,,,
l l
'E t
ja c.e a,c.e L
]
where: [
3'*****
The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is surmarized in Table 6-1, and shows that the crack growth is very small, regardless [
3a,c.e 4
6-2
TABLE 6-1 FATIGUE CRACK GROWTH AT [I
]a,c.e (40 YEARS)
FINAL FLAW (in)
~a,c.e INITIAL FLAW (IN)
[
]a,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 i
1 n
a S
O 6-3
+
v--
---,e,
,..-,----,.----,-------n---------,---,-.,,----m-
-an n,
k i
a.C.e l
i l
I m.
b i
i I
i a
i l
i l
i 1
I FIGURE 6-1 TYPICAL CROSS SECTION OF [
]fa,c.e 4
4 S
- a.c. e
-w 4u>us as Wzo E
e9 E
z s
N 4
a e
z l
CeCwo4eu t
4 Figure 6-2 Reference Fatiq0e Crack Growth Curves for [
e a,c e e
i' 6-5
.m,,.
y
-..y,,,-.------,-2
C8/38/8942 2 a,c, e L
4 L-I-
t i.
Reference.Fati'g'eCrackGrowthlawfori Figure 6-3' u
j.
Ja,e e in a Water Environment at 600F i
f i
6-6 l-g
7.0 ASSESSMENT
OF MARGINS In Reference 1, the maximum design moment was [
]a,c.e in-kips, whereas, the maximum moment as noted in Section 3.0 of this report is [
]a,c e in-kips). The maximum value of J [
3a,c.e in-lb/in2 as detemined by [
]a,c.e analyses compared with the value of [
]a,c e in-lb/in2 in Reference 1.
Furthermore, Section 4.3 shows that for fully aged material of chemistry worse than that existing in Vogtle 2
cast piping J is [
3a,c.e in-lb/in. This value is very near IN the applied J for the maximum moment on the Vogtle piping containing a
[
]a,c.e inch flaw; thus initiation would not be anticipated for Vogtle piping under these conditions. The test results for the worst case material show J-values to near [
]a,c.e in-lb/in3 which is almost a factor of [ ]a,c.e greater than the applied value.
As shown in Section 3.0, a margin of a factor of [ ]a,c.e exists between calculated and ASME Code allowable faulted condition and thermal stresses.
l Referring to Section 4.3, the estimated' tearing modulus for Vogtile Units 1 and 2 cast _SS piping in the fully aged condition is at least [
l
]. a,c.e T,pp for the reference flaw as taken from Reference 13 is [
] a,c.e Consequently, a margin on local stability of at least
[]a,c.e. exists relative-to tearing.
l' InSection4.4,itisseenthata[
] a,c.e flaw has a J value at 2
maximum load of [
]a,c.e in-lb/in which is also enveloped by l
the J,,x of Reference 1 and the value used for testing of aged material.
l T
is again less than [
]"'
In Section 4.4, the critical flaw size using f
[
]a,c.e methods is calculated to be [
] a,c.e inches.
Based f
on the above, the critical flaw size will, of course, exceed ['
] a,c.e f
7-1
In Section 5.0, it is shown that a flaw of [
3,c.e would yield a a
leak rate in excess of [
]a,c.e while for a'[
] a,c.e inch flaw, the leak rate is [
] a,c.e Thus, there is a margin of [ ] a,c, between the flaw size that gives a leak rate of [
]a,c e and the " critical" flaw size of [
]a,c.e In summary, relative to 1.
Loads a.
Vogtle Units 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.
b.
Margins at the critical location of [ la.c.e on faulted conditions and thermal stresses exist relhtive to ASME Code allowable values.
~
2.
Flaw Size A margin of at least. [ ]a c.e exists between the critical flaw and
.a.
the flaw yielding a leak rate of [
]a,c.e b.
A margin exists of at least [ ]a,c.e relative to tearing.
c.
If [
]a,c.e is used as the basis for critical flaw size, the margin for global stability would exceed [
3a,c.e 3.
Leak Rate A margin in excess of [
3a,c.e exists for the reference flaw ([
3a,c.e) between calculated leak rates and the criteria of Regulatory Guide 1.45.
4 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 Vogtle plants as follows:
The loads, material properties, transients, and geometry relative to a.
the Yogtle Units 1 and 2 RCS primary loop are enveloped by the parameters of WCAP-9558, Revision 2 (Reference 1) and WCAP-10456 (Reference 13).
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 nonnal operation.
O Water hammer should not occur in the RCS piping because of system c.
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.
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 Vogtle plants.
8-1
9.0 REFERENCES
1.
WCAP-9558, Rev. 2, " Mechanistic Fracture Evaluation of Reactor Cool. ant Pipe Containing a Postulated Circumferential Through-Wall Crack,"
Westinghouse Proprietary Class 2, June 1981.
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.
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. Yollmer),
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 (D. G.
Eisenhut), Westinghouse Proprietary Class 2, November 10, 1981.
8.
Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnson) 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.
O j
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 Heat No. 156576, at 600*F", )[ R and D Memo No. 83-5P6EVMTL-M1, Westinghouse Proprietary Class 2, March 7,1983.
- 13. WCAP-10456, "The Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping For }[ NSS3," }[ 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 Instabi.11ty, 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 Vessel Technology, Vol.101, Feb.1979.
l a,c.e 17.
18.
I l
9-2
,E a
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- 6p__J, hm--AAa. & W c_.. haw A W - 4 h4 2
A41 4-4L A a 544_424 MeSJ bM4.d.ms < Ad a4
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- 4. '
APPENDIX A h
M ema i E
4,C,e 5
g t --
t 1
O.
I r
t
.g.
P e.
L
(
l h
h i
i f
1 i
- O P
t f
L me M
s t
i f
I i
t I
f 4
(.
I A-1 e
---.a.
1
- a,c.e FIGURE A-1 PIPE WITH A THROUGH-WALL CRACK IN BENDING A-2
-