ML20214V237
| ML20214V237 | |
| Person / Time | |
|---|---|
| Site: | Beaver Valley |
| Issue date: | 03/31/1987 |
| From: | Mutyala B, Palusamy S, Witt F WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19292H353 | List: |
| References | |
| WCAP-11318, NUDOCS 8706120030 | |
| Download: ML20214V237 (67) | |
Text
WESTINGHOUSE PROPRIETARY CLASS 3 WCAP-11318 TECHNICAL JUSTIFICATION FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR BEAVER VALLEY UNIT 1 MARCH 1987 D. H. Roarty S. A. Swamy J. C. Schmertz T. H. Liu Y. S. Lee D. Lindgren VERIFIED:
APPROVED:
M
/
F. 4. Witt
- 5. 5. P usimy,' Manager Strue ral Materials Engineering APPROVED:(([//
7 APPROVED: M M[
i 5.B.'Mutyalf, Manager C. W. Hirst, Manager Piping Desicfn and Mechanical Equipment and Qualification Systems Licensing
\\
l WESTINGHOUSE ELECTRIC CORPORATION Generation Technology Systems Division P. O. Box 2728 Pittsburgh, Pennsylvcnia 15230-2728 ON $[000$34 PDA
TABLE OF CONTENTS Section Title Paoe 1.0
SUMMARY
AND INTRODUCTION 1-1 1.1 Summary 1-1 1.2 Introduction 1-1 1.2.1 Purpose 1-1 1.2.2 Scope 1-2 1.2.3 Objectives 1-2 1.2.4 Background Information 1-2 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT 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 GEONETRY 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 4-5 l
5.0 LEAK RATE PREDICTIONS 5-1 5.1 Introduction 5-1 5.2 General Considerations 5-1 5.3 Calculation Method 5-1 5.4 Leak Rate Calculations 5-2 6.0 FATIGUE CRACK GROWTH ANALYSIS 6-1
7.0 ASSESSMENT
OF MARGINS 7-1 iv l
TABLE OF CONTENTS (Cont'd.)
Section' Title Page
8.0 CONCLUSION
S 8-1
9.0 REFERENCES
9-1 APPENDIX A - Limit Moment A-1
' APPENDIX B - Alternate Toughness Criteria for the B-1 Beaver Valley Unit 1 Cast Primary Loop Components B.1 Introduction B-1 B.2 Chemistry and KCU' Toughness B-1 B.3 The As-Built Beaver Valley Unit 1 Loops B-1 B.4 Alternate Toughness Criteria for the B-2 Beaver Valley Unit 1 Cast Primary Loop Material on a Component-by-Component Basis 1
V
LIST OF TABLES Table Title Page 3-1 Beaver Valley Unit 1 Primary Loop Data Including Faulted Loading Conditions 3-3 3-2 Normal Condition (Dead Weight + Pressure + Thermal)
Loads for Beaver Valley Unit 1 3-4 4-1 Fracture Toughness Criteria Used in the Leak-Before-4-8 Break Evaluation 6-1 Fatigue Crack Growth at (Nozzle Safe-End Region]a,c.e 6-3 (40 Years) 7-1 Summary of J,pp and Leak Rate Results as a Function 7-3 of Crack Length at the Four Critical Locations B-1 Chemical and Physical Properties of Beaver Valley Unit 1 Primary Loop Material - SA 351/CFBM B-4 B-2 Fracture Toughness Criteria for the Cast Primary Piping Components of the Beaver Valley Unit 1 Nuclear Plant B-5 vi
-?
LIST OF FIGURES Figure Title Pm 3-1 Reactor Coolant Pipe 3-5 3-2 Schematic Diagram of Primary Loop Showing Weld Locations 3-6
- Beaver Valley Unit 1 4-1
[
]a,c.e Stress Distribution 4-9 4-2 J vs aa for SA351 CF8M Cast Stainless Steel at 600"F 4-10 4-3 J-Aa Curves at Different Temperatures, Aged Material 4-11
[
]a,c.e (7500 Hours at 400*C) 4-4
" Critical" Flaw Size Prediction - Hot Leg at 4-12 Critical Location 1 4-5
" Critical" Flaw Size Prediction - Crossover Leg at 4-13 Critical Location 5 4-6
" Critical" Flaw Size Prediction - Crossover Leg at 4-14 Critical Location 10-4-7
" Critical" Flaw Size Prediction - Cold Leg at 4-15 Critical Location 11 l
5-1 Analytical Predictions of Critical Flow Rates of 5-4 l
Steam-Water Mixtures 5-2
[
]a,c.e Pressure Ratio as a 5-5 Function of L/D 5-3 Idealized Pressure Drop Profile Through a 5-6 Postulated Crack vii
LIST OF FIGURES (Cont'd.)
Figure Title Page 6-1 Typical Cross-Section of (
Ja,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 Ja,c.e in a Water Environment at 600*F
. A-1 Pipe with a Through-Wall Crack in Bending A-2 B-1 Typical Layout of the Primary Loops for a Westinghouse B-6 Three-Loop Plant With Isolation Valves IdentificakionofHeatswithLocationforColdLeg B-7 B-2 B-3 Identification of Heats with Location for Hot Leg B-8 i
B-4 Identification of Heats with Location for Crossover Leg B-9
)
1 I
viii
1.0
SUMMARY
AND' INTRODUCTION 1.1 Summary The original structural design basis for the Beaver Valley Unit I reactor coolant system primary loop required that the effects of pipe breaks be considered. However such breaks have been shown to be highly unlikely on a generic basis and the Nuclear Regulatory Commission has revised criteria which allow exclusion of dynamic effects from the design basis on a plant specific basis.
In this report the applicability of the generic evaluations to the Beaver Valley Unit 1 piping system is demonstrated by presenting a fracture mechanics evaluation, a determination of leak rates from a through wall crack, a fatigue
- crack growth evaluation and an assessment of margins.
Major emphasis is on the pipes which are~ limiting from the standpoint of toughness. Geometries, loadings and heat chemistries are summarized. Fracture toughness values are established for each pipe using the alternate toughness criteria approach.
Fracture mechanics and leak rate calculations showed that acceptable margins I
i exist between cracks which are stable and those for which detectable leak rates are demonstrated.
j This report demonstrates that the reactor coolant system primary loop pipe breaks need not be considered in the structural design basis of the Beaver Valley Unit 1 plant in accordance with the revised General Design Criterion 4.
1.2 Introduction l
1.2.1 Purpose This report applies to the Beaver Valley Unit 1 Reactor Coolant System (RCS) primary loop piping.
It is intended to demonstrate that for the specific l
parameters of the Beaver Valley plant, RCS primary loop pipe breaks need not be considered in the structural design basis. The approach taken has been accepted by the Nuclear Regulatory Commission (NRC) (Reference 1).
l 1-1
,s
-,n n-n,,,--m.,, -, -,,
,,-w
1.2.2 Scope The existing structural design basis for the RCS primary loop requires that dynamic effects of pipe breaks be evaluated.
In addition, protective measures for the dynamic effects associated with RCS primary loop pipe breaks have been incorporated in the Beaver Valley Unit 1 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 2).
In~ order to demonstrate this applicability of the generic evaluations to the Beaver Valley plant, Westinghouse has performed a fracture mechanics evaluation, a determination of leak rates from a through-wall crack, a fatigue crack growth evaluation, and an assessment of margins.
1.2.3 Objectives In order to validate the elimination of RCS primary loop pipe breaks for the
. Beaver Valley Unit 1 plant, the following objectives must be achieved:
a.
Demonstrate that margin exists between the " critical" crack size and a postulated crack which yields a detectable leak rate.
b.
Demonstrate that there is sufficient margin between the leakage through a postulated crack and the leak detection capability of the Beaver Valley plant.
c.
Demonstrate that fatigue crack growth is negligible.
t 1.2.4 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 3). That Topical Report employed a deterministic fracture mechanics evaluation and a probabilistic analysis to support the elimination i
1-2
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. This material was provided to the NRC along with Letter Report NS-EPR-2519 (Reference 4).
i.
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 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 Beaver Valley (References 5 and 6).
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 Mountains determined the mean probability of a direct LOCA (RCS primary loop pipe break)
-10 to be 10 per reactor year and the mean probability of an indirect LOCA to
~7 be 10 per reactor year. Thus, the results previously obtained by Westinghouse (Reference 3) 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 asymmetric blowdown loads that result from a number of discrete break locations on the PWR primary systems.
The NRC Staff evaluation (Reference 1) concludes that an acceptable. technical basis has been provided so that i
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. In a more formal recognition of LBB methodology applicability for PWRs, the NRC appropriately modified 10CFR50, General Design Criterion 4, " Requirements for Protection Against Dynamic Effects for Postulated Pipe Rupture" (Federal Register / Volume 51, Number 70/ April 11,1986/ Rules and Regulations, pp. 12502-12505).
t l
1-3 l
m
-c--,
-m- - - - - - - -
-,--ww e,
m --
This report provides a fracture mechanics demonstration of primary loop integrity for the Beaver Valley Unit 1 plant consistent with the NRC position for exemption from consideration of dynamic effects.
e f
i i
i 1-4
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 450 reactor years, including five plants each having over 16 years of operation and 15 other plants each with over 11 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 (Reference 7)..
Since some residual stresses and some degree of material susceptibility exist in any stainless steel piping, the potential for stress I
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 a' d external) as well as other materials in the system, n
applicable ASME Code rules, fracture toughness, welding, fabrication, and processing.
The environments known to increase the susceptibility of austenitic stainless steel to stress corrosion are-(Reference 7):
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 occurren'ce 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 2-1
written specifications. External cleaning for Class 1 stainless steel piping includes patch tests to monitor and control chloride and fluoride levels. For 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 conduc.ive to stress corrosion cracking with the major water chemistry control standards being included in the plant l
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.
1 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 lines. The reactor coolant system, including piping and primary components, is designed for normal, upset, emergency, and faulted condition transients.
l 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 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 f
position; pressure is controlled by pressurizer heaters and pressurizer spray I
also within a narrow range for steady-state conditions. The flow characteris-tics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump 2-2
l characteristics, are controlled in the design process. Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the 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 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 similarsto Beaver Valley Unit 1.
Stresses in the elbow below the reactor coolant pump resulting from system vibration have been found to be very small, between 2 and 3 ksi at the highest.
These stresses are well below the fatigue i
endurance limit for the material and would also result in an applied stress intensity factor below the threshold for fatigue crack growth, i
l i
l l
2-3
- - - ----m-
. - - ~- - - -. _ _ _ _.,,, _, _ _ _ _, _ _ _ _. _,. _ _ _ _ _. -,,,,, _...,
3.0 PIPE GE0 METRY AND LOADING The general analytical approach is discussed first. A segment of the primary coolant crossover leg pipe, shown below to be limiting in terms of stresses, is sketched in Figure 3-1.
This segment is postulated to contain a 1 -
circumferential through-wall flaw. The inside diameter and wall thickness of the pipe are 31.2 and 2.88 inches, respectively. The pipe is subjected to a normal operating pressure of 2200 psi. Figure 3-2 identifies the loop weld locations. The material properties and the loads at these locations resulting from deadweight, thermal expansion, and Safe Shutdown Earthquake are indicated
- in Table 3-1.
These loads include effects associated with the reduction of snubbers on the steam generators. As seen from this table, the junction of the crossover leg splitter elbow and the reactor coolant pump inlet nozzle (which is location 10 shown on Figure 3-2) is the worst location for crack stability analysis based on the highest stress due to combined pressure, dead weight, thermal expansion, and SSE (Safe Shutdown Earthquake) loadings. At this location, the axial load (F ) and the bending moment (M ) are 1961 x
b kips (including axial force due to pressure) and 24,748 in-kips, respectively.
This location is a load critical location. However, as seen later, significant degradation of end-of-service life fracture toughnesses due to thermal aging occurs at this location as well as at other pipe and fitting locations.
The highest stressed weld location for which a pipe or fitting suffers such degradation will be referred to as a toughness critical location.
The associated heat of material will be called the toughness critical material. For this plant the highest stressed welds with toughness critical material are located at points 1 in the hot leg, point 5 in the crossover leg, and point 11 in the cold leg.
(
The loads of Table 3-1 are calculated as follows: The axial force F and transverse bending moments, M and M, are chosen for each static load y
2 (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 ' Nys, and Mzs.
Based on s
elastic SSE response spectra analyses, amplified pipe seismic loads, F '
d Myd' Nzd, are obtained. The maximum pipe loads are obtained by combining the static and dynamic load components as follows:
3-1
..---~- -
--w-. -------,,--,--e,-
-w
F, =
lFsl
+
lF I d
/M
+M M
b y
z where:
lMys ! + IN I
M
=
yd y
lM I+lN I
M zd 2
zs The normal operating loads (i.e., algebraic sum of pressure, deadweight, and 100 percent power thermal expansion loading) at the critical locations identified in Table 3-1 (also see Figure 3-2) are given in Table 3-2.
The loads were determined as described above.
The calculated and allowable stresses for ASME III NB-3600 equation 9F (faulted i.e., pressure, deadweight, and SSE) and equation 12 (normal operating thermal stress) at load critical location 10 are as follows:
Calculated Allowable Ratio of Equation Stress Stress Calculated /
Number (ksi)
(ksi)
Allowable 9F 16.5 50.7 0.33 12 9.08 50.7 0.18 At the other locations, the calculated stresses and ratios are even less.
3-2
t TABLE 3-1 BEAVER VALLEY UNIT 1 PRIMARY LOOP DATA INCLUDING FAULTED LOADING CONDITIONS a
Faulted Loads Yield Ultimate Bending Direct. Stress Flow Stress Stress Stress Axial Load Moment (ksi)
Inside Wall Weld Radius Thickness "y
"u
[
],,
(Kips)
(in-Kips)
F, Mb locations (in)
(in)
(ksi)
(ksi)
(ksi)
F, M
"a
- K + Z-b b
I 14.6 2.7 18.7 67.0 42.9 1602 16473 14.1 2
14.6 2.7 18.7 67.0 42.9 1601 1866 6.9 3
14.6 2.7 18.7 67.0 42.9 1600 9372; 10.6 i'
4 15.6 2.7 18.7 67.0 42.9 1728 12944 12.9 b
S 15.6 2.88 19.4 67.0 43.2 1696 23231 15.0 6
15.6 2.88 19.4 67.0 43.2 1551 15866 11.5 7
15.6 2.88 19.4 67.0 43.2 1545 13535 10.6 8
15.6 2.88 19.4 67.0 43.2 1897 16011 12.7 9
15.6 2.88 19.4 67.0 43.2 1892 11925 11.0 b
10 15.6 2.88 19.4 67.0 43.2 1961 24748 16.5 b
11 13.85 2.56 19.4 67.0 43.2 1482 6939 10.2 12 13.85 2.56 19.4 67.0 43.2 1483 5990 9.6 13 13.85 2.56 19.4 67.0 43.2 1486 3977 8.4 14 13.85 2.56 19.4 67.0 43.2 1487 5911 9.6 4
15 13.85 2.56 19.4 67.0 43.2 1455 7462 10.3 a Includes internal pressure bCritical location i
i a
TABLE 3-2 NORMAL CONDITION (DEAD WEIGHT + PRESSURE + THERMAL)
LOADS FOR BEAVER VALLEY UNIT 1 Weld Axial Load Bending Moment F,(Kips)"
_Mb (in-Kips)
Location b
l 1485 15234 2
1485 753 3
1485 7492 4
1649 9996 b
S 1554 7426 6
1516 5439 7
1509 10865 8
1707 5100 9
1707 1398 b
10 1925 22164 b
li 1422 3687 12 1422 3727 13 1422 2841 14 1422 1958 15 1416 2976 l
l
" Includes internal pressure l
bCritical location i
l 3-4
1
)
.a y~ ~
$l li +
a.
+l 9
53 1. _
a N
u.A 3
I I
I I
I I
I i
l I
I I
l l
l
__f L
__q I
I I
I I
I I
I I
l l
.I I
l l
I 1
P = 2,200 psi F = 1,961 kips M = 24,748 in-kips Figure 3-1.
Reactor Coolant Pipe 3-5 l
REACTOR PRESSURE VESSEL NOT leg og e7 i TOR C00UWT Pump STEAM GD OtATOR CROSSOVER LEs l
- 7 1
]
HOT LEG Temperature: 614.'F; Pressure: 2235 psi CROSSOVER LEG Temperature: 547'F; Pressure: 2200 psi COLD LEG Temperature: 547'F; Pressure: 2300 psi Figure 3-2.
Schematic Diagram of Primary Loop Showing Weld Locations - Beaver Valley Unit 1 1
3-6
_= ~..
'4. 0 FRACTURE MECHANICS EVALUATION 4.1 Global Failure Mechanism Determination of the conditions which lead to failure in stainless steel should be done with plastic fracture methodology because of the large amount of deformation accompanying fracture. One method for predicting the failure of ductile material is the plastic instability method, based on traditional plastic limit load concepts, but accounting for strain hardening and taking into account the presence of a flaw. The flawed pipe is predicted to fail when the remaining net section reaches a stress level at which a. plastic hinge is formed. The stress level at which this occurs is termed as the flow stress. The flow stress is generally taken as the average of the yield and ultimate tensile strength of the material at the temperature of interest.
This methodology has been shown to be applicable to ductile piping through a i
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 equilibrium of the section containing the flaw (Figure 4-1) when i
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 limit moment for such a pipe is given by:
1 a,c e g
)
where:
[
)
l
)a,c.e 4-1
.i
[
ja,c.e i
i The analytical model described above accurately accounts for the piping j
internal pressure as well as imposed axial force as they affect the limit moment. Good agreement was found between the analytical predictions and the experimental results (Reference 8).
4 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.
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 from a 4
yc 1
J-integral resistance curve is a material parameter defining the crack i
initiation.
If, for a given load, the calculated J-integral value is shown to be less than the J f the material, then the crack will not initiate.
If Ic the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:
dJ E
fapp*Ha
}"f 7
I 4
i i
4-2
~
t L
__ _ = __ _.,_.
where:
s T,pp = applied tearing modulus j
E = modulus of elasticity of = [
]a,c.e (flow stress) a = crack length ja,c.e In summary, the local crack stability will be established by the two-step criteria:
I l
J<J Ic I
or if J g JIe T,pp < Tmat An additional supplementary criterion is that J < J,,, where J,,, is the maximum value of J obtained from J tests for the material in question.
4.3 Material Properties The primary loop piping material for Beaver Valley Unit 1 is cast stainless steel (SA-351-CF8M) and associated welds. Welds exist as indicated in Figure 3-2.
The tensile and flow properties of the load critical location and the toughness critical locations are given in Table 3-1.
The pre-service fracture toughness of cast materials in terms of J have been found to be very high at 600*F. Typical results are given in Figure 4-2 taken 2
from Reference 9.
J is observed to be over 5000 in-lbs/in. However, Ic 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.
l l
4-3 l
1.
~
. _ _ =
3 To determine the effects of thermal aging on piping integrity, a detailed study was carried out in Reference 9.
In that report, fracture toughness results were presented for a material representative of [
]a,c.e Toughness results were provided for the material in the full service life condition and these properties are also presented in Figure 4-3 of this report for information. The J value for this material at operating temperature was
]a,he and the maximum value of J obtained in the tests
[
was in excess of [
la,c.e The tests of this material were conducted on small specimens and therefore rather short crack extensions occurred, (maximum extension 4.3 mm) so it is expected that higher J values would be sustained for larger specimens. T was [
Ja,c.e at ut operating temperature. The effects of the aging process on'the end-of-service life fracture toughness are discussed in Appendix B.
End of-service life toughnesses for the heats are established using the alternate toughness criteria methodology. By that methodology a heat of material is said to be as good as [
]a,c.e if,it can be demonstrated that its end-of-service fracture toughnesses equal or exceed those of [
Ja,c.e Of the thirty heats examined in Appendix B, fourteen are seen to be not as good as [
la,c.e All of the piping fits into this category except for [
19 in the crossover leg (See Table B-1).
Furthermore, only one of the fittings fits into this 3
category. This is the 31" x 40D elbow in the crossover leg (See also Table B-1). The fracture toughness criteria to be used in the fracture mechanics evaluation, based on the alternate toughness methodology of Appendix B, are given in Table 4-1.
These toughness values are the lowest of all heats occurring at the critical loations 1, 5, 10, and 11, shown on Figure 3-2.
The toughness values used in this report for location 5 are actually from location 6 (an adjacent weld). The actual toughness at location 5 is slightly higher than the toughness at location 6.
Available data on aged stainless steel welds (References 9 and 10) indicate the J values for the worst case welds are of the same order as the aged 3c material. However, the slope of the J-R curve is steeper, and higher J-values 2
have been obtained from fracture 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 4-4
i l'
l than that in the base metal because the yield stress for the weld materials is a
much higher at temperature. Therefore, weld regions are less limiting than the cast material.
It is thus conservative to choose the end-of-service life toughness properties of [
la.c.e as representative of those of the welds.
Also, such pipes and fittings having an end-of-service life calculated room temperature charpy U-notch energy, (KCU) greater than that of [
la,c.e are also conservatively assumed to have the properties of [
Ja,c.e,
In the fracture mechanics analyses that follow, the fracture toughness properties given in Table 4-1 will be used as the criteria against which the applied fracture toughness values will be compared.
4.4 Results of Crack Stability Evaluation Figure 4-6 shows a nlot of the plastic limit moment as a function of through-wall circumferential flaw 1.ength in the crossover leg of the main coolant piping (critical location 10.) This limit moment was calculated for Beaver Valley Unit 1 from data for a pressurized pipe at 2200 psi with an axial force of 1961 kips, operating at 547'F with ASME Code minimum tensile properties.
The maximum applied bending moment of 24748 in-kips can be plotted on this figure and used to determine a critical flaw length, which is shown to be
[
]b,c.e inches. '
In Figures 4-4, 4-5 and 4-7 plots of the plastic limit moment as a function of through wall circumferential flaw length at the remaining critical locations of the hot leg, crossover leg and cold leg, respectively, are given.
These limit moments were calculated as above using the appropriate pressure, forces, (a) In this report all applied J values were conservatively determined by using base metal strength properties.
4-5
and dimensions as given either in Table 3-1 or Figure 3-2 with bending moment as a parameter. The ASME Code minimum properties at 547'F and 614'F were used for the cold and hot legs respectively. Critical flaw lengths were determined as in Figure 4-4 by use of the maximum applied bending moment. The critical flaw length in Figures 4-5 and 4-6 are all seen to exceed the [
]b,c e 7
inches established for critical location 10.
For fracture mechanics evaluations the toughness and load critical locations were evaluated as follows.
In Table 3-1, the outer surface axial stress (o,) at load critical location 10 is seen to be 16.5 ksi. Stresses due to the internal pressure of 2200 psi are as follows (see Reference 11):
c (circumferential stress):
11.9 ksi r (radial stress):
0 The von Mises effective stress, o,ff, (see Reference 12) is given by c,ff = h[(o,- r) *I
- r) +I
~ "c) c a
and is 14.75 ksi.
Thus the effective stress is less than the yield stress and by the Von Mises plasticity theory yielding does no't occur. Also, similar consideration at the other toughness critical locations confirms that yielding does not occur there. Hence, linear elastic fracture nechanics is applicable for analyzing the pipes with hypothesized flaws at the other toughness critical locations.
The analytical method used for the local stability evaluation at these locations is summarized below.
The stress intensity factors corresponding to tension and bending are expressed, respectively, by (see Reference 13) 4-6 h---
I L
sa F (a)
Kt* t g
b ' 'b F I")
Y b
where F (a) and F (a) are stress intensity calibration factors corre-t b
sponding to tension and bending, respectively, a is the half-crack length, a is the half-crack angle, is the remote uniform tensile stress, and t
is the remote fiber stress due to pure bending. Data for F (a) and b
t F (a) are given in Reference 13. The effect of the yielding near the b
crack tip can be incorporated by Irwin's plastic zone correction method (see Reference 14) in which the half-crack length, a, in the above formulas is replaced by the effective crack length, a,ff, defined by 2
1 K
a,f f = a + - q 23 gy for plane stress plastic zone-size corrections, where e is the yield y
strength of the material and K is the total stress intensity factor due to combined tensile and bending loads (i.e., K = Kt +2 K).
Finally, the b
J,pp-value is determined by the relation J,pp = K /E, where E is Young's Modulus.
J,pp was calculated for the four critical locations using crack length as a parameter. The results are presented in Table 7-1 of Chapter 7 wherein J,pp values and leak rates are examined in assessing margin.
As shown in Table 7-1, for J,pp or T,pp less than the local crack stability criteria given in Section 4.2, the critical circumferential flaw b
lengths are at least [
l,c.e inches for the hot leg, [ ]b,c.e inches for the crossover leg, and [
]b,c.e inches for the cold leg.
4-7
TABLE 4-1 1
a FRACTURE TOUGHNESS CRITERIA USED IN THE LEAK-BEFORE-BREAK EVALUATION Location or J
J,,,
Ic 2
2 Description (in-lb/in )
T (in-lb/in )
mat
~
b'C
1 b
S 10 11 aThe lowest of the values for all heats at a given location are given here.
bToughness listed is for location 6 which has a lower toughness than location 5.
4-8
a,c.e r /'/ ////
g
\\ 2a
/
ewtral Aaas
=== N uV Figure 4-1.
[
Ja,c.e Stress Distribution 4-9 l
P
~
b,c.e L
Figure 4-2.
J vs aa for SA351-CF8M Cast Stainless Steel at 600*F 4-10
t
~ b,c.e i
l l
l Figure 4-3.
J-Aa at Different Temperatures for Aged Material
[
Ja,c.e (7500 Hours at 400*C) 4-11
--w.---
y-y p-
l L
l b,c.e v
i FLAW GEOMETRY OD = 34.6 in, t=
2.7 in.
P = 2235 psig F = 1602 kips o = 18.7 ksi y
= 67.0 ksi u
of = 42.9 ksi Temp = 614 F Figure 4-4.
" Critical" Flaw Size Prediction - Hot Leg at Critical Location 1 4-12
b,c e t
l v
i FLAW GEOMETRY OD = 36.96 in.
t=
2.88 in.
P = 2200 psig F = 1696 kips o = 19.4 ksi y
= 67.0 ksi u
of = 43.2 ksi Temp = 547 F 1
I Figure 4-5.
" Critical" Flaw Size Prediction - Crossover Leg at Critical Location 5 4-13
L b,c.e L
l rs l
FLAW GEOMETRY OD = 36.96'in.
t=
2.88 in.
P = 2200 psig F =.1961 kips e = 19.4 ksi y
= 67.0 ksi u
of = 43.2 ksi 0
Temp = 547 F Figure 4-6.
" Critical" Flaw Size Prediction - Crossover Leg at Critical Location 10 l
4-14 l
l L
l b,c.e l
FLAW GEOMETRY OD = 32.82 in.
t=
2.56 in.
P = 2300 psig F = 1482 kips o = 19.4 ksi y
o = 67.0 ksi j
u of = 43.2 ksi 0
Temp = 547 F N
Figure 4-7.
" Critical" Flaw Size Prediction - Cold Leg at Critical Location 11 4-15
i 5.0 LEAK RATE PREDICTIONS 5.1 Introduction Fracture mechanics analysis has shown that postulated through-wall cracks in the primary loop would remain stable and not cause a gross failure of this component. If such a through-wall crack did exist, it would be desirable to detect the leakage such that the plant could be brought to a safe shutdown condition. The purpose of this section is to discuss the method which will be used to predict the flow through such a postulated crack and present the leak rate calculation results for through-wall circumferential cracks.
5.2 General Considerations The flow of hot pressurized water through an opening to a lower back pressure causes flashing which can result in choking.
For long channels where the ratio of the channel length, L, to hydraulic diameter, D, R/D ) is g
greater than [
]a,c.e, both [
Ja,,e must be considered.
In this situation the flow can be described as being single phase through the channel until the local pressure equals the saturation pressure of the fluid. At this point, the flow begins to flash and choking occurs.
Pressure losses due to momentum changes will dominate for [
Ja,c.e,
However, for large L/D values, friction pressure drop will become important H
and must be considered along with the momentum losses due to flashing.
5.3 Calculation Method The basic method used in the leak rate calculations is the method developed by
[
ja,c,e,
The flow rate through a crack was calculated in the following manner. Figure 5-1 from Reference 15 was used to estimate the critical pressure, Pc, for the primary loop enthalpy condition and an assumed flow. Once Pc was found for a i
5-1
given mass flow, the [
Ja,c,e i
was-found from Figure 5-2 taken from Reference 15. For all cases considered, since [
Ja,c.e Therefore, this method will yield the two phase pressure drop due to momentum effects as illustrated in Figure 5-3.
Now using the assumed flow rate, G, the frictional pressure drop can be calculated using APf=[
~
Ja,c.e (5-1) where the friction factor f is determined using the [
Ja,c.e The crack relative roughness, c, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these I
calculations was [
Ja,c.e RMS.
The frictional pressure drop using Equation 5-1 is then calculated for the assumed flow and added to the [
Ja c.e to obtain the total pressure drop from the primary system to the atmosphere. That is, for the primary loop i
Absolute Pressure - 14.7 = [
Ja,c.e(5-2)
I for a given assumed flow G.
If the right-hand side of Equation 5-2 does not agree with the pressure difference between the primary loop and the atmosphere, then the procedure is repeated until Equation 5-2 is satisfied to within an acceptable tolerance and this results in the flow value through the crack. This calculational procedure has been recommended by (
Ja,c,e for this type of [
Ja,c.e calculation.
l 5.4 Leak Rate Calculations Leak rate calculations were made as a function of crack length for all the 4
critical locations previously identified. The normal operating loads of Table 5-2 s
... - ~.
,,-,=,.-v-
+
r-r..%-r---w-p-,-w--r-
---,--.w-3 e,.+w-,
.--w w---,wi
3-2 were applied in these calculations. The crack opening area was estimated using the method of Reference 13 and the leak rate was calculated using the two phase flow formulation described above.
The results are tabulated in Table 7-1 of Chapter 7 wherein J,pp values and leak rates are examined in assessing surgin.
The Beaver Valley Unit 1 plant has an RCS pressure boundary leak detection system which is consistent with the guidelines of Regulatory Guide 1.45 for detecting leakage of 1 gpm in one hour.
For the worst critical location, which is location 5, the largest st-flaw has a factor of about 80 above the 1 gpm criteria of Regulatory Guide 1.45.
For the other critical locations, the leak rate factors are greater than 80.
k 5-3
e a.c.e d
NI E
>t 8
a I
~
~
2 STAGNATHW ENTHALPY (10 gg; Figure 5-1.
Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures 5-4
a,c.e h*I 9*
I w
Iw E
YWt 5
LENGTH / DIAMETER RATIO IUD)
Figure 5-2.
'[
ja,c e Pressure Ratio as a function of L/D 5-5
l L
_ e,c.e a,c.e j
f
-^
Figure 5-3.
Idealized Pressure Drop Profile Through a Postulated Crack 5-6
a 6.0 FATIGUE CRACK GROWTH ANALYSIS To determine the sensitivity of the primary coolant system to the presence of small cracks, a fatigue crack growth analysis was carried out for the [
Ja,c.e region of a typical system (see Location
[
Ja,c e of Figure 3-2).
This region was selected because crack growth calculated here will be typical of that in the entire primary loop.
Crack growths calculated at other locations can be expected to show less than 10%
variation. Thermal aging has been shown not to impact fatigue crack growth (References 9 and 10).
A[
]a,c.e of a plant typical in geometry and operational characteristics to any Westinghouse PWR System.
[
ja,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:
Cross Section A: [
Ja,c.e Cross Section B: [
Ja,c.e Cross Section C:
[
la,c,e Fatigue crack growth rate laws were used [
l Ja,c.e The law for stainless steel was derived from Reference 17, with a very conservative correction for the R I
ratio, which is the ratio of minimum to maximun stress 'during a transient.
For stainless steel, the fatigue crack growth formula is:
l 6-1 l
h=(5.4x10-12) g 4.48inches / cycle eff where K,ff = K,,, (1-R)0.5 i
R=Kmin/Emax
[
ya.c.e a,c.e
[-
)
where: [
] a,c.e 1
The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is summarized in Table 6-1, and shows that the crack growth is very small, regardless [
ja,c.e j
4 h
1 3
s-2
TABLE 6-1 FATIGUE CRACK GROWTH AT [
]a,c.e (40 YEARS)
FINAL FLAW (in)
[
3a,c.e INITIAL FLAW (IN)
[
Ja,c.e
[
3a,c.e
[
3a,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
i
+
4 i
l.
6-3
e I
I t
1 if l
1 J
i i
l I
Figure 6-1.
Typical Cross-Section of [
ya,c,e 6-4
a.c. e IIT u>
.N E
U 6s W
E Es 4
N
<s Iaeuu<su I
t I
Figure 6-2.
Reference Fatigue Crack Growth Curves for (
a,c.e j
6-5
,._--s, -, - -. _ _,.
--n.-
n.-
- e. - -
,,,,,-,n.
e.c..
t a
4 4
i Figure 6-3.
Reference Fatigue Crack Growth Law for (
Ja.c.e in a Water Environment at 600'F 6-6
7.0 ASSESSMENT
OF MARGINS The rasults of the toughness and leak rate calculations for the three critical locations are summarized in Table 7-1.
Margins for these critical locations are discussed below.
)'
At critical locations 1, 5,10, and 11, flaw sizes which yield a factor of 10 greater than the 1 gpm leakage requirements of Regulatory Guide 1.45 are 4.9, 7.15, 4.2, and 6.9 inches, respectively. Flaw lengths at least twice those giving 10 gpm leakage are stable resulting in J,pp or T,pp values less je mat. For locations 1, 10, and 11 J,pp for such large flaws than J or T were found to be less than the appropriate Jje.
At location 5, J,pp exceeded J for an appropriately large flaw si n, but the flaw is seen to je be stable with J,pp less than J,,, and T,pp well less than Tmat.
As shown in Section 3.0, a margin of a factor of not less than 2.5 exists between calculated and ASME Code allowable faulted. conditions and thermal stresses.
In Section 4.4, the " maximum" flaw sizes at load critical location 1 and the other critical locations are calculated using the limit load method and shown to be at least [
]b,c.e inches. Thus, based on the above, the " maximum" j
flaw sizes at these locations will, of course, exceed the stable crack lengths at their respective locations.
In summary, relative to:
1.
Loads t
The J,pp values for Beaver Valley Unit I are enveloped by the J a.
values established from testing of highly aged material.
t 7-1 i
l
b.
Margins at the critical location of at least 2.5 on faulted conditions and thermal stresses exist relative to ASME Code allowable values.
2.
flaw Size a.
A margin of at least two on flav size exists for stable flaw sizes with flow rates which exceed a ieak rate of 10 gal / min.
4 b.
If limit load is used as the basis for critical flaw size, the margin for global stability well exceeds that based upon fracture mechanics.
3.
Leak Rate At the critical locations, leak rates are seen to exceed 10'gpm for all flaws considered which is well in excess of the criterion of Regulatory Guide 1.45. For the larger stable flaws [
Ja,c.e leak rates are seen to be greater than 70 gpm.
t l
1 i
7-2 i
8 TABLE 7-1
SUMMARY
OF J,pp AND LEAK RATE RESULTS AS A FUNCTION OF CRACK LENGTH AT THE FOUR CRITICAL LOCATIONS Crack Leak b
b e
Location Critical Length Rate (Loops)
J T
J,,,
(Inch)
J,pp T,pp (GPM)
Ic mat b,c.e 1
199.7 (Hot Leg) 78.7 10.0 5
82.1 (Crossover 69.3 Leg) 10.0 1
10 81.0 (Crossover 65.0 Leg) 10.0 11 115.5 (Cold Leg) 86.8 10.0 i
~
2
- a. J values have units of in-lb/in,
- b. Values are lowest of all heats in indicated legs.
- c. Crack lengths yielding 10 gpm leakage for locations 1, 5, 10, and 11 are 4.9, 7.15, 4.2, and 6.9 inches respectively.
7-3
l
8.0 CONCLUSION
S This report justifies the elimination of RCS primary loop pipe breaks for the Beaver Valley Unit 1 plant as follows:
l i
a.
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.
i b.
Water hammer should not occur in the RCS piping because of system design, testing, and operational considerations.
c.
The effects of low and high cycle fatigue on the '.ntegrity of the primary piping are negligible.
d.
Adequate margins exist for ASME code allowable faulted and thermal loads.
1 A
e.
Adequate margin exists between the leak rate of stable flaws and the criterion of Reg. Guide 1.45.
f.
Ample margin exists between stable flaws of item e and larger stable l
flaws.
g.
Ample margin exists in the material properties used to demonstrate end of-service life (relative to aging) stability of the critical flaws.
For each critical location a flaw is identified (see Table 7-1) that will be stable throughout reactor life because of the ample margins in e, f, and g i
above and will leak at a detectable rate which will assure a safe plant l
shutdown.
Based on the above, it is concluded that RCS primary loop pipe breaks need not be considered in the structural design basis of the Beaver Valley Unit 1 plant.
4 8-1
F
9.0 REFERENCES
1.
'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.
g 2.
Letter from Westinghouse (E. P. Rahe) to NRC (R. H. Vollmer),
NS-EPR-2768, dated May 11, 1983.
3.
WCAP-9283, "The Integrity of Primary Piping Systems of Westinghouse Nuclear Power Plants During Postulated Seismic Events," March, 1978.
4.
Letter Report NS-EPR-2519, Westinghouse (E. P. Rahe) to NRC (D. G.
Eisenhut), Westinghouse Proprietary Class 2, November 10, 1981.
5.
Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated April 25, 1983.
l 6.
Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated July 25, 1983.
7.
NUREG-0691, " Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors", USNRC, September 1980.
8.
Kanninen, M. F., et. al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks", EPRI NP-192, September 1976.
9.
WCAP-10456, "The Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping for W NSSS," W Proprietary Class 2, November 1983.
9-1 l
+
- 10. Slama, G., Petrequin, P., Masson, S. H., and Mager, T. R., "Effect of Aging on Mechanical Properties of Austenitic Stainless Steel Casting and e
Welds", presented at SMiRT 7 Post Conference Seminar 6 - Assuring Structural Integrity of Steel Reactor Pressure Boundary Components, August 29/30, 1983, Monterey, CA.
- 11. Durelli, A. J., et. al., Introduction to the Theoretical and Experimental
(
Analysis of Stress and Strain, McGraw Hill Book Company, New York,
- ~
(1958), pp 233-236.
- 12. Johnson, W. and Mellor, P. B., Engineering Plasticity, Van Nostrand Relmhold Company, New York, (1973), pp 83-86.
- 13. Tada, H., "The Effects of Shell Corrections on Stress Intensity Factors and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe," Section 11-1, NUREG/CR-3464, September 1983.
14.
Irwin, G. R., " Plastic Zone Near a Crack and Fracture Toughness," Proc.
7th Sagamore Conference, P. IV-63 (1960).
15.
[
3a,c,c
)
16.
[
ja,c.e
- 17. Bamford, W. H., " Fatigue Crack Growth of Stainless Steel Piping in a Pressurized Water floactor Environment," Trans. ASME Journal of Pressure Vessel Technology, Vol.101, Feb.1979.
18.
(
3,c.e 6
4
- 19. [
3a c.e
- 20. Witt, F. J. and Kim, C. C., Toughness Criteria for Thermally Aged Cast Stainless Steel, Westinghouse Proprietary Class 2 Report.WCAP 10931 Revision 1, July 1986.
- 21. Letter: Dominic D. Dilanni, NRC to D. M. Musolf, Northern States Power Company, dated December 22, 1986, Docket Nos. 50-282 and 50-306.
l 9-3
APPENDIX A LIMIT MOMENT I
. y I
)
a,c.e
)
A-1 s
I
- a,c.e l
Figure A-1.
Pipe With a Through-Wall Crack in Bending A-2
APPENDIX B ALTERNATE TOUGHNESS CRITERIA FOR THE BEAVER VALLEY UNIT 1 CAST PRIMARY LOOP COMPONENTS B.1 INTRODUCTION Not all of the individual cast piping components of the Beaver Valley Unit 1 primary loop piping satisfy the original [
]a,c e criteria (Reference 9).
In this appendix, the alternate toughness criteria for thermally aged cast stainless steel developed in Reference 20 will be used to categorize the various individual cast piping components thus establishing criteria based upon which the mechanistic pipe break evaluation may be performed. The
~
procedure of Reference 20 has been approved by the NRC (Reference 21.)
First the chemistry and calculated room temperature charpy U-notch energy, (KCU) values are given followed by an identification of each of the heats of material with a specific loop and location. The criteria for the various individual loop compcnents are tabulated.
B.2 CHEMISTRY AND KCU TOUGHNESS The correlation of Reference 10 which is based on the chemistry of the cast stainless steel piping was used to calculate the associated KCU value.
The chemistry and end-of-service life KCU toughness values are given in Table B-1.
Of the thirty heats of cast stainless steel, fourteen fail to meet the current
[
Ja,c.e criteria. These heats occur in the pipes and fittings of the hot, cold and crossover legs in each of the three reactor loops.
B.3 THE AS-BUILT BEAVER VALLEY UNIT 1 LOOPS Beaver Valley Unit 1 is a three-loop Westinghouse type pressurized water reactor plant. A typical three-loop primary system is sketched in Figure B-1.
The three loops are identified as Loops 1, 2 and 3 in Beaver Valley Unit 1.
Sketches for associating piping component with specific locations and loop B-1
are given in Figures B-2 through B-4.
The individual components are identified by heat numbers. The components which have toughnesses less than that of [
]a,c.e are indicated in Table B-1.
As may be seen, only one pipe is better than [
Ja,c.e, and only one fitting fails to be as good as [
]a,c.e,
B.4 ALTERNATE TOUGHNESS CRITERIA FOR THE BEAVER VALLEY UNIT 1 CAST PRIMARY LOOP MATERIAL ON A COMPONENT-BY-COMPONENT BASIS The alternate toughness criteria for the Beaver Valley Unit 1 cast primary loop material may be obtained by applying the methodology of Reference 20 to Table B-l'.
First, it is observed that sixteen of the 30 heats fall into Category 1 i.e., they are as tough as [
Ja,c.e The remaining heats fall into Category 2 with one in Category 3.
The toughness criteria for all 30 heats are given in Table B-2.
Example toughness calculations for the Category 3 heat are given below.
[
2 19 daJ/cm which falls below that of [
9 Ja,c.e The 5-ferrite content is [
1 By Reference 20, the
[
ja,c.e Thus, for full-embrittlement j
=[
Ja,c.e T
- I3*'
at J,37 =[
]W and KCU < [
Ja,c.e B-2
[
la,c.e is a Category 3 material as defined in Reference 20 and the end-of-service life fracture toughness is [
la,c.e These results are given in Table B-2 for Category 3.
An example calculation for a Category 2 heat is given below. Similar calculations for the remaining twelve Category 2 heats were made.
The example calculation will be made for Heat [
j9 The ferrite content is [
]9 and the end-of-service life KCU is [
]b,c.e 2
daJ/cm. The [
\\
]a,c.e, the heat falls into Category 2.
Thus:
Jjc = [
3a,c.e T
- I mat ja,c.e and J,,x = [
ja,c e B-3
1 O.
^
Jl l
4 b
2m W
J J
4>
E 4.N W6 C3 V E-OMn M4 WM m
N I EWJ L4 O=
EE AW
>=
U~L 8
. "EJ m >
OE 4_
E 4a XWG l
C W
J CD 4>
l 1
B-4
?-
TABLE B-2 FRACTURE TOUGHNESS CRITERIA FOR THE CAST PRIMARY PIPING i
COMPONENTS OF THE BEAVER VALLEY UNIT 1 NUCLEAR PLANT b,c.e.g l
a.
Not applicable b.
Given in the same order as Table B-1 B-5
L Steam Generator E
Hot Leg i
I Reactor Vessel Pump Cold Leg l
l Isolation Valve l
~ Isolation Valve l
Il Figure B-1.
Typical Layout of the Primary Loops for a Westinghouse Three-Loop Plant With Isolation Valves B-6
I
-9 k
w...
l l
~~
h Figure B-2.
Identification of Heats with Location for Cold Leg B-7
g P
i Figure B-3.
Identification of Heats with Location for Hot Leg B-8
k 9
l l
Figure B-4.
Identification of Heats with Location for Crossover Leg B-9