ML20213E055
| ML20213E055 | |
| Person / Time | |
|---|---|
| Site: | North Anna  |
| Issue date: | 08/31/1986 |
| From: | Lee Y, Ma W, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19292G230 | List: |
| References | |
| WCAP-11164, NUDOCS 8611120383 | |
| Download: ML20213E055 (79) | |
Text
,
WESTINGHOUSE CLASS 3 WCAP-11164 TECHNICAL BASES FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTUrr AS A STRUCTURAL DESIGN BASIS FOR NORTH ANNA UNITS 1 & 2 August 1986 Winston K. Ma Y. S. Lee S. A. Swamy T. H. Liu H. F. Clark, Jr.
Verified By: F. J. Witt APPROVED: I APPROVED: "9< '
E. RVJohnson, Manager
/ S. S. P[ usamy, Manager Structural Materials Engineering Structural and Seismic Development APPROVED: / Ad M1 oA AAl-J
/'.J.(cInerney,iIanag[dSystems Mechanical Equipment an Licensing WESTINGHOUSE ELECTRIC CORPORATION NUCLEAR ENERGY SYSTEMS P.O. Box 355 Pittsburgh, Pennsylvania 15230 8611120303 861106 DR ADOCK0500g8
ACKNOWLEDGEMENT The authors wish to acknowledge the contributions to the North Anna 1 and 2 Primary Loop Mechanistic Fracture Evaluation by Mr. J. F. Petsche and Mr. D.
E. Prager.
4 e
w
=
TABLE OF CONTENTS Section Title Pa2'
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 GEOMETRY AND LOADINGS 3-1 3.1 Geometry and Faulted Condition Loads 3-1 3.2 Load Critical Location 3-2 3.3 Toughness Critical Location 3-2 3.4 Normal Operating Loads 3-2 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 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 GRDWTH ANALYSIS 6-1 v
J
I TABLE OF CONTENTS (CONT.)
Section Title Pasg 7.0 ASSESSMENT OF MARGINS 7-l
8.0 CONCLUSION
S 8-l
9.0 REFERENCES
9 -1 .
APPENDIX A - Limit Moment A-1 APPENDIX B - Alternate Toughness Criteria for the North Anna B-1 Plants No. 1 and 2 Primary Loop Components s I
\\
vii
LIST OF TABLES TABLE TITLE PAGE 3 -1 Summary of Dimensions, Physical Properties, 3-3 g and Faulted Loads for Primary Loop L No. 1 of North Anna Plants No. 1 and 2 3-2 Summary of Dimensions, Physical Properties, 3-4 .
and Faulted Loads for Primary Loop '.,
No. 2 of North Anna Plants No.1 and 2 ,
3-3 Sunenary of Dimensions, Physical Properties, 3-5 and Faulted Loads for Primary Loop No. 3 of North Anna Plants No.1 and 2 L
3-4 Normal Operating Loads and Other Data for 3-6 North Anna Plants No. I and 2 Loop No. 1 3-5 Normal Operating Loads and Other Data for 3-7 North Anna Plants No. 1 and 2 Loop No. 2 3-6 Normal Operating Loads and Other Data for 3-8 ..
North Anna Plants No. 1 and 2 Loop No. 3 3-7 Normal Operating Loads at the Load Critical 3-9 Locations for North Anna Plants No. 1 and 2 3-8 Calculated and Allowable Stress 3-10 for ASME NB-3600 Criteria 1x
LIST OF TABLES TABLE. TITLE PAGE 4
4-1 Summary of the Alternate Toughness J gg 4-8 - --
for the Heats of Materials Which Do Not Satisfy the Toughness Criteria of ya .c.e 4-2 Predicted Critical Crack Length at 4-9 Various Locations Based on ,
Limit Load Method.
4-3 Sunenary of Applied J Values vs J Ic 4-10 __
Toughness for North Anna Plants No. 1 and 2 s-5-1 Summary of Leak Rate Calculations 5-4 for North Anna Plants No.1 and 2 6-1 Fatigue Crack Growth at [ 6-3 a
J .c.e (40 years) 7-1 Sununary of Stability and Leak Rate Calculations 7-4 xi
LIST OF FIGURES FIGURE TITLE PAGE 3-1 Schematic Diagram of Primary Loop Showing 3-11 Weld Locations 3-2 Reactor Coolant Pipe 3-12 4 -1 [ ]8 'C # Stress Distribution 4-11 4-2 J vs As for SA351-CF8M Material 4-12 at 600"F Taken From a 23 Inch Diameter Pipe 4-3 J vs Aa for SA351-CF8M Cast Stainless 4-13 Steel at 600"F 4-4 J - Aa Curves at Dif ferent Temperatures 4-14 a
for Aged Material [ J .c.e 4-5 Predicted " Critical" Crack Length at 4-15 Weld Location No.1 for North Anna Plants No. 1 and 2 4-6 " Critical" Crack Length at Weld Location 4-16 No. 2 for North Anna Plants No.1 and 2 4-7 " Critical" Crack Length at Weld Location 4-17 No. 3 for North Anna Plants No. 1 and 2 4-8 " Critical" Crack Length at Weld 4-18 No. 4 for North Anna Plants No.1 and 2 4-9 " Critical" Crack Length at Weld Location No. 5 of 4-19 Loop No. 1 for North Anna Plants No. 1 and 2 xiii
LIST OF FIGURES FIGURE TITLE PAGE 5-1 Analytical Predictions of Critical Flow 5-5 Rates of Steam-Water Mixtures 5-2 [ ]3'C Pressure Ratio 5-6 as a Function of L/D 5-3 Idealized Pressure Drop Profile Through a 5-7 Postulated Crack 6-1 Typical Cross-Section of ( 6-4 g c.e a
6-2 Reference Fatigue Crack Growth Curves 6-5 for [
j a.c.e 6-3 Reference Fatigue Crack Growth Law for 6-6 a
[ J .c.e in a Water Environment at 600'F xv
l.0 INTRODUCTION 1.1 Purpose This report applies to the North Anna Units 1 & 2 plant reactor coolant system primary loop piping. It is intended to demonstrate that specific parameters for the North Anna Units 1 & 2 are enveloped by the generic analysis performed 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 North Anna Plants 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 North Anna Plants.
Westinghouse has performed a comparison of the loads and geometry for the North Anna plants with the envelope parameters used in the generic analyses (Reference 1), a f racture mechanics evaluation, a determination of leak rates from 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 North Anna Units 1 & 2. In order to validate this conclusion the following objectives must be achieved,
- a. Demonstrate that the North Anna 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.
- c. Demonstrate that there is sufficient margin between the leakage through a postulated crack and the leak detection capability of the North Anna Plants.
- d. Demonstrate that fatigue crack growth is negligible.
1.4 Background Information Westinghouse has perfortned 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 l Westinghouse performed additional tests and analyses to justify the i
elimination of RCS primary loop pipe breaks. As a result of this ef fort, 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 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 North Anna Units 1 & 2 (References 8 and 9). The results from the LLNL study were released at a March 28, 1983 ACRS Subconvrittee meeting. These studies, which are applicable to all Westinghouse plants east of the Rocky Muntains, deternined the mean probability of a direct LOCA (RCS primary loop pipe break) to be 10 -10 per reactor year and the mean probability of an 1-2
indirect LOCA to be 10~ per reactor 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 asymmetric blowdown loads that result f rom a number of discrete break locations on the PWR primary systems. The NRC Staff evaluation (Reference 2) concluded 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 f racture mechanics demonstration of the primary coolant loop integrity.
This report will demonstrate the applicability of the Westinghouse generic evaluations to North Anna Units 1 & 2.
1-3
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 hamer, or fatigue (low and high cycle). This operating history totals over 400 reactor-years, including five plants each having over 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 F L
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 inunune 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 corrosion are (Reference 10): oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced fonns 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.
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 rhlorides and fluorides within the specified limits. This is assured by controlling charging flow chemistry and specifying proper wetted surface materials.
2.2 Water Hamer )
Overall, there is a low potential for water hamer in the RCS since its design and operation precludes 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. 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 l valve and pump actuations result in relatively slow transients with no significant effect on the system's dynamic loads. To ensure dynamic system l stability, reactor coolant parameters are stringently controlled. Temperature l
during nonnal 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 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 flow and vibration characteristics of the system. Preoperational testing and 2-2
f,f 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 Fatique 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 has been carried out as part of this study in the form of a f atigue crack growth analysis, as discussed in Section 6.
High cycle f atigue loads in the system would result primarily f rom 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 North Anna Units 1 & 2. 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 applied stress intensity factor below the threshold for f atigue crack growth.
2-3
3.0 PIPE GEOMETRY AND LOADINGS 3.1 GEOMETRY AND FAULTED CONDITION LOADS A typical primary loop pipe schematic for North Anna Units 1 and 2 is shown in Figure 3-1 in which all the weld locations are identified. The dimensions, physical properties, and the loads acting at various weld locations of each loop of North Anna Units 1 & 2 due to pressure, deadweight, thermal expansion, and Safe Shutdown Earthquake Loads (SSE) are tabulated as shown in Tables 3-1 through 3-3. A typical segment of the primary coolant pipe under mechanical l:
loads of F, and Mb is schematically illustrated as shown in Figure 3-2.
The loads shown in Tables 3-1 through 3-3 are calculated as follows.
The axial force F and transverse bending moments, M and Mz, are chosen 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,, M , and M . Based on elastic SSE response spectra analyses, amplified pipe seismic loads, F , M ,Mg , are obtained. The maximum d
pipe loads are obtained by combining the static and dynamic load components as follows:
= F + F F, 3 d M " My + M z b
where y ys "yd
" +
z Zs Zd 3-1
3.2 LOAD CRITICAL LOCATION As seen from Tables 3-1 through 3-3, 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, deadweight, thermal expansion, and SSE (Safe Shutdown Earthquake) loading. This location will be referred to as the load critical location.
3.3 TOUGHNESS CRITICAL LOCATION 11 addition to loads, the fracture toughness of the materials plays an equally it.portant role in performing the crack stability evaluation. As seen later (Appendix B), the lowest end-of-service life fracture toughness occurs at weld locations 2, 3, 4, and S (Figure 3-1). These locations will be referred to as the toughness critical locations.
3.4 NORMAL OPERATING LOADS The normal operating loads which include thermal, pressure, and deadweight ef fects (axial force and bending moment) are provided in Tables 3-4 through 3-6. This information will be used in calculating the leak rates through postulated through-wall flaws.
The calculated and allowable stresses for ASME Code Section III, NB-3600, equation 9 (for faulted conditions, i.e., pressure, deadweight, and SSE), and equation 12 (for thermal stress) are tabulated in Table 3,9 3-2
[ _ _ _ _ _ _ _ _ . _ _ _ _ _
TABLE 3-1 SUP9tARY OF DIMENSIONS, PHYSICAL PROPERTIES, AND FAULTE0 LOADS FOR PRIMARY LOOP NO.1 0F NORTH ANNA PLANTS NO. 1 & 2 Faulted Loads
- Inside Wall Yield Ultimate Bending Direct Stress Weld Radius Thickness Stress Stress Flow Stress. Axial Load Moment (kst) locations (in) (in) o, d' C (kips) (in-kips) F, M o, [ b F, M b 'x " F
- T i 1*
- 14.60 2.43 21.20 65.20 43.20 1871 26322 22.51 2 15.60 2.70 18.70 67.00 42.85 1841 22817 16.45 3 15.60 2.88 19.30 67.00 43.15 1727 20730 14.10 4*** 15.53 2.58 19.30 67.00 43.15 1607 6315 8.84 5 13.85 2.56 19.30 67.00 43.15 1616 12465 13.93 m 6 14.60 2.43 21.20 65.20 43.20 1817 3455 9.69 7 14.60 2.43 21.20 65.20 43.20 1868 15740 16.56 8 15.53 2.58 19.30 67.00 43.15 1612 13702 12.31 9 15.53 2.58 19.30 67.00 43.15 IE07 10883 11.71 10 15.53 2.58 19.30 67.00 43.15 IE06 15492 13.86 11 15.53 2.58 19.30 67.00 43.15 1919 24445 18.45 12 13.85 2.56 19.30 67.00 43.15 1701 13984 15.17 13.85 2.56 19.30 67.00 43.15 1702 12332 14.21 13 13.85 2.56 19.30 67.00 43.15 1710 10731 13.30 14 l
l
- Includes internal pressure
- Load critical location
- Toughness critical location d
)
l TABLE 3-2 SIMtARY OF DIMENSIONS, PHYSICAL PROPERTIES ~. AND FAULTED LOADS FOR PRIMARY LOOP NO. 2 NORTH ANNA PLANTS NO. 1 & 2 s
faulted Loads inside Wall Yield Ultimate Bending Direct Stress Weld Radius Thickness Stress Stress Flow Stress Antal Load Moment (ksi) locations ( f.n) (in) o a, p j'C (l.jps) (in-kips) F, M b
y F, M b "x " F
- T 1** 14.60 2.43 21.20 65.20 43.20 1:371 26100 22.38 2 15.60 2.70 18.70 67.00 43.15 1339 23244 16.63 3 15.60 2.88 19.30 67.00 43.15 1714 20628 14.02 4 15.53 2.58 19.30 67.00 43.15 1500 5886 8.62 5*** 13.85 2.56 19.30 67.00 43.15 1521 ,
10880 13.03 6 14.60 2.43 21.20 65.20 43.20 1373 3529 9.74 7 14.60 2.43 21.20 65.20 43.20 1369 15927 16.67 8 15.53 2.58 19.30 67.00 43.15 1504 12648 11.79 9 15.53 2.58 19.30 67.00 43.15 1304 10671 11.60 10 15.53 2.58 19.30 67.00 43.15 1302 15138 13.68 11 15.53 2.58 19.30 67.00 43.15 1919 24530 18.49 12 13.85 2.56 19.30 67.00 43.15 1719 13954 15.23 13 13.85 2.56 19.30 67.00 43.15 1720 12355 14.30 14 13.85 2.56 19.30 67.00 43.15 1726 9513 12.66
- Includes internal pressure
- Load critical location
- Toughness critical location
TA8LE 3-3
SUMMARY
OF DIMENSIONS, PHYSICAL PROPERTIES.AND FAULTED LOADS FOR PRIMARY LOOP NO. 3 0F NORTH ANNA PLANTS NO. 1 & 2 l'aulted Loads
- Inside Wall Yleid Ultimate 8ending Direct Stress Weld Radius Thickness Stress Stress Flow Stress Axial Load Moment (ksi) locations (in) (in) o y u [ j (l.jps)
F, (in-kips)
My F,
a, = 7 + 7 M
b 1** 14.60 2.43 21.20 65.20 43.20 1806 26551 22.37 2*** 15.60 2.70 18.70 67.00 42.85 1810 21598 15.80 3 15.60 2.88 19.30 67.00 43.15 1725 21253 14.31 4 15.53 2.58 19.30 67.00 43.15 1601 6580 8.94 5 13.85 2.56 19.30 67.00 43.15 1659 12202 13.96 6 14.60 2.43 21.20 65.20 43.20 1824 2944 9.21 7 14.60 2.43 21.20 65.20 43.20 1321 14823 15.85 8 15.53 2.58 19.30 67.00 43.15 1505 14432 12.63 9 15.53 2.58 19.30 67.00 43.15 1306 10861 11.70 10 15.53 2.58 19.30 67.00 43.15 1305 15165 13.70 11 15.53 2.58 19.30 67.00 43.15 1315 23808 18.14 12 13.85 2.56 19.30 67.00 43.15 1756 14881 15.92 13 13.85 2.56 19.30 67.00 43.15 1757 13092 14.88 14 13.85 2.56 19.30 67.00 43.15 1767 10612 13.47
- Includes internal pressure
- Load critical location
- Toughness critical location I ll l l OMM
TABLE 3-4 NORMAL OPERATING LOADS AND OTHER DATA FOR NORTH ANNA PLANTS 1 & 2 LOOP NO. 1 TION NORMAL OPERATING LOADS OTHER NORMAL (SEE FIGURE ODERATING DATA 3-1) Fx (kips) Mb (in-kips) 1 1,509 21,333 1
2 1,604 17,044 PRESSURE (psig) {
1 HOT LEG 2,259 3 1,641 12,610 ]
CROSS-0VER 2,221 I LEG 4 1,544 3,199 COLD LEG 2,305 5 1,314 5,063 6 1,509 1,112 TEMPERATURE (*F)
HOT LEG 620.2 7 1,510 11,234 CROSS-0VER 555.5 LEG 8 1,550 7,703 COLD LEG 555.5 9 1,763 7,297 10 1,763 12,752 11 1,855 17.244 12 1,353 4,156 13 1,353 3,812 ;
14 1,356 .
5,256 3-6 s
TABLE 3-5 NORMAL OPERATING LOADS AND OTHER DATA FOR NORTH ANNA PLANTS 1 & 2 LOOP NO. 2 NORMAL OPERATING LOADS L TION OTHER NORMAL (SEE FIGURE OPERATING DATA 3-1) Fx (kips) Mb(in-kips)
- 1 1,502 21,518 PRESSURE (psig) 2 1,598 17,264 2,259 HOT LEG CROSS-0VER 2,221 3 1,637 12,826 LEG COLD LEG 2,305 4 1,541 3,426 l
5 1,309 4,139 l l
6 1,505 1,092 TEMPERATURE (*F)
H0T LEG 620.2 7 1,505 11,379 CROSS-0VER 555.5 8 1;546 7,909 LEG COLD LEG 555.5 9 1,764 9,247 10 1,764 12,830 11 1,858 17,490 12 1,358 4,320 13 1,358 4,056 14 1,352 4,853 l
l 3 l l
9 TABLE 3-6 ,
NORMAL OPERATING LOADS AND OTHER DATA FOR NORTH ANNA PLANTS 1 & 2 LOOP NO. 3 j
i, TION NORMAL OPERATING LOADS (SEE FIGURE OTHER NORMAL 3-1) OPERATING DATA Fx (kips) Mb(in-kips)
. I 1,487 21,781 2 1,601 16,380 PRESSURE (Psig)
IHOT LEG' 2,259 3 1,641 12,904 CROSS-0VER 2,221 LEG 4 1,544 3,233 COLD LEG 2,305 5 1,317 4,874 6 1,505 382 TEMPERATURE (*F)
HOT LEG 620.2 7 1,505 10,508 CROSS-0VER 555.5 LEG 8 1,549 7,874 COLD LEG 555.5 9 1,767 9,481 10 1,767 12,971 11 1,855 17,276
+
12 1,361 4,364 13 1,361 3,988 14 1,357 5,049 3-8
I i
TABLE 3-7 NORMAL OPERATING LOADS AT THE LOAD CRITICAL LOCATIONS FOR NORTH ANNA PLANTS NO. 1 & 2 LOAD CRITICAL LOCATION NORMAL OPERATING LOADS WELD NO Mb (in-kips)
PLANT LOOP (SeeFigl3-1) Fx (kri)
Y e
No. I 1 1,509 21,333' l No. 1 ,
8 1,502 21,518 No. 2 1 No. 2 i
No. 3 1 1,487 21,781 i
TABLE 3-8 CALCULATED AND ALLOWABLE STRESS FOR ASME NB-3600 CRITERIA STRESS RESULTS (1) APPLICABLE (ksi)
LOAD WELD ASME CONDITION LOCATION NS-3600 CALCULATED ALLOWABLE RATIO 0F 1 EQUATION STRESS STRESS CALC. STRESS TO ALLOWABLE 1 9F 13.13 53.9 .24 i
1
- 2. 9F 19.39 53.9 .36
! FAULTED 3 9F 19.78 57.9 .34
- 4 9F 21.30 55.5 .38 :
i 5 9F 21.60 57.9 .37 1-1
-1 12 12.52 53.9 .23
- i
! 2 12 16.02 53.9 .30 NORMAL
, OPERATING ;
3 12 6.99 57.9 .12 4 12 10.09 55.5 .18 j i
i 1 5 12 6.09 57.9 .11 NOTE (1) TheHighest Stressed Weld Location of all Loops 3-10
REACTOR VESSEL 4
~
HOT LEG D/ - _
= COLD LEG C @
O
%,* p R @
7 COOLANT CIRCULATING CROSS-0VER LEG STEAM GENERATOR V
I d s ,
0 @
HOT LEG Temperature:620.2*F; Pressure: 2,259 psi CROSSOVER LEG Temperature:555.5'F; Pressure: 2,221 psi COLD LEG Temperature:555.5*F; Pressure: 2,305 psi Figure 3-1 Schematic Diagram of Primary Loop Showing Weld Locations 3-11
CRACK LENGTil A
t = 2.43 N inche -'
______.____.__.__L___.______.____
w Q & i 1 + F -w \
l M
h
=29. O hes NOTES: (1) Typical dimensions shown are that of Location 1, see Fiqure 3-1 (2) Loads F and M at various locations, see Table 3-1 through 3-3 ,
FIGURE 3-2 Reactor Coolant Pipe
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 f racture 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 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 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:
v- .
a,c.e 4-1
a,c.e
.O i
The analytical model described above accurately accounts for the piping 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 11).
4.2 Local Failure Mechanism ,
The local mechanism of failure is primarily dominated by the crack tip ,
behavior in terms of crack-tip blunting, initiation, extension, and i'r.;;1y 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 Ic from a J-integral resistance curve is a material parameter defining the crack initiation. If, for a given load, the calculated J-integral value is shown to be less than the J g of the material, then the crack will not initiate. If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:
dJ E T =
app da 2
,f 4-2
'where:
T,,, = applied tearing modulus E = modulus of elasticity a
J .c.e (flow stress) a f=[
a = crack length g
j a.c.e In sungnary, the local crack stability will be established by the two-step criteria:
J<J g or T,,, <T mat 1fJ>J
- Ic In this analysis, a hypothesized circumferential through-wall 7.5-inch long flaw is taken as a reference flaw and is used as a basis for evaluation.
4.3 Material Properties The materials in the North Anna Units 1 and 2 primary loops are cast stainless steel (SA-351-CF8A) and associated welds. The fittings are made of ,
SA-351-CF8M cast stainless steel. The tensile and flow properties of the
~
critical locations on each loop are given in Tables 3-1 through 3-3.
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 Figures 4-2 and 4-3 taken from References 12 and 13. J gg is observed to be over 5000 j
in-lbs/in .2 However, 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 j
determine the effects of thermal aging on piping integrity, a detailed study l
4-3
~ -
was carried out in Reference 13. In that report, fracture toughness results were presented for a material representative of [
]# Toughness results were provided for the material in the full service life condition and these properties are also presented in k value far this material Figure 4-4 of this report for information. The J ;
- at operating temperature was approximately [ J .c.e and the l maximum value of J obtained in the tests was in excess of [
a J c.e The tests for this material were conducted on small specimens and therefore rather short crack extensions resulted (maximum !
extension was 4.3 m). Therefore it is expected that higher J values would be sustained for larger specimens. T was [ ] at operating temperature. As noted in Appendix B, the end-of-service life toughnesses for those heats which fall below that of [ ]a,c.e are established using the alternate toughness criteria methodology of Reference
- 22. The end-of-service life J k values are sununarized in Table 4-1 for both the load critical locations and toughness critical locations.
Available data on aged stainless steel welds (Reference 13 and 14) indicate the J gg values for the worst case welds are of the same order as the aged material but .the slope of the J-R curve is steeper. Higher J-values (in 2
excess of 3000 in-lb/in ) have been obtained from fracture tests of these welds. 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 temperature. Therefore, weld regions are less limiting than the cast material.
In the fracture mechanics analyses that follow, the fracture toughness properties established above will be used as the criteria against which the applied fracture toughness values will be compared.
4.4 Results of Crack Stability Evaluation Crack stability evaluations have been performed at locations 1 through 5 of Figure 3-1. Location 1 is the load critical location and locations 2 through 5 are the toughness critical locations. Figures 4-5 through 4-9 show plots of 4-4
the plastic limit moment as a function of through-wall circumferential flaw length for these critical locations. The limit moments were calculated using the ASME Code minimum properties at operating temperature. The maximum applied bending moments can be plotted on the figures, and used to determine the critical flaw lengths. The predicted critical flaw lengths for the five locations of interest are provided in Table 4-2.
The outer surface axial stress (a,) at the load critical location (loop 1, location 1, figure 3-1) is 22.51 ksi (see Tables 3-1 through 3-3). The internal pressure at that location is 2259 psig. Stresses due to the internal pressure are cal.culated using Lame' equations for thick cylinders.
a c (circumferential stress): 12.53 ksi e radial stress: 0 r
The von Mises effective stress, a,ff, is given by a,ff = (8, -8
- 7) + (a c ~#
r}
- I# a -#
c) and is [ ]"'.
Thus the effective stress is less than the yield stress and by the von Mises plasticity theory, yielding does not occur. Also, similar consideration at the toughness critical location confirms that the effective stress is less than the yield stress. Hence, linear elastic fracture mechanics is applicable for analyzing the pipes wi'th hypothesized flaws. The analytical method used for the local stability evaluation is sunenarized below.
The stress intensity factors corresponding to tension and bending are expressed, respectively, by (see Reference 15)
K sa F tI")
t"#t K
b"#by va FI) b l
l 4-5
where F (a) and F (a) are stress intensity calibration factors t b corresponding to tension and bending, respectively, a is the half-crack length, a is the half-crack angle, a is the remote uniform tensile t
stress, and ab is the remote fiber stress due to pure bending. Data for F (a) and F (a) are given in Reference 15. The effect of the yielding t b near the crack tip can be incorporated by Irwin's plastic zone correction method (see Reference 16) in which the half-crack length, a, in 'these formulas 1 is replaced by the ef fective crack length, a,ff, defined by a,fg = a + b 25 a y 2 for plane stress plastic corrections, where a is the yield strength of the material and K is the total stress intensity due to combined tensile and bending loads. Finally, the J -value g is determined by the relation J g =
K2 /E, where E = Young's Modulus at temperature.
The applied J values are then calculated for various flaw lengths, including the reference flaw length of 7.5 inches at all the locations of interest, as shown in Table 4-3. The corresponding J gg values established in Appendix B are also entered in this table for comparison.
At the load critical location, J,pp was found to be [ ]**C
2 in-lbs/in which exceeds J Ic but is less than J .
Therefore, the applied tearing modulus, T,pp, was calculated using the methodology of reference 15, ar.d found to be [ ]**C which is a factor of over 3 below Tat f[ ] . Thus the flaw under consideration will remain stable. A similar estimate was obtained for a
[ ]a c.e through-wall flaw. The purpose of the evaluation was to investigate the crack stability for a postulated flaw larger in size than the 7.5-inch reference flaw. The maximum applied J was estimated to be a 2
[ J ,c.e in-lbs/in . The applied tearing modulus T,pp, was calculated and found to be [ Ja .c.e which is significantly less than i l 4-6 A
- - - - - - -r- ----r----, -- -- _ _ _ _ _ _
T mat.
Therefore, it is further concluded that a postulated [ ]**C
through-wall flaw at the load critical location 1 will remain stable f rom both a local and global stability standpoint.
At the toughness critical locations, the applied values of J were equated with the corresponding J gg values for a conservative estimate of the critical flaw sizes. The critical flaw sizes for the critical locations are presented
- n Table 4-3.
In summary, crack initiation is not expected to occur in North Anna Units 1 a
ana 2 primary loop even for circumferential flaws about [ J c.e long at the load critical location. At one of the low toughness regions (most critical location) the critical flaw size is found to be [ ].
e 4-7
TABLE 4-1 SUPNARY OF THE ALTERNATE TOUGHNESS Jge FOR THE HEATS OF MATERIALS WHICH DO NOT SATISFY THE TOUGHNESS CRITERIA-0F [ ]a.c.e(FROM APPENDIX B)
- _ a,c.e i
m
.=
TABLE 4-2 PREDICTED CRITICAL CRACK LENGTH AT VARIOUS LOCATIONS BASED ON LIMIT LOAD METHOD MAXIMUM LOADS AT THE LOCATION WELD LOCATION Predicted Axial Force Bending Moment Critical Crack eld. p Fx , (kips) Mb , (in-kips) Length (in)
~
1 3 1,807 26,439 2 1,839 22,898 2
f _
_ ,a,c.e 3 3 1,726 21,254 4 1 1,807 10,884 5 1 1,710 12,466 See Figure 3-1 for Weld Locations 4-9
TA8LE 4-3 SufGERY OF APPLIED J VALUES VS J ge TOUGHNESS FOR NORTH ANNA PLANTS 1 & 2 a,c.e l
l i
?
E 1
~
(1) Weld locations. See Figure B-1 through B-6. -
(2) Jge values of locations 2 through S. See Table B-3 (3) The criterion for Location 1 is Tapplied < Tmat L_________ _ _ _
- /// '////lil j, ,
2a i
N YN))
NEUTRAL AXIS a,c.e FIGURE 4 1 [ 3 STRESS DISTRIBUTION 4-11
r
- - a ,c .e .
l 1
)
l
\
i i
Figure 4-2 J vs aa for SA351-CF8M material at 600*F taken from a 23 inch diameter pipe 4-12
. - a,c.e e
Figure 4-3 J vs aa for SA351-CF8M Cast Stainless Steel at 600*F i
4-13
o a,c.e FIGURE 4-4 J-aa Curves at Oifferent Temperatures for Aged Material
[ ]C (7500 Hours at 400*C) 4-14
-- a,c.e
- i. I v
i FLAW GEOMETRY 00 = 34.06 in t = 2.43 in p = 2,259 psi F = 1,807 kips e
y = 21.20 ksi u = 65.20 ksi o
of = 43.20 ksi Temp = 620.2' F l
FJGURE 4-5 Predicted " Critical" Crack Length at Weld Location No. I for North Anna Plants No. I and 2 4-15
a,c.e L I
$/
/AJL70 .
U l
FLAW GEOMETRY 00 = 36.60 in t = 2.70 in P = 2,259 psi F =
1,839 kips
- y 18.70 ksi
= 67.00 ksi o,
= 43.15 ksi of Temp = 620.2* F l
l 1
FIGURE 4-6 " Critical" Crack Length at Weld Location No. 2 for North Anna Plan'.c No. I and 2 4-16
a,c.e L l
$/
/Al/B U
l FLAW GEOMETRY 00 = 36.96 in t = 2.88 in P
= 2,221 Psi T = 1,726 kips e
y 19.30 ksi o
u 67.00 ksi
'f 43.15 ksi Tamp
- 555.5eF FIGURE 4-7 " Critical" Crack Length at Weld Location No. 3 for North Anna Plants No. I and 2 4-17 l
l l- - - - - _ - _ _ .
a,c.e L l v i FLAW GEOMETRY 00 = 36.22 in t = 2.58 in p = 2,221 psi F = 1.607 kips e
y
= 19.30 ksi e, = 67.00 ksi
=
of 43.15 ksi Temp = 555.5T FIGURE 4-8 " Critical" Crack Length at Weld No. 4 for North Anna Plants No. I and 2 ,
i 4-18
l a,c.e L l M/
f@ \
() !
l l l
FLAW GEOMETRY 00 = 32.82 in t = 2.56 in P = 2,305 psi F = 1,616 kips e "
y 19.30 ksi
'u 67.00 ksi e
r = 43.15 ksi Temp = 555.5T FIGURE 4-9 " Critical" Crack Length at Weld Location No. 5 of Loop No. I for North Anna Plants No. 1 & 2 4-19
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 f ailure of this component. If, in the unlikely event, 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 c rac ks .
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, DH (L/DH ) is greater than [ J a .c.e. both [ ]a,c.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 [ ].
However, fsr large L/0H values, friction pressure drop will become important 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
[
a J .c.e . The flow rate through a crack was calculated in the following manner. Figure 5-1 from Reference 16 was used to estimate the critical pressure, Pc, for the primary loop enthalpy condition and an assumed flow. Once Pc was found for a given mass flow, the [
Ja ,c.e was found from Figure 5-2 taken 5-1
from Reference 17. For all cases considered, since [
a J ,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 A pf = [ ]a,c.e (5-1) 4 where the friction f is determined using the [ ].a.c.e The crack relative roughness, c, was obtained f rom fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [ ]a,c.e RMS.
The f rictional pressure drop using Equation 5-1 is then calculated for the assumed flow and added to the [
a J .c.e to obtain the total pressure drop from the primary system to the atmosphere. That is, for the primary 1:op Absolute Pressure - 14.7 = [ Ja ,c.e (5-2) for a given assumed flow G. If the right-hand side of Equation 5-2 does not agree with the pressure dif ference 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 i c rack. This calculational procedure has been recommended by [
]* # for this type of [
a J .c.e calculation.
5.4 LEAK RATE CALCULATIONS The leak rate calculations for the North Anna Units 1 & 2 are perfonned for the load critical location of the two plants and the toughness critical locations identified in Table B-3 of Appendix B. The normal operating loads used for leak rate calculations are given in Tables 3-4 through 3-6.
5-2
k l
Leak rate estimates are performed by applying the normal operating bending moment in addition to the normal operating axial force. These loads are applied to the 7.5-inch through-wall reference flaws postulated at the critical locations identified in previous sections. In order to determine the sensitivity of leak rate to flaw size, through-wall flaws [ ]
in length are also postulated.
The North Anna plant; have RCS pressure boundary leak detection systems which are consistent with the guidelines of Regulatory Guide 1.45. Thus for the 7.5-inch flaw, a factor in excess of 20 exists between the calculated leak rate and the criteria of Regulatory Guide 1.45. Relative to the [
a J ,c.e These leak rate margins are calculated at the toughness critical location (No. 5 at reactor vessel inlet).
The leak rate margins are significantly higher at other locations.
1 I
)
l l 5-3 l
TABLE 5-1
SUMMARY
OF LEAK RATE CALCULATION FOR NORTH ANNA PLANTS 1 & 2 l
l IDENTIFICATION OF LOATION REFERENCE CRACK SIZE SMALLEST CRACK SIZE DESCRIPTION WELD
- PLANT LOOP OF THE 7.5 in. _
_ a.c.e NO. N0. NO. LOCATION ~
LEAK RATE 4ARGIN W/
GPM RESPECT TO R.G. 1.45 Worst Lead Critical 1 1 1 Location 40.2 40 Lowest J Location IC y' 2 2 3 of Weld 46.0 46
- o No. 2 Lowest J 3 1 2 Location IC 28.4 28 of Weld No. 3 Lowest J 4 1 1 Location ic 31.5 31 of Weld No. 4 Lowest J 5 2 2 L cation IC of Weld 20.2 20 No. 5
- See Figure 3-1 for Weld Locations _
l l
l l
'" a,c,e l
=
_3 i
i t
8 a
5 N
3 2
l STAGNATION ENTHALPY (10 Stu/lb)
Figure 5-1 Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures 5-5
- a,c.e o
9 E
a E
w E
4 e
t 6
LENGTH / DIAMETER RATIO (L/D)
Figure 5-2 [ ]a,c.e Pressure Ratio as a Function of L/D 5-6 l
- ., a ,c ,e a,c.e
~
f l L _
l t _- _ =
Figure 5-3 Idealized Pressure Drop Profile Through a Postulated Crack 5-7 ,
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 [
] region of a typical system [
]' # 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.
A finite element stress analysis was carried out for the inlet nozzle safe-end region of a plant typical in geometry and operational characteristics to any Westinghouse PWR System. The specific system was a plant with piping outside diameter of 33 inches and wall thickness of 2.85 inches. The corresponding dimensions for the North Anna Plants are 33.08 inches in diameter and 2.81 inches wall thickness. These differences are insignificant as far as fatigue crack growth analysis is concerned. 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:
- ,a,c.e 2
Fatigue crack growth rate laws were used [
l JC The law for stainless steel was derived f rom Reference 19, with a very conservative correction for the R I ratio, which is the ratio of minimum to maximum stress during a transient.
For stainless steel, the fatigue crack growth formula is:
6-1
h=(5.4x10-12) g,,,4.48 inches / cycle where K,,, = K ,, (1-R)0.5 i
min max
[
j a,c.e a,c.e
[ ]
' 8'C
where: [ ]
The calculated fatigue crack growth for semi-elliptic surface flaws of circumf erential orientation and various depths is sununarized in Table 6-1 The results show that the crack growth is very small, regardless [
a 3 .c.e 6-2
TABLE 6-1 FATIGUE CRACK GROWTH AT [ ]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 i
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
6-3
- a,c.e Figure 6-1 Typical Cross-Section of [ ]a c.e 6-4 l
- a ,c.e i
j i.
f i
Figure 6-2 Reference Fatigue Crack Growth Curves for [
Ja.c.e 6-5
a.c. e a
l l
J e
u i
Figure 6-3 Reference Fatigue Crack Growth Law for[ ]a.c.e in a Water En iv ronment at 6000F 6-6
, -- ----rww+ - - ---e9wmPr ,-m-W---e-q*w-- T--wvmyN * *-*Pr9 w w - - - + =-w- - '
- ur -ww--u--e----- _--- -----__ _ _ _ .
7.0 ASSESSMENT OF MARGINS 3
In reference 1, the maximum design moment was 45.6x10 in-kips, whereas, the j maximum moment as noted in Section 3.0 of this report is [
Ja .c.e in-kips. At the toughnest critical .
locations the design moment is even lower.
The results of the crack stability evaluation and leak rate calculations for the load critical location and the four toughness critical locations, examined l in this report, are sunenarized in Table 7-1. Margins for these critical locations are discussed below.
At load critical location 1 a 7.5 inch through-wall circumferential flaw is seen to be stable exhibiting a J,,, <J and a T,,, less than T i by over a factor of three. A[ ] long flaw is also shown to be stable at this location. At toughness critical locations 2 through 5 a 7.5 inch long flaw produces a J,,, less than the corresponding J . From Table 7-1 the lowest " critical" flaw size is seen to be [ ] '' inches at location 2.
As shown in Section 3.0, a margin of a factor greater than 2 and 3 exist between calculated and ASME Code allowable faulted conditions and thennal stresses, respectively.
I In Section 4.4 the " critical" flaw sizes at load critical location 1 and i toughness critical locations 2 through 5 are calculated (using the limit load l method) to be at least [ ] incher (see Table 4-2). Thus, based on the
( above, the " critical" flaw sizes at these locations will, of course, exceed j a
[ J .c.e inches.
In Section 5.0, it is shown that at load critical location 1, a flaw of 7.5 inches would yield a leak rate in excess of 98 gpm while for a [ ]a,c.e ;
inch flaw. the leak rate is [ ]a,c.e gpm. Thus, there is a margin of at least 2 between the flaw size that gives a leak rate well exceeding (by a i
! f actor of 14) the criterion of Regulatory Guide 1.45 and the " critical" flaw '
l size of [ ] I 7-1
I At toughness critical location 2, a flaw of [ ]"'C inches would yield a leak rate of [ ]**' gpm while a flaw of 7.5 inches would produce leakage of 46 gpm. Thus, at toughness critical location 2, there is also a margin of at least 2 between the flaw size that gives a leak rate exceeding the 1 gpm criterion of Regulatory Guide 1.45 and the " critical" flaw size of about [
]"'C. Ample margins of similar magnitude or higher are apparent at other toughness critical locations (see Table 7-1).
l l
In summary relative to I l
l
- 1. Loads
- a. North Anna Units 1 and 2 are enveloped both by'the maximum loads and J values in Reference 1 and the J values obtained in tests of aged material.
- b. Margins at the critical locations of at least 2 and 3 on f aulted condition and therwel stresses, respectively, exist relative to ASME Code allowable values.
- 2. Flaw Size
- a. A margin of at least 2 exists between the critical flaw size and the flaw yielding a leak rate exceeding the 1 gal / min criterion of Regulatory Guide 1.45.
- b. If limit load is used as the basis for " critical" flaw size, the margin for global stability would exceed 4 when compared to the reference flaw.
1-2
- 3. Leak Rate At load critical location 1, a margin in excess of 98 exists for the 7.5-inch long flaw between the calculated leak rate and the 1 gpm criterion of Regulatory Guide 1.45. For the toughness critical locations these margins between the calculated leak rates and 1 gpm criterion vary between a factor of 20 and a factor of 46 for the 7.5-inch long postulated flaws.
i 4
e d
i 7-3 i
TABLE 7-1
SUMMARY
OF STAOILITY AND LEAK RATE CALCULATIONS i LOCATION Jge Tmat Jean Crack Size Japp11H Tapp11ed Leak Rate 1 (in-Ib/in2) (in-Ib/in 2) (in-lb/in2) (in ) (in-Ib/in2) gp.
1(a) 750 60 2200
_7. 5 944 -
98.3 a,c.e i
. 2(b) 408.3 5.53 541.9 l _J
~7.5 325 -
46.0 , ,,c,,
[ 3(b) 487.9 20.9 939.6
[7.5 229 -
28.4_' ,,,,,
i 4(b) 296.1 4.44 403.4
'_7 . 5 137 -
31.52a,c.e i
i l
l 5(b) 343.7 0 343.7 i
[7.5 259 -
20.25ac.e
]
(a) Load-critical location
]
(b) Toughness critical location i
l
l I
8.0 CONCLUSION
S i
This report justifies the elimination of RCS primary loop pipe breaks for the North Anna Units 1 and 2 as follows-1
- 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.
- 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 integrity of the primary piping are negligible.
- d. A large margin exists between the leak rate of the reference flaw and the criteria of Reg. Guide 1.45.
- e. Ample margin exists between the reference flaw chosen for leak detectability and the " critical" flaw.
- f. Ample margin exists in the material properties used to demonstrate end-of-service life (relative to aging) stability of the reference flaw.
The reference flaw will be stable throughout reactor life because of the ample margin in d, e, and f 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 need not be considered in the structural design basis of the North Anna Units 1 and 2. l l
1 8 -1
9.0 REFERENCES
- 1. WCAP-9558, Rev. 2 " Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall Crack,"
Westinghouse Proprietary Class 1. June 1981.
- 2. USNRC Generic letter 84-04,
Subject:
" Safety Evaluation of Westinghouse Topical Reports Dealing with Elimination of Postulated Pipe Breaks in RdR 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. 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 HS-EPR-2519. Westinghouse (E. P. Rahe) to NRC (D. G.
Eisenhut), Westinghouse Proprietary Class 2, November 10, 1981.
I 8. Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated April 25, 1983, i
- 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
- - - -,-,--,---,-nn--,--+--,- - , , --------------r-----ma----a a,,-w- ,, .e- - ---- - , , - - -, - - - ,,--e-------,---., e .-r . ,,,-,,--mn-----.-
- 11. Kanninan, M. F., et. al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks", EPRI NP-192, September 1976.
- 12. Landes, J. D., et. al., " Fracture Toughness of 316 Stainless Steel Piping Material at 600*F, Westinghouse R&D Report 79-7D3-PIPRE-R1, May 17,1979 (Westinghouse Proprietary Class 2).
- 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 f ect of Aging on Mechanical Properties of Austenitic Stainless Steel Casting and Welds", presented at SMIRT 7 Post Conference Seminar 6 - Assuring -
l Structural Integrity of Steel Reactor Pressure Boundary Components, August 29/30, 1983, Monterey, CA.
f
- 15. Paris, P. C., Tada. H., "The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulating Circumferential Through-Wall Cracks, "NUREG/CR-3464 September 1983.
- 16. Irwin, G. R., " Plastic Zone Near a Crack and Fracture Toughness," Proc.
7th Sagamore Conf erence, P. IV-63 (1960).
- 17. [
j a.c.e
- 18. [
j a.c.e
- 19. 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.
9-2
- 20. [
j a.c.e
- 21. [
ya .c.e
- 22. Witt, F. J., and Kim, C. C., " Toughness Criteria for Thermally Aged Cast Stainless Steel", Westinghouse Proprietary Class 2 Report WCAP-10931, Rev.1, July 1986, i
J G
i i
l I !
9-3 i
APPENDIX A a,c e LIMIT MOMENT -
M h
6 0
i l
i I
i w
A-1
l l
l t
l
\
t l
l l
l l
a,c.e '
FIGURE A-1 Pipe with a Through-Wall Crack in Bending A-2
4 APPENDIX B ALTERNATE TOUGHNESS CRITERIA FOR THE NORTH ANNA UNITS 1 & 2 PRIMARY LOOP COMPONENTS i
8.1 INTRODUCTION
Not all of the individual piping ccmponents of the North Anna Units 1 and 2 primary loop satisfy the present [ ]a,c.e criteria. In this appendix, the alternate toughness criteria developed in Reference 22 will be used to categorize the various individual piping components thus establishing criteria based upon which the mechanistic pipe break evaluation will be performed.
8.2 CHEMISTRY AND KCU TOUGHNESS The correlation of reference 14 which is based on the chemistry of the cast stainless steel piping was used to calculate the associated KCU values. The chemistry and the end-of-service life KCU toughness values are given in Table 8-1 for these heats which fail to meet the current [ ]a,c.e criteria.
A total of thirteen heats do not exhibit an end-of-service life KCU value of at least [ ]#, thus requiring specific f racture toughness i evaluations as described in Reference 22.
B.3 IDENTIFICATION OF PRIMARY LOOP COMPONENT North Anna Units 1 and 2 are typical 3-loop Westinghouse pressurized water reactor plants. The material heat numbers for the pipe and piping components vary from loop-to-loop, and from plant to plant. The heat numbers associated with each of the loops are identified in Figures B-1 through B-6. The heats of materials which do meet the current [ J criteria are marked with an asterisk (*) in these figures. Alternate toughness criteria are required for the remaining heats. A summary of these heats requiring alternate toughness criteria is provided in Table B-2.
l l
B-1
, , - , , _,, , - , - - , .- ,-- - ---------,,---,,,,-,,,-,----,-,--r---- - .- - - - - , - - - - - - - - .- - . , , - . _ . - - , , , _ - , - -
B.4 ALTERNATE TOUGHNESS CRITERIA FOR THE NORTH ANNA UNIIS 1 & 2 PRIMARY LOOP 4
COMPONENTS The alternate toughness criteria for the North Anna Units 1 & 2 primary loops are obtained by applying the methodology of Reference 22 to Table B-1.
As the first step, the [
ya .c.e a ,c.e Where
[
j a.c.e The next step is to determine the categories identified in Reference 22. In s umma ry,
[ _
j a.c.e The final step is to use Table 2-8 of Reference 22 to determine the end-of-service life toughness. The f racture toughness results for North Anna Units 1 and 2 are presented in Table B-3.
B-2
. . . - _ - - ,-,-.n -,.- , - - , . , , - , , . . - . - , . - - - - , - - - - - - . - . - - - _ , - - - - . - - - ., - - - - . - - - - -
. TABLE B-1 CHEMICAL COMPOSITIONS, END-0F-LIFE SERVICE KCU OF THE HEATS OF THE MATERIALS FORTHEPRIMARYLOOPCOMPONENTSWHICH00NO{'{AJISFYTHE
]
! TOUGHNESSCRITERIAFOR{'
L ]'-
t
..c..
n I
i i Y w
i 1
i e
l
! I, 3
- Identification numbers, see Figures B-1 through B-6 i
TABLE B-2 IDENTIFICATI0fl 0F THE HEATS OF MATERIALS WHERE ALTERNATE TOUGHNESS CRITERIA ARE NEEDED
- a,c.e ao i
NOTE: The blank space means that the teat of the material does satisfy the toughness criteria of [ ]a,c.e
. TABLE 8-3 ALTERNATE FRACTURE TOUGHNESS FOR MATERIAL HEATS WHICH DO NOT SATISFY THE TOUGHNESS CRITERIA FOR [ Ja.c.e
_ a,c.e 1
l
- The lowest J Ic value for each weld location.
- Conservatively selected per Reference 22.
B-5 1 .
- = . - . . .. - _-
FIGURE B-1 PRIMARY LOOP PIPING COMPONENTS WHICH C0 NOT SATISFY
[ 3a.c.e TOUGHNESS CRITERIA PLANT NO.1 LOOP NO. 1
~
a,c,e i
k f
f l
{
i 1
, *The heat for this location does satisfy the toughness criteria for [ 3a,c.e No alternative criteria is needed.
B-6 n - - .--n...,..- . , , , . n - - , - - - - , . ,-n-.--. - -- , -.--, - . , .- - .,- ,. -
FIGURE R-2 PRD1ARY LOOP PIPING COMPONENTS WHICH DO NOT SATISFY [ ]a,c.e TOUGHNESS CRITERIA PLANT NO. 1 LOOP NO. 2 a,c,e s
1
~
- The heat at this location does satisfy the toughness criteria for [ 3a.c.e No alternative criteria is needed.
B-7 4
i
FIGURE'B-3' PRIMARY LOOP PIPING COMPONENTS WHICH DO NOT SATISY [ ]a,c.e TOUCHNESS CRITERIA-
.xt PLANT- NO.1 LOOP N0. 3 i
4 s s
.s i
1 ,
(
i f
i i
- The heat at this location does satisfy the toughness criteria for [ ]a,c.e No alternative criteria is needed.
B-8
FIGURE B-4 PRIMARY 100P PIPING COMPONENTS WHICH DO NOT SATISFY [ ]a,c.e TOUGHENESS CRITERIA PLANT NO. 2 LOOP N0. 1 a,c.e l
- The heat at location does satisfy the toughness criteria for [ 3a,c3 No alternative criteria is needed.
B-9
i
~
FIGURE B-5 PRIMARY LOOP PIPING COMPONENTS WHICH DO NOT SATISFY [ ]a.c.e TOUGHNESS CRITERIA PLANT NO. 2 LOOP NO. 2 ,,c,,
1 l
l l
/
- The heat at this location does satisfy the touchness criteria for [ ]a,c,5 No alternative criteria is needed.
B-10 4
FIGURE B-6 PRIMARYLOOPPIPINGCOMPONENTSWHICHDONOTSATISFY[ ]a c.e i T0llGHNESS CRITERIA PLANT NO. 2 LOOP NO. 3
_.,a,c.e a
- The heat at this location does satisfy the toughness criteria for [ ]a,c.e No alternative criteria is needed.
B-11 .
________ _ ____ ___ _ _