ML20081K353
| ML20081K353 | |
| Person / Time | |
|---|---|
| Site: | Prairie Island  |
| Issue date: | 03/31/1991 |
| From: | Palusamy S, Witt F WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19302E705 | List: |
| References | |
| WCAP-12876, NUDOCS 9106260332 | |
| Download: ML20081K353 (70) | |
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WESTINGHOUSE PROPP.lETARY CLASS 3
(
TECHNICAL JUST!FICATION FOR ELIMINATING PRESSURIZER SURGE LINE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR PRAIRIE ISLAND UNIT 1 March 1991 D. C. Bhowmick S. A. Swamy Y. S. Lee D. E. Prager K. R. Hsu
'erified:
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V F.' J. Wpt
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Struct0ral Mechanics Technology Approved.
6#.w/d erfM//N
- 5. 5. Pelvsamy, Manager' Diagnostics and Monitoring Technology work Performed Under Shep Order:
NYVP-950 WESTINGHOUSE El.ECTRIC CORPORATION Nuclear ar.d Advanced Technology Civision P.O. Box 2728
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Pittsburgh, Pennsylvania 15230-2728 e 1991 Westinghcuse Electric Corp.
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TABLE OF CONTENTS Sectier Title Pace
1.0 INTRODUCTION
1-1
1.1 Background
1-1 1.2 Scope and Objective 1-1 1.3 References 1-3 2.0 OPERATION AND STABILITY OF THE PRESSURIZER SURGE LINE AND THE REACTOR C00LAN' SYSTEM 2-1 2.1 Stress Corrosion Crack'ng 2-1 2.2 Water Hammer 2-3 2.3 Low Cycle and High Cycle Fatigue 2-4 2.4 Summary Evaluation of Surge Line for Potential Degradation During Service 2-4 2.5 References 2-5 3.0 MATERIAL CHARACTERIZATION 3-1 3.1 Pipe and Weld Materials 3-1 3.2 Material Properties 3-1 3.3 References 3-2 4.0 LOADS FOR FRACTURE MECHANICS ANALYilS 4-1 4.1 Loads for Crack Stability Analysis 4-2 4.2 Loads for Leak Rate Evaluation 4-2 4.3 Loading Condition 4-2 4.4 Summary of Loads and Georatry 4-4 4.5 Governing Location; 4-5 e
5173t/C32*01 10 jjj
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TABLE OF CONTENTS (cont.)
Section Title Page 5.0 FRACTURE MECHANICS EVALUATION 5-1 5.1 Global Failure Mr.Nar; ism 5-1 5.2 Leak Rate Predictions 5-2 5.3 Stability Evaluation 5-4 5.4 References 5-5 6,0 ASSESSMENT OF FATIGUE CRACK GROWTH 6-1 6.1 Introduction 6-1 6.2 Initici Flaw Size 6-2 6.3 Results of FCG Analysis 6-2 F.4 References 6-3
7.0 ASSESSMENT
OF MARGINS 7-1
8.0 CONCLUSION
S 8-1 APPENDIX A Limit Moment A-1 4
e a
su3:<c325eito jy l
I LIST OF TABLES Table Title Pace 3-1 Room Tempera +ure Mechanical Properties of the Pressurizer Surge Line Materials 3-3 3-2 Room Temperature ASME Code Minimum Properties 3-4 3-3 Representative Tensile Properties 3-5 3-4 Modulus of Elasticity (E) 3-6 4-1 Types of Loadings 4-6 4-2 Normal and Faulted Loading Cases for Leak-Before Break Evaluations 4-7 4-3 Associated Load Cases for Analyses 4-B 4-4 Summary of LBB Loads and Stresses by Case for Governing Locations 4-9 5-1 Leakage Flaw Size 5-6 5-2 Summary of Critical Flaw Size 5-7 6-1 Fctigue Crack Growth Results for 10% of Wall Initial Ficw Size 6-4 7-1 Leakage Flaw Sizes, Critical Flaw Sizes and Margins 7-2 7-2 LBB Conservatisms 7-3 517h/032591 10
LIST OF FIGURES Figure Title Ptge 3-1 Prairie Island Unit 1 Surge Line layout 3-7 4-1 Prairie Island Unit 1 Surge Surge Line Shoaing the Governing Locations 4-10 5-1 Fully Plastic Stress Distribution 5-8 5-2 Analytical Predictions of Critical Flow Rates of Steam-Water Mittures 5-9 5-3
[
la,c,e Pressure Ratio as a Function of L/0 5-10 5-4 Idealized Pressure Drop Profile through a Postulated Crack 5-11 5-5 Loads Acting on the Model at the Governing Location 5-12 5-6 Critical Flaw Sizu Prediction for Node 1020 Case 0 5-13 5-7 Critical Flaw Size Prediction for Node 1020 Case E 5-14 5-8 Critical Flaw Size Prediction for Node 1020 Case F 5-15 5-9 Critical Flaw Size Prediction for Node 1020 Case G 5-16 5-10 Critical Flaw Size Prediction for Node 1240 Case 0 5-17 5-11 Critical Flaw Size Prediction for Node 1240 Case E 5-18 t
5173s/032591 10
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LIST OF FIGURES (cont.)
Figure Title Page 5-12 Critical Flaw Size Prediction for Node 1240 Case F 5-19 5-1S Critical flaw Size Prediction for Node 1240 Case G 5-20 6-1 Determination of the Effects of Thermal Stratification on Fatigue Crack Growth 6-5 6-2 Fatigue Crack Growth Methodology 6-6 6-3 Fatigue Crack Growth Rate Curve for Austenitic Stainless Steel 6-7 6-4 Fatigue Crack Growth Rate Equation for Austenitic Stainless Steel 6-8 6-5 Fatigue Crack Growth Critical Locations 6-9 6-6 Fatigue Crack Growth Controlling Positions at Each Location 6-10 A-1 Pipe with a Through-Wall Crack in Bending A-3 4
5173s t3nsi ic yjjj l
SECTION 1.0 INTRODUCTION
1.1 Background
The current structural design basis for the pressurizer surge line requires postulating non-mechanistic circumferential and longitudinal pipe breaks.
This results in additional plant hardware (e.g. pipe whip restraints and jet shields) which would mitigate the dynamic consequences of the pipe breaks.
It is, therefore, highly desirable to be realistic in the postulation of pipe breaks for the surge line.
Presented in this report are the descriptions of a mechanistic pipe break evaluation method and the analytical results that can be used for establishing that a circumferential type break will not occur within the pressurizer surge line.
The evaluations considering circumferentially oriented flaws cover longitudinal cases.
The pressurizer surge line is known to be subjected to thermal stratificatior. and the effects of thermal stratification for Prairie Island Unit 1 surge line has been evaluated and documented in WCAP-12839 (Reference 1-1).
The results of the stratification evaluatica as described in WCAP-12839 have been used in the leak-before-break evaluation presented in this report.
1.2 Scoce and Objective The general purpose of this investigation is to demonstrate leak-oefore-break for the pressurizer surge line.
The scope of this work covers the entire pressurizer surge line from the p mary loop nozzle junction to the pressur:.er nozzle junction. A schematic drawing of the piping system is shown sn Section 3.0.
The recommendations and criteria proposed in NUREG 1061 Ve!ame 3 (References 1-2 and 1-3) are used in this evaluation.
The criteria and the resulting steps of the evaluation procedure can be briefly summarized as follows:
1)
Calculate the applied loads.
Identify the location at which the highest stress occurs.
2)
Identify the materials and the associated material properties, smucuso,,o y.3
3)
Postulate a surface flaw at the guverning location.
Determine fatigue crack growth.
Show that a through wall crack will not re.
t.
4)
Postulate a through wall flaw at the governing location.
The size of the flaw should be large enough so that the leakage is assured of detection with margin using the installed leak detecticn equipment when the pipe is subjected to normal operating loads. A margin of 10 is demonstrated between the calculated leak rate and the leak detection capability.
5)
Using r.aximum faulted loads, demonstrate that there is a margin of at least 2 between the leakage size flaw and the critical size flaw.
6)
Review the operating history to ascertain that operating experience has indicated no particular susceptibility to failure from the effects of corrosion, water hammer or low and high cycle fatigue.
7)
For the material actually in the plant provide the material properties justify that the properties used in the evaluation are representative of the plant specific material.
The flaw stability analyses is performed using the methodology described in SRP 3.6.3 ("eference 1-3).
The leak rate is calculated for the normal operating condit: n.
The leak rate prediction model used in this evaluation is an [
la,c e The crack opening area required fo" calculating the leak rates is obtained by subjecting the postulated through wall flaw to normal operating loads (Reference 1-4).
Surface roughness is accounted for in determining the leak rate through the postulated flaw.
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1-2
. The computer codes used in this evaluation for leak rate and fracture mechanics calculations have been validated (bench marked).
1.3 References 1-1 WCAP-12839, Structural Evaluation of the Prairie Island Unit 1 Pressurizer Surge Line, Considering the Effects of Thermal Stratification 1-2 Report of the U.S. Nuclear Regulatory Commistion Piping Review Committee
- Evaluation of Potential for Pipe Breaks, NUREG 1061, Volume 3, November 1984.
1-3 Standard Review Plan; public comments solicited; 3.6.3 Leak-Before-Break Evaluation Procedures; Federal Register /Vol. 52, No.167/ Friday, August 28, 1987/ Notices, pp. 32626-32633.
1-4 NUREG/CR-3464,1983, "The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulated Circumferential Through Wall Cracks."
i
/
5173s/032591 10
}.3 l
SECTION 2.0 OPERATION AND STABILITY OF THE PRESSURIZER SURGE LINE AND THE REACTOR COOLANT SYSTEM 2.1 Stress Corrosion Cracking The Westinghouse reactor coolant system primary loop and connecting Class 1 lines have an operating history that demonstrates the innerent operating stability characteristics of the design.
This includes a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress corrosion cracking).
This operating history totals over 450 reactor years, including five plants each having over 17 years of cperation and 15 other e
plants each with over 12 years of cperation.
In 1978, the United States Nuclear Regulatory Commission (USNRC) formed the second Pipe Crack Study Group.
(The first Pipe Crack Study Group established in 1975 addressed cracking in boiling water reactors only.) One of the objectives of the second Pipe Crack Study Group (PCSG) was to include a review of the potential for stress corrosion cracking in Pressurized Water Reactors (PWR's).
The results of the study performed by the PCSG were presented in NUREG-0531 (Reference 2-1) entitled " Investigation and Evaluation of Stress Corrosion Cracking in Piping of Light Water Reactor Plants."
In that report the PCSG stated:
"lne PCSG has determined that the potential for stress-corrosion cracking in PWR primary system piping is extremely low because tne ingredients that produce IGSCC are not all present.
The use of hydrazine additives and a hydrogen overpressure limit the oxygen in the coolant to very low levels.
Other impurities that might cause stress corrosion cracking, such as halides or caustic, are also rigidly controlled.
Only for brief periods during reactor shutdown when the coolant is exposed to the air and during the subsequent startup are conditions even marginally capable of producing stress-corrosion cracking in the primary systems of PWRs.
smvouw,o g.5
Operating experience in PWRs supports this determination.
To date, no stress-corrosion cracking has been reporttd in the primary piping or safe ends of any PWR."
During 1979, several instances of cracking in PWR feedwater piping led to the establishment of the third PCSG.
The investigations of the PCSG reported in NUREG-Oo91 (Reference 2-2) further confirmed that no occurrences of IGSCC have been reported for PWR primary coolant systems.
As stated above, for the Westinghouse plants there is no history of cracking failure in the reactor coolant system loop or connecting Class 1 piping.
The discussion below further qualifies the PCSG's findings.
Cor stress corrosion cracking (SCC) tu occur in piping, the following three conditions must exist simultaneously:
high tensile stresses, susceptible material, and a corrosive environment.
Since some residual stresses and some degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimized by properly selecting a material immune to SCC as well as preventing the occurrence of a corrosive environment.
The naterial specifications consider compatibility with the system's operating environment (both internal and external) as well as other material in the system, applicable ASME Code rules, fracture toughness, walding, fabrication, and processing.
i The elements of a water environment known to increase the susceptibility of austenitic stainless steel to stress corrosion are:
oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g.,
sulfides, sulfides, and thionates).
Strict pipe cleaning standards prior to nperation 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 preoperatier.al testing, water chemistry is controlled in accordance with written specifications.
Requirements on chlorides, fluorides, conductivity, and pH are included in the acceptance criteria for the piping.
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. During plant operation, the reactor coolant water chemistry is monitored and maintained within very specific 1 % s.
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 l
operating procedures as a condition for plant operation.
For example, during normal power operation, oxygen concentration in the RCS and connecting Class 1 i
lines is expected to be in th9 ppb range by controlling charging flow chem-l istry and maintaining hydrogen in the reactor coolant et specified concentra-tions.
Halogen concentrations are also stringently controlled by maintaining i
concentrations of chlorides and fluorides within the specified limits.
This I
is assured by controlling charging flow chemistry.
Thus during plant opera-tion, the likelihood of stress corrosion cracking is minimized.
2.2 Water Hammer Overall, therc is a low potential for water hammer in the RCS and connecting surge lines since they are designed and operated to preclude the voiding condition in normally filled lines.
The RCS and connecting surge line including piping and components, are designed for normal, upset, emergency, and faulted condition transients.
The design requirements are conservative relative to both the number of transients and their severity.
Relief valve actuation and the associated hydraulic transients following valve opening are considered in the system design.
Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlled.
Temperature during normal operation is maintained within a narrow range by control rod position; pressure is 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 l
system resistance and the reactor coolant pump characteristics are controlled in the design process.
Additionally, Westinghouse has instrumented typical P
reactor coolant systems to verify the flow and vibration characteristics of Ib the system ard connecting surge lines. Preoperational testing and operating experience have verified the Westinghouse approach.
The operating transients 5173s/032591 10 g.3
of the RCS primary piping and connected surge lines 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 111 of the ASME Code. A furtner evaluation of the low cycle fatigue loading is discussed in Section 6.0 as part of this study in the form of a fatigue crack growth analysis.
Pump vibrations during operation would result in high cycle fatigue loads in the piping system. During operation, an alarm signals the exceedance of the RC pump shaft vibration limits.
Field measurements have been made on the reactor coolant loop piping of a number of plants during hot functional testing.
Stresses in the elbow below the RC pump have been found to be very small, between 2 and 3 ksi at the highest.
Recent field measurements on typical PWR plants indicate vibration amplitudes less than 1 ksi. When translated to the connecting surge line, these stresses would be even lower, well below the fatig:e endurance limit for the surge line material and would result in an applied stress intensity factor below the threshcid for fatigue crack growth.
2.4 Summary Evaluation of Surge Line for Potential Degradation During Service There has never been any service cracking or wall thinning identified in the pressurizer surge lines of Westinghouse PWR design.
Sources of such degradation are mitigated by the design, construction, inspection, and operation of the pressurizer surge oiping.
There is no mechanism for water hammer in the pressurizer / surge system.
The pressurizer safety and relief piping system which is connected to the top of the pressurizer could have loading from water hammer events.
However, these loads are effectively mitigated by the cressurizer and have a negligible effect on the surge line.
t 5173 s /C2591 10 p4
k
. Wall thinning by erosion and erosion-corrosion effects will not occur in the surge line due to the low velocity, typically less than 1.0 ft/sec and the material, austenitic stainless steel, which is highly resistant to these degradation mechanisms.
Per NUREG-0691, a study of pipe cracking in PWR piping, only two incidents of wall thinning in stainless steel pipe were reported and these were not in the surge line.
Although it is not clear from the report, the cause of the wall thinning was related to the high water velocity and is therefore clearly not a mechanism which would affect the surge line.
i It is well known that the presstirizer surge lines are subjected to thermal stratification and the effects oi stratification are particularly significant during certain modes of heatup and cooldown operation.
The effects of stratification have been evaluated for the Prairie Island Unit 1 surge line and the loads, accounting for the stratification effects, have been derived in WCAP-12839.
These loads are used in the leak-before-break evaluation described in this report.
The Prairie Island Unit 1 surge line piping and associated fittings are forged
~
product forms (see Section 3) which are not susceptible to toughness degradation due to thermal aging.
Finally, the maximum operating temperature of the pressurizer surge piping, which is about 650*F, is well below the temperature which would cause any creep damage in stainless steel piping.
2.5 References 2-1 Investigation and Evaluation of Stress-Corrosion Cracking in Piping of 1ight Water Reactor Plants, NUREG-0531, U.S. Nuclear Regulatory Commission, February 1979.
2-2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980.
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SECTION 3.0 MATERIAL CHARACTERIZATION 3.1 Pipe and Weld Materials The pipe material of the prer,surizer surge line for the Prairie Island Unit 1 are A376/TP316 and A403/WP316.
These are a wrought product form of the type used for the primary loop piping of several PWR plants.
The surge line is connected to the primary loop nozzle at one end and the other end of the surge
'ine is connected to the pressurizer nozzle.
The surge line system does not include any cast pipe or cast fitting.
The welding processes used is shielded metal arc (SMAW). Weld locations are identified in Figure 3-1.
In the following section the tensile properties of the materials are presented for use in the leak-before-break analyses.
3.2 Material Procerties The room temperature mechanical properties of the Prairie Island Unit 1 surge line materials were obtained frcm the Certified Materials Test Reports and are given in Table 3-1.
The room temperature ASME Code (Reference 3-1) minimum properties are given in Table 3-2.
It is seen tnat the measured properties well exceed those of the Code.
The representative minimum and average tensile properties were established (see Table 3-3).
The material properties at temperatures (135 F, 205*F, 455'F, and 653*F) are required for the leak rate and stability analyses discussed later.
The minimum and average tensile properties were calculated by using the ratio of the ASME Code Section III properties at the tcnperatures of interest stated above.
Table 3-3 shows the tensile proper ties at various temperatures.
The modulus of elasticity values were established at various temperatures from the ASME Code Section III (Table 3-4).
In the lesk-before-break evaluation, the representative minimum properties at temperature are used for the flaw stability evaluations and the representative sm omai ic 3_1
-~
average properties are used for the leak rate predictions.
The minimum ultimate stresses a e used for stability analyses, These properties are summarized in Table 3-3, 3.3 References 3-1 ASME Boiler and Pressure Vessel Code Section III, Division 1 Appendices July 1, 1989.
I 5t?3s/032591 to 3-2
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TABLE 3-1 i
l Room Temperature Mechanical Properties of the Pressurizer Surge Line Materials t
YlELO ULTIMATE ID HEAT NO./ SERIAL NO.
MATERIAL STRENGTH STRENGTH ELONG.
R/A (psi)
(Psi)
(%)
(%)
1 F0701/4280 A376/TP316 41,700 91,500 54.1 74.6 50,700 90,600 51.6 72.6 2
E1577/4307X A376/TP316 47,800 86,500 25.3 53.3 40,800 85,300 49.6 64.3 3
F0701/4280 A376/TP316 47,100 91,500 54.1 74.6 50,700 90,600 51.6 72.6 a
4 D6034/2 A376/TP316 31,800 75,400 88.0 79.9 38,400 76,500 78.5 79.1 5
480418 A403/WP316 40,000 82,000 46.4 69.9 Shop Weld (SW)
- Fabricated by GTAW and SMAW combination Field Weld (FW) - Fabricated by GTAW and SMAW combination 4
sm,,o2m so 3_3
TABLE 3-2 Room Temperature ASME Code Minimum Properties Material Yield Stress Ultimate Stress
]
(psi)
(psi)
A376/TP316 30,000 75,000 A403/WP316 30,000 75,000 4
P e
S173s/032591 10 3-4
TABLE 3-3 Representative Tensile Properties Minimum Temperature Minimum Average Ultimate Material
('F)
Yield (psi)
Yield (psi)
(psi)
A376/TP316 135 30,240 42,130 75,400 205 27,220 37,920 75,320 455 21,800 30,370 72,180 g
653 19,590 27,290 72,180 A403/WP316 135 38,040 38,040 82,000 205 34,240 34,240 81 910 455 27,430 27,430 78,500 653 24,640 24,640 78,500 U 73 s'C32591 10 3-5
l TABLE 3-4 Modulus of Elasticity (E)
Temperature E (ksi)
(*F) 135 27,950 205 27,600 455 26,115 653 25,035 4
inwars" '
3-6
FW:
FIELD WELD SW:
SHOP WELD I
l FW s
i RCL
-)
Sa i
.)
1 PZR
'2j v
h FW FW I
SW O
Note: Piece #5 is 10" x 14" concentric reducer.
Figure 3-1 Prairie Island Unit 1 Surge Line layout sic 2.,ozim io 37
10N 4.0 LOADS..
. MECHANICS ANALYSIS Figure 3-1 shows schematic layout of the surge line for Prairie Island Unit 1 and identifies the weld locations.
The stresses due to axial loads and bending moments were calculated by the following equation:
e=h+f (4-1)
- where,
{'
stress a
=
F axial load
=
M bending moment
=
A
=
metal cross-sectional area 2
section modulus
=
The bending moments for the desired loading combinations were calculated by the following equation:
M * (H
- HZ)
B Y
(4-2)
- where, M
=
g bending moment for required loading My Y component of bending moment
=
M3 2 component of bending moment
=
The axial load and bending moments far crack stability analysis and leak rate predictions are computed by the methods to be explained in Sections 4.1 and 4.2 which follow.
5173s 'C 3259 7 10 4.}
4.1 Loads for Crack Stabil '/ Analysis The faulted loads for the crack stability analysis were calculated by the absolute sum method as follows:
IF I + IF I + IF l + IF I
(4'3)
F
=
DW TH p
SSE IN I + IN YDW Y TH I + I"Y SSE I (4'4)
N Y
Z ZDWI + I"Z TH I+
I (4-5)
IN M
Z SSE a
DW Deadweight
=
Applicable thermal load (normal or stratified)
TH
=
P
=
Load due to internal pressure SSE SSE loading including seismic anchor motion
=
4.2 Loads for Leak Rate Evaluation The normal operating leads for leak rate predictions were calculated by the algebraic sum method as follows:
F FDW+FTH + F (4-6)
=
Y I"i)DW + (N )TH (4-7)
N Y
(M )DW * (N )TH (4-8)
M
=
7 Z
Z The parameters and subscripts are the same as those explained in Section 4.1.
4.3 Loading Conditions Because thermal stratification can cause large stresses at heatup and cooldewn temperatures in the range of 455'F, a re/iew of stresses was used to identify the worst situations for LBB applications.
The loading states so identified are given in Table 4-1.
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Seven loading cases were identified for LBB evaluation as given in Table 4-2.
Cases A, B, C are cases for leak rate calculations with the remaining cases being the corresponding faulted situations for stability evaluations.
The cases postulated for leak-before-break are summarized in Table 4-3.
The cases cf primary interest are the postulation of a detectable leak at normal power conditions [
ga,c,e The combination [
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s m. cuse' '
4-3 l
l The more realistic cases [
ja.c.e
[
]a,c.e The logic for this AT (
J,c,a a
is based on the following:
Actual practice, based on experience of other plants with this type of situation, indicates that the plant operators complete the cooldown as quickly as possible once a l'ak in the primary system is detected.
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Specifications may require cold shutdown within 36 hours4.166667e-4 days <br />0.01 hours <br />5.952381e-5 weeks <br />1.3698e-5 months <br /> but actual practice is that the plant depressurizes the system as soon as possible once a primary system leak is detected.
Therefore, the hot leg is generally on the warmer
~
side of the limits (~200*F) when the pressurizer bubble is quenched. Once the bubble is quenched, the pressurizer is cooled down fairly quickly reducing the AT in the system.
4.4 Summary of Loads and Geometry The load combinations were evaluated at the various weld locations, Normal loras were determined using the algebraic sum method whereas faulted loads were combined using the absolute sum method.
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4.5 Governing Locations All the welds at Prairie Island Unit 1 surgeline are fabricated using the SMAW procedure.
The following governing locations were established for the welds.
The highest stressed weld location was found at Node 1240 (near the pressurizer junction).
Additional calculations were performed for the next highest stressed weld location at Node 1020 (near the hot leg junction).
Figure 4-1 shows the governing locations.
The loads and stresses at these governing locations for all the loading combinations are shown in Tables 4-4.
e l
l O
I s m..uis "
4-5 l
4 TABLE 4-1 Types of Loadings Pressure (P) bead Weight (DW)
Normal Operating Thermal Expansion (TH)
Safe Shutdown Earthquake and Seismic Anchor Motion (SSE)a a,c.e i
l i
i 1
aSSE is used to refer to the absolute sum of these loadings.
l I
i l
l l
l l
4-6 s m, a oito
i
-TABLE 4-2
{
Normal-and Faulted Leading Cases for Leak-Before-Break' Evaluations
'1 -
i CASE A:
This is the normal operating case at 653*F contisting of the l
algebraic sum of the loading components due to P,-DW and TH.
l
_ a,c.e l
CASE B:
s CASE C:
i l
l CASE D:
This is the faulted operating case at 653*F consisting of the absolute sum (every component load is taken as positive) of P, DW, TH and SSE.
l CASE E:
a,c.e l
CASE F:
c h
i i
l CASE G:
I j
l
- f. -
4 s m.io m eiio 47 l~
,n
TABLE'4-3 Associated Load Cases for. Analyses A/D This is heretofore standard leak-before-break evaluation.
_ a,c.e A/F B/E B/F a
B/Ga I
l 8
C/G a
These are judged to be low probability events.
l i
1 I
i s11WDxn91 to 4.g 4
.m,,
m.
~,.
~ _
\\
-TABLE 4-4 s
Summary of LBB-Loads and Stresses by Case for Governing Locations-1 Node Case F (1bs)
S (psi)
M (in-lb)
S (psi).
S (psi)
X X
g B
T i
i 1020 A
134150 4943 835533 13468' 18411 l
1020 a,c.e -
1020 4
1020 0
151864 5596.
1150890 18551 24146 l
1020 a,c.e l
1020 1020 3
i 1240 A
140976 5194 564140 9093 14287-I 1240 a,c.e 1240 j
1240 0
149846 5521 1169853 18856 24377 l
1240 a,c.e
}
1240 4
l 1240 l
e I
i i
f i
I I
i 5171s/030agt :o gg
= - - - -
o PIPE 10" SCHEDULE 140 o MINIMUM WALL THICKNESS 0.875" NEXT HIGHEST STRESSED
\\
ANALYZED WELD LOCATION
\\
(SMAW) 10'20 s
RCL PZR 1240 HIGHEST STRESSED WELD LOCATION (SMAW)
/
l l
l Figure 4-1 Prairie Island Unit 1 Surge Line Showing Governing Locations mwomei o 4-10
l SECTION 5.0 FRACTURE MECHANICS EVALUATION h
i 5.1 Global failure Mechanism l
4 Determination of the conditions which lead to failure in stainless steel f
should be done with plastic fracture methodology because of the large amount j
of deformation accompanying fracture. One method for predicting the failure of ductile material is the [
Ja,c.e method, based on traditional plastic limit load concepts, but accounting for [
)^' and taking into account the presence of a flaw.
The flawed component is predicted to fail when the remaining net section reaches a stress level at which plastic hinge is formed.
The stress level at which this occurs is tarmed the flow stress.
[
j la,c.e This methodology has been i
shovn to be applicable to ductile piping through a large number of experiments and is used here to predict the critical flaw size in the pressurizer surge lina.
The failure criterion has been obtained by requiring equilibrium of.the j
section containing the flaw (Figure 5-1) when loads are applied.
The detailed development is provided in Appendix A for a through wall circumferential flaw l
in a pipe section with internal pressure, axial force, and imposed bending moments.
The limit moment for such a pipe is given by:
[
]a,c,e (5-1) i j
where:
I
[
Ja,c.e i
1
$173s /032591 10 5-1
[
a,c.e
)
(5-2)
The analytical model doscribed above accurately accounts for the 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 5-1).
Flaw stability evaluations, using this analytical model, are presented in Section 5.3.
5.2 Leak Rate Predictions Fracture mechanics analysis shows in general that postulated through wall cracks in the surge line would remain stable and do not cause a gross failure of this component. However, 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 shutdcwn condition.
The purpose of this section is to discuss the method which will be used to predict the flow through such a postulated c.ack and present the leak rate calculation results for through wall circumferential cracks.
5.2.1 General Considerations The flow of hot pressurized water through an opening to a lower back pressure
,(causing choking) is taken into account.
For long channels where the ratio of the channel length, L, to hydraulic diameter, D, (L/D ) is greater than g
g
[
ja,c,e, both [
] ' 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.
$l e
.-At this point, the flow begins to flash and choking occurs.
Pressure losses due to momentum changes will dominate for [
Ja,c.e Howe.wr, for large L/D values, the friction-pressure drop wil? become important and must g
be considered along with the momentum losses due to flashing.
5.2.2 Calculational Method in using the [
l 3a,c.e,
The flow rate through a crack was calculated in the following manner.
Figure 5-2 from Reference 5-2 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 [
]a,c.e was found from Figure 5-3 taken from Reference 5-2.
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
~
5-4.
Now using the assumed flow rate, G, the frictional pressure _ drop can be calculated using aPf=[
Ja,c.e (5-3) where the friction factor f is determined using the [
. )a,c.e.
The crack relative roughness, t, was obtained from fatigue crack data on stainless steel samples.
The relative roughness value used in these calculations was [
la,c,e RMS.
The frictional pressure drop using Equation 5-3 is then calculated for the l
assumed flow and added to the [
Ja,c,e to obtain the total pressure drop from the system under consideration to the atmosphere. Thus, mwomei n.
5-3
l I
Absolute Pressure - 14.7 = [
_]a,c,e (5-4) for a given assumed flow G.
If the right-hand side of Equation 5-4 does not p
agree with the pressure difference between the piping under consideration and the atmosphere, then the procedure is repeated until Equation 5-4 is satisfied to within an acceptable tolerance and this results in the flow value through the crack, 5.2.3 Leak Rate Calculations Leak rate calculations were performed as a function of postulated through-wall crack length for the critical locations previously. identified.
The crack opening area was estimated using the method of Reference 5-3 and the leak rates were calculated using the calculational methods described above.
The leak rates were calculated using the normal operating loads at the governing node identified in section 4.0.
The crack lengths yielding a leak rate of 10 gpm (10 times the leak detection capability of 1.0 gpm) for critical location at the Prairie Island Unit 1 pressurizer surge line are shown in Table 5-1.
The Prairie Island 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.
5.3 Stability Evaluation A typical segment of the pipe under maximum loads of axial force F and bending moment M is schematically illustrated as shown in Figure 5-5.
In order to calculate the critical flaw size, plots of the limit moment versus crack length are generated as shown in Figures 5-6 to 5-13.
The' critical flaw size corresponds to the intersection of this curve and the maximum load line.
The critical flaw size is calculated using the lower bound base metal tensile properties established in section 3.0.
su3.-can n m 5-4
i i
The welds at the location of interest (i.e. the governing location) are SMAW.
l Therefore, "Z" factor correction for SMAW weld was applied (References 5-5 and 5-6) as follows:
j l
Z = 1.15 [1 + 0.013 (0.D. - 4)) (for SMAW)
(5-5) where OD is the outer diameter in inches.
Substituting 00 = 10.75 inches, the j
Z factor was calculated to be 1.25 for SMAW.
The applied loads were increased j
by the Z factors and the plots of limit load versus crack length were generated as shown in Figure 5-6 to 5-13.
Table 5-2 shows the summary of l
critical flaw sizes for Prairie Island Unit 1.
i j
5.4 References j
5-1 Kanninen, M. F. et al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks" EPRI NP-192, September 1976.
5-2 [
)a,c,e i
l 5-3 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 II-1, NUREG/CR-3464, September 1983.
5-4 NRC letter from M. A. Miller to Cecrgia Power Company, J. P. O'Reilly, l
dated September 9, 1987.
5-5 ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640.
1 l
5-6 Standard Review Plan; Public Comment-Solicited; 3.6.3 Leak-Before-Break
[
Evaluation Procedures; Federal Register /Vol. 52, No. 167/ Friday, August 28, 1987/ Notices, pp. 32626-32633.
5 '73s.'C32591 10 55
. =. - -... -.
l TABLE 5-1 l
Leakage Flaw Size Nede Point Load Case Temoerature Crack Lenoth (in.)
(*F)
(for 10 gpm leakage) 1020 a,c.e 1240 1
e I
i 5173s/030891 10 g,
L
TABLE 5-2'
' Summary of Critical Flaw: Size
.4-l Critical
}
Node Point load Case Temoerature Flaw Size (in)
I i
('F) 4-i 1020 a,c,e i
i 1
1240 i
4 um W
h 4'
t i
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l i
l l
r 1
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1 l
l l
l sin. om.mo 3.,
t l
l
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I I-i I
i a,c,e l
1 I
1 Figure 5-1 Fully Plastic Stress Distribution l
5172s <c30asi 10 5-8
-3,
,7
a,c.e Figure 5-2 Analytical Predictions of Critical Flow Rates of Steam-Water hiixtures sua. cami ic g,g
1
-l a
f e
i d.
C, e
i, i
1 1
4 4
t i
t h
I Figure 5-3 (
ya,c,e p,,,,
p I
l l
1102 a. 02 t 291 10 5-10 I
~
e
I 1
3, c, e a :,e
[
~
-: =
=
Figure 5-4 Idealized Pressure Drop Profile Thrtaugh a Postulated Crack SIC 2s 2129110 o-11
7 3
{
/f
'g U_1 L
d t
Y
^
Oz b
0 r
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I l
1 l
i j
I l
I I
4 9
I I
l l
l I
I I
I I
l i I Q7 Figure 5-5.
Loads Acting on the Model at the Governing Location m.,omi in 5-12
d ce i
PIPE OD=19.75 T=
.889 SIGY=19.6 SIGU=72.2 Fa=152.
M=.115E+94 Figure 5-6.
Critical Flaw Size Prediction for Node 1020 Case D SIC 2s,021211 'O
.]
a, c, e IB PIPE OD=19.75 T=
.889 SIGV=19.6 SIGU=72.2 Fa=152.
M=.121E+04 4
Figure 5-7.
Critical Flaw Size Prediction for Node 1020 Case E sio2. o ini io 5-14
a, c, e l
EB PIPE OD=19.75 T=
.889 SIGY=27.2 SIGU=75.3 Fa=34.6 M:915.
Figure 5-8 Critical Flaw Size Prediction for Node 1020 Case F S102s/021291 10
l a,-c, e 4
PIPE OD=19.75 T=
.889 SICY=39.2 SIGU=75.4 Fa=36.4 M=.155E+94 Figure 5-9 Critical Flaw Size Prediction for Node 1020 Case G l
sio2.<o:imi in 5-16 l
l
i 3,
C, e PIPE OD=19.75 T=.889 SIGY=19.6 SICU=72.2 Ib Fa=159.
M=.117E+04 Figure 5-10 Critical Flaw Size Prediction for Node 1240 Case O S102: C2:291 to g.{J
a.C. e PIPE OD=19.75 T=.889 SICY=19.6 SIGU=72.2 4B l
Fa=150.
M=.101E+94 Figure 5-11 Critical Flaw Size Prediction for Node 1240 Case E S102 s>C21291 10
a, c, e
~
~
e PIPE OD=19.75 T=
.889 SIGY=21.8 SICU=72.2 Fa=32.9 M=842.
Figure 5-12 Critical Flaw Size Prediction for Nede 1240 Case F 1
s/0 129 10 5-19
at c, e 1
)
I i
?
i i
J l
i o
4 3
II PIPE OD=19.75 T=
.889 SICY=21.8 SICU=72.2 e
Fa=37.3 M=.172E+94 f
i i
i Fi are 5-13 Critical Flaw Size Prediction for Node 1240 Case G i
l sin.,o2imi io 5-20 i
I
SECTION 6.0 ASSESSMENT OF FATIGUE CRACK GROWTH 6.1 Introduction To determine the sensitivity of the pressurizer surge line to the presence of small cracks when subjected to the transients discussed in WCAP-12839, fatigue crack growth analyses were performed.
This section summarizes the analyses and results.
Figure 6-1 presents a general flow diagram of the overall process.
The methodology consists of seven basic steps as srown in Figure 6-2.
Steps 1 through 4 are discussed in WCAP-12839. Steps 5 through 7 are specific to fatigue crack growth and are discussed in this section, fhere is prc<ently no fatigue crack growth rate curve in the ASME Code for austenitic stainless steels in a water environment. However, a great deal of work has been done (References 6-1 and 6-2) which supports the development of such a curve.
An extensive study was performed by the Materials Property Council Working Group on Refemnce Fatigue Crack Growth concerning the crack growth behavior of tht ee steels in air environments, published in Reference 6-1.
A reference curve 'or stainless steels in air environments, based on this work, is in the 1985 Edition of S,ction XI of the ASME Code.
This curve is shown in Figure 6-3.
A compilation of data for austenitic,tainless steels in a PWR water environment was made by Bamford (Re.erence 6-2), and it was found that the effect of the environment on the crack growth rate was very small.
For this reason it was estimated that th s environmental f actor should be set at 1.0 in the crack growth rate equation from Reference 6-1.
Based on these works (References 6-1 and 6-2) the fatigue crack growth law used in the analyses is as shown in Figure 6-4,
'"3.,c32m '
6-1
6.2 Initial Flaw Size Various initial surface flaws were assumed to exist.
The flaws were assumed to be semi-elliptical with a six-to-one aspect ratio.
The largest initial flaw assumed to exist was one with a depth equal to 10% of the nominal wall thickness, the maximum flaw size that could be found acceptable by Section XI of the ASME Code.
3.3 Results of FCG Analysis Fatigue crack growth analyses were performed at locations 1 and 2 where detailed fracture; mechanics analyses as described in Section 5 were completed.
it should be noted that location 1 is near the reactor coolant icop no::le and location 2 is near the pressurizer no::le, l
/
Results of the fatigue crack growth analysis are presented in Table 6-1 for an initial flaw of 10% icinimum wall thickness.
Conservatisms existing in the fatigue crack growth analysis are listed below.
1.
Plant operational transient data has shown that the conventional design transients contain significant conservatisms
[
ja.c,e 4.
FCG neglects fatigue life prior to initiation
$ 173s 'C32591 10 gp
r-4 6.4 References
(
6-1. James, L. A. and Jones, D.
P., "Fatigt.e Crack Growth Correlations for Austenitic Stainless Steel in Air," in Predictive Capabilities in Environmentally Assisted Cracking, ASME publication PVP-99, December 1985.
6-2. Bamford, W. H., " Fatigue Crack Growth of Stainless Steel Reactor Coolant Piping in a Pressurized Water Reactor Environment," ASME Trans. Journal of Pressure Vessel Technology, Feb. 1979.
k l
9 l
4 im. unei in 6-3 i
TABLE 6-1 FATIGUE CRACK GROWTH RESULTS FOR 10% OF WALL INITIAL FLAW SIZE Initial Initial Final (40 yr)
Final Flaw I:
Location Position Size (in)
(% Wall)
Size (in)
(% Wall) a,c.e las e
$173s C30891 10 y-4
DETERMINATION OF THE EFFECTS OF THERMAL $TR ATfFICA7lON 3 ce 5
~
Figure 6-1 Determination of the Effects of Thermal Stratification on Fatigue Crack Growth s,n. vw,,o g_5
a.c.e
~
k-
=
l Figure 6-2 Fatigue Crack Growth Methodology l
l 1173s'03089' 10 h*6
4 sea.e
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,e fff!
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ll /// l l
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/ h /
.l
- t..
ses sanm p figure 6-3 fatigue Crack Growth Rate Curve for Austenitic Stainless Steel s m. m n ie 6-7
h=CFSEAK.30 3
where dd
= Crack Growth Rate in inches / cycle
-20 C
= 2.42 x 10 F
= Frequency f acter (F = 1.0 for temperature below 800=F)
S
= R ratio correction (S = 1.0 for R = 0; 5 = 1 + 1.8R for 0 < R <.8; and 5 = -43.35 + 57.97R for R > 0.8)
E
= Environmental Factor (E = 1.0 for PWR) 4K
= Range of stress intensity factor, in psi v'in R
= The ratio of the minimum Kg (K! min) to the maximum K7 (K g,, )
Figure 6-4.
Fatigue Crack Growth Equation for Austenitic Stcinless Steel e n. n:es,,o 6-8
westmessust enornieTAQv class a e
LOCATION 1
\\
s RCL PZR K] LOCATION 2 D
N MAXIMUM FCG IN THIS AREA REPORTED Figure 6-5.
Fatigue Crack Growth Critical Locations siewenm is 6-9
a c.e Figure 6-6.
Fatigue Crack Growth Controlling Positions at Each Location eo ex i ie 6-10
SECTION 7.0 ASSESSMENT OF MARGINS
?
in the preceding sections, the leak rate calculations, fracture mechanics analysis and fatigue crack growth assessment were performed.
Margins at the critical location cre summarized below:
In Secton 5.3 using the IWS-3640 approach (i.e.
"Z" f actor approach), the
" critical" flaw sizes at the governing locations are calculated.
In Section 5.2 the crack lengths yielding a leak rate of 10 gpm (10 times the leak detection capability of 1.0 gpm) for the critical locations are calculated.
The leskage size flaws, the instability flaws, and margins are given in Table 7-1.
The margins are the ratio of instability flaw to leakage flaw.
The margins for analysis combination cases A/0, [
Ja.c e well exceed the factor of 2.
The margin for the extremely low probability event defined by [
)"' C d is [
Ja.c.e As stated in Section 4.3, the probability of simultaneous occurrence of SSE and maximum stratification due to shutdown because of leakage is estimated to be very low.
Actual material properties given in the certified materials test reports for the pipes in question are used at the nodes where the factor of two on the leakage flaw size is not satisfied.
The margins of the critical flew size to the leakage flaw size obtained from the actual properties are given in Table 7-1 along with the margins obtained from average and minimum properties for leakage flaws and critical flaw sizes, respectively.
By using the actual properties at the critical location, the margin for B/G was calculated to be 2.1 (see Table 7-1).
ln this evaluation, the leak-before-break methodology is applied conservatively.
The conservatisms used in the avaluation are summar.
Table 7-2.
1 1.
I sm,ww
- 7-1
TABLE 7-1 Leakage flaw Sizes, Critical flaw Si:es and Margins Load Critical flaw Leakage flaw Node Case Size (in)
Size (in)
Marain 1020 A/D 9.12 3.10 2.9 a,c,o 1240 A/D 9.05 3.75 2.4
~
a,c,o I
a These are judged to be low probability events b
By using actual properties
$173s 03089110 7-2
TABLE 7-2 LBB Conservatisms o
factor of 10 on Leak Rate o
Factor of 2 on Leakage Flaw for all cases o
Algebraic Sum of Loads for Leakage o
Absolute Sum of Loads for Stability Average Material Properties for Leakage o
o Minimum Material Properties for Stability sm. urni ie 73
G O
O e
SECTION 8.0 CONCLUSIONS This report justifies the elimination of pressurizer surge line pipe breaks as the structural design basis fo: Prairie Island Unit 1 as follows:
Stress corrosion cracking is precluded by use of fracture resistant a.
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 (primary loop and the attached class 1 auxiliary lines) because of system design, testing, and operational considerations.
The effects of low and high cycle fatigue on the integrity of the c.
surge line were evaluated and shown acceptable.
The effectt of thermal stratification were evaluated and shown acceptable.
d.
Ample margin exists between the leak rate of,," 11 .oble flaws and the criterion of Reg. Guide 1.45.
Amole margin exists between the small stable flaw si os vi item d e.
and the critical flaw sizes.
The postulated reference flaw will be stable because of the ample margins in d, e and will leak at a detectable rate which will assure a safe plant shutdown.
i Based on the above, it is concluded that pressuri:er surge line breaks should not be considered in the structural design basis of Prairie Island Unit 1.
8-1
9 4
)
APPENDIX A LIMIT MOMENT 1
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.r,.,-.-m,,--- - - -.m-,-,-,-,,-
e,,, -,,,,,, - -, -. - - -
,-,-r-,
.--,,,-w,-,,,,..#---.,
APPENDlX A LIMIT 90 MENT
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