ML20235F354
| ML20235F354 | |
| Person / Time | |
|---|---|
| Site: | South Texas  |
| Issue date: | 09/30/1987 |
| From: | Palusamy S, Schmertz J, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19304B507 | List: |
| References | |
| WCAP-11572, NUDOCS 8709290061 | |
| Download: ML20235F354 (124) | |
Text
_ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
WESTINGHOUSE CLASS 3 WCAP- 11572
, t i
TECHNICAL BASES FOR ELIMINATING RUPTURE OF THE ACCUMULATOR LINE AND ITS ATTACHED RHR PIPING FROM THE STRUCTURAL DESIGN BASIS FOR SOUTH TEXAS PROJECT UNITS 1 AND 2 September 1987 i
i l
J. C. Schmertz S. A. Swamy Y. S. Lee
- F. J. Witt D. H. Roarty Verified by: K. C. Chang Approved by:. / #
- 5. 5. Palusamy, Manager Structural Materials Engineering Work Ferformed Under Shop Order HPLJ-2052 WESTINGHOUSE ELECTRIC CORPORATION Generation Technology Systems Division P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 em.em.asmie g92{ggggg{go!498 p PDR
i PREFACE l .
The technical justification for eliminating postulated accumulator line and attached RHR line ruptures from the structural design basis was provided in the following reports:
- 1. WCAP-11383, " Technical. Bases for Eliminating Class 1 Accumulator Line Rupture as the Structural Design Basis for South Texas Project Units 1 and 2," January 1987.
- 2. WCAP-11351, " Technical Bases for Eliminating Accumulator Class l' Low Pressure Line Rupture as the Structural Design Basis for South Texas
- Project Units 1 and 2," March 1987.
Based on a review of the above reports, the NRC staff requested additional information. Responses to these requests were provided in the following reports:
- 3. " Additional Information in Support of the Elimination of Postulated Pipe Ruptures in the Accumulator Lines of South Texas Projects Units 1 and 2," May 1987, submitted as Attachment 1 to letter of May 18, 1987 from M. R. Wisenburg (Houston Lighting & Power Company) to NRC.
l
- 4. " Additional Information in Support of the Elimination of Postulated i Pipe Ruptures in the Accumulator Lines of South Texas Project Units 1 1 and 2 (10-inch and 8-inch RHR Piping Attached to the Accumulator Line)," May 1987, submitted as Attachment 1 to letter of May 29, 1987 from M. R. Wisenburg (Houston Lighting & Power Company) to NRC.
In July 1987, it was identified that the loads used in the leak-before-break
, (LBB) analysis changed due to a Bechtel piping code (ME101) analysis ;
discrepancy. Consequently, an assessment was performed to determine the
,, impact of this change on LBB. The assessment, which demonstrated that the LBB !
conclusions presented in WCAPs 11383 and 11351 remain applicable using the 2545s/0396s/082687.10 jj
1 l
1 revised loads, was transmitted to the NRC staff (submitted as attachment 2 to letter of July 16, 1987 from M. R. Wisenberg of Houston Lighting and Power Company to NRC).
The purpose of this report is to provide leak-before-break evaluation using the revised loads and summarizing the information presented ' previous submittals. Some of the information contained in documents 3 . i 4 is included in the text while the remaining portion is included in the appendices.
I 1
1 l
I 1
)
1 25esm2eesmarss?to jjj
- u. . . . . .
TABLE OF CONTENTS Section Title Page
1.0 INTRODUCTION
1-1 1.1 Background 1-1 1.2 Scope and Objective 1-1 j 1.3 References 1-4 l
2.0 FAILURE CRITERIA FOR FLAWED PIPES 2-1 2.1 General Considerations 2-1 2.2 Global Failure Mechanism 2-1 2.3 Local Failure Mechanism 2-2 2.4 References 2-3 3.0 0FERATION AND STABILITY OF THE ACCUMULATOR LINES 3-1 AND ATTACHED RHR LINES 3.1 Stress Corrosion Cracking 3-1 3.2 Water Hammer 3-3 3.3 Low Cycle and High Cycle Fatigue 3-4 3.4 Potential Degradation During Service 3-4 3.5 References 3-5 4.0 MATERIAL CHARACTERIZATION 4-1 4.1 Pipe, Fittings and Weld Materials 4-1 4.2 Tensile Properties 4-1 4.3 Fracture Toughness Properties 4-3 4.4 References 4-4 5.0 LOADS FOR FRACTURE MECHANICS ANALYSIS 5-1 5.1 Loads for Crack Stability Analysis 5-2 5.2 Loads for Leak Rate Evaluation 5-2
. 5.3 Summary of Loads Geometry and Materials 5-2 2545s 0 96s/082687 10 jy
, 1 n~ )
l TABLE 0F CONTENTS (cont.) )
Section Title Page 6.0 FRACTURE MECHANICS EVALUATION 6 6.1 Global Failure Mechanism 6-1 l
6.2' Leak Rate Predictions 6-3 1 6.2.1 General Considerations 6-3 6.2.2 Calculation Method 6-4 6.2.3 Leak Rate Calculations 6-6 6.2.4 Leak Detection Capability 6-7 6.3 Local Stability Analysis 6-7 6.4 References 6-7 7.0 ASSESSMENT OF FATIGUE CRACK GROWTH 7-1 7.1 Acceptability of Fatigue Crack Growth 7-2 7.2 References 7-4 8.0 ASSESSMENT OF MARGINS 8-1
9.0 CONCLUSION
S 9-1 APPENDIX A Limit Moment A-1 APPENDIX B Fatigue Crack Growth Considerations B-1 B.1. Thermal Transient Stress Analysis B-1 B.2 Fatigue Cra.ek Growth Analysis B-6 B.3 References B-9 I
i l.. APPENDIX C Materials Specification and Fracture Toughness C-1 I~
Properties of the Accumulator Tank, Nozzle, and
. Safe End 2545s/0396s/082687 10 y
TABLE 0F CONTENTS (cont.)-
Section- Title -Page C.1 Materials Specification C--I C.2 Fracture Toughness ~ C-1 APPENDIX D Effects of Thermal Aging on Structural Integrity D-1
.of Welds at the Limiting Location APPENDIX E A Summary Evaluation of_ Potential Degradation During Service E-1
, APPENDIX F Benchmark for Leak Rate Prediction Moded Used for F-1 8 inch and 10 inch lines APPENDIX G Additional Information on the Leak Detection System' G-1
~
APPENDIX H Clarification of Definition of the Accumulator H-1 Line and its Attached RHR Piping as Considered in the LBB Analysis 2545s/0396s 482647 10 yj
r LIST OF FIGURES Figure Title Page 1-1 South Texas Project Accumulator Lines with Attached 1-5 RHR Piping 2-1 Schematic of Generalized Load Deformation Behavior 2-4 4-1 True Stress Strain Curve for SA376 TP316 Stainless 4-18 Steel at 560*F 4-2 Minimum True Stress Strain Curve for SA376 TP316 4-19 Stainless Steel at 120*F 4-3 Minimum True Stress Strain Curve for SA312 TP304L 4-20 Stainless Steel at 120'F 5-1 Schematic Layout of Accumulator and RHR Lines Loop 1 5-10 5-2 Schematic Layout of Accumulator and RHR Lines Loop 2 5-11 5-3 Schematic Layout of Accumulator and RHR Lines Loop 3 5-12 6-1 [ Ja c,e Stress Distribution 6-10 6-2 Loads Acting on the Pipe Model 6-11 6-3 Critical Flaw Size Prediction for 12 Inch Line 6-12
, 6-4 Critical Flaw Size Prediction for 10 Inch Line 6-13
. 6-5 Critical Flaw Size Prediction for 8 Inch Line 6-14 l
2585s/0396s/082687 10 yjj
. LIST OF FIGURES (cont.)
Figure Title. Page 6-6 Analytical Predictions of Critical Flow Rates of 6-15 Steam-Water Mixtures 6-7 -[ .)a,c,e Pressure Ratio as a- '6-16 Function of L/D 6-8 Idealized Pressure Drop Profile through a Postulated 6-17 Crack A-1 Pipe with a Through Wall Crack in Bending A-2 B Comparison of Typical Maximum and. Minimum Stress B-15 Profile Computed by Simplified [
3a,c.e B-2 Schematic of Accumulator Line at [ B-16 ja,c.e B-3 [ .)a,c,e and Minimum Stress Profile B-17 for Transient #10-B-4 [ Ja,c.e Maximum and Minimum Stress B-18 Profile for Transient #11 B-5 [ Ja,c.e Maximum and Minimum B-19 Stress Profile for Transient #12 B-6 [ Ja,c.e Maximum and Minimum B-20 Stress Profile for Transient #14 2545s 9396s/082667.10 yjjj L .'
i LIST OF FIGURES (cont.)
. 1 i
i Figure Title Pace C-1 Cross-Section of the Valve-to-Pipe Weld for C-4 8 Inch Diameter RHR Line
)
H-1 ~ South Texas Project 10 Inch and 8 Inch RHR Lines H-3 Attached to the Accumulator Line i
i l
1 l
1 l
1 2545s T396s $626811C jy
LIST OF TABLES Table'No. Title Page 4-1 Room Temperature Mechanical Properties of the 4-6 12 Inch Accumulator Line Materials and Welds.
. of the South Texas Project. Unit 1 Plant Operating at 560*F 4-2 Room Temperature Mechanical Properties of the 4-7 12: Inch Accumulator Line Materials and Welds of the South Texas Project Unit 1 Plant Operating at 120'F 4-3 Room Temperature Mechanical Properties of the 4-8 10 Inch and 8 Inch RHR Line Materials and Welds of the South Texas Project Unit 1 Plant-Operating at 120*F 4-4 Room Temperature Mechanical Properties of the 4-10 12 Inch Accumulator Line Materials and Welds of the South Texas Project Unit 2 Plant Operating at 560'F
, 4-5 Room Temperature Mechanical Prcperties of the 4-11 12 Inch Accumulator Line Materials and Welds of the South Texas Project Unit 2 Plant l-Operating a.t 120*F f
4-6 Room Temperature Mechanical Properties of the 4-12 10 7.nch and 8 Inch RHR Line Materials and Welds of the South Texas Project Unit 2 Plant
. Operating at 120*F mumee.mmm x
LIST OF TABLES (cont.)
Table No. Title Page 4-7 Room Temperature Tensile Properties of the SA351 4-13 CF8A Cast 45 Degree Nozzle 4-8 Typical Tensile Properties of SA376 TP316 SA351 CFA 4-14 and Welds of Such Material for the Primary Loop 4-9 Comparison of Tensile Properties of Accumulator Lines 4-15 and Their Attached RHR Lines With Those of Typical Wrought Primary Loops and ASME Code Minimum Requirements 4-10 Fracture Toughness Properties Typical of the 4-16 Accumulator Line and Attached RHR Piping 4-11 Chemistry and Ends of Service Life RCU Toughness 4-17 for the Six 45-Degree Nozzles 5-1 Summary of Envelope Loads for 12 Inch Pipe 5-4 5-2 Summary of Envelope Loads for 10 Inch Pipe 5-5 '
5-3 Summary of Envelope Loads for 8 Inch Pipe 5-6 5-4 Loading Components at Governing Locations For 5-7 12 Inch Line 5-5 Loading Compenents at Governing Locations For 5-8 10 Inch Line
. 5-6 Loading Components at Governing Locations For 5-9 8 Inch Line 2546s '0396: '0626871C Xi
V.
L LIST' 0F. TABLES (cont.)
Table No'. Title Page 6-1 Japplied Resuits 6-9 8-l' Comparison of Results vs. Criteria 8-3 B-1 Thermal Transients Considered for_ Fatigue Crack B-11 Growth Evaluation '
B-2 Stresses for the Minor Transients-(PSI) B-12 B-3 Envelope Normal Loads B-13 B-4 Accumulator Line Fatigue Crack Growth Results B-14
)
- )
ms.mn.arauo x$$
, SECTION 1.0
, INTRODUCTION
1.1 Background
The current structural design basis-for the accumulator line and its attached RHR piping 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 these lines. Presented in this report are the descriptions of a mechanistic pipe break evaluation method
.and the analytical results that are used for establishing that a circumferential-type break will not occur. The evaluations considering circumferential1y oriented flaws cover longitudinal cases, and thereby
. eliminate the need for sone of the plant hardware. The scope of the piping covered by the work is shown in Figure 1-1 (Also see Appendix H).
1.2 Scope and Objective The general purpose of this investigation is to demonstrate leak-before-break for the accumulator line and its attached RHR piping. Schematic drawings of the piping system are shown in Section 5.0. The recommendation and criteria l proposed in NUREG 1061 Volume 3 (1-1) are used in this evaluation. These criteria and 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.
1 mwem,amne 11
- 3) Postulate a surface flaw at' the governing location. Determine -
fatigue crack growth. Show that a through-wall crack will not result. -
- 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-detection 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 normal plus SSE 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 anc' high cycle fatigue.
- 7) For.the base and weld metals actually in the plant provide the material properties including toughness and tensile test data. -
Justify'that the properties used in the evaluation are representative of the plant specific material. Evaluate long term effects such as thermal aging where applicable.
- 8) Demonstrate margin on applied load.
The flaw stability criteria proposed for the analysis examines both the global and local stability for a postulated through-wall circumferential flaw. The global ant. lysis is carried, out using the [ 3a,c.e method, based on traditional plastic limit load concepts, but accounting for [
Ja,c.e and taking into account the presence of a flaw. The local stability analysis is carried out using the method described in NUREG/CR 3464 -
(1-2). This method is based on linear elastic fracture mechanics and it can be used up to load levels p-oducing small plastic zone size. For higher loads, -
sw mm,memne 12
i
~
the local stability analysis is carried out by performing a static elastic plastic [- Ja,c.e of a straight piece of the accumulator line pipe containing a through-wall circumferential flaw. The EPRI elastic plastic fracture handbook method was used for this purpose.
The leak rate is calculated for the normal operating condition. The leak rate prediction model used in evaluating the 12 inch line extending from the RCS cold leg injection points to the first check valve is an [
3a,c.e In evaluating the remaining lines, which operata at 120'F, a single phase flow in a tight crack including friction losses associated with crack opening geometries is used. The crack opening area required for calculating the leak ratas is obtained by subjecting the postulated through-wall flaw to normal operating loads. Surface roughness is accounted for in determining the leak rate through the postulated flaw.
As stated earlier, the evaluations described above considering circumferen-tially oriented flaws cover longitudinal cases in pipes and elbows. The likelihood of a split in the elbows is very low because of the fact that the elbows are [ ]C and no flaws are actually anticipated. The prediction methods for failure in elbows are virtually the same as those for
[
Ja,c,e However, the elbows are [ 3a,c.e and, therefore, the probability of any longitudinal flaw existing in the accumulator line and the attached RHR line is much smaller when compared with the circumferential direction. Based on the above, it is judged that circumferential flaws are more limiting than l , longitudinal flaws in elbows and throughout the system.
The computer codes used in this evaluation have been validated (bench marked) as described in References (1-4) and (1-5).
2fr45s 'C 396s /D9348710
}.3
~
1.3 References
~
1-1 ~ Report:of the U.S. Nuclear Regulatory Commission. Piping Review Committee -
Evaluation.of Potential for Pipe Breaks,-NUREG 1061, Volume 3, November 1984.
1-2 -NUREG/CR-3464, 1983, "Th'e Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulated Circumferential Through Wall Cracks."
1-3 Begley, J. A., et. al., " Crack Propagation Investigation Related to the Leak-Before-Break: Concept:for LMFBR-Piping" in Proceedings, Conference on Elastic Plastic Fracture, Institution of Mechanical Engineers, London 1978.
1-4 Swamy, S. A., et. al., " Additional Information in Support of the Elimination of Postulated Pipe Ruptures in the Pressurizer Surge Lines of
. South Texas Project Units 1 and 2" WCAP-11256, September 1986,
~
Westinghouse Proprietary Class 2.
1-5 Swamy, S. A. et. al., " Technical Basis- for Eliminating Pressurizer Surge Line Ruptures as the' Structural Design Basis for South Texas Project Units 1 and 2," WCAP-11256 Supplement 1, November 1986, Westinghouse Proprietary Class 2.
5 mwem,amne 14
.t.
>=
2 N
, E IE (w
so :"
bi R og w-
"E E si Ek 5W EE EE
?,
g "r
f5o
. 5. ~~
7 s
e 7 g ,
g t a r a s 'r y '
'm _
E "
E
$ "e [ Og n si se e- !:$
s 2 ..
m- .. ,g s
m an as g,
5 e m
- Ok g gre EE c;
i au Q**
tr FIGURE 1-1 South Texas Project Accumulator Line with Its Attached RHR Piping 1-5
SECTION 2.0 FAILURE CRITERIA FOR FLAWED PIPES
.2.1 General Considerations' j H .
l' Active research;is being carried out in industry'and universities as well as
~
other research organizations to establish fracture criteria for ductile !
materials. Criteria being investigated include those based on J-integral initiation toughness, equivalent energy,. crack opening displacement, crack opening' stretch, crack opening angle, net-section yield, tearing modulus and void nucleation. Several of these criteria are discussed in an ASTM publication (2-1).
A practical approach based on the ability to obtain material properties and to make calculations using the available tools was used in selecting the criteria ,
for this investigation. The ultimate objective is to show that the accumula-tor line or the attached RHR line containing a conservatively assumed circum-ferential through-wall flaw is stable under the worst combination of postu-lated faulted and operating condition loads within acceptable engineering accuracy. With this viewpoint, two mechanisms of failure, namely, local and global failure mechanisms are considered.
2.2 Global Failure Mechanism For a tough ductile material which is notch insensitive the global failure will be governed by plastic collapse. Extensive literature is available on this subject. A PVRC study (2-2), reviews the literature as well as data from several tests on piping components, and discusses the details of analytical methods, assumptions and methods of correlating experiments and analysis.
e no,5m.amne 2-1
A schematic description of the plastic behavior and the definition of plastic .
load is shown in Figure 2-1. For a given geometry and loading, the plastic load is defined to be the peak load reached in a generalized load versus -
displacement plot and corresponds to the point of instability.
A simplified version of this criterion, namely, net section yield criterion has been successfully used in the prediction of the load carrying capacity of pipes containing gross size through-wall flaws (2-3) and was found to correlate well with experiment. This criterion can be summarized by the following relationship: l Wa < Wp (2-1) where Wa = applied generalized load Wp = calculated generalized plastic load ;
Wp represents the load carrying capacity of the cracked structure and it can be obtained by an elastic plastic finite element analysis or by empirical correlation which is based on the material flow properties as discussed in -
Section 6.1 ,
2.3 Local Failure Mechanism The local mechanism of failure is primarily dominated by the crack tip behavior in terms of crack-tip blunting, initiation, extension and finally crack instability. The material properties and geometry of the pipe, flaw size, shape and loadings are parameters used in the evaluation of local failure.
The s tability will be assumed if the crack does not initiate at all. It has been demonstrated that the initiation toughness, measured in terms of Jjc 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 J Ic f the material, then the crack will not initiate.
If the initiation criterion is not met, one can calculate the tearing modulus es defined by the following relation:
2Sd Ss /C396s /08268710 p_p
t E
T,pp = dJdi y (2-2) where T,pp
= applied tearing modulus E =. modulus of elasticity
= flow stress ='[ ]a,c.e of_
a = crack length e,
y u
= yield and ultimate strength of the material respectively.
'In.. summary, the local crack- stability is established by the two-step criteria:
J<Jlc, or (2-3)
T,pp < Tmat, if J 3 J Ie (2-4) 2.4 References
^
2-1 J.D. Landes, et al., Editors, Elastic-Plastic Fracture, STP-668, ASTM, Philadelphia', PA 19109, November 1977.
2-2 J. C. Gerdeen, "A. Critical Evaluation of Plastic Behavior Data and a Unified Definition of Plastic Loads for Pressure Components," Welding Research Council Bulletin No. 254.
2-3 Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circut.ferential Cracks, EPRI-NP-192, September 1976.
3 Y
254Ss 4396s @9N8710 2-3 l j
e P
W,* PL ASTIC LQ&D 1
I l
l a
l S
N
. d s I 5 l -
l l
l l
AP GENERALIZED DISPLACEMENT l
l l
Figure 2-1 Schematic of Generalized Load-Deformation Sahavior 2SdSs t'318' 10 p.4
SECTION 3.0 OPERATION AND STABILITY OF THE ACCUMULATOR LINES AND ATTACHED RHR LINES 3.1 Strass Corrosion Cracking The Westinghouse reactor coolant system primary loop and connecting Class I lines have an operating history that demonstrates the inherent 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 400 reactor years, including five plants each having over 15 years of operation and 15 other plants each with over 10 years of operation.
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 3-1) entitled " Investigation and Evaluation of Stress Corrosion Cracking in Piping of Light Water Reactor Plants." In that report the PCSG stated:
"The PCSG has determined that the potential for stress-corrosion cracking in PWR primary system piping is extremely low because the 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 ms uncares' '
3-1
l
- of producing stress-corrosic,n cracking in the primary systems of PWRs.
}
) Operating experience in PWRs supports this determination. To date, no stress-corrosion cracking has been reported in the primary piping or safe ends of any PWR."
During 1979, several instances of cracking in PWR feedwater piping led to the j establishment of the third PCSG. The investigations of the PCSG reported in NUREG-0691 (Reference 3-2) further confirmed that no occurrences of IGSCC have been reported for PWR primary coolant systems. l l
1 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 4
discussion below further qualifies the PCSG's findings. ;
For stress corrosion cracking (SCC) to occur in piping, the following three j conditions must exist simultaneously: high tensile stresses, susceptible material, and a corrosive environment. Since some residual stresses and some j 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 I environment. The material 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, welding, fabrication, and processing. l 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, 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. Requirements on chlorides, fluorides, conductivity, and pH are included in the acceptance criteria for the piping.
nn.,cm canano 3-2
l' l
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 and connecting Class 1 lines is expected to be in the ppb range by controlling charging flow chem-istry and maintaining hydrogen in the reactor coolant at specified concentra-tions. . Halogen concentrations are also stringently controlled by maintaining l
concentrations of chlorides and fluorides within the specified limits. .Thus-during plant operation, the likelihood of stress corrosion cracking is minimized.
I
- - '3.2 Water Hammer l
Overall, there is a low potential for water hammer in the RCS and connecting accumulator lines since.they are designed and operated to preclude the voiding condition in normally filled lines. The RCS and connecting accumulator lines
~
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 sprey 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 l
~
reactor coolant systems to verify the flow and vibration characteristics of the system and connecting accumulator lines. Preoperational testing and
,m. m.., _ , x I 33 o
operating experience have verified the Westinghouse approach. The operating transients of the RCS primary piping and connected accumulator lines are such
-that no significant water hammer can occur.
3.3 Low Cycle and'High Cycle Fatigue Low cycle fatigue considerations are accounted for in the design of the piping system through the fatigue usage factor evaluation to show compliance with the
. rules of Section III of the ASME Code. A further evaluation of the low cycle fatigue loading is discussed in Chapter 7 as part of this study in the form of a fatigue crack growth analysis.
High cycle fatigue loads in the system would result primarily from pump l
vibrations during operation. 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 cf 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. When translated to the connecting accumulator lines, these stresses are even lower, well below the fatigue endurance limit for the accumulator line material and would result in an applied stress intensity factor below the threshold for fatigue crack growth.
3.4 Potential Degradation Durino Service There has never been any service cracking found in the accumulator lines or their attached RHR piping lines for Westinghouse PWR designs.
Wall thinning by erosion and erosion-corrosion effects will nut occur in the accumulator line and its attached RHR line due to the low velocity, typically loss than 10 ft/sec and the material, austenitic stainless steel, which is .
highly resistant to these degradation mechanisms.
Vibratory fatigue loads are monitored for the 12-inch accumulator line during the hot-functional testing of the plant and are well below the high cycle fatigue allowables. The vibratory fatigue loads for the 8-inen and 10-inch 2515s ' Chi 6s '08268710 34
l-
~
. pipe attached to .the accumulator' line should be less severe as compared to the 12-inch pipe which.is attached to the primary loop. Also, socket welds, which are susceptible to this type of loading, are not'used in this piping.
There is no mechanism for water' hammer in the accumulator piping system.
Water hammer effects, flow stratification effects, and'other sources of potential. degradation during service are discussed in further detail in Appendix E.
u l
3.5' References i
3-1 Investigation and Evaluation of Stress-Corrosion Cracking in Pioing.of Light Water Reactor Plants, NUREG-0531, U.S. Nuclear Regulatory Commission, February 1979.
l l
3-2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980.
i l 1
l
- l I
i I
rw.icawarut to 35
SECTION 4.0 MATERIAL CHARACTERIZATION 1 1
l 4.1' Pipe, fittinos, and Weld Materials o
The aipe' material of the 12-inch line extending from the RCS cold leg !
f inje: tion point to the first check valve is SA 376-TP316, a wrought product .
forrr of tLe type- used for the primary loop piping of several PWR plants. The fittings are SA403-WP316 which is. wrought and formed pipe of SA182 F316. The ,
pipe materials for the remaining lines, which operate at 120*F, are
]
SA3' 6-TP316, SA312-TP316L, and SA312-TP304L. For these lines, the fittings j are SA403-KP316, SA182-F316L, and SA403-TP316L. The weld wire used in the l shop fabrication is generally low carbon' 316L; in some instances the weld wire has low carbon with high silicon (316LSi). The. welding processes used were gas tungsten arc (GTAW), submerged arc (SAW), gas metal arc (GMAW) and shielded metal arc (SMAW). The field welds used 308L weld wire. For each accumulator. line there is a 45 degree nozzle intersecting the cold leg of the primary. loop. The material of these nozzles is SA351 CF8A, a cast product form.
j In the following section the tensile and fracture toughness properties of the above materials are presented and criteria for use in the leak-before-break analyses are defined.
Material properties for the accumulator tank, nozzle, and safe end are given in Appendix C.
4.2 Tensile Properties Material certifications 9:ere used to establish the tensile properties for the piping, fittings, and welds. The properties are g'.ven in Tables 4-1, 4-2. and 4-3 for Unit 1, and Tables 4-4, 4-5 and 4-6 for Unit 2.
wwn*,amne 43
i l
The properties in Table 4-1 through 4-6 are those at room temperature. In the ,
leak-before-break evaluations presented later, the ASME Section III code minimum properties at operating temperatures are used. The viability of using .
such properties for the lines operating at 120*F, is presented below. j
[
3a,c.e
[
)"*C All the properties presented are seen to exceed the room temperature code minimum properties. Larger margins are noted when comparing the experimental yield stress data with the code minimum properties.
The material certification records are summarized for the 10-inch diameter and 8-inch diameter portion of the RHR system in, tables 4-3 and 4-6. From these records, the lower bound room temperature yield strengths for SA312 TP304L, SA403 h'P316L, and SA182 F516 are noted to be 38,100 psi, 32,420 psi, and 38,127 psi respectively. Clearly, the lowest base metal yield strength at room temperature is 32,420 psi. The corresponding yield strength at 120*F is estimated to be 31,500 psi. The ASME Code minimum yield strength of 29,200 psi for SA376 TP316 is seen to be lower than the lower bound value of 31,500 psi obtained by considering all the base metals in the 10-inch and 8-inch RHR riping.
I Based on this discus-ion it is concluded that the use of ASME Section III Code I
minimum properties is justified for crack stability analytis. [
ja,c,e 2645s C96s'091087 IC 42
For leak rato calculations, the average values of yield strength were used.
These values are [.
jo.c.e Moduli and stress-strain relationships of the materials at the appropriate operating temperature, i.e., 560*F and 120*F are provided in figures 4-1 through 4-3.
4.3 Fracture Toughness Properties 1
[
Ja c.e Lower bound estimates for the fracture toughness of welds, taking thermal aging into account, are discussed in Reference 4-4. [
- a,c.e Forged stainless steel is considered not susceptible to thermal aging for the applications at hand; however, thermal aging embrittlement must be considered for the cast 45' nozzle.
[
3a,c.e By the criteria established in Reference 4-5, the fracture toughness is at least as great as the toughness'of [ Ja,c.e ,
[ .)a,c.e is the same heat which serves as a lower bound for welds as
~
seen in Table 4-10. [
)"'C The fracture criteria are thus no, un, aw io 43
- - a,c.e -
For the line's operating at 120*F, the toughness will be higher than'the toughness'at 600*F. In general, for materials exhibiting ductile fracture over the range of. temperatures of interest, such as the stainless steel-product forms under consideration,.the initiation toughness, J Ic is less at reactor operating temperature (600'F) than at the low temperature of concern (120*F). Tmat usually increases somewhat with temperature however.
Therefore, it is conservative to use the higher temperature J Ie as instability criterion for operation at 120*F. Furthermore, crack instability of the lines at 120*F is determined based on JIc " I 3**
including thermal aging, even though the thermal aging does not occur at the 120*F operating temperature. Therefore, J3e = [ P is a very low value compared to the J Ie values of the actual materials. A discussion on the effects of thermal aging on the integrity of welds at the
- limiting location is given in Appendix D.
4.4 References 4-1 F, .1, Witt et al., " Integrity of the Primary Piping System of Westinghouse Nuclear Power Plants During Postulated Seismic Events,"
WCAP-9283, March 1978.
4-2 S. S. Palusamy, " Tensile and Toughness Properties of Primary Piping Weld Metal for Use in Mechanistic Fracture Evaluation," WCAP 9787, May, 1981 (Westinghouse Proprietary Class 2).
4-3 S. S. Palusamy, et al., " Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall 4
Crack," WCAP-9558, Rev. 2, May 1982, (Westinghouse Proprietary
~
Class 2).
l
- 5,s ecaes,5aasa7 ic 4_4 l
l
l 4-4 W. H. Bamford, et al., "The Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping for Westinghouse Nuclear Steam Supply Systems," WCAP-10456, November,1983 (Westinghouse Proprietary Class 2).
4-5 F. J. Witt and C. C. Kim, " Toughness Criteria for Thermally Aged Cast Stainless Steel," WCAP 10931, Revision 1, July 1986 (Westinghouse Proprietary Class 2).
i I
e 1
1B
, hb ' h4 7 10 4.,5
- N
- OA I E 712555987125555929506 TR . . .
%CA 141535531413045567081 V 767776667677776666744 0N EI R
N OH I C TN 525000005250000000000 AI
%G 830260858306560809505 NR 556655445565455445433 OE LP E
S HA CX 000000000000000000000 NE S 555500505550550555500 I T WS) 4,4,3,2,0,4,0,9,4,4,3,0,72,4,0,15,7,5,4, OEi , ,
2H LR s 450565784506465791485 1 T FT p 666667766666667766677 UF S(
, EO*
HS0 T 6 E5 EH FH TT 000000000000000000000 OTT AG) 000000000000000000000 A MNi 6,2,6,5,0,6,7,4,6,2,6,0,0,5,6,77,0,0,0,4, SF I E s ,
EOG TR p 5821 23805822913893906 I N LT( 888888888888788878798 TSI US RDT ELA PER 1 OWE TS
- R P ES 4 POO SE N FR 000000000000000000000 E LAT FT) 000000000000000000000 L A N OSi 3,7,1,0,0,2,4,4,3,71 0,5,0,2,4,6,1 5,0,4, B CSA s , . ,
A I LL %Op 328907573280017580074 T NAP L( 443456654435556654566 AI 2. E HR1 0I CE Y ETT MAI MN LL LL LL E U 66L 66L 66 RE 666661 1 666666611 66611 UNT L 111 1 1 3311 1111 1 3311133 TI C A 33333RR3333333RR333RR ALE I PPPPPEEEPPPPPPEEEPPEE R J R TTTWW - - - TTTWWW - - - TW - -
ERO E - - - - - - - - - - - - - 99 POR MTP T
A 66633 994.666333 994.63 777005557770005557055 EA M 33344AAA333444AAA34AA TL AAAAAFFFAAAAAAFFFAAFF U SSSSSSSSSSSSSSSSSSSSS MM OU OC RC 7 A 8313688 831 64688A3409 699993350999993339945 R 000881 176008881 1 80867 TE 855557758555557765509 AB PLL55110PLL55511 0L531 EM HU TTTTTTTTTTTTTTTTTTTTT N HHHHHHHHHHHHHHHHHHHHH T gg ggg g C nn nnn n UM ii i ii i DR eeettddd eeetttddd etdd OO pppttl l l pppt ttl l l ptl l RF iiiiioeeiiiiiieoeiiee u, P PPPFFWWWPPPFFFWWWPFWW m
P O . e OO n LN 1111 1 11 1 2222222223333 v wd i
1
N OA
%CA IE TR 155212591150815592.A 455141554305341556/8
. . . 0 U 676767766676667766N4 DN EI R
N OH I C TN 20052500250002000000 AI
%G 32003028305853628040 NR 56565664554445564453 OE F LP
- E 0
H 2 C 1 00000000000000000000 N S 00000000000000000000 I T WS) 5,0,4,4,5,4,8,1 5,7 8,1,9,5,3,8,1 2,9,5, A OEi , , ,
2 LR s 50505097594785697918 1 G FT p 66766657666766657667 N S(
ESI HDT TLA ER EH FWE TT 00000000000000000000 O P AG) 00000000000000000000
- DO MNi 2,5,6,6,2,6,5,7,2,6,0,7,4,2,5,5,7 7,1 ,0, SN I E s ,
EAT TR p 87328278829808178920 I N LT( 87888878897888878789 TSA US RLL EAP PI 2 OR1 TS
. AI FR 00000000000000000000 E LMN FT) 00000000000000000000 L A U OSi 7,0,2,1 7,1 0,4,7,7,5,4,4,7 0,0,4,6,7,0, B CE s , , ,
A I NT %D p 22782825260572125817 T NI C L( 44634346445654546546 ALE 2. E H J 0l CRO Y EOR MTP A L L L L L ELS 6 6 6L 6LL6 RUA 661 66661666166661 661 UMX L 1131 1 1131 11 31 11131 13 TUE A 33R3333R333R3333R33R ACT I PPEPPPPEPPPEEPPPEEPE RC R TW TTTW - TTW - - TWW - - T -
EAH - - - - - - -
P M
T U
E T
A 63 9.6663 9.663 94.633 9 4.2 9 7057770577055700551S E O M 34A3334A334AA344AA3A T S AAFAAAAFAAAFFAAAFFAF SSSSSSSSSSSSSSSSSSSS M E O H O T R 62 F 3981 3198314853598A20 O 91 39991 391 9379913324 R 091000910681508918 - 6 TE 53755537585705537610 AB L51LLL51LP510L551013 EM HU TTTTTTTTTTTTTTTTTTTT N HHHHHHHHHHHHHHHHHHHH T g g g gg C n n n nn UM i i i ii DR etd eeetd eetdd ettdd ed OO ptl ppptl pptl l pttl l pl RF ii e iiii e iii e e iii e e i e P PFWPPPFWPPFWWPFFWWPW P
O .
OO 11 1 22222333333333333 LN
. e ed I l
I E 5596.AA 55.AAA 5 5.A 9. A 5596.AA 592.A TR AA
%CA //4559//35///45/5/3559//456/
U NN7765NN77NNN77N6N7765NN766N DN EI R
N OH I C 0000000000800508000000005000 TN AI 1362880402204622830288116804
%G 6556436666566565466643565446 NR 1 .
OE LP F E 0 0000000000000000000000000000 2 0000000000000000000000000000 1 S 2,9,3,3,1 2,2,0,8,8,7,2,0,0,8,7,1 5,8,3,1 6,2,3,0,1 2,0, WS) , , , ,
T OEi 3465715149351593728 571125791 A LR s 6666786665666656766678766766 H FT p C G S(
N N I I T .
0 A EH 0000000000000000000000000000 1 R TT 0000000000000000000000500000 E AG) 5,1,5,5,7,5,2,9,0,5,9,2,9,5,5,9,7,9,0,5,7,50,5,5,7,7,9, E P MNi H O I E s 251182101751 0175821182121890 T TR p 8888898887888878888889988878 T LT( :
FSN US OLA AL SI P ER : S TS I E1 TT EE 0000000000000000000000000000
_ RAT SR 5000000500005000000000050005 EMI FT) 0,6,0,0,4,6,1 1 5,0,4,1 1 5,0,4,4,1 5,0,4,6,4,0,5,4,6,1 P N FSi , , , , , ,
3 OEU O s 2419509182191821528950128581
- RN Dp 4454674444444444644467544654 4 PI T %L( .
. E L LC E 2. I E
ARJ L 0Y B CHO A I RR i i s
T N P s AH LL L LL L HCS 66 6 66 6L CNA 666611 666666666616661 1 666166 EI X 1 1113311111111 113111331 11 31 1 M E L 3333RR3333333333R333RR333R33
_ 8T A PPPPEEPPPPPPPPPPEPPPEEPPPEEP E I TTWW - TTWWWTTWWW - TWW - - TTW - - T 99.6633366333 9.633 99.663 - - -
9 4.6 RDH R - - - - - - - - - - - - - - -
UNT E 6633
_ TAU T 7700S57700077000570055770557 A O A 3344AA3344433444A344AA334AA3 R S M AAAAFFAAAAAAAAAAFAAAFFAAAFFA E SSSSSSSSSSSSSSSSSSSSSSSSSSSS P lE M i 1
E T
_ T 1 2 2 -
F - - - 2
_ M O 2 2 5 -
O 1 - - 1 - 1 2 31 1 O S - 0 5 - 5 - - - - -
R D R 51 - 1 - 1 2 25 1 L TE - - 76891 - 9081 - 8088 9689 - - 88A -
E AB 11093291021 9102131 09321 1 0331 W EM 8878138879988799187813887188 HU 0053740053300533705374005760 D N 11551X115551155511551X115101 N
A TTTTTTTTTTTTTTTTTTTTTTTTTTTT HHHHHHHHHHHHHHHHHHHHHHHHHHHH 8
] -
_ T1 "
C[ gg 0 ggg ggg gg g
_ U nn 1 n n n 8" nnn nn n DM i i i ii iii ii i OR eettddeettt eetttd ettddeetdde ppttt pptttl pttl pptl p RO ppttl l l l PF iiii eeiiiiiiiiiieiiieeiiieei PPFFWWPPFFFPPFFFWPFFWWPPFWWP
_ P O .
OO 1111 111111122222222222333333
- LN
. o
$o
ii !I N
OA IE
- TR A 5. A
%CA /5/
- U N7N DN EI R^
s N
OH .
IC 008 4 TN
, AI 022
2 0
000 0' 0 1 S 2, 8,7, WS)
T OEi 593 A LRs 656 H FTp C G S(
N N I I T
0 A EH 000 1 R TT 000 E AG) 2,5,9, E P MNi H O IE s 175 T TRp 878 T LT(
) I E l' TS
. TT EE 000 d RAT SR 000
't EMI P N FT)
FSi 1, 0, 4, n OEU O s 921 o RN Dp 444 C PI T %L(
(
3_
L
- LC ARJ E 2.
0Y E
I
- CHO 4 I RR N P E AH
. L HCS B CNA 666 A EI X 111 T M E L 333 8T A PPP E -
I TWW ROH R - - -
UNT E 633 TAU T 700 A O A 344 R S M AAA E SSS .
P E "
M H 8 E T T s F 2 a M O -
O 2 d O S - e R D R 5 t L TE - 08 a E AB 121 c W EM 999 i HU 833 d D N 055 n N i A TTT HHH s s
l e
n
] u T1 C[ " gg ,
U 8nn e DM ii n OR ett i RO ptt l PF iii .
PFF " .
s 0 1 .
P O . ]
s OO 333 1 .
s LN [ ,
e.
llllL
J N
OA IE 3663506
%CA TR 203923550584.AA 40300500.A .
167787465535//66711608/0' 37681 .
U 766667777666NN66788784N64' /744 DN EI R
N OH
.IC 500050000000003007000008000000 TN AI 062991652056642214950041581505
%G 655556566543456666566355336633 NR OE LP E
0000000000005000000000 000000 5555555550052555005550 055500 WS)
S 3,5,3,5,5,9,2,2,2,4,9,6,8,6,5,9,9,6,9,2,7,5,- -
4,5,9,2,5,4, OEi 0561226055848352232078 512085 S LRs 6666666667676666666657 786677 HA FT p CX S(
NE IT 2H EH 000000000000000000000004000000 1T TT 000000000000000000000007000000 UF AG) 6,8,6,0,6,4,5,5,5,6,4,7,4,9,0,24,8,8,5,0,0,4,3,4,5,4,5,0,4, EO* MNi HS0 IEs 278341171306749613278016621706 T 6 TRp 888888878888888888877988898798 E5 LT(
FH US OTT A
SF EOG S I N TS TSI EE 0000000000000000000000 000000 RDT SR 0000000000005000000000 000000 ELA PER FT)
FSi 1,3,1 1,5,5,0,0,0,2,4,6,2,4,1,7,5,4,1,0,5,0, A A 4,6,5,0,0,4, 4 OWE O s 8340041397720229433377//404374
- R P Dp 3444445446565443444436NN674466 4 PDO %L(
E LAT N
N
- 2. E I .
L A 0Y B CSA A I LL T NAP L LL AI LLL L .
HR2 666 6LLL6 66 CE 666666666111666666666166616611 ETT 111111111 3331111111113111311 33 MAI L 333333333RRR333333333R333R33RR MN A PPPPPPPPPEEEPPPPPPPPPEEEEEPPEE E U I TTTTTTWWW - - TTTTTWWWW- - - - - TW - -
- - - - - - - - - - - - - - - - - - - - 99 RE UNT TI C R
E T
666666333 949.666663333 777777000555777770000555557055 94449.63 ALE A 333333444AAA333334444AAAAA34AA R J M AAAAAAAAAFFFAAAAAAAAAFFFFFAAFF ERO SSSSSSSSSSSSSSSSSSSSSSSSSSSSSS POR MTP EA TL U
MM OU 8 OC RC R 4 A TE 1113335438 8- 9663NN4N 34 AB 99999999935312339AB9A 99 EM 0000008881718377092802 9082 HU 5555555357572 - 55580310B0925309 N LLLLLL 55510115LLL915643053L545 68374 67 TTTTTTTTTTTTTTTTTTTTT0309 TT09 HHHHHHHHHHHHHHHHHHHHH3211XHH31
]
T1 C[
U ggg nnn gggg nnnn i n
g
.~ _
DM iii iiii OR RO eeeeeetttdddeeeeettttdddddetdd pppppptttl l l iiiiiiiiieeeiiiiiiiiieeeeeiiee pppppttttl l l ll ptl l 0
PF PPPPPPFFFWWWPPPPPFFFFWWWWWPFWW 1 F
t J
M _
rt P
O . W 4i OO F LN 111111111111222222222222223333 a'
- - TR .
. %CA 7587758773353621581
- U 6744674466766677744, f ON '
EI R
N OH -
I C 0000000003000000000 TN AI 9210921090685096205
%G 5633563355544455633 NR .
OE LP F .E 0 0000000000000000000 2 0000000000000000000 1 .S 6,8,5,7,6,8,5,7,6,8,0,19,2,8,3,8,5,4, WS) ,
T OEi 1981198119678916985 A LR s 6577657766676666577 FT p G . S(
N I
E T -
H A EH 0000000000000000000 T R TT 0000000000000000000 F
E P
AG)
MNi 0, 5, 0, 2, 0, 5, 0, 7, 0, 4, 0, 7, 4, 7, 2, 5, 5, 0, 4, 0SO I E s 3701370138280941706 L TR p 8798879889888788798 S'
EI N AT LT(
O RET EE 000000000000' 0000000 SR 0000000000000000000 PNI FT) 1,0,0,7,1,0,0,7,1 1,0,4,4,6,3,0,0,0,4, I N FSi ,
5 LLU O s 0271027101057891274
- A Dp 4466446644565535466 4 CRT %L( -
. E I OC NTE 2.EI L AAJ 0Y B HLO A CUR T EMP
. MU LL LL L CS 66 66 6LL LL ECA 6611 66116661 6666666 RAX 11331 1 3311131 111 11 1 U E L 33RR33RR333R3333333 T H T. A PPEEPPEEPPPEEEPPPEE AC I TW- - TW - - TTW - - - TWW - -
RNH R - - - - - - - - - - 99 EIT P U E
T 63 99.63 99.663 944.633 7055705577055570055 M2O A 34AA34AA334AAA344AA E1S H AAFFAAFFAAAFFFAAAFF T SSSSSSSSSSSSSSSSSSS E
M l i
O T O
R F O
S 2 2 2 D R 390939093268 A25909 L . TE 9145914599935399145 E AB 0967096700817808967 W EM 5309530955575655309 l U L531L531LL5100L5531 D l N N TTTTTTTTTTT1TTTTTTT A HHII '
HHiHHHHHHHHHHHHH l
T C g g g gg UM n n n nn DR i i i i i OO etdd etdd eetdddettdd RF ptl l ptl l pptl l l pttl l P ii e e ii e e iii e e e iii e e
. PFWWPFWWPPFWWWPFFWW P
O .
OO 1111222233333333333 LN eJ"
' l i! i lll11J N
OA IE
%CA TR AAA 5596.AA 0. A 0460.AA 5596.AA 5086
///4559//4/7118//4559//4539 U
DN NNN7765NN5N7744NN7765NN7665 EI R
N OH IC TN 000000000005000005000005000 AI 443628870056750436288736858
%G 656564356565333665643565443 NR OE LP F E 0 000000000000000000000000000 2 000000000000000000000000000 1 S 0,9,3,3,3,16,4,2,9,7,8,4,4,5,0,3,0,3,1,6,1,3,0,1,9,6, WS) ,
T OEi 150657106046658105571005781 A LR s 666667865655777666678666768 FT p G S(
N I
E T HSA EH 000000009006000000000000000 TLR TT 000000007570000000000500000 AE AG) 9,6,6,5,5,7 5,7,7,2,0,5,9,4,0,9,6,5,5,7,5,5,6,5,7,4,5, FIP MNi ORO I E s 069118225975960091182891802 E TR p 887888987777889878889778889 STT LT(
EAN US IMA T L REP EN S PI 2 TS OL EE 000000008007000000000000000 R T SR 555000005022000550000550000 PRI HN FT)
FSi 1,1, 0, 0, 0, 4, 6,1, 5, 5, 4,1, 9, 4, 0,1, 0, 5, 0, 4, 6, 6, 0, 5, 4, 4, 6, 6 LRU O Dp s 151195086228247118950118570
- A 444546733433666444467444657 4 CHT %L(
E ICC NNE 2.
0Y E
I .
L AIJ B H O A CHR i i I T ECP s s S .
MN LL LLL LL L L I S 66LLLL 866 66 6L6 E A 66666114666L011666611666161 R8X 111113301116333111133111313 U E L 33333RR33331RRR3333RR333R3R TDT A PPPPPEEPPPP3EEEPPPPEEPPPEEE AN I TTTWW- TWWWF - - - TTWW - TTW - - -
RAH R - - - - - - - - - - - - 449 E
PHU T E T
66633 99.23332 999.6633 99.663 777005510008555770055770555 MCO A 33344AA34441AAA3344AA334AAA ENS M AAAAAFFAAAAAFFFAAAAFFAAAFFF TI SSSSSSSSSSSSSSSSSSSSSSSSSSS E .
M0H " -
O1T 8 O 11 1 R F - - - 1 s O 112 12 - 2 a
- - - - - 5 -
S 889 289 19 d D R - - - 76894894N490 - - 8689 - - 8859 e L TE 11103326652B054110932110372 t E AB 888781389309776887813887153 a W EM 000537479004090005374005704 c HU 111551XMEFF141311551X11510X i D N d N TTTTTTTTTTTTTTTTTTTTTTTTTTT n A Hl HHHHHHHHHHHHHHHHHHHHHHHHH l
i s
s l
e
] n T1 gg " gggg gg g u .
C[
U nn 8nnnn nn n e n
DM ii i iii i i i OR eeettddettttdddeettddeetddd i RO pppttll pttttl l l ppttll pptll l l PF iiiiieeiiiiieeeiiiieeiiieee e . -
PPPFFWWPFFFFWWWPPFFWWPPFWWW "
n 0 e 1 n, P a O . ]
OO 111111111111111222222333333 LN 1
[ m -
a'," -
L l
TABLE 4-7 ROOM TEMPERATURE TENSILE PROPERTIES
, OF THE SA351 CF8A CAST 45 DEGREE N0ZZLE Yield Stress Ultimate Strength Elongation Reduction Unit Loop (ksi) (ksi) (%) in Area %
1 1 41.35 86.2 58 70 1 2 41.35 86.2 58 70 1 3 39.10 83.2 59 71 2 1 42.38 86.3 62 69 2 2 38.05 84.8 58 75 2 3 39.15 83.0 62 74 e
m l
2S465 2081787 10 f.
g- .q j i
)
STABLE 4-8' l !~ . . TYPICAL 4 TENSILE PROPERTIES OF.SA376 TP316,'SA351 CF8A.and WELDS OF.
~
SUCH MATERIAL FOR'THE PRIMARY LOOP l Test Temperature Average-Tensile Properties
' Plant- Material (*F) Yield (psi)~ Ultimate (psi)
A SA376 TP316 70 40,900.(48)a' 83,200 (48)
.650 23,500 (19)- 67,900 (19)
E 308 Weld 70 63,900.(3) 87,600 (3)' 8 SA376 TP316 70 47,100 (40) 88,300 (40)' 650- 26,900(22) 69,100 (25) E 308 Weld 70 59,600 (8) 87,200(8) 650 31,500 (1)- 68,800(1) '
'C SA376 TP316 70 46,600(36) 87,300(36) 650 24,200 (18) 66,800(19)
E 308 Weld 70 61,900 (4) 85,400(4) D SA351 CF8A 70 47,300 (14) 84,500 (14) 650 26,000 (4) 70,500(4) Weld 70 61,200(31) 84,500 (32) l i
- a. ( ) indicates the number of test results averaged.
2545s'041787 10 4 }f
l 88 b e l 11
. t . 77 a 9 m 6 oo i tt t o b c l t 5 86 552 U
9 0 87 445 F 7 66 666
- 7 0 6 5
6 a j 9 55 R I 6 88 ES d 2 11 HP l TO e o oo O i t tt DL ) Y b c N i 5 0 50 990 AY s R k 3 6 14 771 SA ( 2 2 35 112 EM NI s I R e LP i t RTS r ab aa b
- OHT e p
0382 . 6 3 s TGN a t AUE o e 9836 07 9 n LOM r t 8888 98 7 e URE P a m MWR m ooo oo o e U I ei tttt tt t r CLU rt b i CAQ ul 4250574 7859 u 0000000000 ACE tU . q I R a 1373495 2559 e 5570050055 EP r 8878878 8778 R 7777877767
. 9 HYM e p
- TTU m 4 M m u FFI e m E OON T i L I 8b aa a a n B SEM m 29 5 i A ES o 3.l 0 4
. M T I OE o 0712 73 2 THD Rd 5454 66 4 e RTO l d E C e oooo oo o o PH i tttt t t t C OTE Y b 1419 E RI M PWS 195.l
. 349 . . M 000 000 A 8078779 8282 S 005 555 a ES 3433455 3336 A 333 222 t LED a I NN d SI A NL n E
TR i i i o H S S t FR L LL L LL c O 666 46L 66 6 46L c D 1 1 1AA6 016 11A1 01 6 i NE 333881 331 3383 331 f OH l PPPFF3 PP3 PPFE PP3 i SC a TTTCCE, TWF TWC TWF t I A i - - - - - - - L - - , - - L r RT r 66311L 2328 631L L2328 e AT e 770556861 080 7056881080 c PA t 334331013413 3431 003413 M a AAAAA333AAAR AAA333AAAR l O M SSSSSEEESSSE SSSEEESSSE a _ C i _ r e t t a . . m n _ e ff n g gg o _ o p ne nn eil e dd ii e _ m ept zpdl l ettd ss ss g o pi t zil eepttl gg gg n _ C iPi oP eWWiii e nn nn a _ / P/FN/W//PFFW ii sss i is R _ e / p/ / p/ p p/ / / / ettdddettd n ccccocoocccc pttl l l pttl i L coccocoocccc A1 AALALLAAAA iii e e e iii e PFFWWWPFFW a iM - i
. iI 1I )} !l
e. c, a e.
~
c,
- _ a e
t a m T
)
2 n E . i N c / I I b L J l R n O i T ( - A ~ - L U M e, U . C ) c, C i e b A s t 00 0 aj p a 00 0 E ( m 5,1 2, - - H;
'se i , -
TG t 50 1 N l 66 6 FI i U ' OP t I r LP e p 0 A 1 CR o
- I H r 4 PR P Y b E TD e G0 0 L E l 00 0 B SH i d 7, 5, 0, -- -
A EC s l - T I A n e 10 5 TT e i 22 4 RT T Y EA P OD RN PA S p S m E e ) N T F 00 00 0 H
- 00 00 0 G t ( 66 66 6 U s O e T T E
R U . T s C ) t A 6 s R 1 t e F 3 E 6 d n f a o 66 h 11 8 A t 33 0 8 s l PP 3 F e a TT E C w i ( o r 66 7 1 L e 77 dd 5 t 33 l l 2 . a AA ee A ,b - M SS WW S [ -
,h
TABLE 4-11
. CHEMISTRY AND END OF SERVICE LIFE KCU TOUGHNESS FOR THE SIX 45-DEGREE N0Z2LES a,c e I
l a N was taken as 0.05 and Cb was taken as 0.0. j I e em
$? $
a.c c Figure 4-1. True Stress Strain Curve for SA376 TP316 Stainless Steel at 550*F l e n o..o. m 4 4-18
a,c.e
~
a, c e i
^
w M we 5
~
2 m b l
- t 1
I
)
Figure 4-2. Minimum True Stress-True Strain Curve for SA376 TP316 Stainless Steel at 120*F 4 l l putsw mte 4-19 l l l
i a ,c c.
. 1
' ~
a,c.e J l 1 i I i U N t M e ~ E r Figure 4-3. Minimum True Stress-True Strain Curve for SA312 TP304L Steinless Steel at 120*F l ws,ww n 4_,0
SECTION 5.0 LOADS FOR FRACTURE MECHANICS ANALYSIS 1 Figures 5-1, 5-2 and 5-3 are schematic layouts of the three accumulator lines with attached RHR piping. The stresses due tt axial loads and bending moments were calculated by the following equation: o=k+y (5.1) where, o = stress F = axial load H = 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=/M 2+H y 2 z (5.2) where, M = bending moment for required loading
~
My = Y component of bending moment M Z
= 2 e mp nent of bending moment l
1 The axial load and bending moments for crack stability analysis and leak rate predictions were computed by the methods to be explained in Sections 5.1 and S.2. t o,za2"" 5-1
i 5.1 Loads for Crack Stability' Analysis The faulted loads for the crack stability analysis were calculated by.the . following equations:
=
-F-(FDW + FTH + Fp l + lFSSE l (5.3)
=
My l(My )DW + @Y)TH I
- I EY)SSE l (5.4)' l
=
M 2 l(M)DW+ Z IMZ )TH I
- IIN )SSEl-2 (5.5)
Where, the subscripts of the above equations represent'the following loading cases, DW '= deadweight TH = normal thermal expansion i SSE = SSE loading including seismic. anchor motion P = load due to internal pressure .. 5.2 Loads for Leak Rate Evaluation .
.The normal operating loads for leak rate predictions were calculated by the ,
following equations: F = FDW + FTH + Fp (5.6) N
- Y (NY )DW + @Y)TH (5.7)
M Z IMZ )DW + @2)TH (5.8) 5.3 Summary of Loads, Geometry and Materials Tables 5-1, 5-2, and 5-3 provide a summary of envelope loads computed for fracture mechanics evaluations in accordance with the methods described in l Section 5.1, and 5.2. The cross-sectional dimensions and materials are also summarized. Load data are tabulated at the highest stressed location (node 104E, loop 1), and the second highest stressed location (Node 795, loop 3), ; for the 12 in. pipe'in Table 5-1. Table 5-2 gives load data for the 10 in. .j pipe for its highest stressed location (node 305, loop 2), and its second l no,m:eene 5-2
, highest stress location (node 303, ioop 2). Table 5-3 gives load data for the 8 in, pipe for its highest stressed location (node 284, loop 2) and its next , highest stressed location (node 139, loop 1). The loading components are provided in tables 5-4, 5-5, and 5-6. Schematic layouts of the accumulator and the attached RHR lines are shown in figure 5-1, 5-2, and 5-3 including the critical locations.
e l . w w w e m ,?n2u n e 53
k 4 1 1 0
- 0 0 9 5 n 1 1 9 6 i :
(
)
_ s p - 5 Fi 7 6 2 k 0 0 0 0 ( 2 2 2 2
)
s E . e 1 DAh 4 4 4 4 II c 7 7 7 7 SD n N i 0 0 0 0 I ( 1 1 1 1
)
M - s ULK e MLCh 5 '5 5 5 I AI c 0 0 0 0 NWH n 0 0 0 0 I Ti M ( 1 1 1 1
~
)
L s ALKe - NLCh 5 5 5 5 E I AI c 2 2 2 2 P MWH n 1 1 1 1 I O Ti . . . P N .( 1 1 1 1 H C E N L I. U D 2 E 1 H 0 0 0 0 C 4 4 4 4 R S 1 1 1 1 O F 1S )
- D E s 5A D e O I h 5 5 5 5 EL S c 7 7 7 7 L tan BE UI i 2 2 2 2 AP OD( 1 1 1 1 TO -
L E V N L E A . I ' 6 6 6 6 F R 61 61 61 61 O E 73 73 73 73 T 3P 3P 3P 3P Y. A AT AT AT AT R M S- S - S - S - A M M P U O S O L 1 1 3 3 E E E DO 4 4 5 5 ON 0 0 9 9 N 1 1 7 7 N g g O n n I d i d i T e l t e l t I t aa t aa D l mr l mr N u re op u re op O a a - C F NO F NO N n n O t o t o I s i s i T e t e t A hda thd a C gac x gac O ioo ei o o L HLL NHlL I*
) 4 1 3 7
- S 6 0 3 7 MNP 4 4 4 3 I I (K
)
S P 4 2 4 2 FI 4 4 4 4 K (
) 5 5 5 5 M]S 5 5 5 5 U aE 9 9 9 9 M[H 8 8 8 8 I LC NLN 0 0 0 0 I AI MW(
E L U 0 0 0 0 D 4 4 4 4 E E 1 1 1 1 P H I C P S H C N ) 5 5 5 5 I E S 7 7 7 7 D .E . 0 I AH 0 0 0 0 1 SI C 1 1 1 1 TDN R U I O O ( F 6 6 6 6 2 S 1 1 1 1
- D L 3 3 3 3 5 A A P P P P O I T T T T E L R - - - -
L E 6 6 6 6
. B E T 7 7 7 7 A P A 3 3 3 3 T O M A A A A L S S S S E
V N P E O 2 2 2 2 O F L O Y R A M M E U D . 5 5 3 3 S OO 0 0 0 0 NN 3 3 3 3 e r g g o n n b i i r t t e a a t r r n N e p e p u O o I O O c T d d I e l e l d D t a t a l N l m l m e O u r u r w . C a o a o F N F N t a n e
. N n n k O t o t o a I s i s i T T e t e t A hd a thd a C gac x gac ]
O i oo eioo a L HLL NHLL [ T*
)
- S 7 4 5 2 MNP 3 7 9 4 I I 2 1 1 1 (K
)
S P 6 5 6 4 FI 2 2 2 2 K (
) 5 5 5 5 M]S 3 3 3 3 UaE 1 1 1 1 M[H 8 8 8 8 I LC NLN 0 0 0 0 I AI MW(
E L U 0 0 0 0 D 6 6 6 6 E E 1 1 1 1 P H I C P S l i C N ) 3 5 5 5 I E S 2 2 2 2 D . E 6 6 6 6 8 I AH SI C 8 8 8 8 R TDN O U I F O ( 3 S 6 6 6 6
- D 1 1 1 1 5 A L 3 3 3 3 O A P P P P E L I T T T T L R - - - -
B E E 6 6 6 6 A P T 7 7 7 7 T O A 3 3 3 3 L M A A A A E S S S S V N E P O 2 2 1 1 F O O L Y R A M M U E S D . 4 4 9 9 OO 8 8 3 3 NN 2 2 1 1 e r g g o n n b i i r t t e a a t r r n N e e u O p p o I O O c T d d I e l e l d D l a t a l N l m l m e O u r u r w C a o a o F N F N t a n e N n n k O t o t o a I s i s i T T e t e t A hda thda C gac x gac ] O i lLoo i eioo a L l NHLL [ T*
t
, gf n( 5 0 8 i 9 0 5 dZ 7 7 7 nM 3 4 - 9 e 5 1 Bt -
) n 3 e
- m dp o ao M oo lL
)
t , b s5 l e9 .- h7 t g gf i - n( 7 7 0 H i 2 2 6 n dY 5 9 - 2 E t o nM 7 4 N xi e 1 2 I et Bt L N a n c e H o m C L o N ( M I 2 1 ) b E l 0 0 6 3 _ H l ( 9 6 2 1 T a 6 8 5 9 _ i e 1 9 2 _ R xc 9 _ O A r 1 F o F S _ N O _ I ) T b A l _ 4C -
- O gf t
_ 5L n( _ EG i 4 2 7 LN dZ 0 2 4
. BI nM 0 7 - 2 AN e 4 0 2 TR Bt 8 E ) n e
V 1 _ O - m _ G p o o M T o A L d ) S a , b T oE l N l4 - E 0 t N t1 gf O s n( P e- i 1 6 3 M h dY 3 9 8 O gn nM 0 - 6 C i o e 1 2 Hi Bt G t n N a e I c m D o o A L M O ( L
)
b I 2 5 6 3 l ( 8 1 2 8 a 2 9 5 6 i e - 6 9 x c 9 A r 1 o F
. t e o l r M t a u h m s + ,
d e d g r s h ap a i e e E c oy e e h r S n LT DW T P S A
'4 "
i
)
b l p t g f n ( 9 i 5 0 8 d Z 2 0 7 n M 9 1 - 6 e 2 2 3 B t - -
, n
) e m
2
- o d p M a o o o )
l L r.. b t , l s 3 - e 0 t h 3 g f g n ( i - i 6 5 4 E H d Y 9 1 4 N n n M 8 2 - 9 I t o e 1 2 2 L s i B t - 2 e t n H N a e C c m N o o I L M w 0 1 R O (
)
b 1 5 6 7 6 F l ( 6 0 6 6 a 4 0 9 5 S i e 2 9 1 N x c 3 O A r I o T F A C O 5 L )
- b 5 G l N -
E I t L N g f B R n ( 1 8 A E i 5 4 3 T V d Z 6 0 7 O n M 2 4 - 9 G _ e - 2 3 _ B t - T n A ) e 2 m S - o T p M N o
~
E o N d L ) O a b P o , l M L 5 - O 0 t C t 3 g f s n ( 2 2 6
~ G e - i 4 9 2 N h d Y 5 5 - 5 I g n n M 2 2 3 D i o e - 2 A H i B t O t n L a e c r o o
-_ L M
(
)
b I l ( 8 2 7 1 a 5 0 6 0 i e 4 0 9 6 x c 2 9 1 A r 3 o F t 0 e o 9 P l r M t t a u fu 7 h m s + . 8 0 d e d g r s h a p a i e e E c % o y e e h r S n 4 L T DW T P S A 7 m*
t g f 7 = i n (
~
8 9 6 d Z 3 9 2 n M 9 3 - 6 e 3 3 1
. B t - -
n
) e 1 m
- o d p M a o o o l L_ )
b t , l s 9 - e 3 t h 1 g f g n ( i - i 2 1 8 H d Y 9 5 6 E n n N 6 5 - 2 N t o e 8 4 I x i B t L e t n N a e i l c m C o o N L M I ( 8
)
R b O l F l ( 8 6 6 2 a 4 4 0 4 S i e - 2 3 N x c 4 1 O A r 2 I o T F A C O 6 L )
- b 5 G l
- N -
. E I -
t L N g f B R n ( A E i 1 1 4 T V d Z 4 1 9 O n M 4 3 - 3
. G e 1 8 1 B t - -
T - n A ) e 2 m S - o T p M N o E o N d L ) O a b P o , l M l 4 - O 8 t C t 2 g f 1 2 3 s n ( 6 8 3 G e - i 1 5 5 N h d Y 4 6 - 5 I g n n M - - D i o e A H i B t ~ O t n L a e c m o o L M (
)
b I l ( 7 8 6 1 a 7 6 0 5 i e 2 2 0 x c 4 1
. A r 2 o
F t
. e o l r M t a u h m s + .
d e d g r s h a p a i e e E c o y e e h r S n L T DW T P S A p* i l 1
)
Node 104E: Highest Load for the 12 in. Pipe i I d/ Y i 104E
/
l 1 l FIGURE 5-1 Schematic Layout of Accumulator Line and the Attached RHR Lines - Loop 1 5-10
1; t l Node 284: Highest Load for 8 in. Pipe' Node 305: Highest Load for 10 in. Pipe L_ 305 Z
^
k a4 3 FIGl'RE 5-2 Schematic Layout of Accumulator Line and the Attached RHR Lines - Loop 2 5-11
i PRIMARY LOOP COLD LEG i ACCUMULATOR TANK FIGURE 5-3 Schematic Layout of Accumulator Line and ; the Attached RHR Lines - Loop 3 I 5-12 !
'SECTION 6.0 FRACTURE MECHANICS EVALUATION 6.1 Global Failure Mechanism Determination of the conditions which lead to failure in stainless steel
'should be done with plastic fracture methodology because of the large amount of deformation accompanying fracture. One method for predicting the failure of ductile material is the [. Ja,c.e method, based on traditional plastic limit load concepts, but accounting for [
.)a,c e 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. I . Ja,c.e This methodology has been shown to be applicable to ductile piping through a large number of experiments and is used here to predict the critical flaw size in the accumulator line. The failure ~ criterion has been obtained by requiring equilibrium of the section containing the flaw (Figure 6-1) 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: ;
"'C
[ ] (6.1) where: I l
)a,c.e m i o m iric 6-1
L [ } ./ Ja.c.e. (6.2) 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 6-1). A typical segment of the pipe under maximum loads of axial . force F and bending moment M is schematically illustrated as shown in figure 6-2. In order to calculate the critical -flaw size, a plot of..the limit moment versus crack length is generated as shown in figures 6-3, 6-4, and 6-5 for 12 in.,10 in., and 8 in. line, respectively. The critical flaw size corresponds to the intersection of_ this curve and the maximum load line. The critical flaw sizes , are calculated using the ASME Code (reference 6-2) minimum tensile properties for SA376 TP316 (wrought) stainless steel. Figures 5-1, 5-2, and 5-3 identify the locations of the critical regions. l 1 If the limit moments are larger than the maximum applied moments for cracks i smaller than the critical flaw sizes corresponding to the maximum applied moments, then the global stability criteria of section 2.2 are satisfied. l Figures 6-3 through 6-5 show the critical flaw size predictions. The critical flaw sizes corresponding to the limit moments are shown below for the 12 in., 10 .in., and 8 in, pipes:
~
a,c.e 12 inch pipe 2545s tot 178710 g.2
'n a c.e 1 L10 inch pipe 8 inch pipe.
If-the ASME Section XI IWB.3640 approach, reference 6-8, is used and if the material strength properties are conservatively assumed to be the same as the base metal properties, the critical flaw sizes'for the weld metal for, the 12 in'.,10 in., and 8 in. pipes are as follows: a,c.e 12 inch pipe 10 inch pipe
;B: inch pipe i 6.2 Leak Rate Predictions fracture mechanics analysis shows that postulated through-wall cracks in the . accumulator line and the attached RHR lines would remain stable and not cause a gross failure of this component. If such a through-wall crack did exist, it would be desirable to detect the leakage such that the plant could be brought to a safe shutdown condition. The purpose of this section is to discuss the method which will be used to predict the flow through such a postulated crack and present the leak rate calculation results for through-wall circumferential cracks.
6.2.1 General Considerations Depending upon the temperature and pressure of the coolant, different methods for calculating leak rates are appropriate. For the 12 inch line extending from the RCS cold leg injection points to the
- first check valve, an ( -
] a,c,e i ms. wnne 6-3
[ .]a,c.e is used. In this model, 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 g , R/D ) is greater than [ Ja.c.e,both'[. Ja, ,e must be considered. In this situation the flow can be described as being single phase through the channel until the local pressure equals the saturation pressure of the fluid. At this point, the flow begins to flash and choking occurs. Pressure losses due to momentum changes will dominate for [ Ja,c e However, for large L/DHvalues, friction pressure drop will become important and must be considered along with the momentum losses due to flashing. For the lines operating et 120'F, the coolant pressure is 665 psi and the fluid inside the pipe is therefore a subcooled liquid. Since the saturation temperature at atmospheric conditions is 212*F, subcooled liquid cannot evaporate when the fluid expands to atmospheric pressure. Hence, for those lines the leak rate calculations are performed based on single phase (liquid only) flow. A benchmark calculation for this model is given in Appendix F. , 6.2.2 Calculation Method . In using the isentrepic equilibrium model, the basic method used in the leak rate calculations is the method developed by [. I 3a,c,e , The flow rate through a crack was calculated in the following manner. Figure ! 6-6 from Reference 6-3 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 ( )*'C'8 was found from Figure 6-7 taken from Reference 6-3. For all cases considered, since [ la,c.e Therefore, this method will yield the two phase pressure drop due to momentum effects as illustrated in Figure 6-8. Now using the assumed flow rate, G, the frictional pressure drop can be m e. * "" " 6-4 l
i l calculated using 2 (L/D - 40)G
< aPf = [f P f 9 c (144)] ' ' (6.3) l where the friction factor f is determined using the [ Ja,c.e L The crack relative roughness, c, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [ Ja,c,e RMS.
The frictional pressure drop using-Equation 6.3 is then calculated for the ' assumed flow and added to the [.
,a,c.e to obtain the total pressure drop from the primary system to the atmosphere. Thus, Absolute Pressure - 14.7 = [ Ja,c.e(6.4) for a given assumed flow G. If the right-hand side of Equation 6.4 does not agree with the pressure difference between the piping under consideration and the atmosphere, then the procedure is repeated until Equation 6.4 is satisfied to within an acceptable tolerance and this results in the flow value through the crack. This calculational procedure has been recommended by [
Ja,c.e for this type of [ Ja,c,e calculation. In using the single phase (liquid only) model, the basic method is to use the simple orifice-type flow formula given by Crane in the Handbook of Hydraulic Resistance Coefficients (reference 6-5). The pressure drop due to friction is included in predicting the leak rate. The leak rate 0 is given by the following equation:
, 0=( ) A ft 3/sec (6.5) 2545s'081787 10 6-5
ja.c.e 6.2.3 Leak Rate Calculations Leak rate calculations were made 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 6-7 and the leak rate were calculated using the calculation methods described above. The leak rates , are calculated using the normal operating loads of axial force F and bending moment M. These are given directly below. . F = 206 kips, M = 1017 in-kips for 12 in. pipe F= 42 kips, M = 401 in-kips for 10 in. pipe F= 25 kips, M = 174 in-kips for 8 in. pipe The crack lengths yielding a leak rate of 10 gpm (10 times the leak detection requirement of 1.0 gpm) are found to be as follows: a,c e Thus " reference" flaw sizes of ( Ja,c.e for respective pipe diameters 12 in., 10 in., and 8 in. are established. . l wwem,wuno s.6 l l i i i
i
)
6.2.4 Leak-Detection-Capability j The leak detection system inside the containment can detect a 1 gpm leak rate
'd - as required by Regulatory Guide 1.45. (See Appendix G). As seen above, a margin of 10 was applied to the leak rate to define the flaw in accordance )
with NUREG 1061, Volume 3.
]
l 6.3 Local Stability Analysis I Crack stability analyses were performed for the 12-inch accumulator piping and the 10 inch and 8 inch RHR piping attached to the accumulator line, using the EPRI elastic plastic fracture handbook method. The minimum yield strength properties established in section 4.0 of this report were used for these calculations. J applied was calculated for twice the reference flaw size using the faulted loadings and was found to be less than the JIc criteria of a,c.e ( J . J applied was also calculated using the reference flaw size with the faulted loads increased by a factor of /2. The J Ic criteria was met for these cases also. The J applied values are summarized in table 6-1. 6.4 References 6-1 Kanninen, M. F. et al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks" EPRI NP-192, September 1976. 6-2 ASME Section Ill, Division 1-Appendices, 1986 Edition, July 1, 1986. 6-3 I 3a,c.e 6-4 [i I
)a,c.e ,
1 m woeneno 6-7
.L 6-5' Crane, D. P., " Handbook of Hydraulic Resistance. Coefficient."
~
6-6 Dillio, C. C., Thermal Engineering Int. Textbook Co'. Scranton. PA p. 273, 1969. 6-7 Tada, H., "The Effects of Shell Corrections on Stress Intensity Factors and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe," Section 11-1, NUREG/CR-3464, September 1983. 6-8 ' Pressure Vesse1 ~ and Pip'ing Codes, Journal of Pressure Vessel Technology,
. Transactions of the ASME, Vol.108, August 1986.
i
)
ms,a.,wune 6-8
TABLE 6-1
.- J RESULTS APPLIED i
J applied U8i"9 Flaw Size Revised Loads Pipe Location Load (in) (in-lb/in2) l
~ '
- 12 104E (Normal + SSE) 7 12 104E /i (Normal + SSE) 10 305 (Normal + SSE) 10 305 O (Normal + SSE) i 8 284 (Normal + SSE) ,
8 284 /2 (Normal + SSE)
- a
- Calculations were performed for larger flaw sizes of 6.2 and 3.1 inches respectively.
t G N ' q l 2546 t3ssi<oe2sa? to go
)
l 1 g - (
+
sem I
- i e
O enN l D l c g b ( e
.e.
e ' M L a M C U. N m .. 7 a N, - 0 b 3 J .O b A O J p E i. 6-10
;l
- l. i
] 41)i!IlI ~
=
k c 9,- ,% 5 7
'c a CF-r 2 1
5 0 ' 0 1 F) m
~
_ _ l e d _. M o _ e _ i p _ . P e _ h _ t
. _ _ n o
F ,L_ i g n t A c s _ - d a o
- - L 2
- - 6 e
- - r u
g
- - i F
- ~
(< i 4 s p
. 0 3
2 0 1
=
7 8 4 P... - 1 8 m 4 5 2 m.C
i i a,c,e L I
)
N) rw stoe:nv 00 =-12.75" t = 1.005" P = 2304 psig F= 207 kips ey = 19.2 ksi ou = 1.8 ksi op = 45.5 ksi Temp = 5600F Figure 6-3 " Critical" Flaw Size Prediction for 12 Inch Line l vs .ua. wena, s.12
- a,c.e S
4 l
. )
FLAK' GEOMETRY OD = 10.75 t = 0.8955 P = 665 psig F = 44 kips c, = 29.16 ksi c = 75.0 ksi
= 52.08 ksi F
Temp = 120*F 1 . l l Figure 6-4 Critical Flaw Size Prediction for 10 Inch Line n o. o ui,se 6-13
i
~
8eCe9 ., I l 1 - T l
)
~
FLAK' GEOMETRY OD = 8.625" t = 0.8135" p = 665 psig l l F = 26 kips cy - 29.16 c = 75.0 ksi cf = 52.08 ksi Te: p = 120*F L - - Figure 6-5 Critical Flaw Size Prediction for 8 Inch Line 6-14 rm.wan u
~
.,a,c,e a
=
I I> 8 s 5 m ( 3
~
~
STAGNATION ENTHALPY (102 8tuhb) Figure 6-6 Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures P v 6-15 1
......-..__-_w
s a ,e ,e h*
/
e E s w s h w f d - w t-6 - LENGTH /DI AMETER RATIO (L/D) Figure 6-7 [ .)"'C Pressure Ratio as a function of L/D 6-16
W a,t,e a,c,e
/ f
- _- 2'
- : L/Dg = 40
*-- L/Dg 40
- 4 L :
~
Figure 6-8 Ideelized Pressure Drep Profile Through a Postulated Cre:k 6-17 i I .. .-
n r,
.SECTION 7.0 m ;
ASSESSMENT. 0F FATIGUE CRACK GROWTH '
, The purpose of the fatigue crack growth (FCG) analysis is'to demonstrate.that Lapostulatedflaw'willnotgrowthroughthewallunderalldesignand operational loadings.
The fatigue crack growth on the South Texas accumulator line with its attached RHR piping was determined by comparison with a generic fatigue crack growth analysis of a similar piping system. In particular,-the 12 inch line extending from the RCS. cold leg injection points to the first check valve l compares reasonably well with the generic analysis, having essentially the same geometry, materials, and fatigue crack growth rate. The 10 inch and 8 inch lines were also reviewed in comparison to-the generic analysis, and found c .to be compatible from the standpoint of materials and geometries (wall thickness)butwithmuchlowerloads. Based on comparing all parameters critical to the. fatigue crack growth analysis, for the 12 inch,10' inch, and 8 i inch lines, it was concluded that the generic analysis'would envelop their
' fatigue crack growths. The details of the generic fatigue' crack growth
, analysis are presented in appendix B. Fatigue crack growth results are summarized in table B-4 of appendix B.
Due to similarities in Westinghouse PWR designs it was possible to perform a generic fatigue crack growth calculation which would be applicable to many projects. A comparison was made of stresses and number of cycles, material, geometry, and types of discontinuities.
'A review of all thermal transient and steady-state stresses indicated a signifi-cant margin in the generic analysis. This was caused primarily by a reduced number of safety injection transients for South Texas because of the 4XL design as compared with the 412 design used in the generic fatigue crack growth calculation. Geometry was essentially identical between the South Texas 12 inch schedule 140 pipe and the generic 10 inch schedule 140 pipe. Both the generic and the South Texas 12 inch line had the same materials for the piping, SA376-TP316 austenitic stainless steel. Although the nozzle materials
'""'""'"' 7-1
l are siightly different, SA-351-CF8A versus SA-182-304N or SA-182-316N for the - generic case, the lower yield strength of the 12 inch piping was still controlling. - In conclusion, the fatigue crack growth calculated for the generic case, as summarized in section B.2.2, is applicable to the South Texas accumulator l lines and attached RHR piping. These results demonstrate that no significant fatigue crack growth will occur over the 40 year plant design life even for the largest postulated flaw. 7.1 Acceptability Fatigue Crack Growth A detailed discussion pertaining to the fatigue crack growth law used in the analysis described in Appendix B and the data used in defining the law are provided in reference (7-1). For the assessment of crack growth acceptabil-ity, the crack growth results of the generic analysis presented in Appendix B are used conservatively and are considered applicable to the South Texas Project Accumulator lines with attached RHR piping. Detailed discussion has - been provided in the previous section. The number of occurrences of a particular transient are equally spread over the lifetime so that the crack growth can be determined at intermediate times during the lifetime of the structure. The cycle input data is used to determine a schedular distribution by dividing each transient into event categories. Specifically, the total cycles of transients are scheduled into five events per year, two events per year, one event per year, one every fourth year, and one every eighth year. The crack size is updated by adding the incremental growth due to a transient to the initial crack size. The updated crack size is used as the initial crack size for the next transient loading. The calculations are repeated to account for all the events during each year. OBE was not included in the fatigue crack growth evaluation of Appendix B because the seismic stress levels were very low. In fact in the South Texas - Project, for SSE plus anchor motion, which has larger loads than OBE, the sm, $m, $a2an e 7-2
maximum nominal axial stress at the governing location in the accumulator line is + 3.7 ksi. This level of stress will result in negligible crack growth, perhaps none, for the small number of cycles due to OBE. t The maximum allowable preservice indication may have a depth of 0.09 in. per IWB-3514.3, Allowable Indication Standard for Austenitic Piping, ASME Code, Section XI - Division 1, 1986 edition. Typical fatigue crack growth results for various initial flaw depths are given in Table B-4 in the appendix to this report. i ja.c.e is conservatively chosen as a basis for examining the NRC criteria (7-3) pertaining to allowable fatigue crack growth. [' Ja,c.e Thus, the first criterion.en flaw depth is satisfied. The worst case transient AK value for the maximum crack depth is [ )**C The flow stress for the base metal at 560*F is 45.5 ksi which can be used to obtain a conservative estimate of the plastic zone size. The expression for plastic zone size, pr , calculation is: [ ] v "p
- I2 oI f AK
) g,,)2 Thus, the plastic zona size is calculated to be ( ).a,c.e The remaining liga ~ ment for the 0.186 in. deep end-of-fatigue-life flaw is 0.819 in. (i .e.1.005 - 0.186) . Thus, the plastic zone size is less than the remaining ligament.
Based on the above, it is concluded that for the South Texas Project Accumulator Lines and attached RHR piping, the fatigue crack growth during service will not be significant. 25.s. ,cm. w 2se' " 7-3
i 7.2 References . l 7-1 Swamy, S. A., et. al., " Additional Information in Support of the - Elimination of Postulated Pipe Ruptures in the Pressurizer Surge Lines of South Texas Project Units 1 and 2" WCAP-11256, September 1986, (Westinghouse Proprietary Class 2). 7-2 Swamy, S. A., et. al., " Technical Bases for Eliminating Pressurizer Surge Line Ruptures as the Structural Design Basis for South Texas Project Units 1 and 2" WCAP-11256 Supplement 1, November 1986 (Westinghouse Proprietary Class 2). 7-3 NRC Letter to HL&P, N. P. Kadambi to J. H. Goldberg, July 10, 1986, Docket Nos. 50-498 and 50-499. I I 1 i an,em, arut ic 74
l SECTION 8.0 ASSESSMENT OF MARGINS 4 ( In the preceding sections, the leak rate calculations, fracture mechanics analyses and fatigue crack growth assessments were performed. Margins are discussed below. { In section 6.1, the critical flaw sizes using limit load methods for the above three different pipe sizes are calculated. Using the IWB-3640 approach, the ! critical flaw sizes for the three different pipe sizes are also as calculated in section 6.1. Those results are as follows: 12 inch pipe Critical flaw size using limit moment: [ Ja,c.e . Critical flaw size using IWB-3640: [ Ja,c.e 10 inch pipe Critical flaw size using limit moment: [ -Ja,c e
~
Critical flaw size using IWB-3640: [ ]a,c.e 8 inch pipe Critical flaw size using limit moment: [ )"' C d Critical flaw size using IWB-3640: [ Ja,c.e In section 6.3, the local crack stability analyses showed that flaw sizes of ( Ja,c.e for the 12" pipe, [ la,c.e for the 10" pipe, and [ Ja,c,e for the 8 inch pipe, were stable when subjected to normal plus SSE loads. Based on the above, the respective critical flaw sizes for normal plus SSE loads will exceed [ ]a,c.e,[ 3a,c.e, and [
)a,c.e, no. ,osinu o 8-1
In section 6.3 it is shown that reference flaws of [ Ja,c.e,[ , Ja.c,e, and [ la.c.e, yielding a leak rate of 10 gpm would be stable for the 12",10", and 8" lines respectively, when subjected to loads - equal to O (Normal + SSE). Thus the required margin of /2 on load has been demonstrated, along with the required margin of 10 in leak rate. From table 6-1 of section 6.3, it is seen that the J, ppg $ values are all less than the lower bound value of J Ie fI 3 II "II three pipe sizes considering the thermal aging effect discussed in section 4.0. In summary, relative to
- 1. Loads The leakage-size crack will not experience unstable crack extension even if very large loads of [2'(Normal + SSE) are applied.
~
- 2. Flaw Size
- a. A margin of at least 2 exists between the critical flaw and the flaw yielding a leak rate of 10 gpm.
- b. If limit load is used as the basit for critical flaw size, a larger margin for global stability would result.
- 3. Leak Rate A margin of 10 exists for all the reference flaws between calculated leak rate and the 1 gpm leak detection criterion of Regulatory Guide 1.45.
A summary comparison of criteria and analytical results is given in tabla 8-1. The criteria are seen to be met for all three different pipe sizes. 296s T3966 '08268710 g.g _ _ _ _ _ . _ _ _ _ _ . . . _ . _ )
TABLE 8-1 l COMPARISON OF RESULTS VS. CRITET:lA ' ' RESULT CRITERION
- 1. NUREG1051 Volumc 3 Met Section 5.2(h) - (Required margin of 2 demonstrated)
Margin on Flaw Size
- 2. NUREG1061 Volume 3 Met Section 5.2(i) - (Required margin of d demonstrated)
Margin on Load
- 3. NUREG 1061 Volume 3 Met Section 5.7 - (Margin of 10 on leak rate Margin on Leak Rate demonstrated)
- 4. NRC criteria on ellowable Het fatigue crack growth ( Ja.c.e
, (af < 60% wall thickness)
- 5. NRC criteria on allowable ~
Met 1 "'C fatigue crack growth (Plastic :cne size < remaining
~
ligament) 1 I ms.mm. arunc g_3
SECTION 9.0 CONCLUSIONS i This report justifies the elimination of accumulator line and attached RHR line pipe breaks as the structural design basis for the South Texas Project Ur.its 1 and 2 as follows:
- 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 (primary loop and the attached auxiliary lines) because of system design, testing, and operational considerations. {
- c. The effects of low and high cycle fatigue on the integrity of the accumulator line and attached RHR line piping are negligible.
- d. Ample margin exists between the leak rate of small stable flaws and
. the requirements of Reg. Guide 1.45. The STP system for detecting unidentified leaks in the RCS conforms to the requirements of Reg. <
Guide 1.45 (FSAR Section 5.2.5).
- e. Ample margin exists between the small stable flaw sizes of item d and the critical flaw.
- f. Amale margin exists in the material properties used to demor, strate I end-of-service life (relative to aging) stability of the critical flaws.
- g. With respect to stability of the reference flew, ample margin exists
. between the maximum postulated loeds and the plant specific faulted loads (i.e. Normal + SSE). i ms.wnne 93 l
1 The reference flaws will be stable throughout reactor-life because of the - ample margins.in d, e, f and g and will leak at a detectable rate which will assure a safe plant shutdown. -
' Based on the above. it is concluded that pipe breaks in the accumulator line and its attached RHR piping should not be considered in the structural design-basis of South Texas Project Units 1 and 2.
h 25456 03Hs/082M710
/
g.g
APPENDIX A LIMIT MDMENT I i a,;,e , t I i
~
ne...aaane A-1 l [ - - , - - - - - - - _ - _ _ _ - - - - _ _ _ - _ - - - _ - _ - - - - - - - - -
Q ->i: u a.c.e
=
FIGURE A-1 Pipe with a Thrwgh-Wall Crack in Bending A-2
- ,. ><>. c>.ee is
APPENDIX B 3 FATIGUE CRACK GROWTH CONSIDERATIONS L B.1.' Thermal Transient Stress Analysis The. thermal transient stress analysis was performed for a typical PWR plant to obtain the'through wall stress profiles for use in the fatigue crack growth analysis of Section B.2. The through wall stress distribution for each transient was calculated for i) the time' corresponding to the maximum inside surface stress.and, ii) the time corresponding to the minimum inside surface stress. These two stress profiles are called the maximum and minimum through well stress distribution, respectively for convenience. The constant stresses due to pressure, deadweight and thermal expansion (at normal operating temperature, 550*F) loadings were superimposed on the through wall cyclical stresses to.obtain the total maximum and minimum stress profile for each transient. Linear through wall stress distributions were calculated by conservative simplified methods for all minor transients.- More accurate nonlinear through wall stress distributions were developed for severe transients by [- Ja,c.e B.1.1 Critical Location for Faticue Crack Growth Analysis The accumulator line design thermal transients (Section B.1.2),1-D analysis data on accumulator line thermal transient stresses (based on ASME Section III NB3500 rules) and the geometry were reviewed to select the worst location for the fatigue crack growth analysis. ['
.Ja,c.e This location is selected as the worst location based on the following considerations:
i) the fatioue usage factor is highest. the stress due to thermal expansion is high.
~
ii) iii) the effect of discontinuity due to undercut at weld will tend to increase the cyclical thermal transient loads. me.-omsu o B-1
'd' , e iv)' ;the review of data shows.that-the 1-D thermal-transient stresses _in the . accumulator .line piping section are' generally higher near the [
.)a,c e ,
B.1.2 Design-Transients The transient con'ditions selected for this evaluation are based on conservative' estimates of the magnitude and-the frequency of the temperature fluctuations resulting from various operating conditions.in the plant. These are representative of the conditions which are considered-to occur during plant operation. The _ fatigue evaluation based on these transients provides. confidence that the component is appropriate for its application over the design life of the plant. All the normal operating _and upset' thermal transients, in accordance with design specification and the applicable system design criteria document (B-1), were considered for this evaluation. Out of these, only [. Ja c.e These transients were selected on the basis of the'following criteria: ,
+a,c e .
(B.1) i (B.2) where,
+a,.c e e
_. _i B.I.3 Simplified Stress Analysis The simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients. . I: . 43 4c 2599s.cci7st it g.g
i [. - Ja,c.e The inside' surface stress was calculated by- the. following. l equation which is similar.to the transient portion of ASME Section III NB3600, j Eq. 11: , n ; i Sg = [. Ja,c.e (B.3) where,
+a,c.e i
}
1 [ Ja,c.e The maximum and minimum inside surface stresses were searched from the S$ values calculated for each time step of the transient solution.
~
The outside surface stresses corresponding to maximum and minimum inside stresses were calculated by the following equations: Sol
- I 3 (B 7)+** '
{ a....o.a.0,2 5 I 3 (B 8)+ '*' w
- g.3
- =_ --_ ___ __ __ s where,
+a,c.e,-
6
-L _
e,c.e The material properties for the accumulator pipe [ ] and the RCL Ja,c.e The values of E and a, at the normal operating temperature, provide a conservative estimation of the through well thermal transient stresses as compared to room temperature properties. The following
- values were conservatively used, which represent the highest of the [
Ja c.e materials:
+a,c.e enum*
i The maximum and minimum linear through wall stress distribution for each thermal transient was obtained by [ 3a c.e The simplified analysis discussed in this section was performed for all minor thermal transients of [' Ja,c.e Nonlinear through wall stress profiles were developed for the remaining severe transients as explained in Section B.1.4. The inside and outside surface stresses calculated by simplified methods for the minor transients are shown in Table B-2. [
.)ace
,. This figure shows that the
~
simplified method provides more conservative crack growth. mv.-mune B-4 L
B.1.4 Nonlinear Stress Dist-ibution for Severe Transients s l' Ja,c.e As mentioned earlier, the accumulator line section near the [ la,c.e is the worst location for fatigue crack growth analysis. A schematic of the accumulator line geometry at this location, is shown in Figure B-2. l
)a,c.e B.1.5 OBE Loads The stresses due to OBE loads were neglected in the fatigue crack growth analysis since these loads are not expected to contribute significantly to crack growth due to small number of cycles.
me.-osu r ie B-5
B.l.6 Total Stress for Fatigue Crack Growth , The total through wall stress at a section was obtained by superimposing the i pressure load stresses and the stresses due to deadweight and thermal expansion (normal operating case) on the thermal transient stresses (of Table B-2 or the nonlinear stress distributions discussed in Section B.1.4). Thus, the total stress for fatigue crack growth at any point is given by the following equation: Total Thermal Stress Due Stress for Transient to Due to Fatigue = + DW + + Internal (B.9) Crack Growth Thermal Pressure Expansion The envelope thermal expansion, deadweight and pressure loads for calculating the total stresses of Equation B.9 are summarized in Table B-3. B.2 Faticue Crack Growth Analysis The fatigue crack growth analysis was performed to determine the effect of the design thermal transients in Table B-1. The analysis was performed for the critical cross section of the model which is identified in Figure B-2. A range of crack depths was postulated, and each was subjected to the transients in Table B-1. B.2.1 Analysis Procedure The fatigue crack growth analyses presented herein were conducted in the same menner as suggested by Section XI, Appendix A of the ASME Boiler and Pressure Vessel Code. The analysis procedure involves assuming an initial flew exists G
.m.-m m ic B-6
at some point and predicting the growth of that flaw due to an imposed series of stress transients. The growth of a crack per loading cycle is dependent on the range of applied stress intensity factor AKy , by the following E relation: h=CoAK" y (B.2.1) where "Co" and the exponent "n" are material properties, and AKy is defined later, in Equation (B.2.3). For inert environments these material properties are constants, but for some water environments they are dependent on the. level of mean stress present during the cycle. This can be accounted for by adjusting the value of "Co" and "n" by a function of the ratio of minimum to maximum stress for any given transient, as will be discussed later. Fatigue crack growth properties of stainless steel in a pressurized water environment have been used in the analysis. , The input required for a fatigue crack growth analysis is basically the information necessary to calculate the parameter AK), which depends on crack and structure geometry and the range of applied stresses in the area where the crack exists. Once AK y is calculated, the growth due to that particular cycle can be calculated by Equation (B.2.1). This increment of growth is then added to the original crack size, the AK) adjusted, and the analysis proceeds to the next transient. The procedure is continued in this manner until all the transients have been analyzed. 1 The crack tip stress intensity factors (K )y to be used in the crack growth analysic were calculated using an expression which applies for a semi-elliptic surface flaw in a cylindrical geometry [B-4). The stress intensity factor expression was taken frcm Reference B-4 and was calculated using the actual stress profiles at the critical section. The maximum and minimum stress profiles corresponding to each transient were input, and each profile was fit by a third order polynomial: e(x)=A+^1h^2({} 0 A(()3 3 (B.2.2) an.-ee"" " B-7
~
The stress. intensity factor K3(9)'was calculated at the deepest point of the crack using the following expression: ,
+a,c.e (B.2.3) 6I N
Calculation of the fatigue crack growth for each cycle was then carried out using the reference fatigue crack growth rate law determined from consideration of the available data for stainless steel in a pressurized water environment. This law allows for the effect of mean stress or R ratio (KImin/Elmax) n the growth rates. The reference crack growth law for stainless steel in a pressurized water environment was taken from a collection of data (B-5) since no code curve is available, and it is defined by the following equation: da HR , [ ) a,c.e (B.2.4) n .-mnue B-8
r
.. where Kdf
- IKImax)11-R)
K
' ! min R=
Imax-l h=crackgrowt'h rate in micro-inches / cycle B.2.2 Results Fatigue crack growth' analyses were carried out for the critical cross section. Analysis was completed for a range of postulated flaw sizes oriented circumferential1y, .and the results are presented in Table B-4. The postulated flaws are assumed to be six times as long as they are deep. Even for the largest postulated flaw of [t Ja.c.e the result'shows that the flaw growth through the wall will not occur during the 40 year design. life of the plant. For smaller flaws, the. 8 'C flaw growth is significantly lower. For example, a postulated [ 3 inch deep flaw will grow to [ Ja.c.e which is less than-[- Ja,c.e the wall thickness. These results also confirm operating plant experience. There have been no leaks observed in Westinghouse PWR accumulator lines. B.3 REFERENCES a,c.e B-1
-L B-2 ASME Section III, Division I-Appendices,1983 Edition, July 1,1983.
B-3 WECAN -- Westinghouse Electric Computer Analysis, User's Manut.1 -- Volumes 1, II, III and IV, Westinghouse Center, Pittsburgh, PA, Third Edition, 1982.
'e 2569s-Os174710 g,g
1-B-4 McGowan, J. J. and Raymund, M., " Stress Intensity Factor' Solutions.for , Internal Longitudinal Semi-Elliptical Surface Flaws in a Cylinder Under Arbitrary Loadings", Fracture Mec~nanic's ASTM STP 677, 1979, pp. 365-380, f: B-5 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, February 1979. 1 e l l c l l 4 l
. I l
me..mune B-10 i 1
TABLE B-1 o THERMAL TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH EVALUATION Trans. No. of No. Description Occurrences
+a,c.e 9 'D .
L. O n....o. ire, se B-11 I ___---___________-___j
N. ' M se - .. LD D + 2Wgg o- 1. I 5* :
- n. W MC w a=
2W -
$ >8 m
W W E e-N E
-Q Z~
m m Z
.-=
c. w N E w m B B LD D
-2 ~EW
< ~ CE:
E O >=- N W W t CO W mO .- W w= J ~ ED CD X Q: *="
< w QD vQ W #
Z r= 8 m W w M W W W W W E W E
>= A M -
>< W
<Q Z= l' m
Z em a
*U U
9 z - N. mo b s xz e a
< l cx 1 2 B-12 L___ ___ __ u ------
TABLE B-3 ENVELOPE NORMAL LOADS 3 CONDITION - a,c.e Normal Operating _.1 l d We
,* 7 ,
TABLE B-4 - ACCUMULATOR LINE FATIGUE CRACK GROWTH RESULTS Wall Thickness = [ ] +a,c.e INITIAL CRACK LENGTH AFTER YEAR CRACK LENGTH 10 20 30 40 (IN.)
+a.c.e 4
kne " w 9 e me.-oe nsno B-14
S. w. e.
- s.t.e W
4
~
Figure B-1 Comparison of Typical Maximum and Minimum Stress Profile Computed by Simplified [ ] +a,c.e am. tuu smen: B-15
0 l l
~ "
*a.c.e 2.
i .
-. +a,c.e Accumulator Pipe Figure B-2 Schematic of Accumulated Line At
[
;r + a,c.e a m . m s,t u n 'ic B 16
i 4 e a c.e d
+a,c e c
Figure B-3 [ ] and Minimum Stress Profile for Transient #10 zn suns uou, e B-17 L -- - _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ __-___________________________________________________________a
*a,c,c.
5
-)
\
1
+ +a.c,e Figste B-4 [ ] Maximum and Minimum Stress Profile ;
for Transient #11 nie.t>o. w e" " B-18 _g
a.c.e
~
4 f 1 J 4 0 2 r4 1 -
} - +a,:,e Figure B-5 [ ] P.eximum and Minimum Stress Profile for Transient #12
- w. no. cnu se g.19 u
. \
s.c . 1 L
/
s,
+a.t,e T
Figure B-6 f 3 Maximum and Minimum Stress Profile for Transient #14
- 3. :2,3. c23. it B-20
APPENDIX C MATERIALS SPECIFICATION AND FRACTURE TOUGHNESS l d PROPERTIES OF THE ACCUMULATOR TANK, N0ZZLE, AND SAFE END C.1 Materials Specification The accumulator nozzle is SA 350 LF-8 with an ultimate tensile strength of ( between 70 and 95 ksi and a 36 ksi yield strength. The tank material is SA j 264 containing SA 537 C1 1 (carbon steel)' u =70 ksi (min.), o y=50 ksi (min.), and SA 240-TP304 (stainless steel bonded to the carbon steel),
=75 ksi, o =30 ksi. The safe end is made of SA 312-TP304 (stainless u y steel pipe): o u= 5 ksi, o y=30 ksi.
C.2 Fracture Toughness The certified material test report (CMTR) for the nozzle providos the following test results: Potential Test Result RT NDT 4 Drop Weight No Break at 0*F (Both heats) -10* Charpy Test Energy: 55, 53, 52, (1st heat) at 60'F 53, 50, 51 (2nd heat), f t-lbs O' Lateral Expansions: 50, 48, 46 (1st heat) 48, 45, 46 (2nd heat), mils me, orno w c.1
Therefore the RT NDT f the nozzle is 0*F. The drop weight test for the tank material was not performed, however, Charpy V-Notch tests at 60*F were made ~ and the test results are as follows: Potential Test Heat Results RT NDT V Notch 1 Energy 53, 53, 52 ft-lb Test at lateral 60*F Expansion 48, 49, 49 mils l 2 Energy 51, 50, 51 ft-lb Lateral Expansion 49, 50, 50 mils 0*F I 3 Energy 50, 50, 51 ft-lb Lateral Expansion 49, 48, 50 mils 4 Energy 55, 53, 53 ft-lb i i Lateral j Expansion 49, 48, 49 mils The V-Notch test results of the tank material show very similar results with those of nozzle material. Thus RT NDT f the tank material is considered to be 0*F. The toughness of ferritic steel is taken from Appendix A, Section XI of the ASME Code as the lesser of Kyc = 33.2 + 2.806 Exp (0.02 (T-RTNOT + 100)) 4.1) or Klc = 200 ksi /in. (upper shelf toughness) . l l me,-wnno c.g
Substituting the operating temperature, 120* and RTNDT = 0'F in equation
. (C.1) the toughness of ferritic steel at 120*F is found to be 200 (si /In~. This 2
is equivalent to JIc = 1333 in-lb/in where J Ie is assumed to be given by J3c = KIc /E where E = 30 x 10 6 p,$, Tests of actual SA 350 LF-2 forging material have shown J Ic values of [ Ja,c.e in-lb/in2 at 50*F and [ la,c.e in-lb/in2 at 75'F. The Ja,c.e Since the tank and nozzle operate at 120*F with RTNDT was [ . an RT f 0*F, they operate at or near the upper shelf fracture toughness NDT temperature. Thus the actual J yc appropriate to the 120'F operating temperature is near the higher test result [ 3a,c.e , Thus the accumulator nozzle and the accumulator tank material exhibit a 2 greater toughness level as compared to the JIc = { ) in-lb/in used as the criterion in the leak-before-break evaluation. for the 8" diameter RHR line, the valve to pipe weld at the limiting location 8 is a field weld. The field welding was performed using the automatic gas tungsten arc process (GTAW). Justification for the selection of limiting toughness properties, used in the analysis, is provided in Appendix D. A sketch of the valve to pipe weld at limiting location 284 is shown in figure C-1. References C-1 HL&P Letter to NRC, M. R. Wisenburg to U.S. NRC, May 18, 1987, ST-HL-AE-2165. NN5
- M I7&71(
l i _ __ . . o
a,c.e- - 1 i i i i t 1 Figure C-1. Crcis-Section cf the Valve-te-Pipe Weld f or 8" diameter RER line. I 1 268%t '396s 7041217 it C-4
)
._. .-- - _ - - _ _ _ ___---_--____-____-___a
APPENDIX 0 l EFFECTS OF THERMAL AGING ON STRUCTURAL INTEGRITY OF WELDS AT THE LIMITING LOCATIONS D.1 Effects of Thermal Aging on Structural Integrity of Welds at the Limiting Locations The limiting weld location for the 12 ir ' line is location 104E as identified in figure 5-1. The limiting weld location for the 10 inch portion of the system is location 305 identified in figure 5-2. The limiting weld location for the 8 inch line is location 284 as identified in figure 5-2. All of these welds are field welds. The field welding was performed using the automatic gas tungsten are process (GTAW). The shop vernacular for GTAW is TIG. All the welding was performed to applicable QA procedures. Traditionally nuclear pipe welds were made by first performing root passes using GTAW followed by either submerged arc welding (SAW) for shop welds or shielded metal are welding (SMAW) for field and shop welds to completion. Thus, typical welding data have been obtained on SAW and SMAW. For exemple in table 3 of reference D-1, welds A, B, C, and G are [ Ja,c e while the remaining (D, E, and F) are [ Ja,c,e (reference D-2). Specifically in figure 2.4-8 weld B of reference D-3 is [ )**C while weld D is [ Ja,c,e . The general argument of reference D-3 th6t welds are not limiting compared to [ Ja,c.e when thermal aging is considered is based on the conclusions reached in reference D-1 and will not be repeated here. Interestingly, the minimum KCU for the thermally aged welds B Ja,c.e noted above was [ Ja,c.e da]/cm 2 or [ Ja c.e and E [ greater after 10,000 hrs of aging at 400*C (D-2) which exceeds the 2 [ Ja.c.e daJ/cm criterion of reference D-4. Roughnesses of GTAW have been evaluated by several investigators, one of the more recent being reference D-5. This reference considers the fracture toughness of several welds including a 30SL GTAW (i.e., TIG weld). Significantly the toughness of the GTAW approached the toughness of the 304 n "' * "" ' D-1
base material inv'estigated. The J Ic exceeded 3000 in-lb/in2 at 550*F with aTmat f 500. The Charpy V-Notch energy at 550*F exceeded the machine capacity (239 ft-lbs). At room temperature the Charpy V-Notch energy was 140
- ft-lbs which corresponds to a KCU energy of around 90 ft-lbs (24 da]/cm2 ),
This value is well above the KCU value at room temperature of around 7 da]/cm2 noted for the unaged welds [ 3a,c.e discussed in reference D-3. The ferrite content in GTAW is typically less than 10%. Thus it is not expected that thermal aging degradation will significantly. reduce the fracture toughness values. It is reasonable then to assume-that the GTAW is at least as good as [ Ja c.e having a J Ie f( ) in-lb/in 2and a T mat f[ ] after 40 years of service. References D-1 Slama, G. , Petrequin, P. , Masson, S. H., and Mager, T. R. , "Effect of Aging on Mechanical Properties of Austenitic Stainless Steel Castings and Welds," presented at SMIRT 7 Post Conference Seminar 6 - Assuring Structural Integrity of Steel ' Reactor Pressure Boundary Components, , August 29/30, 1963, Monterey, CA. D-2 PWS 3-4, Effect of Aging on the Mechanical Characteristics of Austenitic-Ferritic Welds, Internal ID No. E EM DC 0 260, FRAMATOME (Industrial Property). D-3 WCAP-10456, "The Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping for W NSSS," November 1983 (Westinghouse i Proprietary Class 2). O m e,-os n o w n.g
I h APPENDIX E
.. A
SUMMARY
EVALUATION OF POTENTIAL DEGRADATION DURING SERVICE E.1 Potential Degradation During Service In the Westinghouse PWR design.there has never been any service cracking identified in the accumulator lines or their attached RHR piping. Only one incident'of wall thinning has been. identified in RHR lines of Westinghouse PWR design. Sources of such degradation are mitigated by the design, l construction, inspection, and operation of the RHR lines. l Vibratory fatigue loads are monitored'for the 12-inch accumulator line during L 'the hot-functional testing of the plant and are well below the high cycle 1 l= fatigue allowables. The vibratory fatigue loads for the 8-inch and 10-inch
- pipe attached to the accumulator line should be less severe as compared to the.
12-inch pipe which is attached to the primary loop. Also, socket welds, which
. are susceptible to this type of loading, are not used in this piping.
- There is no mechanism for water hammer in the accumulator piping system and based on a review of references E-1 through E-4 only one incident of water hammer has been reported in a PWR RHR system. This incident was a result of
- incorrect valve line-up preceeding a pump start. The only damage sustained was to several pipe supports. It can therefore be concluded that water hammer in the RHR system is highly unlikely and not a viable mechanism for system degradation.
Wall thinning by erosion and erosion-corrosion effects will not occur in the accumulator line and its attached RHR line due to the low velocity, typicaliy less than 10 ft/see and the material, austenitic stainless steel, whien is highly resistant to these degradation mechanisms. Per NUREG-0691 [E-5), a
< study of pipe cracking in PWR piping, only two incidents of wall thinning in stainless steel pipe were reported. One incident was related to the RHR system. However, this occurred in the pump recirculation path which has nu..umno E-1
- w. _ - _ _ - _ _ _ _-___ __-
f1 4 higher flow velocity and is more susceptible to other contributing factors
-such as cavitation, than the RHR piping near the accumulator discharge line.
Therefore, wall thinning is not a significant concern in the portion of the system being addressed-in this evaluation. Flow stratificat' ion, where low flow conditions permit cold and hot water'to
. separate into distinct layers, can cause significant thermal fatigue loadings. This was an important issue in PWR feedwater piping where temperature differences of 300*F were not uncommon under certain operational conditions. Stratification is believed to be'important where low flow conditions and a temperature differential exist. This'is not an issue in the accumulator line or its attached RHR line, where typically there is no flow during normal plant operation. "uring RHR operation the flow causes sufficient mixing to eliminate stratification.
The normal operating temperature of the accumulator line and its attached RHR piping which is about 120*F, is well below the temperature which would cause any creep damage in stainless steel piping. Finally, pipe degradation or failure from indirect causes such as fires, missiles, and component support failure is prevented by designing, l fabricating, and inspecting reactor compartments, components, and supports, to NRC criteria that reduce to a low probability the likelihood of the events unacceptably impacting safety related components. STP complies with the criteria in Standard Review Plans 3.4.1, 3.5.1.2, 3.9.3, 3.9.6, and 9.5.1 as discussed in these same sections in the FSAR. References E-1 Utter, R. A., Et. al., " Evaluation of Water Hammer Events in Light Water Reactor Plants," NUREG/CR-2781, published July 1982. E-2 " Report of the U.S. Nuclear Regulatory Commission Piping Review Committee, Evaluation of Other Dynamic Loads and Load Combinations," NUREG-1061 Volume 4, Published December 1984. nu,-omenc E-2 L-_-__-________________-
E-3 Chapman, R. L., et. al., " Compilation of Data Concerning Known and Suspected Water Hammer Events in Nuclear Power Plants, CY 1969-May 1981," NUREG/CR-2059, Published April 1982. E-4 " Evaluation of Water Hammer Occurrence in Nuclear Power Plants," NUREG-0929 Revision 1, Published March 1984. E-5 " Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors," NUREG-0691, Published September 1980. 8 O nee,.m anc E-3 L
APPENDIX F a BENCHMARK FOR LEAK RATE PREDICTION
. MODEL USED FOR 8 INCH AND 10 INCH IINES F.1 Benchmark for Leak Rate Prediction Model L_ad for 8 Inch and 10 Inch Lines The flow of fluid through postulated flaws in the 8" and 10" RHR piping at 120*F would be a single phase flow. The velocity of flow under the single phase flow condition is given by (reference F-1):
V = (2 gap /Kp)1/2 (F.1) where g is the acceleration due to gravity 32.2 ft/sec 2
- AP is the pressure difference between inlet and exit of a crack.
p is the density of water at atmospheric conditions. k is the friction factor including loss of head due to sudden expansion, contraction and loss of head in the flow path. Mathematically, k is given as k = ahf Cont + ahfExp + f(L/D h) (F.2) where D h is the hydraulic diameter (see references F-1 and F-2.) If the inlet and outlet of the crack are connected with very large environment such as Ain ^1*0 and Aout /A240, then ahfcontraction and Ahf expansion in
- equation (F.2) are 0.5 and 1.0 respectively (F-2). Therefore equation (F.2) can be rewritten as l
l 1 me..c.'m 'o F-1
h k = 1.5 + f (CD/Dh) (F.3) , The friction factor f is obtained from reference [F-3) by using the hydraulic ' diameter, Dh , (Dh = 4A/2 (Cl + Cw)) and crack depth, CD (n te CD =L). l The velocity given in eq. (F.1) is found by using the k value given in equation (F.3). The mass flow rate G is obtained by using G = pVA (F.4) Comparison With Experimental Results An experiment was performed to obtain the leak rate through a crack in a stainless steel pipe. Details of the experiment are provided in reference F-5. A summary of the experimental data was also provided in reference F-4. The crack geometry as taken from reference F-4 is defined below: Crack length, C) = 63.6mm = 2.50 in. , Crack depth, CD = 7.3mm = 0.2874 in. (CD * ') Crack width, Cy = 0.46mm = 0.01811 in. Stagnation pressure, P, = 162 Bar = 2381.4 psig Stagnation temperature, To= 332*C = 630*F Hydraulic diameter, Dh = 0.03596 in. CD /Db = 8.0 Since CD /Dh < 12, the flow behavior will be single phase. The roughness e = 300 x 10 -6 in. (reference F-6) c/Dh = 0.00834 f = 0.039 at Reynold No = 2 x 10 4 By using eq. (F.3)
- 1 k = 1.811 ,
i l
"*d'"*"'
F-2
i 1
'Therefore per equation (F.1)
V = 442.06 ft/sec 4- j i' and mass flow' rate G per equation F-4 is: G = p Q = pVA = 3.95 Kg/sec: calculated value .J l Measured value as taken from reference 4 is: G = 4.09 Kg/see
Conclusion:
calculated and measured values show good agreement. References F-1. D. C. Dillio, " Thermal Engineering," International Text Book Company, Second Edition, 1969. F-2 1. E. Idelchick, " Hand Book of Hydraulic Resistance" - Coefficients of 8 Local Resistance and Friction - Translated from Russian - Published for U. S. Atomic Energy Commission and National Science Foundation, Washington, D.C. - AEC-TV-6630. F-3 L. F. Moody, " Friction-Factor Correlation," Trans. ASME, Vol. 66, No. 8, or Ref. [1].
+
F-4 WCAP-11256 (Supplement 1), " Technical Bases for Eliminating Pressurizer Surge Line Ruptures as the Structural Design Basis for South Texas Project Units-1 and 2," Westinghouse Proprietary Class 2, Nov. 1986, P. 2-13.
< F-5 Chouard, P., Richard, P., "Taux de Fuite dans les Fissures Traversantes:
I. . Recherche d'un Modele Analytique a Partir des Resultats Experimentaux,"
~CEA Laboratoire de Mecanique Appliquee, Note Technique DRE/STRE/LMA85/671.
i m.,4. , ,,, , e r-3
F-6 WCAP-9558, " Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Pestulated Circumferential Through Wall Crack," Westinghouse Proprietary Class 2, May, 1982. 8 9 me,-ose se p.4
k
. i
-APPENDIX G e' ADDITIONAL INFORMATION ON THE LEAK DETECTION SYSTEM
-l !
G-1 ' Additional-Information on the Leak Detection System
. All LBB candidate lines at STP are locateiinside containment. The STP' leakage detection criterion' includes a detected unidentified leak rate of 1.0-
.gpm and, in accordance with NUREG-1061, Volume 3, a margin of 10 was applied to .the leak rate to define the accumulator line'(and attached RHR line)
. leakage size flaw used in the stability analysis. The basis for the 1.0 gpm
-leak rate.is the presence (inside containment) of diverse and redundant leakage detection systems to measure containment noble gas radioactivity, airborne particulate radioactivity, and containment sump level' and flow rate.
These systems are designed to alarm in the control room at a setpoint equivalent of less than or equal to 1 gpm. Indication of containment humidity is also provided in the control room. These methods are in compliance with Regulatory Guide 1.45 as discussed in FSAR Subsection 5.2.5. In addition to the above control room alarms and indications, technical specification 4.4.6.2.1 requires monitoring of containment gaseous or particulate radioactivity and normal sump inventory and discharge at least once per 12 hours. This section of the technical specification also requires performance of a reactor coolant system inventory balance at least once per 72 hours. Experience at other Westinghouse plants indicates a normal background unidentified average leakage rate of between 0.1 gpm and 0.3 gpm, and it has also been demonstrated with pressurized pipe tests that leak rates above 0.1
.gpm can be readily detected visually. The undefined leakage rate at STP is expected to be similar to other plants. Experience at similar plants and the results of these tests indicate that a 1.0 gpm leak rate can be reliably .,
detected and located during plant operation. m wouino G-1
For the low pressure accumulator line piping located from the second check valve off the RCL and continuing to the accumulator tank, additional leak detection is available in the form of accumulator tank level and pressure indication and alarms in the control room. For each accumulator tank there are provided redundant level and pressure indication which alarm on both high and low pressure and high and low level in the control room. The tank range from high level to low level is 300 gallons. As required by technical specification 4.5.1.1 a check to demonstrate operability tust be completed every 12 hours. Tnis ensures that over a 12-hour period a leak rate of 0.42 gpm can be identified and isolated to a specific accumulator piping system. A detailed description of these accumulator monitoring systems is discussed in FSAR Subsection 6.3.5. G-2 Operator Action Detected leaks will be repaired within the system limiting conditions for operation established in either technical specifications or administrative . procedures. When leckage is detected in reactor coolant pressure boundary piping, technical specification 3.4.6.2 requires that the plant be in hot standby within six hours and in cold shutdown with the next thirty hours. Repair would be required before restart. G-2
1 i APPENDIX H DEFINITION OF THE ACCUMULATOR LINE AND ITS ATTACHED RHR PIPING AS CONSIDERED IN THE LBB ANALYSIS H.1 Definition of the Accumulator Line and the Attached RHR Piping as Considered in the LBB Analysis The loads for the entire span of the 12-inch diameter pipe from the primary loop junction which is essentially an anchor point to the accumulator tank nozzle junction (anchor) were reviewed to establish the critical location (see figure H-1). The maximum stress location based on normal and SSE loads occurs in the section betweer the primary loop nozzle and the first valve. The temperature and pressure in this section are 560*F and 2304 psi respectively. The yield strength of the material in this section at operating temperature is significantly lower than the yield strength in the rest of the piping (where the temperature is 120*F), The bousdary between Class 1 and Class 2 piping and the location where the piping schedule changes from schedule 140 to e schedule 40 are identified in figure H-1. With reference to figures 5-1 through 5-3 it should be noted that the branch piping is the Residual Heat Removal (RHR) piping. There is no low energy portion in the 12-inch piping between anchor to anchor as shown in figure H-1, since the pressure and temperature exceed 275 psi and 200 F respectively. In figures 5-1 through 5-3 the 10-inch branch piping attached to the accumula-ter piping is the Residual Heat Removal (RHP) piping. The 10-inch pipe is connected to the 8-inch pipe using a reducer (see figure H-1). The 8-inch portion of the RHR piping af ter the first check valve is not high energy. The l 8-inch piping terminates at the heat exchanger. The scope of piping stress analysis included the junction of the 12-inch accumulator lines (anchor point) to the anchor point on the 8-inch piping. The anchor point is at the heat exchanger nozzle except for the loop 1 RHR line for which an anchor exists f between the first check valve and the the heat exchanger, ms -ww 1e 91 J
References
=
H-1 HL&P Letter to NRC M. R. Wisenburg to V. S. Nuclear Regulatory Commission,
~
May 18, 1987, ST-HL-AE-2165. h 2t%:-08176710 H.2
s 11 _y T 5.- E I
%'E v
5O ; I
~
a
$ R.
W e= E-et 2: HE p: EE O gr* i R7 r
, 4 a
~
4=f = m C
'o 2 e
y-_ n, r ~ q a%.
,r m
7 Em E "o" eo 25 [ Ns
$2 32 32 2 .g '
" " .. .g 3 i EI
- a. .- ."* i e s *t h \
1 E. E EE
+l E5 g*
' N 4r Figure H-1 South Texas Project 10-Inch and 8-Inch RHR li"'
Attached to the Accumulator Line H "4}}