ML20207D743
| ML20207D743 | |
| Person / Time | |
|---|---|
| Site: | South Texas  |
| Issue date: | 06/30/1988 |
| From: | Coslow B, Swamy S, Witt F WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19292J233 | List: |
| References | |
| WCAP-11774, NUDOCS 8808150461 | |
| Download: ML20207D743 (85) | |
Text
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WESTINGHOUSE CLASS 3 WCAP 11774
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I TECHNICAL BASES FOR ELIMINATING RESIDUAL l
HEAT REMOVAL (RHR) LINE RUPTURE FROM THE i
STRUCTURAL DESIGN BASIS FOR SOUTH TEXAS l
PROJECT UNITS 1 AND 2 June 1968 l
l B. J. Coslow I
S. A. Swamy 1
F. J. Witt I
Y. S. Lee Verified by:
(d C. _f18 GC.Schmertz
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Approved by:
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- 5. 5. Pilusamy,' Manafer Structural Materials Engineering Work Performed Under Shop Order TREJ-950 WESTINGHOUSE ELECTRIC CORPORATION l
l Power Systems Division l
P.O. Box 2728 Pittsourgh, Pennsylvania 15230-2728 l
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TABLE OF CONTENTS Section Title Page
1.0 INTRODUCTION
1-1
1.1 Background
1-1 1.2 Scope and Objective 1-1 1.3 References 1-4 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 Wechanism 2-2 2.4 References 2-3 3.0 OPERATION AND STABILITY OF THE RHR LINE AND THE REACTOR COOLANT SYSTEM 3-1 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 Assessment of Pipe Degrad?. tion or Failure from Indirect Causas 3-6 3.6 References 3-6 4.0 MATERIAL CHARACTERIZATION 4-1 4.1 Pipe, Fittings and Weld Waterials 4-1 4.2 Tensile Properties 4-1 4.3 Fracture Toughness Properties 4-3 4.4 References 4-3 1
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 iii
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N TABLEOFCONTENTS(Cont'd.)
Section Title Page 5.3 Sumary of Loads, Geometry and Materials 5-2 5.4 Governing Location 5-3 5.5 References 5-3 6.0 FRACTURE MECHANICS EVALUATION 6-1 6.1 Global Failure Mechanism 6-1 6.2 Leak Rate Predictions 6-2 6.2.1 General Considerations 6-2 6.2.2 Calculation Method 6-3 6.2.3 Leak Rate Calculations 6-4 6.2.4. Leak Detection Capability, Administration 6-4 Procedures and Technical Specification Requirements 6.3 Stability Evaluation Using the "Z" Factor 6-5 Approach 6.4 Local Stability Analysis 6-7 6.5 References 6-7
7.0 ASSESSMENT
OF FATIGUE CRACK GROWTH 7-1 7.1 Acceptability of Fatigue Crack Growth 7-2 7.2 References 7-3
8.0 ASSESSMENT
OF MARGINS 8-1
9.0 CONCLUSION
S 9-1 APPENDIX A Limit Moment A-1 l
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TABLE OF CONTENTS (Cont'd.)
Section Title Page APPENDIX B Fatigue Crack Growth Considerations B-1 B.1 Thermal Transient Stress Analysis B-1 B.1.1 Critical Location for Fatigue Crack Growth Analysis 8-1 B.1.2 Design Transients B-2 B.1.3 Simplified Stress Analysis B-2 B.1.4 OBE Loads B-4 B.1.5 Total Stress for Fatigue Crack Growth B-4 B.2 Fatigue Crack Growth Analysis B-5 B.2.1 Analysis Procedure B-5 B.2.2 Results B-8 B.3 References B-8 e
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y LIST OF TABLES Table No.
Title Page 4-1 Mechanical Properties of STP Unit-1, Loop-1 RHR Line 4-5 at 70'F as Determined by Testing 4-2 Mechanical Properties of STP Unit-1, Loop-2 RHR Line 4-6 at 70*F as Determined by Testing 4-3 Mechanical Properties of STP Unit-1, Loop-3 RHR Line 4-7 at 70'F as Determined by Testing 4-4 Mechanical Properties of STP Unit-2, Loop-1 RHR Line 4-8 at 70*F as Determined by Testing 4-5 Mechanical Properties of STP Unit-2, Loop-2 RHR Line 4-9 at 70*F as Determined by Testing 4-6 Mechanical Froperties of STP Unit-2, Loop-3 RHR Line 4-10 at 70'F as Determined by Testing 4-7 Typical Tensile Properties of SA376 TP316, SA351 CF8A 4-11 and Welds of Such Material for the Primary Loop 4-8 Fracture Toughness Properties Typical of RHR Line 4-12 5-1 Summary of Envelope Loads 5-4 i
5-2 Loading Components at Governing Locations 5-5 i
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8-1 Comparison of Results vs. Criteria 8-3 vii
N LIST OF TABLES (Cont'd.)
Table No.
Title Pm B-1 Thermal Transients Considered for Fatigue Crack B-9 Growth Evaluation B-2 Transient Stresses for RHR Line B-10 B-3 Envelope Normal Loads B-11 i
B-4 RHR Line Fatigue Crack Growth Results B-12 l
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LIST OF FIGURES
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F'1ure Titli Page 1-1 South Texas RHR Line - Scope of Evaluation 1-5 2-1 Schematic of Generalized Load-Deformation Behavior 2-4 4-1 True Stress True Strain Curve for SA376 TP316 4-13 Stainless Steel at 623'F (minimum properties) 4-2 True Stress True Strain Curve for SA376 TP316 4-14 Stainless Steel at 623*F (average properties) 4-3 J Versus T Curves for SMAW Welds 4-15 5-1 Schematic Layout of RHR Lines Loop 1 (Units 1 & 2) 5-6 5-2 Schematic Layout of RHR Lines Loop 2 (Units 1 & 2) 5-7 5-3 Schematic Layout of RHR Lines Loop 3 (Units 1 & 2) 5-8 6-1
[
Ja,c.e Stress Distribution 6-9 6-2 Analytical Predictions of Critical Flow Rates of 6-10 Steam-Water Mixtures 6-3
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Ja,c.e Pressure Ratio as a 6-11 Function of L/0 6-4 Idealized Pressure Drop Profile through a Postulated 6-12 Crack
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LIST OF FIGURES (Cont'd.)
Figure Title Page 6-5 Loads Acting on the Pipe Model at the Governing 6-13 Location 6-6 Critical Flaw Size Prediction for the Base Metal 6-14 Using Limit load Approach l
6-7 Z-Factor Calculation for SMAW Weld to Demonstrate 6-15 l
Margin on Flaw Size 6-8 E Z-Factor Calculations for SMAW Weld to 6-16 Demonstrate Margin on Flaw Size l
l A-1 Pipe with a Through Wall Crack in Bending A-2 I
l B-1 Schematic of RHR Line at [
B-13 ya,c.e X
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l SECTION 1.0 INTRODUCTION
1.1 Background
The current structural design basis for the RHR line requires postulating non-mechanistic circumferential and longitudinal pipe breaks.
This results in additional plant hardware (e.g. pipe whip restraints and jet shields) which would mitigate the dynainic consequences of the pipe breaks.
It is, therefore, highly desirable to be realistic in the postulation of pipe breaks for these lines, and thereby eliminate the need for some of the plant hardware.
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 circumfarential type break will not occur. The evaluations considering circumferentially oriented flaws envelop longitudinal cases, as discussed in section 1.2.
As-built loads are used in the evaluation.
1.2 Scope and Objective The purpose of this investigation is to demonstrate leak-before-break for the high energy portion of the RHR line. The piping covered by this evaluation is shown in figure 1-1, and includes the high energy piping from the primary loop junction up to the first isolation valve. Down stream of the isolation valve, the piping is classified as moderate energy piping. Schematic drawings of the piping system are shown in section 5.0.
The recommendations and criteria 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.
1-1
'N 3)
Postulate a surface flaw at the governing location with the least favorable combination of stress and material properties. Determine fatigue crack growth. Show that a through-wall crack will not result.
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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
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detection with margin using the installed leak detection equipment f
when the pipe is subjected to normal operating loads. A margin of 10 is demonstrated between the calculated leak rate and the leak dotection capability. The associated flaw is called the leakage size flaw.
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 and 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 of at least I.4 on applied load for the leakage size flaw.
The flaw stability criteria used in this analysis address both the global and local stability for a postulated through-wall circumferential flaw.
The global analysis is carried out using the (
la,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 (1-1).
1-2
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o The local stability analysis is carried out using elastic plastic fracture mechanic analysis procedures.
In this report the EPRI elastic plastic fracture handbook (1-2) method was used for the local stability analysis.
The leak rate is calculated for the as-built normal operating condition. The leak rate prediction model used in this evaluation is an [
]a,c.e The crack opening area required for calculating the leak rates is obtained by subjecting the postulated through-wall flaw to normal operating loads (1-3). Surface roughness is accounted for in determinirg 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 [
]a,c.e and no flaws are actually anticipated.
The prediction methods for failure in elbows are virtually the same as those for
[
)a,c.e Therefore, the probability of any longitudinal flaw existing in the RHR line is much smaller when compared with the circumferential direction.
Based on j
the above, it is judged that circumferential flaws are more limiting than longitudinal flaws in elbows and throughout' the system.
Several computer codes are used in the evaluations. The main-frame computer programs are und6r Configuration Control which has requirements conforming to Standard Review Plan 3.9.1.
The fracture mechanics calculations are independently verified. The computer codes used in this evaluation have been validated (benchmarked)asdescribedinreferences(1-5)and(1-6).
1-3
1 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 Kumar, V., German, M. D..snd Shih, C. P. "An Engineering Approach for Elastic-Plastic Fracture Analysis," EPRI Report NP-1931, Project 1237-1, Electric Power Research Institute, July 1981.
1-3 NUREG/CR-3464, 1983, "The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulated Circumferential Through Wall Cracks."
l 1-4 Begley, J.A., et. al., "Crack Propagation Investigation Related to the leak-Before-Break Concept for LMFBR Piping" in Proceedings, Conference on l
Elastic Plastic Fracture, Institution of Hechanical Engineers, Londei l
1978.
1-5 Swamy, S. A., et al., "Additional Information in Support of the Elimination of Postulated Pios Ruptures in the Pressurizer Surge Lines of South Texas Project Units 1 and 2" WCAP-11256, September 1986, Westinghouse Proprietary Class 2.
1-6 Swamy, S. A. et al., "Techaical 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
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Proprietary Class 2.
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~0 12" SCH 140 ISOLATION VALVE
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s IP 1P Moderate Energy Piping '
I Figure 1-1 South Texas RHR Line Sccce of Evaluation 1-5 m
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SECTION 2.0 FAILURE CRI!ZRIA FOR FLAWED PIPES 2.1 General Considerations Active research is being carried out in industry and universities as well as other research organizations to establish fracture criteria for duccile materials. Criteria being investigated include those based on s'-integral initiation toughness, equivalent ene",, crack opening displacement, cract opening stretch, crack opening angio, 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 ari to make calculations using the available tools was used in selecting the criteria for this investigation.
The ultimate objective is to show that ths RHR line containing a conservatively assumed circumferential through-wall flaw is stable under the worst combination of postulated 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 gicbal failure will be governed by plastic coll:pse.
Extensive literature is available on tnis subject. A Pressure Vessel Research Committee 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 analycis.
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, thn plastic load is defined to be the peak load reached in a generalized load versus displacement plot and corresponds to the point of instability.
2-1
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A simplified version of this criterion, namely, not 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:
Wa < Wp (2-1) where Wa = applied generalized load Wp = calculated generalized plastic load Wp represents the load carrying ca,nacity 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 preperties as discussed in Section 6.1 2.3 Local Failure Mechanism The local mechanism of failure is prime.rily 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.
l The stability will be assumed if the crack does not initiate at all.
It has been demonstrated that the initiation toughness, measured in terms of JIc l
from a J-integrai resistance curve, is a material parameter defining the crack initiation.
If, for a given load, the calculated J-integral value is shown to l
be less than J f the material, then the crack will not initiate.
Ic l
If the initiation criterion is not met, one can calculate the tearing modulus i
as defined by the following relation:
dJ E T,pp = g, (2-2) l l
2-2
/
l o
where T,pp applied tearing modulus
=
modulus of elasticity E
=
of flow stress = (o +
u)/2
=
y crack length a
=
yield and ultimate strength of the material
=
c,y u
respectively.
Stability then is said to exist when ductile tearing occurs if T,pp is less than Tmat, the experimentally determined tearing modulus. Hence, the local crack stability is established by the two-step criteria:
J<Jge, or (2-3)
T,pp < Tmat, if J 3 J (2-4) ge 2.4 References 2-1 J.D. Landes, et al., Editors, Elastic-Plastic Fracture, STP-658, ASTM, Philadelphia, PA 19109, November 1977.
2-2 J. C. Gardeen, "A Critical Evaluat'on 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 Circumferential Cracks, EPRI-NP-192, September 1976.
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- p W,=ft.ASTIC LOAD I
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GENERALI2ED DISPLACEMENT I
l Figure 2 1 Schematic of Generalized Load-Deformation Behavicr f
2-4
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1 SECTION 3.0 OPERATION AND STABILITY OF THE RHR l!NE AhD THE REACTOR COOLANT SYSTEM i
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3.1 Stress Corrosion Cracking 1
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The Westinghouse reactor coolant system primary loop and connecting Class 1 lines have an operating history that demonstrates the inherent stability I
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 17 years of operation and 15 other plants each l
with over 12 years of operation.
l In 1978, the United StatGs Nuclear Regulatory Commission (USNRC) formed the l
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 l
of the potential for stress corrosion cracking in Pressurized Water Reactors l
l (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 st-ted:
"The PCSG has determined that the potential for stress-corrosion cracking in PWR primry 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
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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 whan the coolant is exposed to the air I
and during the subsequent startup are conditions even marginally capable 3-1
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1 of producing stress-corrosion 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 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.
As stated above, for the Westinghouse plants there is no history of stress corrosion cracking failure in the reactor coolant loop or connecting Class 1 piping. The discussion below further qualifies the PCSG's findings.
For stress corrosion cracking (SCC) to occur in piping, the following three conditions must exist simultaneously: high tensile stresses, susceptible material, and a corrosive environment. Since some residual stresses and some degree of material susceptibility exist in any stainless steel piping, the potential for stress corrosion is minimized by properly selecting a material
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imune to SCC as well as preventing the occurrence of a corrosive environment. The material specifications consider compatibility with the system's operating environment (both internal and external) as well as other material in the system, applicable ASME Code rules, fracture toughness, welding, fabrication, and processing.
The elements of a water environment known to increase the susceptibility of austenitic stainlees steel to strers 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.
3-2
p 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 wa.or 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 concentrations of chlorides and fluorides within the specified limits.
Thus during plant operation, the likelihood of stress corrosion cracking is minimized.
3.2 Water Hammer Overall, there is a low potential for water hammer in the RCS and connecting RHR lines since they are designed and operated to preclude the voiding condition in normally filled lines. The RCS and connecting RHR 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 openir.g are considered in the system design. Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlled. Temperature during normal operation is maintained within a narrow range by control rod position; pressure is controlled by pressurizer heaters and pressurizer spray also within a narrow range for steady-state conditions.
The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics are controlled in the design process. Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the flow and vibration characteristics of the system and connecting RHR lines. Preoperational testing and operating 3-3
X experience have verified the Westinghouse approach. The operating transients of the RCS prirr.ary piping and connected RHR lines are such that no significant water hammer can occur.
4 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 and Appendix B 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 vibrations during operation. During operation, an alarm signals the exceedance of the RC pump shaft vibration limits. Field measurements have l
been made on the reactor coolant loop piping of a number of plants during hot functional testing. Stresses in the elbow below the RC pump have been found to be very small, between 2 and 3 ksi at the highest. When translated to the connecting RHR lines, these stresses are even lower, well below the fatigue endurance limit for the RHR line material and would result in an applied stress intensity factor below the threshold for fatigue crack growth.
3.4 Potential Degradation During Service In the Westinghouse PWR design there has never been any service cracking identified in the RHR piping. Only one incident of wall thinning has been identified in RHR lines of Westinghouse PWR design. However, this is of no concern in the present application as dascribed later in this section.
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Sources of such degradation are mitigated by the design, construction, inspection, and operation of the RHR lines.
Based on a review of references 3-3 through 3-6 only one incident of water hammer has been reported in a PWR RHR system.
This incident was a result of incorrect valve line up preceding a pump start. The only damage sustained was l
l 3-4 l
i to several pipe supports. Therefore it is concluded that water hammer in the RHR system is unlikely to affect piping integrity or to cause pipe system degradation.
Wall thinning by erosion and erosion-corrosion effects will not occur in the RHR line due to the low velocity, typically less than 10 ft/see and the material, austenitic stainless steel, which is highly resistant to these degradation mechanisms. Per NUREG-0691 (3-2), a study of -Ne 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 higher flow velocity and is more susceptible to other contributing factors such as cavitation, than the RHR piping near the primary loop. Therefore, wall thinning is not a significant concern in the portion of the system being addressed in this evaluation.
Flow stratification, 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 RHR line, whnre typically there is no flow during normal plant operation.
During RHR operation the flow causes sufficient mixing to eliminate stratification.
On December 9,1987, while Alabama Power and Light Company's Farley Nuclea.
Power Station Unit 2 was at 33% power (during a restart following a refueling cutage) a total unidentified RCS leak of 0.7 gpm was detected (3-7).
The source of the leak was identified as a through-wall crack in a weld joint between an elbow and a horizontal pipe section in the 6 inch safety injection (SI) line attached to the cold leg of reactor coolant loop B.
In order to identify the mechanism which caused the crack, instrumentation was placed on the replaced piping and on a reference line to monitor temperatures and accelerations during heatup and operation. The results of this study indicated the existence of thermal cycling and stratification in the vicinity 3-5
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of the cracked weld.
The observed thermal cycling and stratification, due to the leakage past the single isolation valve between the charging system and the safety injection system, have been postulated to have caused the through-wall crack of the 6 inch SI line.
This is the first such reported instance of this mechanism resulting in a through-wall crack on Class 1 pipe.
At South Texas Projects Units 1 and 2, the RHR system configuration does not include a potential leakage path. Therefore, thermal stratification is not expected to occur in the horizontal sections of the RHR lines.
The maximum normal operating temperature of the RHR piping is about 623*F.
This is well below the temperature which would cause any creep damage in stainless steel piping.
3.5 Assessment of Pipe Degradation or Failure from Indirect Ceusa Pipe degradation or failure from indirect causes such as fires, missiles, and component support failure is prevented by designing, 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. South Texas Project 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.
It can be concluded by review of these sections of the South Texas FSAR that the required measures are taken to preclude the degradation or failure from outside sources of piping in the plant and that the methods utilized are consistent with those given in the Standard Review Plan.
3.6 References 3-1 Investigation and Evaluation of Stress-Corrosion Cracking in Piping of Light Water Reactor Plants, NUREG-0531, U.S. Nuclear Regulatory Commission, February 1979.
e 3-6
3-2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980.
3-3 Utter, R. A., et. al., "Evaluation of Water Hammer Events in Light Water Reactor Plants," NUREG/CR-2781, published July 1982.
3-4 "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.
3-5 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.
3-6 "Evaluation of Water Hammer Occurrence in Nuclear Power Plants,"
NUREG-0929 Revision 1, Published March 1984.
3-7 NRC Bulletin No. 88-08, Thermal Stresses in Piping Connected to Reactor Coolant Systems, Nuclear Regulatory Commission, June 22, 1988.
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SECTION 4.0 MATERIAL CHARACTERIZATION 4.1 Pipe, Fittings and Weld Waterials The pipe material of the RHR line is SA376-TP316, a wrought product form of the type used for the primary loop piping of several PWR plants. The RHR line is connected to the primary loop nozzle.
The other end of the RHR line is connected to an isolation valve.
The piping layout includes fittings such as elbows. These elbows are SA403-WP316 steel which is wrought and formed pipe of SA182-TP316 material. The weld wire used in the shop fabrication is generally of low carbon 316L. The field weld used 308L weld wire. The welding processes used are gas tungsten arc (GTAW), submerged arc (SAW) and shieldedmetalarc(SMAW).
In the following sections the tensile and fracture toughness properties of these materials are presented and criteria for use in the leak-before-break analyses are defined.
4.2 Tensile Properties The material certifications provided by reference 4-5 for the RHR lines of the Units 1 and 2 were used to establish the tensile properties for the piping, fittings and welds. These properties are given in tables 4-1 through 4-6.
The properties in these tables are at room temperature.
In the leak-before-break evaluation presented in this report, the minimum properties at operating temperature (623*F) are used for the flaw stability evaluation and average properties are used for the leak rate predictions. The viability of using such properties for the RHR line is presented below.
1 As noted in tables 4-1 through 4-6, the specific room temperature properties of the RHR line heats compare favorably with the properties of similar material of the primary loops (see table 4-7).
[
ja.c.e 4-1
l
(
ja.c.e In brief, the following material properties were used in the analyses set forth in this report.
Minimum Properties for Flaw Stability Analysis (623*F)
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a,c.e Average Properties for Leak Rate Calculations (623*F)
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a,c.e 9
e 4-2 j
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4.3 Fracture Toughness Properties Series of fracture toughness tests on SA376-TP316 pipe material and E308 welds are reported in references 4-2 and 4-3.
Data from these tests at 600*F are summarized in table 4-8.
As seen from this table, the lowest J value for yc 2
the weld materials observed from the J-R curves was [
]a,c.e in-lb/in.
The T
corresponding to this value of J, was over (
Ja c.e The sample mat y
].
The la,c% of [
yielding the lowest slope of the J-R curve had a T t
2 in-lb/in. While the data was found to be around [
corresponding Jge on welds cited above show superior toughness properties, they were not used in the leak-before-break evaluation.
Instead, lower bound toughness data for the SMAW welds (noting that the weld at the governing location was GTAW) were used as follows:
The data on SMAW welds based on testing IT specimens presented in reference 4-4 can be considered as lower bound toughnoss data. The resulting test data are as follows:
2 959 in-lb/in J
=
gc 2
J,,,-
3000 in-lb/in T,,g = 140 (corresponding to J,,x)
The J versus T for the data is plotted in figure 4-3.
The nozzles connecting the RHR Lines with the primary loop are made of forged stainless steel. Forged stainless steel is considered not susceptible to thermal aging for applications at hand.
i 4.4 References 4-1 Nuclear Systems Materials Handbook, Part I - Structural Materials, Group 1 - High Alloy Steels, Section 4. ERDA Report TID 26666 November 1975 Revision.
4-3
7 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 Crack," WCAP-9558, Rev. 2, May 1982, (Westinghouse Proprietary Class 2).
4-4 Toughness of Austenitic Stainless Steel Pipe Welds, EPRI-NP-4768, Electric Power Research Institute, October 1986.
4-5 Project letter ST-YB-W4-1921, from A. Matiuk (Bechtel Energy Corporation) to J. W. Irons (Westinghouse Electric Corporation),
dated December 16, 1987.
S l
l l
l l
l t
(
l 4
i 4-4
/
1 TABLE 4-1 l
MATERIAL PROPERTIES OF STP UNIT-1, LOOP 1 RHR LINE AT 70*F AS DETERMINED BY TESTING
.2% OFFSET ULTIMATE YIELD
% ELONG
% RED HEAT STRENGTH STRESS PER IN PRODUCT NO.
MATERIAL (PSI)
(PSI)
INCH AREA Elbow 55894 SA403(182)/WP316 79,000 50,500 45.00 70,50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Pipe L5093 SA376/TP316 88,200 42,700 53.20 64.10 Pipe L5093 SA376/TP316 88,200 42,700 53.20 64.10 Pipe L5093 SA376/TP316 88,200 42.700 53.20 64.10 Weld 17138 SFA5.9-ER316L 88,700 65,400 48.00 65.90 Weld 17133 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9-ER316L 88,700 65,400 48.00 65.90 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9-ER316L 88,700 65,400 48.00 65.90 Wald 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9-ER316L 88,700 65,400 48.00 65.90 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Pipe L5093 SA376/TP316 88,700 42,700 53.20 64.10 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Elbow 55894 SA403(182)/WP316 79,000 50,500 45.00 70.50 Pipe P8608 SA376-TP316 85,600 43,300 58.50 71.70 Pipe P8608 SA376-TP316 85,600 43,300 58.50 71.70 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9-ER316L 88,700 65,400 48.00 65.90 Wald 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFAS.9-ER316L 83,600 67,200 50.00 65.50 t
4-5
[
's TABLE 4-2 MATERIAL PROPERTIES OF STP UNIT-1, LOOP 2 RHR LINE AT 70*F AS DETERMINED BY TESTING
.2% OFFSET ULT!WATE YIELD
% ELONG
% RED HEAT STRENGTH STRESS PER IN PRODUCT NO.
MATERIAL (PSI)
(PSI)
INCH AREA Elbow 55895 SA403(182)/WP316 81,500 51,000 56.00 74.50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Weld 17138 SFA5.9-ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFA5.9-ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFAS.9-ER316L 83,600 67,200 50.00 65.50 Weld 0575 SFA5.4/E316L 80,400 57,400 45.00 63.80 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Wald 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Pipe P8607 SA376/TP316 89,800 43,900 53.00 71.70 4
4-6 s
[
TABLE 4-3 MATERIAL PROPERTIES OF STP UNIT-1, LOOP 3 RHR LINE AT 70*F AS DETERNINED BY TESTING
.2% OFFSET ULTIMATE YIELD
% ELONG
% RED HEAT STRENGTH STRESS PER IN PRODUCT NO.
MATERIAL (PSI)
(PSI)
INCH AREA Elbow 55895 SA403(182)/WP316 81,500 51,000 56.00 74.50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Weld 0575 SFA5.4/E316L 80,400 57,400 45.00 63.80 Weld 17138 SFA5.9/ER316L 83,600 67,200 30.00 65.50 Weld 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9/ER316L 83,500 67,200 50.00 65.50 Weld 0683A SFA5.4/E316L 79,700 58,600 40.00 66.20 Weld 0575 SFA5.4/E316L 80,400 57,400 45.00 63.80 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9/ER316L 86,'700 62,600 36.00 65.40 Wold 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Pipe P8607 SA376/TP316 89,800 43,900 53.00 71.70 Elbow 55894 SA403(102)/WP316 79,000 50,500 45.00 70.50 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Weld 17138 SFA5.9-ER316L 86,700 62.600 36.00 65.40 i
Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 4-7
TABLE 4-4 MATERIAL PROPERTIES OF STP UNIT-2, LOOP 1 RHR LINE AT 70'F AS DETERMINED BY TESTING
.2% OFFSET ULTIMATE YIELD
% ELONG
% RED HEAT STRENGTH STRESS PER IN PRODUCT NO.
MATERIAL (PSI)
(PSI)
INCH AREA Elbow 55894 SA403(182)/WP316 79,000 50,500 45.00 70.50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Elbow 43897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Pipe L5091 SA376-TP316 82,600 38,100*
60.50 71.20 Pipe P8607 SA376-TP316 89,800 43,900 53.00 71.70 Pipe L5091 SA376-TP316 88,600 44,100 52.00 67.30 Wald 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Wald 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFAS.9/ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Weld 0575 SFA5.4E316L 80,400 57,400 45.00 63.80 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFA5.9/ER316L 83.600 67,200 50.00 65.50 Wald 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Pipe L5091 SA376/TP316 87,'800 43,300 56.00 66.00 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Elbow 55893 SA403(182)/WP316 81,500 49,000 62.00 75.50 Pipe P8607 A376/TP316 89,800 43,900 53.00 71.70 Pipe L5091 SA376/TP316 87,800 43,300 56.00 66.00 Weld 0575 SFAS.4/E316L 80,400 57,400 45.00 63.80 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50
- The lowest room temperature yield stress 4-8
~
7 TABLE 4-5 MATERIAL PROPERTIES OF STP UNIT-2, LOOP 2 RHR LINE AT 70'F AS DETERMINED BY TESTING
.2% OFFSET ULTIMATE YIELD
% ELONG
% RED HEAT STRENGTH STRESS PER IN PRODUCT NO.
MATERIAL (PSI)
(PSI)
INCH AREA Elbow 55895 SA403(182)/WP316 81,500 51,000 56.00 74.50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50
)
Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 84,600 40,500 59.50 68.20 Weld 17138 SFA5.9-ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9-ER316L 86,700 62,600 36.00 65.40 Wald 0575 SFA5.4/E316L 80,400 57,400 45.00 63.80 Wald 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Wald 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Pipe P8607 SA376/TP316 89,800 43,800 53.00 71.70 4-9
X TA8LE 4-6
~
MATERIAL PROPERTIES OF STP UNIT-2, LOOP 3 RHR LINE AT 70'F AS DETERMINED BY TESTING
.2% OFFSET ULTIMATE YIELD
% ELONG
% RED HEAT STRENGTH STRESS PER IN PRODUCT NO.
MATERIAL (PSI)
(PSI)
INCH AREA Elbow 53894 SA403(182)/WP316 77,500 43,000 65.00 76,50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Elbow 53897 SA403(182)/WP316 81,500 48,500 56.50 74.50 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Wald 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Weld 0683A SFA5.4/E316L 79,700 58,600 40.00 66.20 Wald 17138 SFAS 9/ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Wald 17138 SFA5.9/ER316L 86,700 62,600 36.00 65.40 Wald 17138 SFAS.9/ER316L 83,600 67,200 50.00 65.50 Wald 0575 SFA5.4/E316L 80,400 57,400 45.00 63.80 Weld 0683A SFA5.4/E316L 79,700 58,600 40.00 66.20 Weld 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFA5.9/ER316L 86,~700 62,600 36.00 65.40 Wald 17138 SFA5.9/ER316L 83,600 67,200 50.00 65.50 Weld 17138 SFAS.9/ER316L 86,700 62,600 36.00 65.40 Weld 17138 SFAS.9/ER316L 83,600 67,200 50.00 65.50 Pipe L5091 SA376/TP3167 87,800 43,300 56.00 66.00 Elbow 55894 SA403(182)/WP316 81,500 49,000 62.00 75.50 Pipe L5093 SA376/TP316 81,400 44,500 61.00 77.30 Weld 17138 SFA5.9-ER316L 86,700 62.600 36.00 65.40 Weld 17138 SFA5.9-ER316L 83,600 67,200 50.00 65.50 4-10
TABLE 4-7
~
TYPICAL TENSILE PROPERTIES OF SA376 TP316, SA351 CF8A and WELDS OF SUCH MATERIAL FOR THE PRIMARY LOOP Test Temperature Avorage 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 Wald 70 63,900(3) 87,600 (3)
B SA376 TP316 70 47,100(40) 88,300 (40) 650 26,900(22) 69,100 (25)
E 308 Wald 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)
O SA351 CF8A 70 47,300(14) 84,500 (14) 650 26,000(4) 70,500 (4)
Wald 70 61,200(31) 84,500 (32)
- a. (
) indicates the number of test results averaged.
4-11
~
TABLE 4-8 FRACTURE TOUGHNESS PROPERTIES TYPICAL OF THE RHR LINE Test Temp.
Tensile Properties (psi)
J T
Ic mat 2
Material
(*F)
Yield Ultimate (in-lb/in )
- a,c.e SA376 TP316 600 21,700 65,500 SA376 TP316 600 20,500 60,100 b
b Weld (E308 and E316) 600 45,000 61,200 f
Weld 600 h
SA351 CF8A 600 i_
[
ja,c.e i
- b. Lowest of 6 tests.
U a,c.e 4
4 O
O 9
O
/
- l. '
.4
~
a,c.e a
9 6
i
=
Figure 4-1.
True Stress-True Strain Curve for SA376 TP 316 Stainless Steel at 623*F (minimum properties) 4-13 i
,,. _ _,~
._,,e
m
. e-t a.c.e 1.
1 4
True Stress-True Strain Curve for SA 376 TP 316 Stainless Steel Figure 4-2.
(Average Properties) l 4-14 M
/
3000 l
I I
2000 i
N e l
E I
a
'T I
3 1000 l
I
-2 I
~
l I
0 i
O 100 140 200 T
1 Figure 4-3.
J versus T Curve for SMAW Welds i
4-15
p SECTION 5.0 LOADS FOR FRACTURE MECHANICS ANALYSIS Figures 5-1 through 5-3 are schematic layouts of the RHR lines attached to the reactor coolant loops 1, 2, and 3 of South Texas Units 1 and 2.
Highest and next highest stressed node locations are shown by the layout in figure 5-2.
The stresses due to axial loads and bending mousntr. were calcule*.ed by the following equation:
o=k+f (S.1)
- where, stress o
=
axial load F
=
bending moment M
=
metal cross-sectional area A
=
section modulus Z
=
The bending moments for the desired loading combinations were calculated by the following equation:
M=/M 2g (5.2) y Z
- where, bending moment for required loading M
=
Y component of bending moment M
=
y Z component of bcnding moment M
=
Z The axial load and bending mments for crack stability analysis and leak rate i
predictions were computed by the methods to be explained in Sections 5.1 and 5.2.
4 51
N 5.1 Leads for Crack Stability Analysis The faulted loads for the crack stability analysis were calculated by the following equations:.
l (5.3)
TH + F l + lFSSE IF0W+F F
=
p l(M )DW + (N )THI + IIN )SSEl (5.4)
M
=
y Y
Y y
IIN )DW + IN )THI + I(N )SSEl (5.5)
M Z
Z Z
Z Where, the subscripts of the above equations <,e.snt t'1e following loading
- cases, deadweight DW
=
normal thermal expansion TH
=
SSE loading including seismic anchoi mo SSE
=
load due to internal pressure P
=
5.2 Leads for Leak Rate Evaluation The normal operating loads for leak rate predictions were calculated by the following equations:
d (5.6)
F0W+FTH + Fp F
=
(M )0W + (N )TH (5.7)
N
=
y Y
Y (N )CW + (N )TH (5.8)
M Z
Z Z
5.3 Sumary of Loads, Geometry and Materials Table 5-1 provides a summary of envelope loads computed for fracture mechanics evaluations in accordance with the methods described in section 5.1, and 5.2 (references 5-1 and 5-2). The cross-sectional dimensions and materials are also summarized.
Load data are tabulated at the highest stressed location (node 135, loop 2), and the second highest stressed location (node 130E, loop 2). The loading components are provided in table 5-2.
5-2
~
l All loads used in the LBB evaluation are for the as-built configuration. Unit 1 and Unit 2 have the same as-built normal loads at the governing location ti.'
the as-built normal plus SSE loads are slightly higher for Unit 2.
The Unit 2 loads are used in the stability evaluations dit ussed in Sectica 6.0.
5.4 Governing Location The normal plus SSE axial stresses along the RHR line starting from the primary icop nozzle junction up to the first isolation valve were compared.
The maximum stress occurs at node 135 on the loop 2 RHR line.
The welding process at this location is GTAW. This location is identified in figure 5-2.
Detailed fracture mechanics analyses were performed at this highest stressed location.
5.5 Weferences 9
5-1 Project Letter ST-YB-WN-1921, from A. Matiuk (Bechtel Energy Corporation) to J.W. Irons (Westinghouse Electric Corporation) dated December 16, 1587.
5-2 Project te,tter ST-YB-WN-1937, from A. Natiuk (Bechtel Energy Corporation) toJ.W. Irons (WestinghouseElectricCorporation)datedApril 15, 1280.
1 0
5-3
(
f I
TABLE 5-1 SIM ARY OF ENVELOPE LOADS NOMINAL-NINIMUN OUTSIDE WALL WALL INSIDE' h
b NODE DIA THICK THICK DIA F
M j
LOCATION CONDITION NO LOOP MATERIAL (inches) _ SCHEDULE (inches) (inches) (inches) (kips)(in-kips)
Highest Faulted 135 2
SA376 12.75 140 1.125 1.005 10.74 198 862 Load
-TP3166 Location No mal Operating 135 2
5A376 12.75 140 1.125 1.005 10.74 1%
395
-TP316 Next Faulted 130E 2
SA376 12.75 140 1.125 1.005 10.74 198 793 T
Highest
-IP316
^
Load Location Normal 130E 2
SA376 12.75 140 1.125 1.005 10.74 196 3%
Operatii.g
-TP316
- Based on minimum wall thickness of 1.005 in, bThe stress based on the loads is 13.9 ksi using the minimum wall thickness.
/
TABLE 5-2 LOADING COMPONENTS AT GOVERNING LOCATIONS Highest Load Next Highest load (Location - 135, Loop 2)
(Location - 130E, Loop 2)~
j Load Axial Bending Bending Axial Bending Bending Type Force (lbs) Woment NY (ft-lb) Woment NZ (ft-lb)
Force (lb)
Woment NY (ft-lb) Woment NZ (ft-lb)
Dead
-746 626 1027
-646 580 527 Weight l
Thermal 1950 30,156
-12,683 1,948 30.351
-12,114 l
194,828 Pressure 194,828 En SSE +
1792 36,808 12,691 1793 31,378 10,476 Anch. Hot.
I l
Hot leg loop 1
~
l 1
Moderate Energy Piping 4
V' l
Figure 5-1 Schematic Layout of the RHR Lines - Loop 1. Units 1 and 2 l
5-6 l'
/.
e
. E:
s u
- l -
Hot leg Loop 2 D
Node 135 tioderate Energy Piping Node 130E e
Figure 5-2 Schematic Layout of the RHR Lines - Loop 2, Units 1 and 2 5-7 3
_ i.--.. -. -,. _ _. -. - -
- ~
l
'b Moderate 7
Energy Piping Hot leg loop 3 j'
I
(
i Figure 5-3.
Schematic Layout of RHR Lines Loop 3. Units 1 and 2 5-8
/
SECTION 6.0 FRACTURE MECHANICS EVALUATION 6.1 Global Failure Mechanism Oetermination 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 [
la,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 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 called as the flow stress.
(
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 RHR line. The failure criterion has been obtained by requiring equilibrium of the section containing the flaw i
(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:
a,c.e
(
)
(6.1)
(
la,c e (6.2) where:
ja,c.e 6-1
c >
[
ja,c.e u) (f1 w stress) of
= 0.5 (a +
y 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).
Flaw stability evaluations using this analytical model, are presented in section 6.3.
6.2 Leak Rate Predictions The purpose of this section is to discuss the method which will be used to predict the flow through a postulated crack and present the leak rate calculation results for postulated through-wall circumferential cracks in the RHR line.
6.2.1 General Considerations The flow of hot pressurized water through an opening to a lower back pressure (causing choking) is taken into account. For long channels where the ratio of the channel length, L, to hydraulic diameter, D, (L/D ) is greater than g
g
(
Ja.c.e, both (
Ja,c.e must be considered.
In this situation the flow can be described as being single phase through the channel until the local pressure equals the saturation pressure of the fluid.
At this point, the flow begins to flash and choking occurs.
Pressure losses due to momentum changes will dominate for (
Ja,c.e However, for large L/D values, friction pressure drop will become important and must be H
considered along with the momentum losses due to flashing.
6-2
}
/
6.2.2 Calculation Method
~
In using the isentropic equilibrium model, the basic method used in the leak rate calculations is the method developed by [
l l
ja,c,e,
The flow rate through a crack was calculated in the following manner.
Figure 6-2 from reference 6-2 was used to estimate the critical pressure, Pc, for the l
i primary loop enthalpy condition and an assumed flow. Once Pc was found for a given mass ficw, the (
]C'8 was found from figure 6-3 taken from reference 6-2.
For all cases considered, I
since (
la,c.e Therefore, this method will yield the two phase pressure drcp due to momentum effects as illustrated in figure 6-4.
Now using the assumed flow rate, G, the frictional pressure drop can be calculated using APf=(
la,c,e (6.3) where the friction factor f is determined using the (
Ja,c.e
~
The crack relative roughness, e, 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 (
Ja.c.e to obtain the total pressure drop from the primary system l
to the atmosphere.
- Thus, Absolute Pressure - 14.7 = (
Ja,c e (6.4)
O 6-3
1 i
f 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 (
)]a,c.e for this type of (
la,c.e calculation.
6.2.3 Leak Rate Calculations Leak rate calculations were made as a function of postulated through-wall crack length for the critical location previously identified.
The crack opening area was estimated using the method of reference 6-4 and the leak rates were calculated using the calculation methods described above.
The leak rates were calculated using average material properties with the normal operating loads of axial force F = 196 kips and bending moment M = 395 in-kips.
The crack length yielding a leak rate of 10 gpm (10 times the leak detection requirement of 1.0 gpm) is found to be (
Ja,c.e,
Thus the reference flaw size of (
la,c,e is established.
6.2.4 Leak Detection Capability, Administrative Procedures and Technical Specification Requirements All LBB candidate lines at South Texas Project are located inside the containment.
The STP leakage detection criterion includes a detected uniden-tified 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 leakage size flaw used in the stability analysis. The basis for the 1.0 gpm leak rate is the presence (inside the 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 i
equal to 1 gpm.
Indication of containment humidity is also provided in the 6-4
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 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br />.
This section of the technical specification also requires performance of a reactor coolant system inventory balance at least once per 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />.
Detected leaks will be repaired within the system limiting conditions for operation established in either technical specifications or administrative procedures. When leakage 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 within the next thirty hours.
Repair would be required before restart.
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 at one location 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 indicated that a 1.0 gpm leak rate can be reliably detected and located during plant operation.
6.3 Stability Evaluation Using the "Z" Factor Approach A typical segment of the pipe under maximum loads of axial force F and bending moment M is schematically illustrated as shown in figure 6-5.
In order to calculate the critical flaw size, a plot of the limit moment versus crack length is generated as shown in figure 6-6.
The critical flaw size corresponds to the intersection of this curve and the maximum load line. The critical flaw size is calculated using the lower bound base metal tensile 6-5
properties established in section 4.0.
From figure 6-6 the critical flaw size is seen to be C.
.]a,c.e for the base metal.
The we6d at the governing location is a field weld using a GTAW process.
Other field welds on the RHR lines are either GTAW or SMAW.
It has been established (reference 6-5) that GTAW welds exhibit superior toughness properties as compared to SMAW welds.
For this eealuation, a "Z" factor correction for SMAW welds was conservatively applied even though the actual weld of interest is a qTAW weld.
The "Z" factor was applied (reference 6-5) as follows:
Z = 1.15 [1 + 0.013 (0.0. - 4))
(6.5) where OD is the outer diameter of the pipe in inches. Substituting 00 = 12.75 inches, the Z factor was calculated to be 1.2808. The applied loads were increased by the Z factor and the plot of limit load versus crack length was regenerated as shown in figure 6-7.
The lower bound base metal tensile properties (section 4.0) were used for this purpose. From figure 6-7, the
~
critical flaw size is seen to be (
Ja,c.e long. Noting that the flaw yielding a leakage of 10 gpm (i.e. leakage size flaw) was calculated to be (
la,c,e long, a factor of 2.8 exists between the leakage size flaw and the critical flaw. Thus, a margin of greater than 2 on flaw size is in evidence.
In order to determine the margin on applied loads (normal plus SSE), the applied loads were increased by a factor of 1.79 (i.e.1.4 Z) and the plot of limit load versus crack length was generated as shown in figure 6-8.
Again the lower bound base metal tensile properties were used for this purpose.
From figure 6-8 the critical flaw size is seen to be (
la,c.e long which is larger than the [
}a,c.e inches long leakago size flaw. Thus a i
margin on load of at least 1.4 times normal plus SSE loads is demonstrated.
i 6-6
/
I 6.4 Local Stability Analysis
~
In this section the local stability analysis is performed to show that unstable crack extension will not result when postulated through wall flaws are subjected to maximum plant loads.
At the critical location identified in section 5.0, the (normal plus SSE) outer surface axial stress; o,, is seen to be 13.9 ksi based on the minimum wall thickness.
The (normal plus SSE) axial force and bending moment are F = 198 kips and M = 862 in-kips.
The minimum yield strength for flaw stability analysis is [
Ja,c.e ksi (see section 4). The EPRI elastic plastic fracture handbook method is used to calculate the J using the normal plus SSE loads. The J was applied applied calculated for a (
Ja,c.e long postulated through wall flaw (which is 2 times the reference flaw size) and was found to be (
Ja,c,e 2
which is lower than J f 959 in-1b/in,
Ic In addition, for a leakage size flaw, i.e. the reference flaw (
Ja,c,e long, the normal plus SSE load was increased by a factor of /2.
The J-T analysis gave an applied J of (
Ja,c e which is also less than J Ic 2
of 959 in-lb/in.
Thus crack stability is demonstrated for both the above cases.
6.5 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 (
ja,c.e 9
6-7
6-3 (
ja c.e,
6-4 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.
6-5 ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640.
5 O
G 6-8
(
e O
I o.
O i
I i
1 1
1 C
Q 3
.Q
.e.
bw f
.n.
I t
'y s
e 4
"o 1
L ue Q
U.
m e
1 r-e w
M i
I g
G N
L
- 3
.G b
J s
e 6-9 m
e a c.e 5
E l!
8 d
N 5
Figure 6-2.
Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures 6-10
a,c.e 9>
I w
Iw Ea h
Iu LENGTH /DI AMETER RATIO (L/D)
Figure 6-3 (
Ja,c.e Pressure Ratio as a Function of L/0 6-11
I
'N l
a,c.e a,c.e
~
f L/c = q g
" LlDu MM L
Figure 6-4 Idealized Pressure Orop Profile Through a Postulated Crack 6-12 i
\\
l
\\
l
. nom) M 1.125,I(min) I 1.
5 e
~~~~~~--
k<
D l + P -.-
I lg Q
cn 6
4 12.75" P = 2235 psi F=
198 kips M=
862 in-kips l
Figure 6-5 Loads Acting on the Pipe Modal at the Goveraing Location l
l 1
i
'N a,c.e 00 =.12.75" t = 1.005" P = 2235 psig F = 198 kips o = 22.1 ksi y
= 66.4 ksi u
op = 44.3 ksi Temp = 623*F l
l Figure 6-6.
Critical Flaw Size Prediction for the Base Metal Using Limit Load Approach l
6-14
~
?
l i
\\
a,c.e 00 = 12.75" r
t = 1.005" P = 2235 psig F* = 254 kips o = 22.1 ksi y
= 66.4 ksi u
op = 44.3 ksi Temp = 623*F
- Z Factor of 1.28 was applied to the load resulting in applied force of 254 kips.
Figure 6-7 Z-Factor Calculations for SMAW Welds to Demonstrate Margin on Flaw Size 6-15
N a,c e 00 = 12.75" t = 1.005" P = 2235 psig F* = 354 kips o = 22.1 ksi y
o = 66.4 ksi u
op = 44.3 ksi Temp = 623*F N
- d 2 Factor of 1.79 was applied to the load.
THis resulted in applied force of 354 kips.
Figure 6-8 6 Z-Factor Calculations for SMAW Welds to Demonstrate Margin on Loads 6-16
,/-
SECTION 7.0 ASSESSMENT OF FATIGUE CRACK GROWTH The fatigue crack growth on the South Texas Project RHR line was determined by l
comparison with a generic fatigue crack growth analysis of a similar piping system.
The details of the generic fatigue crack growth analysis are presented in Appendix B.
By comparing all parameters critical to the fatigue crack growth analysis, between STP Units 1 and 2 RHR lines and generic RHR iines, it was concluded that the generic analysis would envelop the fatigue crack growth of the STP RHR lines.
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 strasses and number of cycles, material, geometry, and types of discontinuities.
The following summarizes the parameters which were compared:
a,c.e
]a,c.e Critical location Pipe Outer Diameter 12.75" 12.75" Thickness 1.005" 1.005" Material Austenitic Stainless Steel Austenitic Stainless Steel Normal Temperature 617'F 623*F Normal Pressure 2235 psig 2235 psig Normal Operating 20.3 ksi 9.2 ksi Stress (Press,DW, Thermal Exp.)
Thermal Transients See Appendix B Thermal transient loadings are nearly identical for the two projects.
i 7-1
h 1
This comparison demonstrates the many similarities between the STP RHR line and the generic RHR line evaluation. The normal operating stress for the STP RHR lines is significantly lower. This will result in reduced crack growth.
7.1 Acceptability of 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 STP RHR lines.
Detailed discussion in support of this assumption has been provided in the previous section.
The maximum allowable preservice indication may have a depth of 0.1 in. per IWB-3514.3, Allowable Indication Standard for Austenitic Piping, ASME Code,Section XI - Division 1, 1985 edition.
From the genoric analysis described in Appendix B, the final flaw size (40 year life) will be 0.1514 in, for a postu-lated initial flaw size of 0.15 in, subjected to the design transients.
[
Ja,c.e Thus, the first criterion on flaw depth is satisfied.
Secondly, the worst caso transient AK value fer the maximum crack depth is a
(
J,c,e The flow stress for the base metal at 623*F is 44.3 ksi which can be used to obtain a conservative estimate of the plastic zone size.
The expression for the plastic zone size, r, calculation is:
(reference 7-2) p
= H ( 6Eog )2 rp e
7-2
f From the preceding equation, the plastic zone size is calculated to be (
]a,c.e The remaining ligament for the 0.1514 in, deep end-of-fatigue-life flaw is 0.854 in. (i.e.1.005 - 0.1514). Thus, the plastic zone size is far less than the remaining ligament.
Based on the above, it is concluded that for the STP RHR Lines, the fatigue crack growth during service will not be significant.
7.2 References 7-1 Bamford, W. H. "Fatigae Crack Growth of Stainless Steel Reactor Coolant Piping in a Pressurized Water Reactor Environment," ASME Trans. Journal of Pressure Vessel Technology, February 1979.
7-2 Rice, J. R., ASTM STP,1967, Volume 415, p. 247.
9 o
7-3
/
I SECTION 8.0 ASSESSMENT OF MARGINS In the preceding sections, the leak rate calculations, fracture mechanics analysis and fatigue crack growth assessment were performed. Margins at the critical location are sumarized below:
In section 6.3 the "critical" flaw size using the limit load method is calcu-la n d to be (
Ja,c.e long.
Using the IWB-3640 approach (i.e.
"Z" factor approach), the critical flaw size at the governing location weld is found to be (
.]a,c.e long.
In section 6.4 it is demonstrated, using the J-T approach that a postulated [
la c.e long through-wall flaw will remain stable when subjected to normal plus SSE loads.
Based on the above, the critical flaw size will exceed (
) a,c.e In section 6.2 it is shown that at the critical location, a flaw of (
]a,c.e would yield a leak rate of 10 gpm.
Thus, there is a margin of at least 2 on flaw size and a margia of 10 with respect to the plant leak detection capahiiity of 1 gpm.
In sections 6.3 and 6.4 it is shown that the reference flaw (
Ja,c.e yielding a leak rate of 10 gpm would be stable when subjected to a load equal te 1.4 (normal + SSE). Specifically, using the IWB-3640 approach (section 6.3) a [
la c e long through-wall flaw was shown to be stable when subjected tc 1.4 Z (normal + SSE) loads. Also, based on the local stability analysis (section 6.4) the leakage size flaw of (
la,c,e inches was shown to be stable when subjected to /2 (normal + SSE) loads.
e 8-1
\\s In summary, relative to 1.
Loads The leakage-size crack will not experience unstable crack extension even if very large loads of /li(normal plus SSE) are applied.
2.
Flaw Size A margin of at least 2 exists between the critical flaw and the flaw a.
yielding a leak rate of 10 gpm.
b.
If limit load is used as the basis for critical flaw size, an even larger margin would result.
3.
Leak Rate A margin of 10 exists for the reference flaw (
Ja,c,e between calculated leak rate and the 1 gpm leak detection criteria of Regulatory Guide 1.45.
A summary comparison of criteria and analytical results is given in table 8-1.
The criteria are seen to be met.
e 8-2
TABLE 8-1 COMPARISON OF RESULTS VS. CRITERIA
~
CRITERION RESULT 1.
NUREG 1061 Volume 3 Met Section 5.2(h) -
(Required margin of 2 demonstrated)
Margin on Flaw Size 2.
NUREG 1061 Volume 3 Met Section 5.2(i) -
(Required margin of /2 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 allowable Met fatigue crack growth (af < 60% wall thickness and
~
plastic zene size < remaining ligament)
O e
8-3
/
SECTION 9.0 CONCLUSIONS This report justifies the elimiaation of RHR lir.a pipe breaks from the structural design basis for the South Texas Units 1 and 2 as follows:
a.
Stress corrosion cracking is precluded by use of fracture resis, tant materials in the piping system and controls on roactor coolant chemistry, temperature, pressure, and flow during normal ope.ation, 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 irtegrity of the 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 deter. ting unidentified leais 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.
With respect to stability of the reference flaw, ample margin exists between the maxinum postulated loads and the plant specific as-built faulted loads (i.e. normal + SSE).
The reference flaw will be stable because of the ampia margins in d, e, v.d f and will leak at a detectable rate which will assure a safe plant shutdown.
1 Based on the above, it is concluded that pipe breaks in the RHR piring need not be considered in the structural design basis of South Texas Project Units 1 and 2.
9-1
/
APPENDIX A
{
LIMIT MOMENT t
~
(
)
4 l
l l
l l
4 a,C,9 1
1 s
A-1
i t
i s
- a,c, L
i 9
Figure A-1 Pipe with a Through-Wall Crack in Bending A-2
[
i APPENDIX B FATIGUE CRACK GROWTH CONSIDERATION 8.1 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 Ssction 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 wall stress distribution, respectively, for convenience.
The constant stresses due to pressure, deadweight and thermal expansion (at normal operating temperature, 617'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 transients.
3.1.1 Critical Location for Fatigue Crack Growth Analysis The RHR line design thermal transients (Section B.1,.2), 1-0 analysis data on RHR line thermal transient stresses (based on ASME Section III NB-3600 rules) and the geometry were reviewed to select the worst location for the fatigue crack growth analysis.
(
determined to be the most critical location for the fatigue crack growth Ja,c.e This location is selected as the worst location based on the following considerations:
i) the fatigue usage factor is highest.
ii) the effect of discontinuity due to undercut at weld will tend to increase the cyclical thermal transient loads.
iii)
'he review of date shows that the 1-D thermal transient stresses in the RHR line piping section are generally higher near the (
ja.c.e B-1 1
\\
B.1.2 Design Transients The transient conditions selected for this evaluation are based on conser-vative 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 t,o occur during plant operation. The fatigue evaluation based on these tra.nsier;ts provides confidence that the component is appropriate for its application uver the design life of the plant. All the normal cperating and upset thermal transients, in accordance with the design specification and the applicable system design criteria document (B-1), were considered for this evaluation.
Out of these, (
3a,c.e B.1.3 Simplified Stress Analysis The simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients.
[
.]a,c.e The inside surface stress was calculated by the following equation which is similar to the transient portion of ASME Section III NB-3600, Eq. 11:
Sg=[
))a,c.e (B.1)
- where, S$ = inside surface stress a,c.e (B.2)
(B.3)
B-2
s' a,c.e 1
l 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 = [
]
8'">'
(B.5) s02 = (
J **
(B.6)
- where, m.a C,9 l
mm B-3
\\s The material properties for the RHR pipe (
)]a,c.e and the RCL
(
Ja,c.e The values of E and a, at the normal operating temperature, provide a conservative estimation of the through wall thermal transient stresses as compared to room temperature properties.
The following values were conservatively used, which represent the highest of the (
lac.e materials:
a,c.e The maximum and minimum linear through wall stress distribution for each thermal transient was obtained by [
Ja,c.e The simplified analysis discussed in this section was performed for all thermal transients of table B-1.
The inside and outside surface stresses calculated by simplified methods for the transients are shown in table B-2.
B.1.4 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 the small number of cycles.
I B.1.5 Total Stress fo. Fatique Crack Growth The total through wall stress at a section was obtained by superimposing the l
pressure load stresses and the stresses due to deadweight and thermal expansion (normal operating case) on the thermal transient stresses (of table i
B-2).
Thus, the total stress for fatigue crack growth at any poir.t is given by the following equation:
j t
l B-4
Total Stress Thermal Strass Due Stress for Transient to Oue to Stress
+
OW +
+
Internal (B.7)
Fatigue
=
Crack Growth Thermal Pressure Expansion The envelope normal loads from thermal expansion, deadweight and pressure for calculating the total stresses of equation B.7 are sumarized in table B-3.
B.2 Fatigue 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-1.
A range of crack depths was postulated, and each was subjected to the transients in table B-1, and envelope normal loads in table B-3.
B.2.1 Analysis Procedure The fatigue crack growth analyses presented herein were conducted in the same manner as suggested by Section XI, Appendix A of the ASME Boiler and Pressure Vessel Code. The analysis procedure involves assuming an initial flaw exists 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 AK, by the following g
relation:
h=CoaK" (B.2.1) g is whore "Co" and the exponent "n" are material properties, and aKg definad 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 nean stress present during the cycle.
This can be accounted l
f B-5
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 environnent 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 g
crack and structure geometry and the range of applied stresses in the area where the crack' exists. Once AK is calculated, the growth due to that g
particular cycle can be calculated by equation (B.2.1).
This increment of adjusted, and the growth is then added to the original crack size, the AKg analysis proceeds to the next transient. The procedure is continued in this manner until all the transients have been analyzed.
The crack tip stress intensity factors (K ) to be used in the crack growth g
analysis were calculated using an expression which applies for a semi-elliptic surface flaw in a cylindrical geometry (reference B-3).
The stress intensity factor expression was taken from Reference B-3 and was
~
calculated using the actual stress profiles at the critice.1 section.
The maximum and minimum stress profiles corresponding to each transient were input, and each profile was fit by a third order polynomial:
o (x) = A + A1 j+ A ({} + A ({[
(8.2.2) 0 2
3 The stress intensity factor K (4) was calculated at the deepest point of g
the crack using the following expression:
~
a,c.e (B.2.3) m M
B-6
a,c.e 1
Calculation of the fatigue crack growth for each cycle was then carried out using the reference fatigue crack growth rate law determined from considera-tion of the available data for stainless steel in a pressurized water envircement.
This law allows for the effect of mean stress or R ratio Imin/K,,x) on the growth rates.
(K g
The referen:e crack growth law for stainless steel in a pressurized water environment was taken from a collection of data (B-4) since no code curve is available, and it is defined by the following equation:
h={
] a,c.e (B.2.4) where K,ff = (K,,,) (1-R)1/2 g
E Imin g, K,,,
g h = crack growth rate in micro-inches / cycle l
l O
B-7
N 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 (
la.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 flaw growth is significantly lower.
For example, a postulated (
)"'C
inch deep flaw will grow less than (
la,c.e.
These results also confirm operating plent experience.
B.3 REFERENCES B-1 Westinghouse System Standard Design Criteria 1.3, "Nuclear Steam Supply System Design Transients," Revision 2, April 15,1974, Westinghouse Proprietary.
B-2 ASME Section III, Division 1-Appendices, 1983 Edition, July 1, 1983.
l B-3 McGowan, J. J. and Raymund, M., "Stress Intensity Factor Solutions for Internal Longitudinal Semi-Elliptical Surface Flaws in a Cylinder Under Arbitrary Loadings", Fracture Mechanics ASTM STP 677, 1979, pp. 365-380.
t
(
B-4 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.
l 1
l t
A B-8
TABLE B-1 THERMAL TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH EVALVATION Trans.
No. of No.
Description Occurrences 8,C,9
+
1 M
O I-B-9
q-M P f.
TABLE B-2 TRANSIENT' STRESSES FOR RHR LINE (psi)
Transient Maximum Corresponding Minimum Corresponding No, Inside Stress Outside Stress Inside Stress Outside Stress a,c.e P
l O
B-10
TABLE B-3 ENVELOPE NORMAL LOADS i
Condition Normal Operating 4
a 9
l 1
k 0
k B-11
\\
TABLE B-4 RHR LINE FATIGUE CRACK GROWTH RESULTS Section Thickness [
]* a,c, INITIAL CRACK DEPTH AFTER YEAR CRACK OEPTH (in) 10 20 30 40 a,c.e
- This is conservatively taken as minimum thickness of the counter bore region l
l l
l l
l l
B-12 i
e a,c.e RER PIPE Figure B-1 Schematic of RHR Line At (
) a,c.e B-13
.