ML20246K178
| ML20246K178 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 05/31/1989 |
| From: | Lee Y, Schmertz J, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19302D838 | List: |
| References | |
| TXX-89249, WCAP-12268, NUDOCS 8905170255 | |
| Download: ML20246K178 (74) | |
Text
- _ - _
WESTINGHOUSE CLASS 3 t-WCAP-12268' i
.. j TECHNICAL BASES FOR ELIMINATING RL'PTURE OF THE ACCUMULATOR INJECTION N0ZZLES AS A STRUCTURAL DESIGN BASIS FOR COMANCHE PEAK UNIT 1 May 1989 J. C. Schmertz Y. S. Lee S. A. Swamy F. J. Witt Verified by: T v~ /
C. TT Y ng /
Approved by: .
wIh/M F. 5. E47 u'safhy, Manag'er Struefural Materials Engineering Work Performed Under Shop Order TKUJ-2443 s
s WESTINGHOUSE ELECTRIC CORPORATION Nuclear and Advanced Technology Division P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 g51J@h$8 5 A
- __ _ _ _ _ _ _ __ _ __-_ ____ ___-_________-_________ - D
TABLE OF CONTENTS Section Title Pace s.
1.0 INTRODUCTION
1-1 1.1 Background 1-1 1.2 Scope and Objective 1-1 1.3 References 1-3 2.0 FAILURE CRITERIA FOR FLAWED ACCUMULATOR INJECTION 2-1 N0ZZLES 2.1 General Considerations 2-1 2.2 Global Failure Mechanism 2-1 2.3 Lecal Failure Mechanism 2-2 2.4 References 2-3 3.0 OPERATION AND STABILITY OF THE ACCUMULATOR INJEJTION 3-1 N0ZZLES-3.1 Stress Corrosion Cracking 3-1 3.2 Water Hammer 3-3 3.3 Low Cycle and High Cycle Fatigue 3-4 3.4 Summary Evaluation of Accumulator Line for Potential Degradation During Service 3-4 3.6 References 3-5 4.0 MATERIAL CHARACTERIZATION OF THE ACCUMULATOR 4-1 INJECTION N02ZLES 4.1 Nozzle Material and Welding Process 4-1 4.2 Tensile Properties 4-1 4.3 Fracture Toughness Properties 4-2 4.4 References 4-3 s
5.0 LOADS FOR FRACTURE MECHANICS ANALYSIS 5-1 5.1 Loads for Crack Stability Analysis 5-2 5.2 Loads for Leak Rate Evaluation 5-2 1 5.3 Summary of Loads Geometry and Materials 5-2 5.4 Governing Location 5-2 nu.una io 3$
]
TABLE OF CONTENTS (cont.)
- - Section Title Page
- ~
S.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 6.3 Stability Evaluation 6-4 6.4 Local Stability Analysis 6-5 6.5 References 6-6 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 ' CONCLUSIONS 9-1 APPENDIX A Limit Moment A-1 APPENDIX B Fatigue Crack Growth Considerations B-1 B.1 1hermal Transient Stress Analysis B-2 B.1.1 Critical Location for Fatigue Crack B-2 Growth Analysis i
s 3732s/042489 10 jjj
b TABLE OF CONTENTS (cont.)
'~
Section Title Page
~
B.I.2 Design Transients B-3 B.1.3 Simplified Stress Analysis B-3 ,
B.1.4 Non-linear Stress Distribution for B-6 Severe Transients B.1.5 OBE Loads B-6 B.I.6 Total ftress for Fatigue Crack Growth B-7 B.2 Fatigue Crack Growth Analysis B-7 B.2.1 Analysis Procedure B-7 B.2.2 Results B-10 B.3 References B-10 e
i J
s i,
LIST OF FIGURES Figure Title Page 2-1 Schematic of Generalized Load Deformation Behavior 2-4 4-1 True Stress Strain Curve For SA351-CF8A Stainless 4-6 Steel at 557'F 5-1 Schematic Layout of Accumulator Injection Nozzle 5-5 with Section of Accumulator Line 6-1 [ Ja,c.e Stress Distribution 6-7 6-2 Analytical Predictions of Critical Flow Rates of 6-8 Steam Water Mixtures 6-3 Critical or Choked Pressure Ratio As a Function 6-9 of L/D 6-4 Idealized Pressure Drop Profile Through a Postulated 6-10 Crack 6-5 Loads Acting on the Model at the Governing 6-11 Location j 6-6 Critical Flaw Size Prediction Using Limit Load 6-12 Approach k
6-7 Critical Flaw Size Prediction Including Application 6-13 of 2-Factor I
i 3732s 543389 10 y
i LIST OF FIGURES (cont.)
a-Figure Title Page 1
A-1 Pipe with a Through Wall Crack in Bending A-2 I
B-1 Comparison of Typical Maximum and Minimum Stress Profile Computed by Simplified [
i ja c.e B-16 l
1 l
, , B-2 Schematic of Accumulator Line at [ '
ja,c.e B-17 B-3 [ Ja,c.e and Minimum Stress Profile for Transient #10 B-18
._ B-4 [ la,c e Maximum and Minimum Stress
. Profile for Transient #11 B-19 j 8
B-5 .[ 1 ' ** Maximum and Minimum Stress Profile for Transient #12 B-20 B-6 [ ]8'C Maximum and Minimum Stress Profile for Transient #14 B-21
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LIST OF TABLES Table No. Title Page 4-1 Available Mechanical Properties of the Accumulator Injection Nozzles at 650*F 4-4 4-2 Chemistry and End of Service Life KCU Toughness for Accumulator Injection Nozzles 4-5 5-1 Summary of Envelope Loads for Accumulator Injection Nozzles 5-4 8-1 Comparison of Results vs. Criteria 8-3 B-1 Thermal Transients Considered for Fatigue Crack Growth Evaluation B-12 B-2 Stresses for the Minor Transients (PSI) B-13 B-3 Envelope Normal Loads B-14 B-4 Accumulator Line Fatigue Crack Growth Results B-15 I
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WESTIN0H3usE PROPRIETARY CLASS 2 e
SECTION 1.0
,, INTRODUCTION
_. 1.1 Background The current structural design basis for the accumulator line requires postulating non-mechanistic circumferential and longitudinal pipe breaks.
This results in additional plant hardware (e.g. pipe whip restraints and jet shields) which would mitigate the dynamic consequences of the pipe breaks. It is, therefore, highly desirable to be realistic in the postulation of pipe
-breaks for 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 of that method that are I used for establishing Leak-Before-Break (LBB) for the four accumulator injection nozzles. The evaluations considering circumferentially oriented
. flaws cover longitudinal cases. The scope of piping covered by this report only addresses the accumulator injection nozzles. Special attention is directed to the accumulator injection nozzles because they are m.de of cast austenitic stainless steel, which is susceptible to thermal aging. Resolution of the thermal aging issue for these nozzles is shown in this report.
Justification for LBB for the balance of the accumulator lines is already provided in reference 1-1.
1.2 Scope and Objective The general purpose of this investigation is to demonstrate leak-be. fore-break I for the accumulator injection nozzles. A schematic drawing of the piping system at the injection nozzles is shown in Section 5.0. The recommendation l and criteria proposed in NUREG 1061 Volume 3 (1-2) 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 injection nozzle at which the highest stress occurs.
8'22.,ourn to 11 <
- 2) Identify the materials and the associated material properties.
- 3) Postulate a surface flaw at those nozzles having the least favorable
'~
combination of stress and material properties. Determine fatigue crack growth. Show that a through-wall crack will not result.
- 4) Postulate a through-wall flaw at the governing nozzle. 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 nozzle is subjected to normal operating loads. A margin of '
10 is demonstrated between the calculated leak rate and the leak
, detection capability.
- 5) Using ncrmal 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.
The flaw stability criteria proposed for the analysis examines both the global and local stability for a postulated through wall circumferential flaw. The global analysis is carried out using the [ ]a,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 EPRI elastic plastic fracture handbook method.
1 The leak rate is ce' ulated for the normal operating condition. The leak rate prediction model used in this evaluation is an [
3732s/042649 to
}g
]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. Surface roughness is accounted
~
for in determining the leak rate through the postulated flaw.
Several computer codes are used in the evaluations. The main-frame computer programs are under Configuration Control which has requirements conforming to Standard Review Plan 3.9.1. The fractcre mechanics calculations are independently verified.
1.3 References 1-1 CPSES-1 WHIPJET Program Report.
1-2 Report of the U.S. Nuclear Regulatory Commission Piping Review Committee
~
- Evaluation of Potential for Pipe Breaks, NUREG 1061, Volume 3, Novec6er
~
1984.
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.
9 I
i mumm ie 1-3
l
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i SECTION 2.0 FAILURE CRITERIA FOR FLAWED ACCUMULATOR INJECTION N0ZZLES
~
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 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 injection nozzle 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 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.
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.
3732s/0428891C g.}
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:
Wa < Wp (2-1) where Wa = applied generalized load 1 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 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 I size, shape and loadings are parameters used in the evaluation of local failure.
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 J lc 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;c of the material, then the crack will not initiate.
1 If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:
, T,pp = h (2-2) m>= * *~ w 2-2
where T,pp
applied tearing modulus modulus of elasticity E
of -
=
' flow stress = (o y + u)/2 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<J Ic, r (2-3)
T,pp < Tmat, if J 1Jic (2-4) 2.4 References 2-1 J.D. Landes, et al., Editors, Elastic-Plastic Fracture, STP-668, ASTM, Philadelphia, PA 19109, November 1977.
l 2-2 J. C. Gerdeen, "A Critical Evaluation of Plastic Behaviu Data and a l Unified Definition of Plastic Loads for Pressure Components," Welding Research Council Bulletin No. 254.
'2-3 Hechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks, EPRI-NP-192, September 1976. I l
l l
1 I I
l
)
3732s/042889 10 2-3
i Wp = PLASTIC LOAD w p
l l
l c2 /
l 8
l
. S 5 l !
d I 5 I 5
e t I t I
I l
l l )
AP GENERALIZED DISPLACEMENT I
Figure 2-1 Schematic of Generalized Load-Deformation Behavior l I
l n u ..w a a io 2-4 l
j
i SECTION 3.0
)
OPERATION AND STABILITY OF THE ACCUMULATOR INJECTION N0ZZLES 3.1 Stress Corrosion Cracking )
I The Westinghouse reactor coolant system primary loop and connecting Class I lines (which include the accumulator injection nozzles) have an operating history that demonstrates the inherent operating stability characteristics of j the design. This includes a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress corrosion cracking). This i 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
~
seccnd 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 l
3732s/0428891C 3-1
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 cracking failure in the reactor coclant system loop or connecting Class 1 piping. The discussion below further qcalifies 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 immune to SCC as well as preventing the occurrence of a corrosive environment. The material specifications consider compatibility with the system's operating environment (both internal 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 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 j into service, the piping is cleaned internally and externally. During flushes and preoperational testing, water chemistry is controlled in accordance with l written specifications. Requirements on chlorides, fluorides, conductivity, and pH are included in the acceptance criteria for the piping. j l
,,,,,-..m 32
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 concentrations of chlorides and fluorides within the specifud 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 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 de:igned 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 i
relatively slow transients with no significant effect on the system dynamic '
loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlleo. 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 l
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 l the system and connecting accumulator lines. Preoperational, testing and I
vu..a.. w 3_3 {
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 faticue considerations are accounted for in the design of the piping system through the fatigue usage factor evaluation to show compliance with the ruics 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 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 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 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 Summary Evaluation of Accumulator Line for Potential Degradation During Service There has never been any service cracking or wall thinning identified in the accumulator lines of Westinghouse PWR design. Sources of such degradation are mitigated by the design, construction, inspection, and operation of the j accumulator lines.
There is no mechanism for water hammer in the accumulator pining system.
s Wall thinning by erosion and erosion-corrosion effects will not occur in the accumulator line (including the injection nozzles) due to the low velocity, typically less than 10 ft/see and the material, austenitic stainless steel,
(
i 3732vb4288310 3-4
which is highly resistive to these degradation mechanisms. Per NUREG-0691, a study of pipe cracking in PWR piping, only two incidents of wall thinning in j stainless steel pipe were reported and these were not in the accumulator I
line. Although it is not clear from the report, the cause of the wall-thinning was related to the high water velocity and is therefore not a mechanism which would affect the accumulator injection nozzles. l J
l Flow stratification, where low flow conditions permit cold and hot water to l
separate into distinct layers, can cause significant thermal fatigue I loadings. This was an important issue in PWR feedwater piping where temperature differences of 300*F were not uncommon under certain operational j 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 accumulator injection nozzles) where typically there is no flow during normal plant operation. During RHR operation the flow causes sufficient mixing to eliminate stratification.
Finally, the maximum operating temperature of the accumulator injection nozzles, which is about 560'F, is well below the temperature which would cause any creep damage in a stainless steel component.
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.
3-2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980.
m wu me,c 35
i SECTION 4.0 l .
MATERIAL CHARACTERIZATION OF THE ACCUMULATOR INJECTION N0ZZLES l
E 4.1 Nozzle Material and Welding Process l
The injection nozzle material is SA351-CF8A, a cast product form of the type used for primary loop piping of several PWR plants. The welding process used was shielded metal arc (SMAW).
l In the following section the tensile and fracture toughness properties for the above material are presented, and criteria for use in the leak-before-break analysis are defined.
4.2 Tensile Properties The material certifications for the nozzles were used to establish the tensile properties. These properties are given in table 4-1.
The properties in table 4-1 are those at 650*F. In the leak-before-break evaluation presented later, the minimum properties at operating temperature 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 accumulator injection nozzles is presented below.
From table 4-1 the average yield strength value of SA351-CF8A [
e j
i 1
1 mm..ui. io 41 j
_ . . . .. .. - - -- - A
e
]a,c.e The modulus of elasticity was obtained from the Nuclear Systems Materials Handbook (reference 4-1) for consistency with the stress-strain diagram which was also obtained from that reference. The stress strain curve (minimumproperties)isshowninfigure4-1. This curve is used in the crack stability analyses. In brief, the following material properties are the ones used in the analysis described in this report.
Minimum Properties for Flaw Stability Analysis a,c e
_ L Average Properties for Leak Rate' Calculations 1
_, _ l a,c.e i 4.3 Fracture Toughness Properties l
Because the accumulator injection nozzle is a cast product operating at 557'F, thermal aging will take place. [
]a,c.e the end of service life Charpy U-notch energy (KCU) following the procedure of reference 4-2. 1
[- k
),a,c.e By the criteria established in reference 4-3, the fracture toughness of the j SA351-CF5A is at least as great as the toughness of ( ],a,c,e the '
benchmark material of reference 4-3.
sm.msm ie 4-2 4 l
I
i Available data on aged stainless steel welds (References 4-4 and 4-5) indicate ;
the J Ic values for the worst case welds are of the same order as the aged
>^
[ Ja,c,e material. However, the slope of the J-R curve is steeper, and higher J-values have been obtained from fracture tests (in excess of 3000 in-lb/in2). The applied value of the J-integral for a flaw in the weld l
regions will be lower than that in the base metal because the yield stress for the weld materials is much higher at temperature. Therefore, weld regions are i less limiting than the cast material.
Therefore, the toughness values for LBB evaluation are established as those of
[ 3a,c.e.
a,c.e 4.4 References 4-1 Nuclear Systems Materials Handbook, ERDA Report TID 26666, November 1975, Part I, Group 1, Section 4.
4-2 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).
4-3 Toughness of Austenitic Stainless Steel Pipe Welds, EPRI NP-4768, Electric Power Research Institute, October 1986.
]
4-4 WCAP-10456, "The Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping for W NSSS," W Proprietary Class 2, November 1$83. l 4-5 Slama, G., Petrequin, P., Masson, S. H., and Mager, T. R., "Effect of
' .1 Aging on Mechanical Properties of Austenitic Stainless Steel Casting and '
Welds", presented at SMiRT 7 Post Conference Seminar 6 - Assuring Structural Integrity of Steel Reactor Pressure Boundary Components, I August 29/30, 1983, Monterey, CA.
3732s/042889 to 43
j t
N O A I E T R A A A A
% C A / / / /
U N N N N D N E I R
N O H I C T N
- A I A A A A
%G / / / /
N R N N N N O E L P E F E H
T 0 5
F 6 E H O T T )
T A G i S A M N s E I E p A A A A I S T R ( / / / /
T E L T N N N N R L U S E Z
. P Z O 0 1 R N S
- P T S 4 N E E L O S R E A I F T ) 2 1 5 4 L C T F S i 1 6 8 6 B I C O s 9 7 8 4 A N E D p 0 1 5 4 T A J % L ( 2 2 2 2 H N 2 E C I I E 0 Y
, M R O
E T L A A A A A B L 8 8 8 8 A U L F F F F L M A C C C C I U I - - - -
A C R 1 1 1 1 V C E 5 5 5 5 A A T 3 3 3 3 A A A A A M S S S S 1 1 1 2 0 0 3 3 9 9 9 9 1 1 1 1 R / / / /
T E 1 6 8 8 A B 7 7 7 7 E i 4 2 2 2 2 H U 3 3 3 3 N - - - -
3 3 3 3 e l
T b a
% C U M l e e e e l l l l
i D R z z z z a O O z z z z v R F P
o o o o a
, N N N N t
o P N O .
A OO /
L N 1 2 3 4 N
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cr-_--_ _. _. . _ _ _ _ -
TABLE 4-2
- CHEMISTRY AND END OF SERVICE LIFE KCU TOUGHNESS FOR ACCUMULATOR INJECTION N0Z2LES
=
m 8.C,e
-- J m
l i
3732s/042889 10 4-5 l
a,c.e 4
Figure 4-1 True Stress Strain Curve for SA351-CF8A Stainless Steel at 557'F i
sisasmsvu n 45
h SECTION 5.0 LOADS FOR FRACTURE MECHANICS ANALYSIS Figure 5-l'is a schematic layout of the an accumulator line including an accumulator injection nozzle.
The stresses due to axial loads and bending moments were calculated by the following equation:
a=f+{ (5.1) where, o = stress F = axial load M = bending moment
. A =
metal cross-sectional area
=
Z section modulus The bending moments for the desired loading combinations were calculated by the following equation:
M= My 2g 2 (5.2)
Z where, M =
bending moment for required loading My = Y component of bending moment h!
Z
= 2 component of bending moment C
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 5.2.
mamme se 5-1
5.1 Loads for Crack Stability Analysis The faulted loads for the crack stability analysis were calculated by the following equations:
F =
IF DW I + IF TH I + lF pl + IFSSE) (5.3)
My =
l(My )DWl + l(My )THl + l(M y)33g l (5.4)
N
- Z IINZ )DWI
- I(MZ)TH I + IINZ)SSEl (5.5)
Where, the subscripts of the above equations represent the following loading cases, DW = deadweight TH = normal thermal expansion 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 IEZ)DW * (N Z)TH (5.8) 5.3 Summary of Loads, Geometry and Materials Table 5-1 provides a summary of envelope loads computed for fracture mechanics evaluations in accordante with the methods described in sections 5.1 and 5.2.
The cross-sectional dimentions are also summarized, a
5.4 Governina Nozzle s
The normal plus SSE axial stresses for the four injection nozzles were compared. The maximum stress occurs at the loop 1 nozzle. The nozzle with sn w u m sic 5-2
the weakest material properties is also the loop 1 nozzle (see table 4-1). On this basis, the loop 1 injection nozzle is the limiting nozzle. Detailed fracture mechanics analysis were performed at this highest stressed nozzle.
e l
1 l
l l
l l
l l
l l
1 3732s ecd2889 tc 5-3
TABLE 5-1
SUMMARY
OF ENVELOPE LOADS FOR ACCUMULATOR INJECTION NO22LES I l .- 1 i
~ '
NOMINAL MINIMUM-
,. OUTSIDE WALL WALL INSIDE DIA THICK- THICK- DIA F M i LOOP CONDITION- '(inches) .(inches) (inches) (inches)(kips) (in-kips) l t 1
l 1 Faulted 10.75 1.000 0.8955 8.959 156 1180
.l Normal 10.75 1.000 0.8955 8.959 140 730 !
Operating i 2 Faulted 10.75 1.000 0.8955 8.959 147 770 Normal 10.75 1.000 0.8955 8.959 141 340 Operating f i
3 Faulted 10.75 1.000 0.8955 8.959 150 905 Normal 10.75 1.000 0.8955 8.959 141 372 Operating '
1 4 Faulted 10.75 1.000 0.8955 8.959 150 795 Normal 10.75 1.000 0.8955 8.959 141 522 Operating a
b 2732s/042689 10 -
5-4
l e
45* INJECTION ,
/
N0ZZLE /
/
COLD LEG LOOP 1
. i s
Figure 5-1. Schematic Laycut of Accumulator Injection Nozzle with Section of Accumulator Line sm.wass ic 5-5 ,
i o
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 [ 3a,c.e method, based on
' traditional plas' tic limit load concepts, but accounting for [
la,c.e and taking into account the presence of a flaw. The flawed component is predicted to fail when the remaining net section reaches a stress level.at which a plastic hinge is formed. The stress level at which this occurs is termed 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 accumulator injection nozzle.
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 section with internal pressure, axial force, and imposed bending moments. The limit moment for such a pipe is given by:
a,c.e
[ ] (6.1)
I where:
[
3a,c e
% J I
am. a.e io 6-1
[
]a,c,e (6.2)
The analytical model described above accurately accounts for the internal pressure as well as imposed axial force as they affect the limit moment. Good agreement was found between the analytical predictions and the experimental.
results(reference 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 accumulator 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]a,c.e, channel bothlength, [ L, to hydraulic diameter,H D ',c(l/0
.e H ) is greater than
[ ]a 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 [ ]a,c.e However, for large L/Dg values, friction pressure drop will become important and must be
, considered along with the momentum losses due to flashing.
3732s/0428891C g.g
l 1
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 [
.o pressure loss upstream of the choked exit plane, as suggested by Griffith ja,c.e ,
The flow rate through a crack was calculated in the folloviing manner. Figure 6-2 from reference 6-2 was used to est kate the critical pressure, Pc, for the l primary loop enthalpy condition and an assumed flow. Once Pc was found for a given mass flow, the [ Ja,c.e was found from figure 6-3 taken from reference 6-2. For all cases considered, since [ Ja,c.e Therefore, this method will yield the two phase pressure drop due to momentum effects as illustrated in figure 6-4. Now using the assumed flow rate, G, the frictional pressure drop can be calculated using
~
Ja,c.e aPf=[ (6.3) where the friction factor f is determined using the [ )">C
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 [
la,c.e to obtain the total pressure drop from the system under consideration to the atmcsphere. Thus, Absolute Pressure - 14.7 = [ , Ja,c.e(6.4) r i
me,w2nno 6-3
i 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 I
to within an acceptable tolerance and this results in the flow value through I
the crack. This calculational procedure has been recommended by [
la,c.e for this type of [
Ja,c.e calculation.
6.2.3 Leak Rate Calculations Leak rate calculations were made as a function of postulated through-wall l crack length for the critical location previously identified. The crack l 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 l rates were calculated using the normal operating loads of axial force F and bending moment M at the governing nozzle identified in section 5.0. These loads are given directly below.
. F = 140 kips, M = 730 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 long.
Thus the reference flaw size of [ ,)a,c,e is established.
6.3 Stability Evaluation A typical segment of the nozzle under maximum loads of axial force F and f
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 figures 6-6. The critical flaw size corresponds to the inter:oction of this curve and the maximum load line. The critical flaw size is calculated using the lower bound base metal tensile
, properties established in section 4.0. From figure 6-6 the critical flaw size ,
is seen to be [. ,)a,c.e for the base metal. 1 I
mu m n 6-4 j
\
WESTIN2 house PROPRIETARY CLASS 2 l
i 1
l The weld at the location of interest (i.e. the governing location) is a SMAW )
weld. Therefore, a "Z" factor correction for SMAW welds was applied
.. (references 6-5 and 6-6) as follows: j Z = 1.15 [1 + 0.013 (0.D. - 4)) (6.5) where OD is the outer diameter of the nozzle in inches. Substituting l OD = 10.75 inches, the Z factor was calculated to be 1.251. The applied loads -
l were increased by the Z factor and the plot of limit load versus crack length was regenerated as shown in figure 6-7. A flow stress of 51 ksi (reference 6-7)'was used. From figure 6-7, the critical flaw size is seen to be
[9.0 inches]8'C long. Noting that the flaw yielding a leakage of 10 gpm (i.e. leakage size flaw) was calculated to be [2.95 inches)a,c.e long, a factor of 3 exists between the leakage size flaw and the critical flaw. Thus, a margin of greater than 2 on flaw size is in evidence.
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 nozzle identified in section 5.0, the (normal plus SSE) outer surface axial stress; a, is seen to be 24.3 ksi based on the minimum wall thickness. The (normal plus SSE) axial force and bending moment are Fx = 156 kips and Mb = 1180 in-kips.
The minimum yield strength for flaw stability analysis is [21.9)a,c,e ksi (see section 4). The EPRI elastic plastic fracture handbook method is used to calculate the J applied using the normal plus SSE loads. The J applied was calculated for a [5.9 inches)a,c.e long postulated through wall flaw (which
, is 2 times the reference flaw size) and was found to be [2067 in-lb/in2)a,c.e ,
2 Since the J,pp-value is greater than the JIC-value of 750 in-lb/in , the tearing modulus was evaluated. The applied tearing modulus. T was applied ,
found to be 33. Both are well below the allowable of + *'#
]
~
)giveninsection4.0ofthisreport. +
a,c,e Therefore, unstable crack propagation will not result.
m w oscueio 6-5
6.5 References l
,, 6-1 Kanninen, M. F. et al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks" EPRI NP-192,
, September 1976.
)
1 6-2 [Fauske, H. K., " Critical Two-Phase, Steam Water Flows," Proceedings of the Heat Transfer and Fluid Mechanics Institute, Stanford, California, Standford University Press, 1961.]a,c.e )
6-3 [Griffith, P., " Choked Two-Phase Flow," Notes from an MIT Summer Course, 1973. )a , 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 NRC letter from M. A. Miller to Georgia Power Company, J. P. O'Reilly.
dated September 9,1987.
6-6 ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640.
6-7 Leak-Before-Break Evaluation Procedure, Standard Review Plan 3.6.3.
0 3
mwr noe ,o 6-6 v
l
)
l
.e
{
t i
1 i
l l
a,c,e
- l
, Figure 6-1 Fully Plastic Stress Distribution r
6-7
i
~
, i
{
_ 8.c s t 1
i in Figure 6-2 Analytical Predictions of Critical Flow Rates of Steam-Water 1 1
Mixtures {
(
mwunse to l 59 <
t a c.:
r A
h figure 6-3 [ ya,c e Pressure Ratio as a function of L/D 3732stytygg10 6-9
.- V 1
l a,c.e 1
a,c.e
/ y
~
1
-__- _- :=
i
' I l
- L/Dg = 40 4
- --- L/Cg *M e o : L :
9 Figure 6-4. Idealized Pressure Drop Profile Through a Postulated Crack ,
1 I
an2, mares in 6-10
_ _ . _ _ . _______u
< 1
' .w 4
f, 6 6 ,
sg p# l 4 4l
- 9 b 1 i-y .S
- O N .
ed .
1 E .
er= m a .E .E i o O .
l $$3 N M *-4
{
18 ' H !!
- a. w x '
I l
- i l
I i l
.' l l l
l e i j a L -- f J l l l l I I
l I
I l
l l l l ) 1 Figure 6-5. Loads Acting on the Model at the Governing Nozzle vn.mme so g_y3 l
+ a,c.e 0D = 10.75 in.
..~
t = 0.8955 in.
F = 156 kips
, oy = 21.9 ksi ou = 65.2 ksi.
l' i
s Figure 6-6. Critical Flaw Size Prediction Using Limit Load Approach me.mme io 6-12 1
.i l a,c.e 1 l- OD = 10.75 in.
l t = 0.8955 in.
F = 195 kips of = 51.0 ksi
/
1 1
Figure 6-7. Critical Flaw Size Prediction Including Application of Z-Factor l l
i m .
- m eto 6-13 l 1
1
l SECTION 7.0
, ASSESSMENT OF FATIGUE CRACK GROWTH .
,,. The fatigue crack growth on the Comanche Peak Unit 1 accumulator injection nozzle was determined by 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 Comanche Peak Unit 1 and generic, it was concluded that the generic analysis would envelop the fatigue crack growth of the governing accumulator nozzle identified in section 5.0.
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.
. The following summarizes the parameters which were compared:
Comanche Peak Generic Cold Leg Nozzle Accumulator Critical Location to Pipe Weld Injection Nozzle Pipe Outer Diameter 10.75" 10.75" Thickness .895" .896" Material Austenitic Stainless Steel Austenitic Stainless Steel Normai Temperature 550*F 557'F Normal Pressure 2300 psia 2250 psia Normal Operating 10.1 ksi 16.6 ksi Stress (Pres + Dwt Thermal Exp.)
Thermal Transients See Appendix B
- Thermal transient loadings are nearly identical for the two projects. ,
i l
k' 3732s/042789 10 7.}
I
l
{
l This comparison demonstra.tes the many similarities between the Comanche Peak Unit 1 accumulator insection nozzles and the generic nozzle evaluation. The
,. only uncertainty ir. this comparison is the higher level of pressure plus thermal expansion stress for Comanche Peak Unit 1. Since this is essentially i
,,. a mean stress (steady state) this difference will have only a minor impact on the fatigue crack growth calculation by increasing the R ratio. 'It is judged that this would increase the fatigue crack growth by less than [ ).a,c.e j Applying ( Ja,c,e increase to the generic fatigue crack growth data (from appendix B) results in a final flaw depth of approximately [ Ja,c,e for an initial flaw of ( ).a c,e These results demonstrate that no l
significant fatigue crack growth will occur over the 40 year plant desigr, life ,
even for the largest postulated flaw.
7.1 Acceptability Fatiaue 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 acceptability, the crack growth results of the generic analysis presented in Appendix B ar9 used conservatively and are considered applicable to the Comanche Peak Unit 1 accumulator injection nozzles. Detailed discussion in support of this assumption has been provided in the previous section.
The maximum allo.vable 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. Estimated fatigue crack growth results are given in the previous section of this report. (
Ja,c.e is conservatively chosen as a basis for examining the acceptability of fatigue crack growth. [First 0.224 in, is 25% of the
, Ja,c,e Thus, the first criterion on flaw depth is satisfied, s
mwumoc 72
.The: worst case transient AK value'for the maximum. crack depth is (' la c,e LThe: 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.
e The expression for plastic zone size,pr , calculation is: [ Ja,c.e r = E ( AK P
flow )2 Thus, the plastic zone size is calculated to be [ . 'Ja,c,e The remaining ligament for the 0.224 in. deep end of-fatigue-life. flaw is 0.671 in. (i.e.. 895 - 0.224). Thus, the plastic zone size is less than the remaining ligament.
Based on the above, it is concluded that for the Comanche Peak Unit 1 accumulator injection nozzles, the fatigue crack growth during service will not be significant. -
7.2 ' References 7.1 Bamford, W.'H. " Fatigue Crack Growth of Stainless Steel Reactor Coolent 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.
. I i
l i
1 me. win. ie 7-3
l l
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 i critical location are summarized below:
In Secten 6.3 the " critical" flaw size using limit load method is calculated to be [ Ja,c e long. Using the IWB-3640 approach (i.e. "Z" factor approach), the critical flaw size at the governing location is found to be
[ Ja,c,e long. In section 6.4 it is demonstrated that a postulated
[ Ja,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 of course exceed [ ).a,c,e In Section 6.2 it is shown that at the critical location, a flaw of [
Ja,c,e would yield a leak rate of 10 gpm. Thus, there is a margin of at least 2 on flaw size and a margin of 10 with respect to the plant leak detection capability of 1 gpm.
In summary, relative to
- 1. Flaw Size
- a. A margin of at least 2 exists between the critical flaw and the flaw yielding a leak rate of 10 gpm.
{
- 2. If limit load is used as the basis for critical flaw size, larger l margin for global stability would result.
l
- 2. Leal Rate i A margin of 10 exists for the reference flaw [ -
Ja c.e betw2en
. calculated leak rate and the criteria of Regulatory Guide 1.45.
am,mm'
8-1 3
A summary comparison of criter'a and analytical results is given in Table 8-1. The criteria are seen to be met. !
l l
1 l
1 1
l m2.azns ,o g.g l l
l
WESTIN!H2USE PROPRIETARY CLASS 2 TABLE 8-1 COMPARISON OF RESULTS VS.' CRITERIA CRITERION RESULT ,
/
- 1. NUREG-1061 Volume 3 Net Section 5.2(h) - (Required margin of 2 demonstrated)
Margin on Flaw Size
- 2. NUREG-1061 Volume 3 Met Section 5.7 - (Margin of 10 on leak rate Margin on Leak Rate demonstrated)
- 3. NRC criteria on allowable Met fatigue crack growth (Actual depth of fatigue crack
< 60% wall thickness)
(Plastic zone size < remaining ligament) i
\
4 inwonau io 8-3
SECTION
9.0 CONCLUSION
S !
.s
'. This report justifies the applica' tion of " Leak-Before-Break" technology to the
{
accumulator injection nozzles for Comanche Peak Unit 1 as follows:
- a. Stress corrosion cracking is precluded by use of fracture resistant materials and controls on reactor coolant chemistry, temperature, pressure, and flow during normal operation.
- b. Water hammer should not occur in the RCS piping (primary loop and the attached class 1 auxiliary lines) because of system design, testing, and operational considerations. l
- c. The effects of low and high cycle fatigue on the integrity of the accumulator line were evaluated and shown acceptable.
, d. ' Ample margin exists between the leak rate of small stable flaws and the criterion of Reg. Guide 1.45.
e.- Ample margin exists between the small stable flaw sizes of item d and the critical flaw. I i
l The postulated reference flaw will be stable because of the ample margins in d and e and will leak at a detectable rate which will assure a safe plant ;
shutdown.
l Based on the above, it is concluded that the dynamic effects associated with accumulator injection nozzle postulated breaks should not be considered in the i design basis of Comanche Peak Unit 1.
i L
4 mm.mewo -
g.1 j
APPENDIX A LIMIT MDMENT (The internal stress system at the crack plane has to be in equilibrium with the applied loading, i.e., the hydrostatic pressure P, axial force F, and the i
bending momentb M . The angle 6 which identifies the point of stress inversion follows from the equilibrium of horizontal forces (See Figure A-1).
This is:
2
({ a + B) R ,tof -({-s)R,tof={R g P + F/2 Solving for B, 2 pp mR B=j+4ftof m The external bending moment at the instant of failure follows from the equilibrium of moments, which is most easily taken around the axis 1-1. Thus ;
Mb can be determined from l c e Mb*2 fR ,t (,/ 7+6 cosed6 2
+
g/
7-6 cosede) l or i
a,c.e Mb*2 f R,2t (2cos6 - sina) )
i un,mme to A-1
. 1! i ; I\
!lj e.
c, a -
a g
n i
d n
e
- B n
i k
c a
r C
l l
. a W
h g
u
. o r
h T
a h
t i
w e
p i
P 1
A e
. r u
g '
i F
t
- 'l,
>*ur _
ill lll. llt1I
l l
l APPENDIX B FATIGUE CRACK GROWTH CONSIDERATIONS 1
I i
1 1
l I
l I
4 O
3732s/042789 10 8-1 1
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 c 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 wall 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 stress re' ort design thermal transients (Section B.1.2),
1-0 analysis data on accumulator line thermal transient stresses (based on ASME Section III NB3600 rules) and the geometry were reviewed to select the worst location for the fatigue crack growth analysis. [
3a,c,e This location is selected as the worst location based on the following considerations:
i) the fatigue usage factor is highest.
ii) the stress due to thermal expansion is high.
iii) the effect of discontinuity due to undercut at weld will tend to increase the cyclical thermal transient loads.
, iv) the review of data shows that the 1-D thermal transient stresses in the accumulator line piping section are generally higher near the [
Ja,c,e 3732s/D4278910 .
g.g
r..
I-B.I.2 Design Transients The transient conditions selected for this evaluation are based on conservative estimates of the magnitude and the frequency of the temperature
.r 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 [.
]a,c.e These transients were selected on the basis of the following criteria:
+a,c,e l -
l l
l where,
+a,c,e B.I.3 Simolified Stress Analysis The simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients.
[
Ja,c.e The inside surface stress was calculated by the following
, equation which is similar to the transient portion of ASME Section III NB3600, Eq. 11:
S$=[ Ja,c.e (B.3) u n ,m m ese 93
where,
,, S4 = inside surface stress l
+a,c.e l
l P
['
]a,c e The maximum and minimum inside surface stresses were searched from the gS 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:
S01
- I ) (B.7)+a,c e, 502
- I ) (B.8)+a,c.e, 3731s/042780 10
4
)
where,
=.
- 50'1 utside surface' stress at time t,,x.
= i
_ 02 S
utside surface stress at time t min
. p. +a c,e, '
~
The material properties for the accumulator pipe (SA376 TP316) and the RCL
.[.
la.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 [
Ja,c.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 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. [
i
!' ]a,c.e This figure shows that the
, sir.plified method provides more conservative crack growth.
3732s/042789 10 g.5 t . . . .
l B.1.4 Nonlinear Stress Distribution for Severe Transients
[
Ja,c.e As mentioned earlier, the
, accumulator line section near the [ o i. Ja,c.e is the wo $! location for fatigue crack growth analysis. A schematic of the ac;umslator line geometry at this location, is shown in Figure B-2. [
~
ja,c.e B.1.5 OBE Loads The stresses due to DBE loads were neglected in the fatigue crack growth analysis since these loads are not expected to contrib.ute significantly to crack growth due to small number of cycles.
k 3732s/Dd278910 g,g
B.I.6 _ Total Stress for Fatigue Crack Growth The total through wall stress at a section was obtained by superimposing the 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 g owth at any point is given by the following equction:
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 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-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 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 b
im,Smne B-7
)
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 relation:
-h=CoaK" y (B.2.1) where "Co" and the exponent "n" are material properties, and AK; 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 ,y 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 AKy adjusted, and the 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 )y to be used in the crack growth analysis 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 from Reference 8-1 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:
o (x) = A0 + Ay{+A({}+A({)3 2 3 (B.2.2)
- n. ms to B-8
The stress intensity factor K;(e) was calculated at the deepest point of the crack using the following expression:
+a,c.e 2
(2.2.3)
_ +a,c.e
~
where 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/Kimu) on 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:
h=[ '
.. ) a,c.e (B.2.4) l l
l 3732s/D4276910 gg
I l
where K,ff =-(Ky ,,x) (1-R)1/2 K
Imin R=K Imax h=crackgrowthrateinmicro-inches / cycle J
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 circumferentially, 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 [-
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 flaw growth is significantly lower. For example, a postulated [ ,)a,c.e 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.
~
B.3 REFERENCES B-1 a,c.e B-2 ASME Section III, Division I-Appendices, 1983 Edition, July 1, 1983.
B-3 WECAN -- Westinghouse Electric Computer Analysis, User's Manual -- Volumes 1, II, III and IV, Westinghouse Research and Development Center, Pittsburgh, PA, Third Edition, 1982.
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 Mechanics ASTM STP 677, 1979, pp. 365-380.
inwunis io B-10
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.
J a
\
snwurmic B-11
TABLE B-1 1:
THERMAL TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH EVALUATION
,. Trans. No. of No. Description Occurrences
+a,c e 1
4 9
M b
is snwunse to y_3g
,-w- --- yvw ,,-w- ----vu y C.
N @
W$a +
M l I
l W
,e M C:s W -
MM i ce >
i O
- J U 1
M M
W 2
M w
EW
~ Cl3 i E -
M
^ Z CL w
N
+
E w
M M M OM E EW 4 w E
= mW M C W w -
W ~ CIC M
. (
- W h
W 8 E
8 W
M W M M M M W ul Cat M &
H M M -
>< W
< Q E -
M E
e
- LJ U
% W e Z
w w .
M h E E E
- I l 5
e B-13
I 1
1 TABLE B-3
- ENVELOPE NORMAL LOADS CONDITION a,c.e l
Normal Operating l
L 4
l
- l 1
i l
ana,=2ne ic B-14
_ _ _ _ _ _ _ -- _ _---- _ )
4
(
t 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
e 3722s1042849 10 B-15
.c see.e 1
l
\
l l
P l
.s Figure B-1 Comparison of Typical Maximum and Minimum Stress Profile Computed by Simplified [ ] +a,c.e mumsuo ,o g.3g
,o . .
- . i
- - 1. +a c.e j i
- 2. '
- -+ +a,c.e Accumulator Pipe e
s figure B-2 Schematic of Accumulator Line At
{ 3Y + a ,c .e 2J2?s TW3s130M7 it B-17
i aesee i
= .
e I
1
+a,c.e j 4
Figure B-3 [ ] and Minimum Stress Profile for Transient #10 1
22;tar t343643084710 B-18
-_--------__---_____j
1.
1.
's,c,L
,r- ,
1 l l
1 l
r c
A +a,c.e Figure B-4 [ ] Maximum and Minimum Stress Profile for Transient #11 sw, tm,umnt g_19
,,.. , ._ a.see l
l ..
l_ j, 1
l l
l l
l l' t k
= -
e 'j
- i
- s. .
9 +a.c.e Figure B-5 [ ] Naximum and Minimum Stress !
Profile for Transient #12 l
am. mu usut in B-20 ;
l l
1 1
l c' b y
il
_ 1 j
\
l 1
e
-s i
+ +a.c.e Figure B-6 [ ] Maximum and Minimum Stress Profile for Transient 0 4 B-21