ML20246E799
| ML20246E799 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 04/30/1989 |
| From: | Bamford W, Roidt M, Strauch P WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19302D819 | List: |
| References | |
| WCAP-12259, NUDOCS 8905120027 | |
| Download: ML20246E799 (61) | |
Text
- _- . _ _ _ _ _ _ -_ __ , . _ _ . _
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EVALUATION OF THERMAL STRATIFICATION FOR THE COMANCHE PEAK UNIT 1 RESIDUAL HEAT REMOVAL LINES
.s W. H. Bamford P. Strauch M. A. Roidt April 1989 Reviewed by: E8 m E. R.Wohnson .
Approved by: D' Approved $y:1< A t ~ E7-R. B. 'Patel, . Manager 5. 5. Palusamy, Manager Systems Structural Analysis Structural Materials and Development Engineering WESTINGHOUSE ELECTRIC CORPORATION Nuclear and Advanced Technology Division P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728
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36d4s/05038910
TABLE OF CONTENTS '
'Sec't ion ; Title Page
-s:
1.0 BACKGROUND
AND TRANSIENT ANALYSIS 1-1 1.1_ Description of the Genkai Phenomenon 1-1 1.2 Comparison of Genkai and Comanche Peak 1-2 1.3- Overall Assessment Approach 1 1.4 Development of Bounding 1ransients 1-4 2.0 STRESS' ANALYSES 2-1 2.1 Piping System Structural Analysis 2-1 l
2.2 Local Stress Due t'o Non-linear Thermal Gradient 2-3 I
e 2.3 Stress Results 2-4
- 3.0 ASME SECTION III FATIGUE USAGE FACTOR EVALUATION - 3-1 3.1 Code and Criteria 3-1 3.2- Previous Design Methods 3-1 3.3 Analysis for Thermal Stratification 3-1 3.4 Fatigue Usage Results 3-3 4.0 REASSESSMENT OF LEAK BEFORE BREAK 4-1 4.1 Background 4-1 4.2 Methodology 4-1 4.3' Material and Fracture Toughness Properties 4-1 d 4.4 Loading Conditions 4-2 4.5 Results 4-4 4.6 Conclusions 4-4 4.7 References 4-5
- 5.0
SUMMARY
AND CONCLUSIONS 5-1 APPENDIX A: COMANCHE PEAK THERMAL STRATIFICATION ANALYSIS A-1 APPENDIX B: COMPUTER CODES B-1 j
3644s/050389 10 jj
- __ __ _____D
SECTION
1.0 BACKGROUND
AND TRANSIENT ANALYSIS
- :With the-discovery of a crack in the Genkai Unit 1 RHR suction line in June of 1988, attention has been focused on the possibility of stratification occurring in the RHR suction lines. This incident led to the issuance of NRC Bulletin 88-08, Supplement 3: " Thermal Stresses in Piping Connected to Reactor Coolant Systems," on April 11, 1989. This report addresses this issue for the Comanche Peak Unit.1 RHR suction lines..
The assessment begins with a detailed description of the Genkai pnenomenon, and a comparison of the Genkai and Comanche Peak Unit 1 RHR suction line configurations. This comparison is very important, because it will support
- conclusions on the likelihood of such a transient occurring at Comanche Peak.
A review is then provided of the overall approach used to assure that the Comanche Peak RHR suction line design basis is unaffected by stratification, should it' occur. Detailed thermal, stress and fatigue analyses have been performed to assess the effects of bounding stratification events. The results are then used to qualify the RHR suction lines for leak-before-break.
1.1 Description of the Genkai Phenomenon Around 9 a.m. on June 6, 1988, there was an increase in the water flow into the drain sump in the reactor containment vessel of Kyushu Electric Power Company's Genkai Nuclear Power Plant Unit 1 (PWR, 599 MWe), then operating at rated output, and a pool of water was seen on the floor. Manual operation for shutdown of the reactor was therefore started at 1:15 p.m. on the same day, and the reactor was brought to complete shutdown by 5:20 p.m.
The fluid which leaked was the primary coolant, containing radioactivity. The water flow to the drain sump, which normally is approximately one liter per
~
hour, rose to about fif ty liters per hour as of 11 a.m. on June 6. Although the amount of leakage was below Kyushu Electric's safety regulation level, the 364ds/05038910
power company shut down the reactor as a precaution to investigate the source of the leakage. The amount of coolant that leaked out ultimately reached
~
about 1,100 liters, but there was no outside radioactivity. The leakage occurred in the RHR and SI lines attached to loop A. The plant started
~
initial commercial operation in October 1975. Results of the inspection conducted by Kyushu Electric Power Company showed that the point of leakage was in the unisolable section of the branch line from the main primary coolant piping as shown schematically in figure 1-1. It was confirmed that the coolant leaked from a pinhole with a diameter of about one millimeter, near the welded section (elbow to straight horizontal pipe) of the stainless pipe (SS 316TP), which has an outer diameter of 8.6 inches and thickness of 0.81 inches. The cause of cracking was determined to be high cycle thermal fatigue resulting from valve leakage. The section of the pipe was replaced and the
- plant was returned to power.
The cyclic loading which caused the cracking is believed to have occurred in the following way: The isolation valve developed a leak, which allowed hot water from the main loop to flow down the vertical leg, and through the valve. The actual leakage flow was small, and stratified at the top of the pipe, since it was hotter than the bulk water in the horizontal section of the pipe between the elbow and isolation valve. The warm water reached the valve, warming it, and when the valve reached about 385'F, it closed, stopping the leakage flow. Once the stratified flow was cut off, the valve temperature again cooled down and the leak recurred. As the stratified hot water reached the valve again, it closed again, and the cycle repeated. This led to a severe fatigue cycling which initiated and propagated the crack. This scenario was demonstrated in
- laboratory tests.
l
1.2 Comparison of Genkai and Comanche Peak Although the junction of the RHR lines with the main loop is basically the same in every Westinghouse plant, there are differences between Genkai and Comanche Peak, which supports the conclusion that the Genkai transient is unlikely to occur at Comanche Peak. A comparison of Genkai and Comanche Peak geometries is provided in table 1-1. Since the first isolation valve in both the Comanche Peak RHR lines is located at about the same distance from the hot leg as the Genkai isolation valve which led to the cracking incident, the question arises as to what would prevent such an event at Comanche Peak. The answer lies in the difference in pipe diameter between the two plants. The study of the Genkai cracking incident showed that the hot water frc,m the main loop was unable to reach the bottom of the vertical pipe because
- turbulence was limited to about five feet from the main loop junction. For Comanche Peak, the hot water will reach the bottom of the vertical pipe, because the pipe is larger in diameter (12 vs. 8 inches) and the vertical distance is much shorter (5.4 vs. 9 feet). The isolation valves in the two Comanche Peak RHR lines are nearly the same distance away from the vertical leg (2.3 and 2.0 feet for loop 1 and 4, as compared to 2-3 feet at Genkai).
The schematic layouts for RHR lines of loops 1 and 4 are shown in figures 1-2 and 1-3. When a pipe (such as the RHR suction line) is connected to a larger diameter pipe (in this case the hot leg) with high velocity turbulent flow, the turbulence will penetrate down the smaller pipe. The distance of penetration depends on the flow velocity (Reynolds number) in the larger pipe, and the relative diameters. A number of experiments have been performed at Westinghouse and Mitsubishi Heavy Industries to obtain this information, and
- the results are summarized in figure 1-4.
[0n the basis of these experiments, as presented in figure 1-4, main loop fluid velocity of 45 feet per second results in the turbulent penetration of hot water into the RHR line of at least 15 diameters. The presence of elbows has been included in the experimental data, with results obtained for up to nu, une in 1-3
two 90 degree elbows. Figure 1-4 shows that if the isolation valve is located within 15 diameters of the main loop, turbulence will occur up to the valve, and stratification is not possible. The closest isolation valve at Comanche Peak is 13 diameters from the hot leg, so stratification near the valve is very unlikely.]a,c.e The Comanche Peak isolation valves are closed using limit switches. A detail drawing of the Comanche Peak RHR isolation valve is shown in figure 1-S. 1.3 Overall Assessment Approach There are a number of actions which have been taken at Comanche Peak to ensure that no cracking will occur in the RHR lines. This includes volumetric inspection of the lines to assure that no cracks are present, as well as s inspection and testing of the isolation valves to ensure no leakage. Cracking is unlikely ever to occur at the Comanche Peak plant, due to systems
- differences with the Genkai plant.
Even though these precautions are sufficient to ensure that cyclic stratification leading to cracking will not occur in the Comanche Peak RHR lines, an assessment was carried out to quantify and evaluate the effects of postulated stratification on pipe integrity. 1.4 Development of Bounding Transients The development of a stratification transient for the Comanche Peak RHR suction lines began with the temperature profile applicable to the Genkai plant, as obtained experimentally by Mitsubishi Heavy Industries. The transient at Genkai was due to intermittent valve leakage, which provided a path for hot water to be drawn into the RHR line from the main loop. At the bottom of the vertical pipe, a stratified flow was established, with hot water
- filling the top 10 percent of the pipe.
To establish a stratified flow transient for the Comanche Peak plant, a simplified, yet conservative, representation of the Genkai temperature profile l
was assumed to exist in the horizontal portion of each line, after the isolation valve. The same portion of the pipe as at Genkai was assumed to be filled with 'eakage flow at Comanche Peak, which corresponds to a smaller flow rate in the pipe because the diameter is larger. The water at the bottom of the vertical elbow and up to the first valve was assumed to be at hot leg temperature. The water was assumed to stratify after reaching the first isolation valve. The water in the bottom of the pipe was stagnant, and was assumed to cool by a conduction-limited mechanism. The stratified flow at the top of the pipe does not cool as quickly because of its flow, ac shown in figure 1-6. This creates a rather large temperature difference between the top and bottom of the pipe, which is maximized at about four feet from the valve. As the flow continues, it gradually loses heat to the stagnant bulk fluid, and by a distance of 15 downstream of the vrive the temperature difference is very small. By 20 feet, fluid at both the top and bottom of the
~ pipe are within 10-15 degrees of ambient temperature of 120*F as shown in figure 1-6. The development of this temperature profile is provided in detail - in Appendix A.
A
- w. am to 1-5
TABLE 1-1 COMPARISON OF GENKAI AND COMANCHE PEAK PLANTS GENKAI COMANCHE PEAK Line Size 8 inch 12 inch Vertical Drop From RCS 9 feet 5.4 feet Distance to Isolation Valve 2-3 feet 2.3, 2 feet From Vertical Drop Total Length of Pipe, RCS to First 14-15 feet 11.6, 10.7 feet s Isolation Valve Length of Vertical Piping 5 feet Entire Length Experiencing Turbulence m am.eca io 1-6
N.
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k y1 4 L Figure 1-1. Sketch of the Cracking Location in the Genkai Unit 1 RHR Suction Line nu. mens o 1-7
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1 Figure 1-4. Experimental Results of Turbulent Penetration in a Pipe Connected to a High Flow larger Diameter Pipe i wwescus in 1-10
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1 i l SECTION 2.0 STRESS ANALYSES Flow diagram figure 2-1 describes the procedure to determine the effects of
. thermal stratification on the RHR suction lines based on the transient developed in soction 1.0. [
-ja,c,e Section 2.1 Addresses the structural or global effect of stratification Section 2.2 Addresses the local stress effects due to the nonlinear portion of the temperature profile 2.1 Piping System Structural Analysis 2.1.1 Introduction The thermal stratification computer analysis of the piping system to determine the pipe displacement, support reaction loads as well as moment and force loads in the piping is referred to as the piping system structural analysis.
These loads are used as input to the fatigue evaluation. The thermal stratification condition consists of both axial and top-to-bottom variations in the pipe metal temperature, as described in section 1.0. The model consists of straight pipe and elbow elements for the ANSYS computer code. [- l Ja,c,e These studies verified the suitability of the ANSYS computer code for the thermal stratification analysis. [ ja,c.e 3644s '0$038910 2-1
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. 2.1.2 Discussion The piping layout for the RHR suction line analyzed (loop 4) is shown in figu,e 2-2. [
Ja,c e The analyses for loop 4 will also be conservatively applicable to loop 1. The piping analysis'model consists of straight pipe and elbows as shown in figure 2-3. The elements provide the capability to load the piping
.with a top-to-bottom temperature gradient. Spring-damper elements were used at rigid support locations. Since only thermal expansion analysis was performed on the structural model, spring hangers were not included.
In order to determine the adequacy of the ANSYS structural model for stratification analysis, a benchmark analysis was performed by loading the model with the temperature profile for the normal operating condition (617'F from the RCS hot leg to valve 1-8702B, and 120*F beyond the valve, with no top-to-bottom temperature gradient in the pipe), and' comparing the results with the piping analysis results for the normal operating load case. Based on this comparison, the model was determined to be acceptable for thermal stratification analysis. Two thermal expansion loadings were applied to the ANSYS structural model in order to determine the effects of cyclic valve leakage on support loads, pipe loading and movements, and nozzle loads. The first case assumed no leakage, therefore no top-to-bottom temperature gradient. The axial temperature gradient along the line was determined by heat transfer analysis (Appendix A). The line was assumed to be at the RCS temperature between the hot leg connection and the valve due to recirculating flow in the RHR line. ( 3644sec5038910 2-2 I
The second analysis case assumed [
< ]a,c.e The axial temperature distribution of this leakage was also determined by heat transfer ~
analysis (Appendix A). The axial temperature distribution of the remaining
. stagnant water was assumed to be the same as the distribution of the no-leakage case. Therefore a top-to-bottom temperature gradient was input for this case as a step change at the leakage flow / stagnant water interface (figure 2-7).
For the ANSYS code an [ Ja,c,e transverse deflections in a free pipe as the desired step change profiles. 2.2 Local Stress Due to Non-Linear Thermal Gradient 2.2.1 Explanation of Local Stress Figure 2-5 shows the local axial stress components in a beam with a sharply nonlinear metal temperature gradient. Local axial stresses develop due to the restraint of axial expansion or contraction. This restraint is provided by the material in the adjacent beam cross section. For a linear top-to-bottom temperature gradient, the local axial stress would not exi t. [ i n ja,c.e 2.2.2 Superposition of Local and Structural Stresses I . For the purpose of this discussion, the stress resulting from the global structural analysis (section 2.1) will be referred to as " structural stress." am. :s:"' " 2-3
[ i Ja,c.e Local and structural stresses may be
..- superimposed to obtain the total stress. This is true because linear elastic analyses are performed and the two stresses are independent of one another.
Figure 2-6, presents the results of a test case that was performed to demonstrate the validity of superposition. As shown-in the figure, the super-position of local'and structural stress is valid. [ ja,c e 2.2.3 Finite Element Model of Pipe for Local Stress The pipe finite element model is shown'in figure 2-7, along with thermal boundary conditions. The entire cross section was used for modeling and analysis. [ ja,c.e I 2.3 Stress Results l The temperature and stres,s results for the local finite element model are presented in the color plots on in figures 2-8, 2-9 and 2-10. [ 3a,c.e The high
~
temperature region is very localized at the top of the pipe, as expected, and l the pipe wall temperature quickly drops to the stagnant water temperature of about 260'F, for the majority of the circumference. The axial stresses from i 364ds/05C389 to p.4
l this stratified flow are shown in figure 2-9, anc' it can easily be seen that the highest stress (red region) is near the hot-cold water interface, and is a positive 17.4 ksi. Compressive stresses are found at both the top and bottom of the pipe. The stress intensity (next figure, 2-10) was highest at the top [~ of the pipe, at a value of 27.63 ksi. These stresses were combined with the global stresses discussed in section 2.1 to obtain the total stresses from the postulated stratification event. This stress range was computed at six different locations along the line, and used in the fatigue analysis to be discussed in section 3. l l l 3644s 05C3691 2-5
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Figure 2-1. Determination of the Effects of Thermal Stratification 3644s/060383 to 2-6
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OT LEG LOOP 4 l 4.0' i l o 4.7' Y b X :
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.2' Figure 2-2. Piping System Isometric Drawing - RHR Loop 4 3644e109036910 _
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SECTION 3.0 ASME SECTION III FATIGUE USAGE FACTOR EVALUATION
. 3.1 Code and Criteria Fatigue usage factors for the Comanche Peak Unit 1 RHR suction line in loop 4 were evaluated based on the requirements of the ASME B & PV Code, Section III, Subsection NB-3600, for piping components. The fatigue evaluation required for level A and B service limits in NB-3653 is summarized in table 3-1. ASME III fatigue usage factors were calculated for [ Ja,c.e points in the RHR line piping.
3.2 Previous Design Methods Previous evaluations of RHR suction line piping fatigue used the NB-3653 techniques but with thermal transients defined by Westinghouse design specifications, assuming the fluid flows to sweep the RHR line piping with an axisymmetric temperature loading on the pipe inside wall, and that no stratified flow due to valve leakage occurred. These evaluations produced typical usage factors of approximately [ ]a,c.e at elbows and bends, [ la,c,e at valve ends, and [ Ja,c.e at the RCL hot leg nozzle safe end. It must be noted that these usage factors are conservative since, in the design process, calculations are carried to the point where results meet code requirements, and are not further refined to reduce the usage factor. 3.3 Analysis for Thermal Stratification With the thermal transient redefined to account for thermal stratification as described in section 1.0, the stresses in the piping components were established (section 2) and new fatigue usage factors were calculated. Due to the non-axisymmetric nature of the stratification leading, stresses due to all loadings were obtained from finite element analysis and then combined on a stress component basis. w w n m to 3-1
Stresses in the pipe wall due to internal pressure and thermal stratification loading were obtained from the WECAN 2-D analysis of a 12 inch, schedule 140 pipe. [ l ja,c.e Two types of stress were calculated - Sn(Eg10),todetermineelastic-plastic penalty factors, K,, and Sp (Eq 11) peak stress. For most components in the RHR line (girth butt welds, elbows, bends) no gross structural discontinuities are present. As a result, the code-defined "Q" stress (NB-3200), or C 3ElaaT, abTb lin Eq (10) of NB-3600 is zero. Therefore, for these components, the Eq. (10) stress s are due to pressure and moment. Peak stresses, including the total surface stress from all loadings - pressure, moment, stratification - were then calculated for the transient. (
,)a,c.e This evaluation uses the S n and Sp stresses calculated for the transient to determine usage factors at selected locations in the pipe cross section.
Using a standard ASME method, the cumulative damage calculation is performed according to NB-3222.4(e)(5). The inside and outside pipe wall usage factors were evaluated at [ la,c.e along the pipe model. This includes:
- 1) Calculating the S nand S pranges, K,, and Salt f r every combination of the transient loads.
- 2) For each value of Salt, use the design fatigue curve to determine the maximum number of cycles which would be allowable if this type of wwescue to 3-2 4
i cycle were the only one acting. These values, Ny , N 2 ...N n ' were determined from Code figures 1-9.2.1 and I-9.2.2, for austenitic stainless steels.
- 3) Using the actual cycles of the transient loadset, calculate the-additional contribution to the usage factor Uy, from the postulated stratification. If N $is greater than 10 11 cycles, the value of U$ is taken as zero.
- 4) The cumulative usage factor, Ucum, is calculated as Ucum *U1+
U design. The code allcwable value is 1.0. 3.4 Fatigue Usage Results A stress analysis was completed for the stratified flow condition, including local thermal stresses and global piping stresses resulting from the stratification. Deadweight and pressure stresses were constant, so they were not included since they would not contribute to the alternating stress. The transient analyzed alternated from stratified flow to stagnant flow. For stratified flow, the two curves of figure 1-6 were used, while for stagnant flow only the bottom curve was used for the bulk temperature. The criteria used are shown in table 3-1. The analysis was performed on loop 4 because it is most stiffly supported and was therefore expected to have the highest stresses. The fatigue results are shown in table 3-2, for the [ Ja.c,e most important locations. The alternating stress is such that it is below the endurance limit for the material for all the locations. Therefore, the cumulative usage factor for the design is the same whether or not the stratification is included. In addition to the usage factor determination, a comparison was made with the
- l. stress limits for equation 12, in Section III of the ASME Code. All the
, results are far below the 3S, limit of 50 ksi as shown in table 3-3. For
- t. information, a nozzle load comparison has been included, so the forces and 364J s <0$038910 3-3
moments on the two nozzles in the RHR suction line can be compared with and without the postulated leak to the original analysis cases without stratification called " Design Normal Operation". [ t
,)a,c.e s
9
TABLE 3-1 CODE / CRITERIA f
. o ASME B&PV Code, Sec. III, 1986 Edition
- NB3600 NB3200 o Level A/B Service Limits Primary Plus Secondary Stress Intensity 5 3Sm (Eq. 10)
Simplified Elastic-Plastic Analysis (when Eq. 10 > 3 S,) Expansion Stress, Se5 35m (Eq.12) - Global Analysis
- Primary Plus Secondary Excluding Thermal Bending < 3Sm
- (Eq. 13)
Elastic-Plastic Penalty Factor 1.0 $ K ,1 3.333 Peak Stress (Eq. 11)/ Cumulative Usage Factor (Ucum) S
- Kep S /2 (Eq. 14) alt Design Fatigue Curve Ucum 5 1.0 e
3644s /05C38910 3-5
TABLE 3-2 FATIGUE RESULTS - COMANCHE PEAK UNIT 1, LOOP 4 ALTERNATING STRESS CUF (Design) CUF (Total) LOCATION
- - a,c.e
~
- 1. Node 2020 7.5 ksi
- 2. Node 2090 4.4
- 3. Node 2180 8.3
- 4. Node 2265 4.4
- 5. Node 2391 8.5
- 6. Node 2530 3.1 W -
The alternating stress is below the endurance limit stress of 16.8 ksi and therefore does not contribute to the cumulative usage factor. sr as44rosc1so io 3-6
TABLE 3-3 ASME SECTION III EQUATION 12 STRESS INTENSITY is . Node Eq.12 Stress (ksi) 3 Sm Limit
~ ~
C'* 50.1 2020 2090 50.1 2180 50.1 2265 50.1 2391 50.1 2530 50.1
' The equation 12 stresses above are the maximum range stresses between leakage or non-leakage car,es with zero load set, or the range between the leakage and non-leakage cases.
DM g Eq. 12: SE=C2T <3S, C2 = M ment Loading Stress Index D,= Outer Diameter (12.75 in.) M = Moment Range Due to Thermal Expansion 4 I = Moment of Inertia (701 in ) S ,= Design Stress Intensity Limit (16.70 ksi at 650*F) m> w.,esen. e 37
TABLE 3-4 N0ZZLE LOAD COMPARISON (KIPS, IN-KIPS) Loop Nozzle Node 9424 g Fy Fz Fr Mx My Mz Mr Design Normal Operation 5.14 -1.63 -9.67 11.1 623.4 -746.3 441.0 1067.7 a,c.e 8.45 -2.50 -12.1 15.0 348.6 -984.9 681.1 1247.2 7.79 -1.57 -10.2 12.9 698.8 -840.9 633.0 1263.3 e 3644 s '0$C38910 3-8
o i SECTION 4.0 REASSESSMENT OF LEAK-BEFORE-BREAK t - 4.1 Background ~ A leak-before-break evaluation was performed for the 12 inch RHR line of the Comanche Peak nuclear power plant in April 1988 (reference 1). TV Electric subsequently submitted reference 1 to the Nuclear Regulatory Commission (NRC) as a support document for implementor.g leak-before-break for auxiliary high energy lines. The recent concern for thermal stratification in RHR lines has prompted the analyses presented in this report. Specifically, thermal stratification has been shown to impact normal operating loads. In this section, leak-before-break for the RHR lines is reassessed taking into account thermal stratification. The leak-before-break methodology is reviewed, the analyses are'summarired, and conclusions are drawn.
~
4.2. Methodology The steps of the leak-before-break methodology are reviewed in table 4-1. Items '2 and 3 are addressed in sections 2 and 1 of this report, respectively. This section addresses the remaining items. The conservatism used in this section are listed in table 4-2. 4.3 Material and Fracture Touchness Properties ) i Applicable material properties were obtained from the code minimum properties [2] for the materials, which are given in table 4-3. The material is SA376 TP304, a wrought product form, of the type used in the primary loop of many 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 shielded metal arc (SMAW). 36446 #5C38910 4.}
-l The stress strain curve required for the stability analyses is given in figure-4-1. The' curve at 617'F was obtained by application of the Nuclear Systems-Materials Handbook (reference 3). In brief, the following material properties 'r.
were used in the analyses set forth in this report. 1 L Minimum Properties for Flaw Stability Analysis (617'F) a,c.e Average Properties for Leak Rate Calculations (617*F) a,c.e
.~
b Fracture toughness properties are given in table 4-4 taken from reference 4 through 7. Conservative estimates of toughness were chosen by using the material footnoted by d. 4.4 Loading Conditions The stresses due to axial loads and bending moments were calculated by the j following equation: l o={+y e 1 36Ms/050389 to 4-2
s g l 1 'where, . p , b- o = stress F = axial load .; j; M. = bending moment. A = metal cross-sectional area ) Z = section modulus ! i I ,
.The bending moments for the desired loading combinations were calculated by the following equation: .; ! l M = [M y 2+M7 Y where, )
M.. = . bending moment for required loading
]
My = Y component of bending moment
-M = .Z component of bending moment Z
The faulted loads for the crack stability analysis were calculated by the j following equat' ions: l l F = lFDW + F TH + Fp l + IFSSE I (5.3) , My = l(My )DW + INY )THl + l(M y)SSEl (5.4) M Z IINZ)DW + (MZ )THI +'I(N Z)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 1 auwosesse so 4-3
The normal operating loads for leak rate pred'. ! %as were calculated by the following equations: F = FDW + FTH + Fp (5.6)
; N Y (NY )DW + (NY)TH (5.7)
M 7 (Hz )DW + INZ)TH (5.8) Table 4-6 provides a summary of envelope loads computed for fracture mechanics evaluations. The cross-sectional dimensions and materials are also summarized. Load data are tabulated at the highest stressed location (node 21S0, loop 4), and the second highest stressed location (node 2020, loop 4). Also included is the highest stress location in the portion of the line where
., stratification was postulated, node 2391.
4.5 Results L
.. Comanche Peak Unit 1 employs an administrative shutdown specification of 1.0 gpm unidentified leakage in response to Regulatory Guide 1.45. The leakage size flaw then is the one giving 10 gpm. Leakage flaws were calculated using the methodology of section 5.0 of reference 8. The results are given in table 4-7.
Critical flaw sizes were obtained using the limit load procedure as outlined in Section XI paragraph IWB 3640 [2] and in SRP 3.6.3 (reference 9), accounting for the welds by using the appropriate 2-formula. The instability flaw sizes so determined are also given in table 4-7. Stability margins on leakage flaws in excess of 2 are again demonstrated, i 4.6 Conclusions i ) I l Considering the results of the prior sections and this section, the LBB
, criteria outlined in table 4-1 have been met and thus LBB has been demonstrated for the Comanche Peak Unit 1 RHR line considering thermal . stratification; specifically, o LBB exists at operating temperature without stratification.
o LBB exists at operating temperature with stratification. l mwosens io 44 l 1
4 14'.71 - References l' . .Whipjet Program' Final Report, Comanene Peak Electric Station, Unit 1, L R. L. Cloud and Associates, Inc. April 1988. Y 2 .- -1986 ASME Boiler and Pressure Vessel Code, Nuclear Power Plant p Components, Section III Division 1, Appendices, and Section XI, Division 1. l
- 3. Nuclear Systems Materials Handbook, Part I - Structura_1 Materials, Group ,
1 - High Alloy Steels, Section 2, ERDA Report TID 26666, November 1975 Revision.
- 4. Kanninen, M. F., et al., " Instability Predictions of Circumferential1y Cracked 304 Stainless Steel Pipes Under Dynamic Loading," EPRI-NP-2347, April 1982. '
- 5. Bamford, W. H., and Bush A. J., " Fracture Behavior of Stainless Steel,"
in Elastic Plastic Fracture, ASTM STP 668, 1979. l,
- 6. Palusamy, S. S., Tensile and Toughness Properties of Primary Piping Weld Metal for Use-in Mechanistic Fracture Evaluation, WCAP-9787, May, 1981 (Westinghouse Proprietary Class 2).
l
- 7. Bamford, et al., The Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping for Westinghouse Nuclear Steam Supply Systems, WCAP-10456, November, 1983 (Westinghouse Proprietary Class 2). j i
- 8. Witt, F. J., et al., Technical Justification for Eliminating Large Primary Loop Pipe Rupture as the Structural Design Basis for Beaver Valley Unit 2 after Reduction of Snubbers, WCAP-11923, September 1988 (Westinghouse Proprietary Class 2), i m
)
- 9. USNRC Standard Review Plan 3.6.3, Leak-Before-Break Evaluation o Procedures, NUREG-0800.
L au, mens io 4-5
t
+ ,
L TABLE 4-1 I- :c -STEPS _IN-A LEAK-BEFORE-BREAK ANALYSIS-- L. (1) Establish material-properties including fracture toughness' values
. (2), Perform stress analysis of the structure
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(3). Review operating history of the. structure (4) Select locations for postulating flaws t (5)L Determine a flaw size giving a detectable. leak rate
~(6) . Establish. stability of.the-selected flaw L (7) Establish adequate margins in terms of leak rate detection, flaw size and load.
(8) Show that a flaw indication acceptable by inspection remains small throughout service life.
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o TABLE 4 LBB CONSERVATISM-t li l 1 o factor of 10'on Leak: i Rate '
. , ,- o Factor ~of 2 on Leakage Flaw'
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o ' Algebraic Sum-of. Loads for Leakage o Factor of 1.4 on Loads for Stability o . Average Material Properties for' Leakage-o Minimum Material-Properties for. Stability o i' J.. i i am..eusn io . 47
TABLE 4-3 RHR LINE MATERIALS AND WELOS OF THE COMANCHE PEAK UNIT 1 PLANT OUTSIDE WALL OPERATING
- DIAMETER THICKNESS BASE TEMP PRESS LOOP NO. LINE NUMBER (IN) (IN) MATERIAL (F) (PSI) 1 12-RC-1-007-2501R-1 12.750 1.125 SA376 TYPE 316 617 -2235 1 12-RH-1-001-2501R-1 12.750 1.125 SA376 TYPE 316 350 2235 1 12-RH-1-901-2501R-1 12.750 1.125 SA376 TYPE 316 350 400 1 12-RH-1-003-0601R-2 12.750 1.000* SA376 TYPE 316** 350 400-4 12-RC-1-069-2501R-1 12.750 1.125 SA376 TYPE 316 -617 2235 4 12-RH-1-002-2501R-1 12.750 1.125 SA376 TYPE 316 350 2235 4 12-RH-1-900-2501R-1 12.750 '1.125 SA376 TYPE 316 350 400 4 12-RH-1-004-0601R-2 12.750 1.000 SA376 TYPE 316** 350 400
- Portions of line are Schedule 40S; does not affect LBB calculations
** Small portions of line contains SA316 TYPE 304 stainless steel ASME Code Minimum Strength Requirements (psi)
Yield Ultimate f Strength Strength Pipe SA376 TP316 30,000 75,000 i Weld E308 - 80,000 Nozzle SA182 F316N 35,000 80,000 m 43 l L__-___-_______________________
l TABLE 4-4 I FRACTURE-TOUGHNESS' PROPERTIES FOR 304 STAINLESS STEELS AND WELDS-
' J Test Ie 2 T '
. Material Temperature (*F) .(in-lb/in )- mat Reference-
,a,c,e
- SA376 TP304 550 4 V
SA376-TP304 600 5 Weld 600 6 d Weld 600 7 m
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_ _ _ _ _ __ _ _ _ l c:
. TABLE'4'-5 TYPES OF LOADINGS Pressure (P)'
1: Dead Weight-(DW) Normal Operating Thermal. Expansion (TH) Safe Shutdown Earthquake and Seismic Anchor Motion (SSE)a.
- a,c.e -
Gw' aS SE is used to' refer to'the' absolute sum of these loadings. a L, 1
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iw a u gescanio 4-10 i
L TABLE 4-6
SUMMARY
.OF LDADS AND STRESSES AT THE CRITICAL LOCATIONS I
.- 00 = 12 in., Minimum Wall Thickness = 1.246 in.
l
$ Force Dead Moment Total i Node- Case F (1bs) Weight (lb) .M (in-lbs) Stress (psi) !
s,c.e j.. 2180 with leak 243550 130 i n 2180 without leak 245550 130' i 2020 with leak 228990 2990 ' 2020 without leak 229770 2990 ; 2391 with leak 221710 130 2391 without leak 223610 130 :
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- - _ _ _ _ _ _ _ _ _ - _ _ _ s
TABLE 4-7 ) RESULTS
SUMMARY
LEAK-BEFORE-BREAK ASSESSMENT COMANCHE PEAK UNIT 1 q i
.-. Temperature Temperature Distribution Distribution l) w/o Leak with Leak (inches) (inches) Comment N0DE 2391
- - a,c,e
- Leakage Size Flaw Factor of 10 with respect to leak detection capability met.
Critical Flaw Based Factor of 2 on flaw on IWB-3640 Method size - met.
& Critical Flaw with- Factor of 1.4 on 1.4 Factor on Load Based - - load - met.
on IWB-3640 Approach
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N00E 2180 a,c.e Leakage Size Flaw Factor of 10 with respect to leak detection capability ! met. ! Critical Flaw Based on Factor of 2 on IWB-3640 Method flaw size - met. Critical Flaw with Factor of 1.4 on 1.4 Factor on Load Based - - load - met. on IWB-3640 Approach NODE 2020 a,c.e Leakage Size Flaw Factor of 10 with respect to leak detection capability
- met.
Critical Flaw Based on Factor of 2 on flaw IWB-3640 Method size - met.
.- Critical Flaw with 1.4 Factor of 1.4 on Factor on Load Based on - -
load - met. IWB-3640 Approach nuuescue to 4-12
r_._ _ l
. 1 1
l
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a,c.e l l l l l ' l 1 l ! ) l 1 l l l I l 1 _ l l l n Figure 4-1. Minimum True Stress-True Strain Curve for Type 316 Stainless Steel at 617*F nu,<csaas ' 4-13
SECTION 5
SUMMARY
AND CONCLUSIONS
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'A detailed evaluation of the residual heat removal lines for the Comanche Peak Unit 1 plant has been completed. The evaluation has included all the design basis transients for the system, as well as a postulated thermal i
stratification transient. The thermal stratification transient has been included in response to concerns raised by a cracking incident which occurred at a plant in Japan. Although geometrical differences between this plant and the Comanche Peak Unit 1 make such a cracking incident very unlikely, the stratification transient was postulated for completeness.
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i,c,e l 4, A leak before break evaluation showed that the RHR lines meet all the necessary requirements and margins, whether stratification is present or not. j Thus we have shown that the integrity of the RHR lines are assured, whether or not stratification were to occur. I i i W O 3644s/05C38910 5-1
h L APPENDIX A COMANCHE PEAK THERMAL STRATIFICATION ANALYSIS
.A It is of interest to estimate maximum temperature differences between
.. stratified layers of fluid in horizontal piping layouts for the purpose of
' studying the propensity of a given configuration .for developing high stresses at the inner radii of the pipes. The through-wall crack developed at GENKAI and the cursory similarity between that plant and the Comanche Peak unit would indicate an evaluation of the possibility of a similar occurrence at the latter unit.
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APPENDIX B COMPUTER CODES I This section lists and summarizes the computer codes used in the analysis of L, stratification in the Comanche Peak Unit 1 RHR lines:
- 1. WECAN
- 2. ANSYS B.1 WECAN B.1.1 Description WECAN is a Westinghouse-developed, general purpose finite element program. It contains universally accepted two-dimensional and three-dimensional I* isoparametric elements that can be used in many different types of finite i'
element analyses. Quadrilateral and triangular structural elements are used for plane strain, plane stress, and axisymmetric analyses. Brick and wedge structural elements are used for three-dimensional analyses. Companion heat conduction elements are used for steady state heat conduction analyses and transient heat conduction analyses. B.I.2 Feature Used The temperatures obtained from a static heat conduction analysis, or at a specific time in a transient heat conduction analysis, can be automatically input to a static structural analysis where the hect conduction elements are replaced by corresponding structural elements. Pressure and external loads can also be include in the WECAN structural analysis. Such coupled I thermal-stress analyses are a standard application used extensively on an industry wide basis. O-I 18 nu.,esme in B-1 [ _ _ _ _ _ _ _ _
B.1.3 Program Verification Both the WECAN program and input for the WECAN . verification problems, ? currently numbering over four hundred, are maintained under configuration control. Verification problems include coupled thermal-stress analyses for f the quadrilateral, triangular, brick, and wedge isoparametric elements. These problems are an integral part of the WECAN quality assurance procedures. When a change is made to WECAN, as part of the reverification process, the configured inputs for the coupled thermal-stress verification problems are used to reverify WECAN for coupled thermal-stress analyses. B.2 ANSYS B.2.1 Description ANSYS is a public domain, general purpose finite element code, s B.2.2 Feature Used The ANSYS elements used for the analysis of stratification effects in the RHR lines are STIF 20 (straight pipe), STIF 60 (elbow and bends) and STIF14 (spring-damper for supports). B.2.3 Program Verification As described in section 2.1, the application of ANSYS for stratification has been independently verified by comparison to WESTDYN (Westinghouse piping analysis code) and WECAN (finite element code, section 8.1). The results from ANSYS are also verified against closed form solutions for simple beat. configurations. 4 i}}