ML18100A842
| ML18100A842 | |
| Person / Time | |
|---|---|
| Site: | Salem  |
| Issue date: | 12/31/1993 |
| From: | Bricenash R, Strauch P WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML18100A840 | List: |
| References | |
| IEB-88-008, IEB-88-8, WCAP-13899, NUDOCS 9402020252 | |
| Download: ML18100A842 (59) | |
Text
EVALUATION OF SALEM UNITS 1 AND 2 CHARGING, ALTERNATE CHARGING AND AUXILIARY SPRAY PIPING PER NRC BULLETIN 88-08 DECEMBER 1993 9402020252 PDR ADDCK Q
WESTINGHOUSE CI.ASS 3 (Non-Proprietary)
WCAP-13899 EVALUATION OF SALEM UNITS 1 AND 2 CHARGING, ALTERNATE CHARGING AND AUXILIARY SPRAY PIPING PER NRC BULLETIN 88-08 DECEMBER 1993 P. L. Strauch R. L. Brice-Nash b
- Reviewed by:
Reviewed by:~
7'1J" M.A. Gray Approved by: //'.1-,d.w,,,;r--
S. A. Swamy,~anager Structural Mechanics Technology
.JI-, ~ Approved by: ~-~ r/!: * / f.,bf T. H. Liu R. B. Patel, M~
System Structural Analysis and Development Work performed for Public Service Electric and Gas Company under Shop Orders PZGP-145 and 964.
WESTINGHOUSE ELECTRIC CORPORATION Nuclear and Advanced Technology Division P. 0. Box 2728 Pittsburgh, Pennsylvania 15230-2728
© 1993 Westinghouse Electric Corporation
TABLE OF CONTENTS SECTION TITLE PAGE
1.0 INTRODUCTION AND BACKGROUND
1-1 2.0 ISOLATION VALVE LEAKAGE TRANSIENT DEVELOPMENT 2-1 3.0 STRESS AND FATIGUE EVALUATION 3-1
4.0 CONCLUSION
S AND RECOMMENDATIONS 4-1
5.0 REFERENCES
5-1
1.0 INTRODUCTION AND BACKGROUND
NRC Bulletin 88-08 (Reference 1) was issued on June 22, 1988 as a result of a pipe cracking incident at Farley Unit 2, and subsequent evaluations which confirmed that valve leakage caused the failure. The purpose of the bulletin was to request that licensees review their reactor coolant systems (RCS's) to identify any connected, unisolable piping that could be subjected to adverse thermal stresses, and take action to ensure that such piping will not be subjected to unacceptable thermal stresses.
Three specific actions were requested by Bulletin 88-08:
- 1) Review systems connected to the RCS to determine whether unisolable sections of piping connected to the RCS can be subjected to stresses from temperature stratification or temperature oscillations that could be induced by leaking valves, and that were not evaluated in the design analysis of the piping.
- 2) For any unisolable sections of piping connected to the RCS that may have been subjected to excessive thermal stresses, examine nondestructively the welds, heat-affected zones and high stress locations, including geometric discontinuities, in that piping to provide assurance that there are no existing flaws.
- 3) Plan and implement a program to provide continuing assurance that unfsolable sections of all piping connected to the RCS will not be subjected to combined cyclic and static thermal and other stresses that could cause failure during the remaining life of the unit. This assurance may be provided by redesigning and modifying these sections of piping to withstand combined stresses caused by various loads including temporal and spatial distributions of temperature resulting from leakage across valve seats; instrumenting this piping to detect adverse temperature distributions and establishing appropriate limits on temperature distributions; or providing a means for ensuring that pressure upstream from block valves which might leak is monitored and does not exceed RCS pressure.
The NRC was prompted to issue Supplement 1 to Bulletin 88-08 on June 24, 1988, following a pipe cracking incident at Tihange Unit 1 in Belgium. This crack was in the base metal of an elbow, and not in the weld or heat-affected zone, as at Farley. The purpose of this supplement was to emphasize the need for sufficient examinations of unisolable piping connected to the RCS to ensure that there are no rejectable crack or flaw indications, and that examinations of high stress locations should include the base metal, as appropriate.
Supplement 2 of Bulletin 88-08 was issued on August 4, 1988. The experience at Farley Unit 2 and Tihange Unit 1 indicated that the ultrasonic testing procedures used were unable to reliably detect thermal fatigue cracks in stainless steel piping. Therefore, the purpose of this supplement was to emphasize the need for enhanced ultrasonic techniques and experienced examination personnel to detect such cracks.
1-1
- Supplement 3 of Bulletin 88-08 was issued as a result of a cracking incident in the residual heat removal suction piping at a foreign reactor on June 6, 1988. This incident was different than the Farley incident, in that this event involved "hot" leakage exiting the RCS, whereas the previous incidents involved "cold" leakage entering the RCS. Also, this event involved periodic leakage through the valve packing gland, whereas the Farley incident involved steady leakage through the isolation valve (upstream to downstream). Therefore, the purpose of this supplement was to alert utilities that periodic valve seat leakage through packing glands could result in unacceptable thermal stresses.
The incidents resulting in the issuance of NRC Bulletin 88-08 and its supplements have several common factors:
- 1) the thermal loading was stratified
- 2) the root cause was related to leakage through an isolation valve
- 3) the failures occurred in unisolable piping
- 4) the damage resulted from thermal fatigue.
To comply with the requested actions of NRC Bulletin 88-08, United States utilities and many utilities in other countries have initiated programs to identify and inspect susceptible piping, and also provide for "continuing assurance" of piping integrity. These methods have included redesign/modification, temperature or pressure monitoring, inspection methods (periodic valve inspection, leak testing or nondestructive testing), operational modifications, and structural integrity analysis. In addition to these programs, the Electric Power Research Institute's TASCS (ThermAI Stratification, Cycling and Striping) Program was initiated to study the various phenomena associated with valve leakage, and develop methods to predict and evaluate the impact on structural integrity. Westinghouse is the prime contractor for the EPRI TASCS Program. Results of this study are included in Reference 2.
In response to NRC Bulletin 88-08, Public Service Electric and Gas Company (PSE&G) performed a systems review of the lines at Salem Units 1 and 2 to identify possible areas where valve leakage may jeopardize piping integrity. The following unisolable sections of auxiliary piping connected to the RCS were identified in the review (Reference 5):
- 1) Charging/Safety Injection (SI) subsystem piping downstream of the Boron Injection Tank (BIT) check valve (SJl 7's)
- 2) Chemical and Volume Control System (CVCs) alternate charging piping downstream of check valve CV80
- 3) CVCS auxiliary spray piping downstream of check valve CV76.
To address action 2 of NRC Bulletin 88-08, nondestructive examinations of areas where thermal cycling could potentially induce fatigue cracking were performed at Salem Units 1 and 2. No recordable indications were found (Reference 5).
To provide for continuing assurance of boron injection piping integrity over the life of the Salem Units, PSE&G has chosen to implement a design change. Details on this change are not included as part of this report .
- 1-2
- To provide for continuing assurance of piping integrity for the alternate charging and auxiliary spray piping, PSE&G has chosen to perform fatigue analysis based on the requirements of the ASME B&PV Code, 1986 Edition,Section III, Subsection NB-3653, for piping components. (The code of record for the design analysis is B31.1, 1967). This analysis includes both design transients and valve leakage transients. The purpose of this report is to describe the methodology, inputs and conclusions of the isolation valve leakage analysis. The normal charging flowpath, which has flow during normal operation and is therefore not susceptible to the adverse stresses discussed, has been included in the evaluation of the alternate charging line. This has been done to allow PSE&G personnel the option of using the alternate charging flowpath, while isolating the normal charging flowpath .
- 1-3
2.0 ISOLATION VALVE LEAKAGE TRANSIENT DEVELOPMENT Schematic diagrams for the Salem Units 1 and 2 normal charging, alternate charging, and auxiliary spray lines are provided in Figures 2-1 and 2-2. As shown, the common source for these lines is the regenerative heat exchanger, which has a discharge temperature of approximately 490°F under normal operating conditions. Isolation valves 1CV79 and 2CV79 are closed during normal operation, thus isolating the Units 1 and 2 alternate charging lines from charging pressure. Likewise, isolation valves 1CV75 and 2CV75 isolate the Units 1 and 2 auxiliary spray lines during normal operation. The Units 1 and 2 normal charging isolation valves 1CV77 and 2CV77 are open during normal operation.
Should any of the closed isolation valves leak during normal operation, the leakage could significantly cool from the regenerative heat exchanger discharge temperature of 490°F to as low as ambient temperature (assumed to be 100°F), before entering the unisolable piping, whicl} is approximately 540°F. (The unisolable piping is defined as the section of pipe between the RCS cold leg, or main spray line, and the adjacent check valve). This cooling is primarily. dependent on the leak rate, the length of piping, and the pipe insulation. For example, smaller leak rates flowing through longer lengths of piping with inadequate insulation will result in significantly cooler leakage. This, in tum, will yield higher stratification loadings in the unisolable piping sections.
Piping lengths from the regenerative heat exchanger to the unisolable piping are provided in Figures 2-1 and 2-2 for the charging and auxiliary spray lines. As shown, the Unit 2 normal charging line has a longer length (96 feet) than the other three charging lines. Therefore, a rounded-up length of 100 feet was_selected to envelop the Units 1 and 2 charging and alternate charging lines. Similarly, the Unit 2 auxiliary spray line length of 170 feet was selected to envelop the Unit 1 auxiliary spray line.
The Units 1 and 2 charging and auxiliary spray lines are covered with varyillg lengths, types and thicknesses of insulation. To generate conservative transient loadings, the thinnest insulation with the highest thermal conductivity was used. For the charging lines, [
].,'"" was used. In the auxiliary spray line transient development, [:
]a.c.e was used.
Figure 2-3 illustrates the terminology used in the heat transfer discussions which follow.
2.1 Heat Transfer Coefficients To evaluate the structural integrity of a pipe with a TASCS loading, the pipe wall temperature is required. To obtain the pipe wall temperature, heat transfer coefficients must be determined. These heat transfer coefficients for the water-to-pipe (h 1:J, and insulation-to-air (h45) were calculated for the Salem Units 1 and 2 charging and auxiliary spray piping, using the methods described below.
2-1
2.1.1 Pipe Inner Surface Heat Transfer Coefficient For the heat transfer analyses, the following heat transfer coefficient was applied to the entire pipe inner surface for the steady state solution (Reference 2).
a,c,e This coefficient (h1:z) is for free convection between the water and the pipe metal. Due to the low velocities involved in leakage flows, free convection has been calculated to be significantly higher than forced convection, and therefore controls the heat transfer.
2.1.2 Insulation Outer Surface Heat Transfer Coefficient The heat transfer from the surface of the insulation follows a parallel path to the air (through convection and radiation), whereas the heat transfer from the water and through the pipe metal and insulation follows a series path. The heat transfer coefficient for the insulation outer surface (h45) is comprised of free convection and radiation to the ambient air:
The free convection heat transfer coefficient is defined as follows:
[ ]~~
- 2-2
where:
a,c,e The heat transfer is assumed to be free convection from the insulation surface.
The radiation heat transfer coefficient is defined as follows:
where:
h4Sr = radiation heat transfer coefficient CJ = Stefan-Boltzmann constant (1.714E-9 Btu/hr-ft2-R4)
E = surface emissivity Temperatures for the radiation heat transfer calculation must be in degrees Rankine.
Values for the convection and radiation heat transfer film coefficients of [i
]'c;e were calculated for the charging and auxiliary spray piping.
2.2 Heat Transfer from Conduction To determine the temperature distribution for TASCS evaluations, it is often required to perform heat transfer calculations for flows which are not stratified or for stagnant pipe sections. Generally, this is required to detennine the temperature distribution in a pipe line as a boundary condition to the section of pipe with a TASCS loading.
One example of this is a case in which a leak is initially at a high temperature but possibly several hundred feet away from the unisolable pipe where a TASCS evaluation is required. It must be determined if the leak can cool before it reaches the unisolable pipe. High leak rates will remain hot, and small leaks will cool down. Another example is the determination of the piping temperature distribution in the vicinity of the unisolable piping. Since the leakage flow rates in TASCS evaluations are generally low, leakage may significantly heat up before entering the unisolable piping. Therefore, the boundary temperature distribution must be quantified prior to performing leakage heatup calculations. Both of these examples were used in the evaluation of the Salem Units 1 and 2 charging and auxiliary spray piping.
2-3
- 2.2.1 Flow without Stratification This section provides closed form solutions to determine the axial temperature distribution in the Salem Units 1 and 2 charging and auxiliary spray piping, assuming that leakage flows from the regenerative heat exchanger, through the closed isolation valves and toward the unisolable piping. From this, the temperature of the leakage near the entrance to the unisolable piping is determined. (This temperature will be used in the calculation of leakage flow heat up in Section 2.4). Assuming that the ambient temperature is 100°F, the cooling of leakage flow which is initially hot (490°F) is determined along the pipe length. It is assumed that stratification is not significant for this case since the pipe, at a given cross-section, will tend to reach a near-uniform temperature distribution in the steady state. The axial a,c,e temperature distribution is determined by (Reference 2):
- 2-4
- Methods for calculating the heat transfer coefficients, h12, and h45 were described in Section 2.1. Values of [
auxiliary spray lines, respectively.
]8'"'e were calculated for h 12, for the charging and Figures 2-4 through 2-9 provide the axial temperature distributions from the regenerative heat exchanger to the unisolable piping for the Salem Units 1 and 2 charging and auxiliary spray piping, at arbitrary leakage flowrates of 0.1, 0.25 and 0.5 gpm. These temperature distributions will be used in Section 3.0 to determine the global structural loadings and stresses. Figures 2-10 and 2-11 provide leakage temperatures as a function of flowrate in the vicinity of the unisolable piping for the charging and auxiliary spray piping. These temperatures will be used as "initial" leakage temperatures in the leakage heat up calculation discussed in Section 2.4.
2.2.2 Conduction This section provides the methodology used in the calculation of the axial temperature distribution of the piping in the vicinity of the unisolable piping. The result is a one-dimensional closed form solution which ignores the effects of conduction through the fluid (axially), which is negligible when compared to conduction through the pipe metal.
The axial temperature distribution is calculated using the following (Reference 2):
a,c,e 2-S
For the charging and auxiliary spray piping, it is assumed that the unisolable piping, including the adjacent check valve, is maintained at 540°F due to the short lengths of unisolable piping, convective currents and, for the charging lines, loop turbulence. The temperature beyond the check valve is assumed to be governed by the conduction equation given above. Figures 2-12 and 2-13 provide axial temperature distributions for the piping in the vicinity of the unisolable piping. These distributions are used to calculate the heatup of the leakage flow in Section 2.4.
2.3 Height or a Stratified Flow When thermal stratification occurs, the pipe is partially filled with hot water and partially filled with cold water. In some cases, the interface is very small and the gradient is very large. In the other extreme, the transition between the hot and cold fluid can occur over the entire pipe cross-section (see Figure 2-14). The causes of each of these two cases are related in a complex fashion to the flow rate, temperature difference, length of flow, pipe slope, pipe material and temperature (insulation characteristics), entrance conditions and exit conditions.
Once the interface height (H) has been identified, the velocity and other fluid parameters can be calculated for given volumetric flow rates. This is important because calculation of the heat transfer and stability of a stratified flow are dependent on the flow velocity.
The calculation of the stratification interface height (H) is based on the variable, yc* the critical depth, where [ ]a.c.e (Reference 2). The method to calculate yc follows.
The critical depth (yJ is calculated using an iterative solution of the following equation (Reference 2):
a,c,e Note: the same geometric parameters, A, W, and a are used to characterize H, with the subscript "y" eliminated.
2-6
This methodology was used to estimate the height of the interface of the stratified flow in the unisolable sections of the charging and auxiliary spray piping. This approach inherently assumes that the interface is well defined, which is generally a required assumption in evaluating a stratified flow. This approach also assumes that there is flow involved in the stratified loading. In the case of the Salem evaluation, this flow is the leakage from the regenerative heat exchanger. (There are cases where a stratified condition can be established without a sustained flow, such as the stratification resulting from free convection currents).
Stratification heights for the Salem Units 1 and 2 charging and auxiliary spray piping are provided in Figures 2-16 and 2-17.
2.4 Heat Transfer of a Leakage Flow This section provides the methodology used to determine the heatup of the cold stratified leak flow from the regenerative heat exchanger in the hot ambient piping immediately upstream of the unisolable piping.
A stratified flow will eventually reach equilibrium (in terms of temperature) with the ambient fluid because of heat transfer between the hot and cold layers. For high flow rates, the pipe will reach equilibrium at the temperature of the flowing fluid. For low flow rates, the flowing fluid will equilibrate with the ambient fluid in the pipe. Therefore, it is useful to determine the rate at which a small leak flow transfers heat from its environment (pipe and fluid), which would reduce the stratification temperature difference and, hence, the stress in a pipe with a stratified flow.
Using the depth of the leak flow as determined in the previous section, the calculation of the axial temperature distribution, T(x) of the leak flow is determined by the following equation (Reference 2):
a,c,e
- 2-7
- where: a,c,e *
[
The heat transfer coefficient (h 1:z) is defined by the following equation (Reference 2):
r~
where:
a,c,e In the calculations for the Salem Units 1 and 2 charging and auxiliary spray piping, since the ambient fluid cools down from 540°F by conduction (see Figures 2-12 and 2-13), a variable ambient fluid temperature was used, applying a finite difference approach.
This method was bench-marked in the TASCS program (Reference 2), using horizontal stainless steel pipes, with saturated water. All tests were conducted for a cold leak into hot fluid. This method is therefore applicable to the Salem charging and auxiliary spray piping.
Figures 2-18 and 2-19 provide the charging and auxiliary spray leakage temperatures at the entrance to the unisolable piping (i.e., at the check valve outlet weld), following the heatup near the unisolable piping. (The eooldown curves of Figures 2-10 and 2-11 are also included to indicate the amount of this heatup, which is the difference between the two curves). As shown, the leakage temperature reaches a minimum value at [ ]8'c.e for both the charging and auxiliary spray piping. The stratification differential temperature in the unisolable piping is then calculated as 540°F minus the leakage temperature, as shown in Figures 2-20 and 2-21.
2-8
Figures 2-22 and 2-23 define the leakage stratification loadings used in the heat transfer and stress evaluations of the charging and auxiliary spray piping in Section 3.0. [
.]a.c:,e Stress cycling resulting from isolation valve leakage entering the unisolable piping is assumed to result from turbulence from the RCS cold leg loop piping for the normal and alternate charging lines, and from the main spray piping, during spray operation, for the auxiliary spray line. This turbulence can mix the stratified leakage, which may then be reestablished, completing the cycle. Stress cycling may also result from the combination of leakage transients with plant design transients. This is discussed in detail in Section 3.0.
2-9
- SALEM UNIT 1 CHARGING SYSTEM FROM REGENERATIVE HEAT EXCHANGER TO COLD LEGS AND SPRAY PIPING (ALTERNATE CHARGING) 3*.1RC1091 27.5.* ID - 1086 LOOP4 1CV79 1 CV275 1CV80 COLD LEG XA-65 FA-33 FA-33 3*.1025
!! (NORMAL CHARGING)
'l REGEN.
3* -1024 3*.1RC1077 27.5* ID -1071 c: HEAT
- 0 LOOP3 w m EXCHANGER I
1CV77 1 CV274 1CV78 COLD LEG
=
I\)
I (RHX)
XA-65 FA-33 FA-33 2*-1020 (AUXILIARY SPRAY) 2*-1RC1281 4*.1201 (PZR SPRAY LINE)
DISTANCES FROM RHX TO UNISOLABLE PIPING:
1CV75 1CV76 ALTERNATE CHARGING - 43 FEET
.NORMAL CHARGING - 88 FEET 3/4-1021 XA-26 FA-45 AUXILIARY SPRAY -161 FEET 1CV272 1CV273 FA-14 FA-41
SALEM UNIT 2 CHARGING SYSTEM FROM REGENERATIVE HEAT EXCHANGER TO COLD LEGS AND SPRAY PIPING (ALTERNATE CHARGING) 3*-115e 27.5" ID -1011 LOOP4 COLD LEG 2CV79 2CV275 2CV80 XA-65 FA-33 FA-33 3*-11s7 REGEN. (NORMAL CHARGING) 27.s* 1D -1oee 3*-1os5 3*-1oss HEAT LOOP3 EXCHANGER COLD LEG (RHX) . 2CV77 2CV274 2CV78 XA-65 FA-33 FA-33 CD N
I DISTANCES FROM RHX TO UNISOLABLE PIPING:
(AUXILIARY SPRAY) 2*-1155
~
4*-2PS1000 ALTERNATE CHARGING- 68 FEET (PZR SPRAY LINE)
NORMAL CHARGING - 96 FEET 2CV75 2CV76 AUXILIARY SPRAY -170 FEET XA-26 FA-45 3/4-1219 2CV272 2CV273 FA-14 FA-41
- HEAT TRANSFER TERMINOLOGY
\
T 1= WATER TEMPERATURE T 2= PIPE INNER WALL TEMPERATURE T 3= PIPE OUTER WALL TEMPERATURE T 4= INSULATION OUTER WALL TEMPERATURE Ts= AMBIENT AIR TEMPERATURE FIGURE 2*3 2-12
SALEM CHARGING LINE TEMPERATURE DIST.
a,c,e FIGURE 2*4 2-13
SALEM CHARGIN*G LINE TEMPERATURE DIST.
a,c,e FIGURE 2-5 2-14
SALEM CHARGING LINE TEMPERATURE DIST.
a,c,e FIGURE 2*6 2-15
SALEM AUX SPRAY LINE TEMP. DISTRIBUTION a,c,e FIGURE 2*7 2-16
SALEM AUX SPRAY LINE TEMP. DISTRIBUTION a,c,e FIGURE 2*8 2-17
SALEM AUX SPRAY LINE TEMP. DISTRIBUTION a,c,e FIGURE 2*9
- 2-18
CHARGING LINE LEAKAGE TEMPERATURE a,c,e
. FIGURE 2*10
- 2-19
AUX. SPRAY LINE LEAKAGE TEMPERATURE a,c,e FIGURE 2*11 2-20
CHARGING LINE TEMPERATURE DISTRIBUTION a,c,e DIRECTION OF LEAKAGE FLOW FIGURE 2*12 2-21
AUX. SPRAY LINE TEMP. DISTRl:BUTION a,c,e
)(-~~~~~-.~~~--~~--1l DIRECTION OF LEAKAGE FLOW FIGURE 2*13 2-22
Interface Thickness Interface Thickness Cooler FIGURE 2*14 : VARIATION OF STRATIFICATION INTERFACE 2-23
a,c,e FIGURE 2-15: THERMAL STRATIFICATION INTERFACE°TERMINOLOGY
- 2-24
SALEM CHARGING LINE LEAKAGE HEIGHT a,c,e FIGURE 2*16 2-25
SALEM AUX. SPRAY LINE LEAKAGE HEIGHT a,c,e FIGURE2-17 2-26
~------
CHARGING LINE LEAKAGE TEMPERATURE a,c,e FIGURE 2*18 2-27
AUX. SPRAY LINE LEAKAGE TEMPERATURE a,c,e FIGURE 2*19 2-28
CHARGING LINE STRATIFICATION DELTA T a,c,e '
FIGURE 2*20
- 2-29
AUX. SPRAY LINE STRATIFICATION DELTA T a,c,e FIGURE 2-21 2-30
CHARGING LINE STRATIFICATION LOADING a,c,e FIGURE 2*22 2-31
- AUXILIARY SPRAY LINE STRATIFICATION LOADING a,c,e FIGURE 2*23 2-32
- 3.0 STRESS AND FATIGUE EVALUATION To evaluate the effects of valve leakage into the unisolable piping, it was necessary to consider two stress effects on the piping: a "local" effect and a "global" effect. Local stresses are obtained by modeling a section of the pipe and imposing the leakage stratification transients as defined in Section 2 and shown in Figures 2-22 and 2-23. Two cases were analyzed, one with no end restraints and the other with full rotational restraints. The resulting radial, circumferential and axial stresses are defined as local stresses. For the case without end restraints, the piping will "bow" and expand along the pipe axially. For the second case, the piping will expand axially but not "bow". For both the charging and auxiliary spray lines, it was determined that the local stresses would be bounded by the two cases.
Global stresses result from the effects of supports and pipe geometry not permitting the piping to expand and deflect as it would in the unrestrained condition. Leakage depth, stratification temperature difference and the extent of stratified pipe are important parameters in the determination of global piping stress. A leakage depth at the pipe centerline will maximize the effect of global bending. Global stresses will typically increase as the length of stratified piping increases.
3.1 Local Stress Evaluation To determine the magnitude of local stresses resulting from the postulated stratified leakage within the pipe, a thermal solution is determined and is used as input to a stress solution.
This was performed using a two-dimensional (20) model of one-half of the pipe cross section, for each pipe size, with symmetric boundary conditions for nodes on the plane of symmetry. The Westinghouse general purpose finite element program WECAN (Ref. 3) was used to obtain both the thermal and the stress solutions. The local stresses for the postulated leakage were obtained by replacing the heat transfer elements used in the thermal analysis model with 20 isoparametric generalized plane strain elements. [
For the charging line, the leakage transient [
- 3*1
[
)'"'e. The leakage transient for the auxiliary spray line is the same as the charging line transient, except [
To estimate transient cycles, a time history analysis was performed [
Time history plots of stress intensity with respect to time, at the highest stressed locations in the charging and auxiliary spray models, are shown in Figures 3-1 through 3-4. [
The resulting maximum temperature and stress intensity contours for both the charging line pipe and the auxiliary spray pipe, at the highest stressed locations in the finite element models, are shown in Figures 3-5 through 3-8 for the free moment cases, and Figures 3-9 through 3-12 for the fixed moment cases. These stresses are evaluated as peak stresses in the fatigue analysis.
3.2 Global Stress Evaluation Global stresses are obtained by performing thermal stress analysis of the piping system with stratification postulated in the section of pipe between the reactor coolant loop a~d the adjacent check valve for the "out of service" charging lines, and between the main spray pipe and the adjacent check valve for the auxiliary spray line. Turbulence was assumed to extend into the charging lines from flow in the reactor coolant loop. Turbulence limits the length of stratified piping by causing the hot and cold fluid to mix. This turbulence may be quantified as shown in the following equation (Reference 2):
~ ~ a,c,e 3-2
- where, for the charging lines:
a,c,e Based on this equation, the maximum distance at which cycling from reactor coolant loop turbulence may occur is [
For the auxiliary spray line, there are periods of time when the spray valves are open, and other times when the spray valves are closed, and only trickle flow is present in the main spray line. The distance from the spray line to the check valve in the auxiliary spray line is very short (less than six inches). Global stratification effects will therefore be negligible whether or not the spray valves are open.
For both the charging lines and the auxiliary spray line, [
r-e The global moment stresses were combined with the local stresses in the fatigue evaluation, to determine fatigue usage for the postulated leakage loadings.
3.3 Fatigue Evaluation The fatigue evaluation of the charging, alternate charging and auxiliary spray lines for postulated leakage was based on the requirements of the ASME B&PV Code, 1986 Edition, Section Ill, Subsection NB-3653, for piping components. Fatigue usage was calculated [.
3-3
L
- 3.3.1 Charging Lines The normal and alternate charging line [ ]a.c.e were evaluated by considering the following transients:
a,c,e The postulated isolation valve leakage stress cycles can only occur in the charging or alternate charging line when each is out of service. Both lines were conservatively assumed to be out of service 60% of the time. The plant availability factor was assumed to be 80%. Since the plant was originally qualified to B31.1 requirements, it was assumed that seismic loadings were the maximum at which primary stresses met B31.1 allowables.
Each of the loadsets was allowed to combine according to the requirements of the ASME Code until all cycles were exhausted. Altern!lting stress, allowable cycles, actual cycles, and usage factor were calculated for each transient loadset combination. The design fatigue curves used in the calculation of usage factor were from the 1986 ASME Code Figures 3-4
- 1-9.2.1 and 1-9.2.2, for austenitic stainless steels. These curves were used because they include high cycle fatigue considerations. The total usage factor is the sum of usage factors for each loadset combination, which was shown to be less than 1.0 for the life of the plant.
3.3.2 Pressurizer Auxiliary Spray Line The Units 1 and 2 pressurizer auxiliary spray line [ ]a.c.e were evaluated using the same method as for the charging lines in Section 3.3.1, including the following items: a,c,e The postulated leakage cycles can occur during normal plant operation, and during heatup and cooldown when the auxiliary spray line is not in service. The plant availability factor was assumed to be 80%. Seismic loadings were assumed to be 10,000 in-lbs for each moment component.
Each of the loadsets was allowed to. combine according to the requirements of the ASME Code until all cycles were exhausted. Alternating stress, allowable cycles, actual cycles, and usage factor were calculated for each transient loadset combination. The design fatigue curves used in the calculation of usage factor were from the 1986 ASME Code Figures
- 3-5
1-9.2.1 and 1-9.2.2, for austenitic stainless steels. These curves were used because they include high cycle fatigue considerations. The total usage factor is the sum of usage factors for each loadset, which was shown to be less than 1.0 for twenty-four (24) years of operation.
This includes six occurrences of the inadvertent auxiliary spray transient.
3-6
a,c,e FIGURE 3-1: CHARGING PIPE STRESS INTENSITY, FREE MOMENT CASE 3-7
a,c,~
FIGURE 3-2: AUXILIARY SPRAY PIPE STRESS INTENSITY, FREE MOMENT CASE 3-8
a,c,~
FIGURE 3-3: CHARGING PIPE STRESS INTENSITY, FIXED MOMENT CASE 3-9
a,c,e I
FIGURE 3-4: AUXILIARY SPRAY PIPE STRESS INTENSITY, FIXED MOMENT CASE 3-10
a,c,e FIGURE 3-5: CHARGING TEMPERATURE, FREE MOMENT CASE 3-11
a,c,e FIGURE 3-6: CHARGING STRESS INTENSITY, FREE MOMENT CASE 3-12
a,c,e i
- 1 I
FIGURE 3-7: AUXILIARY SPRAY TEMPERATURE, FREE MOMENT CASE 3-13
a,c,e FIGURE 3-8: AUXILIARY SPRAY STRESS INTENSITY, FREE MOMENT CASE 3-14
a,c,_e FIGURE 3-9: CHARGING TEMPERATURE, FIXED MOMENT CASE 3-15
FIGURE 3-10: CHARGING STRESS INTENSITY, FIXED MOMENT CASE
- 3-16
a,c,e FIGURE 3-11: AUXILIARY SPRAY TEMPERATURE, FIXED MOMENT CASE
- 3-17
a,c,e FIGURE 3-12: AUXILIARY SPRAY STRESS INTENSITY, FIXED MOMENT CASE 3-18
4.0 CONCLUSION
S AND RECOMMENDATIONS An evaluation of the Salem Units 1 and 2 charging, alternate charging and auxiliary spray piping considering the effect of postulated isolation valve leakage transients and design transients on fatigue usage has been performed. The conclusion of this evaluation for the normal and alternate charging lines is that the cumulative fatigue usage**for design transients and postulated isolation valve leakage transients is less than 1.0 for the life of the plant. The conclusion for the Units 1 and 2 auxiliary spray lines is that the cumulative fatigue usage for design transients and postulated isolation valve leakage transients is less than 1.0 for 24 calendar years. This assumes a worst case scenario of continuous isolation valve leakage at the critical leakage flowrate [* r*e. Also, six occurrences of the inadvertent auxiliary spray transient are included in the evaluation. Determination of past inadvertent auxiliary spray events and tracking of future events could extend this period beyond 24 years.
Since the Salem Units 1 and 2 normal and alternate charging piping cumulative fatigue usage is less than the allowable of 1.0, no additional actions are required to satisfy the requirements of NRC Bulletin 88-08. However, continuing assurance of piping integrity for the Units 1 and 2 auxiliary spray lines will be required after 24 calendar years for each unit. This assurance may be provided by installing temperature monitoring instrumentation for detection of piping thermal stratification and/or cycling due to valve leakage. Temperature sensors should preferably be resistance temperature detectors (RTD's), located between the connection to the main spray piping (2"-6000# sockolet) and the check valve outlet (CV76), on the top and the bottom of pipe. (If there is insufficient space to locate the RTD's between the sockolet and valve, then the RTD's may be installed immediately upstream of the check valves). After RTD installation, temperatu.res should be recorded during normal plant operation over a period of 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> to determine the baseline temperature history. Baseline temperatures should meet the following criteria:
- The maximum top-to-bottom temperature difference should not exceed 50°F.
- Top and bottom temperature time histories should be in-phase.
- Peak-to-peak temperature fluctuations should not exceed 60°F.
Monitoring should be performed at the beginning of power operation, after startup from a refueling outage, and at least at six-month intervals thereafter, between refueling outages.
During each monitoring period, temperature readings should be recorded continuously for a 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> period.
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Actions should be taken to correct valve leakage if the following conditions occur:
- The maximum temperature difference between the top and the bottom of the pipe exceeds 50°F.
- Top and bottom temperature histories are in-phase but the peak-to-peak fluctuations of the top or bottom temperatures exceed 60°F.
- Top and bottom temperature histories are out-of-phase and the bottom peak-to-peak temperature fluctuations exceed 50°F.
- Temperature histories do not correspond to the initially recorded baseline histories.
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- t S.O REFERENCES
- 1. United States Nuclear Regulatory Commission Bulletin 88-08, "Thermal Stresses in Piping Connected to Reactor Coolant Systems", 6/22/88; Supplement 1, 6/24/88; Supplement 2, 8/4/88; and Supplement 3, 4/11/89.
- 2. "Thermal Stratification, Cycling and Striping (TASCS) Preliminary Final Report",
Prepared by Westinghouse Electric Corporation for Electric Power Research Institute, Research Project 3153-02, Dated October 1993.
- 3. Westinghouse general purpose finite element program WECAN/PLUS, Version Release 90-2 (20402403020), Westinghouse Proprietary.
- 4. Westinghouse Systems Standard 1.3.X, Revision 0, September 1978, "Nuclear Steam Supply System Auxiliary Equipment Design Transierits for All Standard Plants", and Westinghouse Systems Standard 1.3.F, Revision 0, March 1978, "Nuclear Steam Supply System Reactor Coolant System Design Transients", Westinghouse Proprietary.
- 5. Westinghouse Report "Salem Units 1 & 2 NRC Bulletin 88-08 Evaluation Summary Report", December 1992.
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