ML20076J954

Pressurizer Safety & Relief Line Evaluation Summary Rept

ML20076J954
Person / Time
Site:	Beaver Valley
Issue date:	06/30/1983
From:	Chang K, Ching Ng, Laura Smith DUQUESNE LIGHT CO., WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML20076J951	List: ML20076J948 ML20076J954
References
RTR-NUREG-0737, RTR-NUREG-737, TASK-2.D.1, TASK-TM NUDOCS 8307070119
Download: ML20076J954 (71)

Text

_ - - - _ _ _ _ . . . - - .

WESTINGHOUSE PROPRIETARY CLASS 3 PRESSURIZER SAFETY AND RELIEF LINE EVALUATION

SUMMARY

REPORT j DUQUESNE LIGHT COMPANY BEAVER VALLEY - UNIT 1 4

C. K. Ng j L. C. Smith June, 1983 i

Approved: / // _

K.C.Ckk,3 nager Systems Structural An ysis i

k[hDCkO P

0766s:10-1 m

TABLE OF CONTENTS Section Title 1 INTRODUCTION 2 PIPE STRESS CRITERIA 2.1 Pipe Stress Calculation 2.2 Load Combinations 3 LOADING CONDITIONS ANALYZED 3.1 Loading 3.1.1 Therral Expansion 3.1.2 Pressure 3.1. 3 Weight 3.1.4 Seismic 3.1.5 Safety and Relief Valve Thrust 3.2 Design Conditions 3.2.1 Design Pressure 3.2.2 Design Temperature 3.3 Plant Operating Conditions 3.3.1 Normal Conditions 3.3.2 Upset Conditions 3.3.3 Emergency Conditions 3.3.4 Faulted Conditions 0766s:10-2

TABLE OF CONTENTS (Cont)

Sec tion Title 4 ANALYTICAL METHODS AND MODELS 4.1 Introduction 4.2 Static Analysis 4.3 Dyr.amic Analysis 4.4 Seismic Analysis 4.5 Pressurizer Safety and Relief Line Analysis 4.5.1 Plant Hydraulic Model 4.5.2 Comparison to EPRI Test Results 4.5.3 Valve Thrust Analysis 5 METHOD OF STRESS EVALUATION 5.1 Introduction 5.2 Primary Stress Evaluation 5.2.1 Design Conditions 5.2.2 Upset Conditions 5.2.3 Emergency Conditions 5.2.4 Faulted Conditions 5.3 Secondary Stress Evaluation ,.

6 RESULTS 6.1 Evaluation Prior to EPRI Test Program 6.2 Evaluation Subsequent to EPRI Test Program 6.2.1 Thermal Hydraulic Results 6.2.2 Structural Results l

6.3 Summary of Results and Conclusions i

l 7 N0Dl'FICATIONS J,1-Elevation of Loop Seal Temperature 7,2 Pipe Support Modification -

7.3 Conclusion -

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. SECTION 1 INTRODUCTION The pressurizer safety and relief valve (PSARV) discharge piping system for pressurized water reactors, located on top of the pressurizer, providos overpressure protection for the reactor coolant system. A water seal is often maintained upstream of each pressurizer safety and relief valve to prevent a steam interface at the valve seat. This water seal practically eliminates the possibility of valve leakage. Whil e this arrangement maximizes the plant availability, the water slug, driven by high system pressure upon actuation of the valves, generates severe hydraulic shock loads on the piping and supports.

Under NUREG 0737,Section II.D.1, " Performance Testing of BWR and PWR Relief and Safety Valves", all operating plant licensees and applicants are required to conduct testing to qualify the reactor coolant system relief and safety valves under expected operating conditions for design-basis transients and accidents. In addition to the qualification of valves, the functionability and structural integrity of the as-built discharge piping and supports must also be demonstrated on a plant specific basis.

In response to these requirements, a program for the performance testing of PWR safety and relief valves was formulated by EPRI. The primary objective of the Test Program was to provide full scale test data con-firming the functionability of the reactor coolant system power operated relief valves and safety valves for expected operating and accident conditions. The second objective of the program was to obtain suffi-cient piping thermal hydraulic load data to permit confirmation of models which may be utilized for plant unique analysis of safety and relief valve discharge piping systems.

This report is the response of the Duquesne Light Company to the US NRC plant-specific submittal request for piping evaluation and is applicable to the Beaver Valley - Unit 1 pressurizer safety and relief valve discharge piping system.

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- SECTION 2 PIPE STRESS CRITERIA 2.1 PIPE STRESS CALCULATION The piping between the pressurizer nozzles and the pressurizer relief tank was analyzed according to the requirements of the appropriate equations of the ANSI B31.1-1967 Code (hereafter referred to as the Code). These equations establish limits for stresses from sustained loads, sustained plus occasional loads (including earthquake), thermal expansion loads, and sustained plus thermal expansion loads. The allowable stresses for use with the equations were determined in accordance with the requirements of the Code.

2.2 LOAD COMBINATIONS In order to evaluate the pressurizer safety and relief valve piping, appropriate load combinations and acceptance criteria were developed.

The load combinations and acceptance criteria are identical to those recommended by the piping subcommittee of the PWR PSARV test program and are outlined in Tables 2-1 and 2-2. Definitions of the load abbrevia-tions are provided in Table 2-3.

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. TABLE 2-1 LOAD COMBINATIONS AND ACCEPTANCE CRITERIA FOR PRESSURIZER SAFETY AND RELIEF VALVE PIPING AND SUPPORTS - UPSTREAM OF VALVES Piping Plant / System Allowable Stress Combination Operating Condition Load Combination Intensity 1 Normal N 1.0 S h 2 Upset N + OBE + S0T U 1*2 S h 3 Emergency N + SOT 1*O E h 4 Faulted N + MS/FWPB or DBPB 2.4 S h

+ SSE + SOTp 5 Faul ted N + LOCA + SSE + S0Tp 2.4 S h NOTES: (1) See Table 2-3 for SOT definitions and other load abbreviations.

(2) The bounding number of valves (and discharge sequence if setpoints are significantly different) for the applicable system operating transient defined in Table 2-3 should be used.

(3) Verification of functional capability is not required, but allowable loads and accelerations for the safety-relief valves must be met.

(4) Use SRSS for combining dynamic load responses.

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, TABLE 2-2 LOAD COMBINATIONS AND ACCEPTANCE CRITERIA FOR PRESSURIZER SAFETY AND RELIEF VALVE PIPING AND SUPPORTS - SEISMICALLY DESIGNED DOWNSTREAM PORTION Piping Plant / System Allowable Stress Combination Operating Condition Load Combination Intensity 1 Normal N 1.0 S h 2 Upset N + SOT g 1.2 S h 3 Upset N + OBE + S0T U 1.8 S h 4 Emergency N + SOT I'0 b E h 5 Faul ted N + MS/FWPB or DBPB 2.4 S h

+ SSE + SOT p 6 Faulted N + LOCA + SSE + SOT p 2.4 S h NOTES: (1) This table is applicable to the seismically designed portion of downstream non-Category I piping (and supports) necessary to isolate the Category I portion from the non-seismically designed piping response, and to assure acceptable valve loading on the discharge nozzle.

(2) See Table 2-3 for S0T definitions and other load abbreviations.

(3) The bounding number of valves (and discharge sequence if setpoints are significantly different) for the applicable system operating transient defined in Table 2-3 should be used.

(4) Verification of functional capability is not required, but allowable loads and accelerations for the safety-relief valves must be met.

(5) Use SRSS for combining dynamic load responses.

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TABLE 2-3 DEFINIT'ONS OF LOAD ABBREVIATIONS N = Sustained loads during normal plant operation S0T = System operating transient S0TU = Relief valve discharge transientIII S0TE = Safety valve discharge transient (1), (2)

S0Ty = Maximum of SOT U and SOT E ; or transition flow OBE = Operating basis earthquake SSE = Safe shutdown earthquake MS/FWPB = Main steam or feedwater pipe break DBPB = Design basis pipe break LOCA = Loss-of-coolant accident Sh = Basic material allowable stress at maximum (hot) temperature (1) May also include transition flow, if determined that required operating procedures could lead to this condition.

(2) Although certain nuclear steam supply systems design transients (for example, loss of load) which are classified as upset condi-tions may actuate the safety valves, the extremely low number of actual safety valve actuations in operating pressurizer water reactors justifies the emergency condition from the ASME design philosophy and a stress analysis viewpoint. However, if actuation of safety valves would occur, a limitation must be placed to shut down the plant for examination of system integrity after an appro-priate number of actuations. This number can be determined on a plant specific basis.

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- SECTION 3 LOADING CONDITIONS ANALYZED 3.1 LOADING The piping stress analyses described in this section consider all pertinent loadings. These loadings result from thermal expansion, pressure, weight, earthquake, and safety valve and relief valve operation.

3.1.1 THERMAL EXPANSION The thermal growth of the reactor coolant loop equipment and all connected piping is considered in the thermal analysis of this system.

The modulus of elasticity (E), the coefficient of thermal expansion at the metal temperature (a), the external movements transmitted to the piping as described above, and the temperature rise above the anbient temperature (AT for various operating modes), define the required input data to perform the flexibility analysis for thermal expansion.

Because of the many possible operating modes, the system may experience many different thermal loadings. The temperatures used in the expansion analysis are based on all available information and include pertinent valve opening cases.

To provide the necessary high degree of integrity for the piping, the transient conditions selected for secondary stress evaluation are based on conservative estimates of the magnitude and anticipated frequency of occurrence of the temperature and pressure transients resulting from the possible operating conditions.

The transients selected are conservative representations of transients for design purposes, and are used as a basis for piping secondary stress !

evaluation to provide assurance that the piping is acceptable for its application over the design life of the plant.

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~

For purposes of piping evaluation, the number of transient occurrences is based on a plant design life of 40 years.

3.1.2 PRESSURE Pressure loading in this report is either design pressure or operating pressure. The design pressure is used in the calculation of longitu-dinal pressure stress in accordance with the Code. The range of oper-ating pressure is used in calculating various stress intensities, as applicable.

3.1. 3 WE IGHT To meet the requirements of the Code, a weight analysis is performed by applying a 1.0 g uniformly distributed load downward on the complete piping system. The distributed weight characteristics of the piping system are specified as a function of its properties. This method provides a distributed loading to the piping system as a function of the weight of the pipe, insulation, and contained fluid during normal oper-ating conditions.

3.1. 4 SEISMIC Seismic motion of the earth is treated as a random process. Certain assumptions reflecting the characteristics of typical earthquakes are made so these characteristics can be readily employed in a dynamic response spectrum analysis.

Piping rarely experiences the actual seismic motion at ground elevation, i since it is supported by components attached to the containment build-ing. Although a band of frequencies is associated with the ground earthquake motion, the building itself acts as a filter to this environ-ment and will effectively transmit those frequencies corresponding to its own natural modes of vibration.

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The forcing functions for the piping seismic analyses are derived from dynamic response analyses of the containment building when subjected to seismic ground motion. These forcing functions are in the form of floor response spectra. Response spectra are obtained by determining the maximum response of a single mass-spring-damper oscillator to a base motion time history. This single mass-spring-damper oscillator system represents a single natural vibration mode of the piping system. A plot of the maximum responses versus the natural frequencies of the oscil-lator forms the response spectrum for that particular base motion.

The intensity and character of the earthquake motion producing forced vibration of the equipment mounted within the containment building are specified in terms of the floor response spectrum curves at various elevations within the containment building.

The seismic floor response spectrum curves corresponding to the highest elevation at which the component or piping is attached to the containment building are used in the piping analysis.

Seismic loads must be known to calculate the resultant moment (Mi )

used in the design equations. The plant operating condition (full load) is the condition under which the specified earthquake is assumed to occur.

3.1.5 SAFETY AND RELIEF VALVE THRUST The pressurizer safety and relief valve discharge piping system provides overpressure protection for the RCS. The three spring-loaded safety valves and three power-operated relief valves, located on top of the pressurizer, are designed to prevent system pressure from exceeding design pressure by more than 10 percent and 100 psi, respectively. A water seal is often maintained upstream of each valve to minimize leaka ge . Condensate accumulation on the inlet side of each valve prevents any leakage of hydrogen gas or steam through the valves.

0766 s:10-11 If the pressure exceeds the set point and the valves open, the water slug from the loop seal discharges. The water slug, driven by high system pressure, generates transient thrust forces at each location where a change in flow direction occurs.

The safety and relief lines are analyzed for various cases of thrust loadings to ensure the primary and secondary stress limits are not exceeded.

3.2 DESIGN CONDITIONS The design conditions are the pressures, temperatures, and various mechanical loads applicable to the design of nuclear power plant piping.

l 3.2.1 DESIGN PRESSURE The specified internal and external design pressures are not less than the maximum difference in pressure between the inside and outside of the component, which exists under the spect fied normal operating condi-tions. The design pressures are used in the computations made to show compliance with the Code.

3.2.2 DESIGN TEMPERATURE The spect fled design temperature is not less than the actual maximum metal temperature existing under the specified normal operating condi-tions for each area of the component considered. It is used in computa-tions involving the design pressure and coincidental design mechanical loads.

3.3 PLANT OPERATING CONDITIONS 3.3.1 NORMAL CONDITIONS A normal condition is any condition in the course of system startup, design power range operation, hot standby, and system shutdown, other than upset, faulted, emergency, or testing conditions.

0766s:10-12

l 3.3.2 UPSET CONDITIONS An upset condition is any deviation from normal conditions anticipated to occur often enough that design should include a capability to with-stand the condition without operational impairment. Upset conditions include those transients resulting from any single operator error or control malfunction, transients caused by a fault in a system component requiring its isolation from the system, and transients due to loss of load or power. Upset conditions include any abnormal incidents not resulting in a forced outage and also forced outages for which the corrective action does not include any repair of mechanical damage.

3.3.3 EMERGENCY CONDITIONS Emergency conditions are defined as those deviations from normal conditions which require shutdown for correction of the conditions or repair of damage in the system. The conditions have a low probability of occurrence but are included to prov.de assurance that no gross loss of structural integrity will result as a concomitant effect of any damage developed in the system. The total number of postulated occur-

! rences for such events shall not cause more than 25 stress cycles.

3.3.4 FAULTED CONDITIONS Faulted conditions are those combinations of conditions associated with extremely low probability - postulated events whose consequences are such that tie integrity and operability of the nuclear energy system may be impaired to the extent that considerations of public health and safety are involved.

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SECTION 4 ANALYTICAL METHODS AND MODELS

4.1 INTRODUCTION

The analytical methods used to obtain a piping deflection solution consist of the transfer matrix method and stiffness matrix formulation.

The complexity of the piping system requires the use of a computer to obtain the displacements, forces, and stresses in the piping and support members. To obtain these results, accurate and adequate mathematical representations (analytical models) of the systems are required. The modeling considerations depend upon the degree of accuracy desired and the manner in which the results will subsequently be interpreted and evaluated. All static and dynamic analyses are performed using the WESTDYN computer program. This program, described in WCAP-8252, was reviewed and approved by the U.S. NRC (NRC letter, April 7,1981 from R. L. Tedesco to T. M. Anderson).

The integrated piping / supports system model is the basic system model used to compute loadings on components, component and piping supports, and piping. The system model includes the stiffness and mass charac-teristics of the piping system. The deflection solution of the entire system is obtained and then internal menber forces and piping stresses are calculated.

4.2 STATIC ANALYSIS The piping system models, constructed for the WESTDYN computer program, are represented by an ordered set of data which numerically describes the physical system.

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-I4-

The spatial geometric description of the piping model is based upon the isometric piping drawings and equipment drawings. Node point coordi-nates and incremental lengths of the members are determined from these drawings. Node point coordinates are put on network cards. Incremental member lengths are put on element cards. The geometrical properties along with the modulus of elasticity, E, the coefficient of thermal expansion, a, the average temperature change from the ambient tempera-ture, AT, and the weight per unit length, w, are specified for each el emen t. The supports are represented by stiffness matrices which define restraint characteristics of the supports. Sketches for various  !

parts of the safety and relief valve discharge piping are shown in figures in Section 6.

The static solutions for deadweight and thermal loading conditions are obtained by using the WESTDYN computer program. The WESTDYN computer program is based on the use of transfer matrices which relate a twelve-element vector [B] consisting of deflections (three displacements and three rotations) and loads (three forces and three moments) at one location to a similar vector at another location. The fundamental transfer matrix for an element is determined from its geometric and elastic properties. If thermal effects and boundary forces are included, a modified transfer relationship is defined as follows:

-T 11 T

~

'A o '

'6

'A i '

12 t 4 > +4 y=4 >

T T F ,f) ,F,

- 21 22- 1 g, t 9 l or TBgg+R1=B1 where the T matrix is the fundamental transfer matrix as described above, and the R vector includes thermal effects and body forces. This B vector for the element is a function of geometry, temperature, coeffi-cient of thermal expansion, weight per unit length, lumped masses, and externally applied loads.

0766s:10-15 . _ _ _ _ _ - _

• ' The overall transfer relationship for a series of elements (a section) can be written as follows:

B1=TBgo+R1 B2=T821+R2=TTB21g+TR21+R2 B3=TB32+R 3=TTTB32gg+TTR321+TR32+R3 or

~

B =l

[n3 T

• Bg +

n E

n h w T r '* R r-1 +R n n r r=2 r (1 -

j A network model is made up of a number of sections, each having an over-all transfer relationship formed from its group of elements. The linear elastic properties of a section are used to define the characteristic stiffness matrix for the section. Using the transfer relationship for a section, the loads required to suppress all deflections at the ends of the section arising from the thermal and boundary forces for the section are obtained. These loads are incorporated in the overall load vector.

After all the sections have been defined in this manner, the overall stiffness matrix, K, and associated load vector needed to suppress the deflection of all the network points is determined. By inverting the stiffness matrix, the flexibility matrix is determined. The flexibility matrix is multiplied by the negative of the load vector to determine the network point deflections due to the thermal and boundary force effects. Using the general transfer relationship, the deflections and internal forces are then determined at all node points in the syste'n.

The support loads, F, are also computed by multiplying the stiffness matrix, K, by the displacement vector, 6, at the support point.

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l

, 4.3 DYNAMIC ANALYSIS The models used in the static analyses are modified for use in the dynamic analyses by including the mass characteristics of the piping and equipment.

4.4 SEISMIC ANALYSIS The lumping of the distributed mass of the piping systems is accom-plished by locating the total mass at points in the system which will appropriately represent the response of the distributed system. Effects of the equipment motion, that is, the pressurizer, on the piping system are obtained by modeling the mass and the stiffness characteristics of the equipment in the overall system model.

The supports are again represented by stiffness matrices in the system model for the dynamic analysis. Mechanical shock suppressors which resist rapid motions are now considered in the analysis. The solution for the seismic disturbance employs the response spectra method. This method employs the lumped mass technique, linear elastic properties, and the principle of modal superposition.

From the mathematical description of the system, an overall stiffness matrix [K] is developed from the individual element stiffness matrices using the transfer matrix [K ] Rassociated with mass degrees-of-freedom only. From the mass matrix and the reduced stiffness matrix, the natural frequencies and the normal modes are determined. The modal participation factor matrix is computed and combined with the appro-priate response spectra value to give the modal amplitude for each mode. Since the modal amplitude is shock direction dependent, the total modal amplitude is obtained conservatively by either the absolute sum or the SRSS of the contributions for each direction of shock. The modal amplitudes are then converted to displacements in the global coordinate system and applied to the corresponding mass point. From these data the 7 forces, moments, deflections, rotation, support reactions, and piping stresses are calculated for all significant modes.

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The seismic reponse from each earthquake component is computed by conbining the contributions of the significant modes.

4.5 PRESSURIZER SAFETY AND RELIEF LINE ANALYSIS l

4.5.1 PLANT HYDRAULIC MODEL When the pressurizer pressure reaches the set pressure (2,500 psia for a safety valve and 2,350 psia for a relief valve) and the valve opens, the high pressure steam in the pressurizer forces the water in the water loop seal through the valve and down the piping system to the pressur-izer relief tank. For the pressurizer safety and relief piping system, analytical hydraulic models, as shown in Figures 4-1 and 4-2, were developed to represent the conditions described above.

The computer code ITCHVALVE was used to perform the transient hydraulic analysis for the system. This program uses the Method of Characteris-tics approach to generate fluid parameters as a function of time. One-dimensional fluid flow calculations applying both the implicit and explicit characteristic methods are performed. Using this approach the piping network is input as a series of single pipes. The network is generally joined together at one or more places by two or three-way j unctions. Each of the single pipes has associated with it friction factors, angles of elevation, and flow areas.

0766s:10-18

Conservation equations can be converted to the following characterisitic equa tions:

1 '

dz

g = V+c i

dP dV qc 2 g + pc g = c(F + pgcoso) ah

• E h=V-c 9

h - pc h = -c(F + pgcoso)

1. E 1

! 2 - ahlap c

= ah 1 57 z = variable of length measurement

t = time V = fluid velocity c = sonic velocity p = pressure p = fluid density F = flow resistance g = gravi ty i o = angle off vertical J = conversion factor for converting pressure units to j equivalent heat units l h = enthal py l q = rate of heat generation per unit pipe length I

i The computer program possesses special provisions to allow analysis of valve opening and closing situations.

0766s:10-19 -

Fluid acceleration inside the pipe generates reaction forces on all segments of the line that are bounded at either end by an elbow or bend. Reaction forces resulting from fluid pressure and momentum variations are calculated. These forces can be expressed in terms of the fluid properties available from the transient hydraulic analysis performed using program ITCHVALVE. The momentum equation can be expressed in vector form as:

F =

pVdv + h pV(V

• ndA) ev y From this equation, the total force on the pipe can be derived:

r y (1 - cos ai) aW r 2 (1 - cos a2) aW pipe *{ sin at af Bend 1 { sin a2 Bend 2

+- straight $ at dl 9

cJpipe A = piping flow area v = volume F = force r = radius of curvature of appropriate elbow a = angle of appropriate elbow W = mass acceleration g

e

= gravitational constant All other terms are previously defined.

Unbalanced forces are calculated for each straight segment of pipe from the pressurizer to the relief tank using program FORFUN. The time-histories of these forces are stored on tape to be used for the subse-quent structural analysis of the pressurizer safety and relief lines.

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I 4.5.2 COMPARIS0N TO EPRI TEST RESULTS Piping load data has been generated from the tests conducted by EPRI at the Combustion Engineering Test Facility. Pertinent tests simulating dynamic opening of the safety valves for representative commercial upstream environments were carried out. The resulting downstream piping  !

loadings and responses were measured. Upstream environments for particular valve opening cases of importance, which envelope the commercial scenarios, are:

A. Cold water discharge followed by steam - steam between the pressure source and the loop seal - cold loop seal between the steam and the valve, B. Hot water discharge followed by steam - steam between the pressure source and the loop seal - hot loop seal between the steam and the valve.

C. Steam discharge - steam between the pressure source and the valve, Specific thermal hydraulic and structural analyses have been completed for the Combustion Engineering Test Configuration. Figure 4-3 illus-trates the placement of force measurement sensors at the test site.

Figures 4-4, 4-5 and 4-6 illustrate a comparison of the thermal hydrau-lically calculated results using the ITCHVALVE and FORFUN computer programs versus experimental results for Test 908, the cold water discharge followed by steam case. Figure 4-4 shows the pressure time histories for PT9, which is located just downstream of the valve.

Figures 4-5 and 4-6 illustrate, respectively, the force time histories of the horizontal run (WE28/WE29) and the long vertical run (WE32/WE33) immediately downstream of the safety valve. Significant structural damping in the third segment after the valve was noticed at the test and was verified by structural analyses. Consequently, a comparison of force WE30/WE31 was not presented here. No useable test data for sensor WE34/WE35 was available for Test 908.

1

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' Figures 4-7 through 4-11 illustrate a comparison of calculated versus experimental results for Test 917, the hot water discharge followed by steam case. Figure 4-7 shows the pressure time histories for PT9.

Fi gures 4-8, 4-9, 4-10 and 4-11 illustrate, respectively, the thermal hydraulically calculated and the experimentally determined force time histories for (WE28/WE29), (WE32/WE33), (WE30/WE31) and (WE34/WE35).

Blowdown forces were included in the total analytically calculated force for WE34/WE35 as this section of piping vents to the atmosphere.

Al though not presented here, comparisons were also made to the test data available for safety valve discharge without a loop seal (steam di scharge) .

The application of the ITCHVALVE and FORFUN computer programs for cal-culating the fluid-induced loads on the piping downstream of the safety l

and relief valves has been demonstrated. Al though not presented here, the capability has also been shown by direct comparison to the solutions of classical problems.

The application of the structural computer programs (discussed in Section 4.5.3) for calculating the system response has also been demon stra ted. Structural models representative of the Combustion Engineering Test Configuration were developed. Fi gures 4-12, 4-13 and 4-14 illustrate, respectively, a comparison of the structural analysis resul ts and the experimental resul ts for locations (WE28/WE29),

i (WE32/WE33) and (WE30/WE31) for test 908. No useable test data for sensor (WE34/WE35) was available. Figures 4-15, 4-16, 4-17 and 4-18 show for test 917, respectively, the structural analysis results versus the test results for locations (WE28/WE29), (WE32/WE33), (WE30/WE31) and (WE34/WE35).

l 4.5.3 VALVE THRUST ANALYSIS The safety and relief lines were modeled statically and dynamically as described in Sections 4.1 through 4.3. The mathematical model used for dynamic analyses was modified for the valve thrust analysis to represent I

the safety and relief valve discharge. The time-history hydraulic 1

0766s:10-22

forces determined by FORFUN were applied to the piping system lump mass points. The dynamic solution for the valve thrust was obtained by using a modi fied-predictor-corrector-integration technique and normal mode theory.

The time-history solution was found using program FIXFM3. The input to this program consists of natural frequencies, normal modes, and applied forces. The natural frequencies and normal modes for the modified pres-surizer safety and relief line dynamic model were determined with the WESTDYN program. The time-history displacement response was stored on magnetic tape for later use in computing the total system response due to the valve thrust conditions. The time-history displacements of the FIXFM3 program were used as input to the WESDYN2 program to determine the time-history internal forces and deflections at each end of the piping elements. For this calculation, the displacements were treated as imposed deflections on the pressurizer safety and relief line masses. The solution was stored on tape for later use in the piping stress evaluation and piping support load determination.

The time-history internal forces and displacements of the WESDYN2 program were used as input to the POSDYN2 program to determine the

+ maximum forces, moments, and displacements that exist at each end of the piping elements and the maximum loads for piping supports. The results from program POSDYN2 are saved on TAPE 14 for future use in piping stress analysis and support load determination.

l 0766s:10-23

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c.

time (seconds)

FIGURE 4-4 : Co-parison of the EPRI Fressure Time-History for P109 from Test 908 with the ITCHVALVE Pre-dicted Pressure Time-History i

l 1.0EL A

v4 \

I

) <

' i

/ i e / l

/ \ / g 0.0 j /

. -~~'

I 3

i I I I

l

- I I I I

-1.0E4 -F

_ i i w I 'i E I l

e t ,

1

\ l

\ l 1

-2.0E4 -

tests ITCHVALVE

-3.0E4 0.05 0.15 0.25 Time (seconds) l l

FIGURE 4-5: COMPARISON OF THE EPRI FORCE TIME-HISTORY FOR 1

WE28 and WE29 FROM TEST 90S WITH THE ITCHVALVE PREDICTED FORCE TIME-HISTORY l

l l__ __ _, ,_ - ... -.

1.0E5 ,

P t

A wA E

0

_ J

--[

w

-1.0E5 I i

l

-2.0E5 t

tests

- ITCEYALVE

-3.0E5 0.1 0.2 0.3 0.4 0.5 time (se::endsl FIGURE 4-6:00TARIS0N OF THE EPRI FORCE TIME-HISTORY FOR WE12 AND WE33 FROM TEST 908 WITH THE ITCHVALVE PREDICTED FORCE TIME-HISTnRy l

I 500. >

r i < \

r g

2- \

I J 400 , . f [\g \

I I'

q

\

\

)

\

\

\

\

~ 300 .

\

G a I

\

\

~

l \

E s

=

1 ^ -

I

~ - -

C \

I t I - ---

~  % ---

200a . I t

I I

I Test l

I 100 .

---

ITCHVAL\T j

I I

I J

. /%

- 0.1 0.2 o,'3 0.4 0.5 time (seconds) i FIGURE 4-7 : Comparison of the EPRI Pressure Time-History from PT09 from Test 917 with the ITCHVALVE Predicted Pressure Time-History

1 4000 i -

2 l

I 2000 m, -

i d pb i I , [

/ 9

.; 0.0 L ..c,j " li' i

~

$. ,O jf fs f sl i

,I '

E l l j k' C ' ' .

' -2000 '

i i .}

-4000 y!;

b'

-6000 tests

---

ITCHVALYq

-8000 0.0 0.1 0.2 0.3 0.4 0.5 0.6 time (seconds)

FIGURE 48: Comparison of the EPRI Force Time-History for WE28 and WE29 from Test 917 with the ITCHVALVE Predicted Force Time-History t

~ ~' ' ~

2.0E4 l

r.

.I 's. .,

1.0E4 , , ,-

l I '

r i I[

'. , i.

7 .!

E .* *

\ .* U R f\ m , e. t O. w . , . gj v vv v 5

w -

/

1 I

i a

\ $

-1.0E4 ,

s tests ITCMVALVE

-2.0E4 0.0 0.1 0.2 0.3 0.4 0.5 time (seconds)

FIGURE 4-9 : Comparison of the EPRI Force Time-History I

i for WE32 and WE33 from Test 917 with the ITCHVALVE Predicted Force Time-History

3.0E tests ITCHVALVE_

2.0E4 .\

t

\

9 8 \

I \

l , \

l t

~

1.0E4 ,

I h\

i  ! h ,'s C. *

\

/

, t 0.

,._ -' \ D

> /

\ /

)

-1.0E4 ,

-2.0E4 0.0 0.1 0.2 0.3 0.4 0.5 i

time (seconds)

FIGUPI 4-10: Comparison of the EPRI Force Time-History For E30 and E31 From Test 917 with the ITCHVALVE Predicted Force Time-History

e 2.0E4 l

i tests

- -- - ITCHVALVE L

l \

C I \

y l \

= I t w i \

U / \

f 1.0E4 s ,

' \

t

\, ,

)

I \

l N I

I I '

l /

I i

i I

I 0.0 s' O.0 01 0' 0.3 o,4 0.5 I

time (seconds)

! FIGURE 4-11: Comparison of the EPRI Force Time-History For WE34 and WE35 from Test 917 with the ITCHVALVE Predicted Force Time-History l

l l

l l

l 20,0 I

10.0 gl I, M 0.0 --

'i ! ii i!6 A : ,_ _ h j ij sj 's ~ '-

3'

{

8 i l h

' l J1 l '

-10,0 '  ; l i

Y N Tests

---

FIXN

-20.0 (Structural Analysis) f

-26.1; 0.05 0.15 0.25 0.35 0.45 Time (Sec.)

FIGURE 4-12: Comparison of the EPRI Force Time-History for WE28 and WE29 from Test 905 with the FIXFF.3 Predicted Force Time-History G

i l

l

\

I 111.01 ' '

100.0 I i

I i

50.0 g'" !,

i' l'

,', a 4

0.0 e *'"

'., 5.I Tra > % ]\ r A

4 g -

o ,

j8 g 's I!pWI

_ -50.0 i E il I l

E !i

$ -1 00.0 -

w

-150.0 Tests L

-200.0 -

(Structural Analysis)

FIXFM3 h'[

-246.92 ' '

O.0 O.1 0.2 0.3 0.4 Time (SEC)

Figure 4-13: Comparison of the EPRI Force Time-History For WE32 l

l and WE33 From Test 908 With the FIXFM3 Predicted Force Time-History 1

l l

l 90.896 3,1 ill -

75.0 i.g f 6

g I ,1 l 'i

'),'I 8 I I 50.0  :

I  !! I 'g I I i e j  ! e 25.0 f i,l ,

,' ', e  !

M

,1 1 I ,i i I I e i i i g i G ' I I i l

^'

i vv- I;

! lI ,

p; 8

lh8 i

I lll :

l I

I li I

8 l

" -25.0 s t i 31 i I

t I ..

-50.0 p . -

4 I l8 1

-75.0 Tests tl

---

FIXFM3 I (Structural Analysis) ll'>

-98.324 0.0 0.1 0.2 0.3 0.4 Time (SEC) 1 i

F*.gure 4-14: Comparison of the EPRI Force Time-History For WE30 and WE31 From Test 908 With the FIXFM3 Predicted l Force Time-History I

l i

5.0 4.0 T

'b 3.0 . ,

2.0 ,'

l u. '

ly o

e I ,. .

s s e l '. n. ,' 8 e . ', l 1.0 -

Ii

/ , ,

J I

~

i l v:~ "\

n  : \

00 9 ;p t;  ; r ~J 1 is i i i

-1.0 i'. .k ,

If l l

) j j fI I f l , b I

m

& -2.0

I f d

i 'p (

s i l

E -3.0  !

2 '

ll'

.i

\

. -4.0 l

?

-5.0 i

i d 1 -6.0 1

Tests -7.0

---

FIXFM3 mc ural Analys s)

-8.0 0.0 0.1 0.2 0.3 0.4 0.495 Time (SEC) ,

Figure 4-15 Comparison of the EPRI Force Time-History For WE28 and

WE29 From Test 917 With the FIXFM3 Predicted Force Time-History l

1 4

l l i

12.956

[,

,\ -

I' t ,

10.0 j j. . a I n I l '\

5.0

! !A ll  !

i  ?' ( jh i l II il 0 f I a }l pl

,f 83 \\ s k \ l ,_ \t \ bsh Til dsf A

q

?-

E

\\\ 1't v

\ j V ,,q' %l ,

i[\

1 0 a (s\ di

-5.0 if l

j l

i I

\ / Tests

\ I

' i ---- FIXFM3

{ ,o (Structural Analysis)

-1.0. 0 i

r e o

-13.265 0.0 0.1 0.2 0.3 0.4 0.495 Time (SEC)

Figure 4-16: Comparison of the EPRI Force Time-History For WE32 and WE33 From Test 917 With the FIXFM3 Predicted Force Time-History

25.853 ' '

25.0 1 Is II

I 20.0 +

h '

1 I l 15.0* j n s

\

10.0 t s,

i \

r \

5.0 , -g I t I i s O ~#

e E 0.0

' ' ' ' ' ' ' ^

r'v o j svim - -

u i ,

j

- l tu f ' p

O \  !

' -5.0 h -i  :/-

Y

-10.0 g

\l Tests

-15.0 -

FIXFM3 Y (Structural Anal-

-20.0 O.0 0.1 0.2 0.3 0.4 0.495 Time (SEC)

Figure 4-17: Comparison of the EPRI Force Time-History For WE30 and WE31 From Test 917 With the FIXFF3 Predicted Force Time-History

I l

i 14.58E o

/ f\

l i\

i l

12.5

! i1 i I

, ti i

\

l \

l \

f I 10.0 I

' i 1 \'

i (

l t t

, l

', \

7.5 i i s

G I \

c- t D

- l g ,

8 5*0 -

5 I / \j Tests N

---

FIXFM3 (Structural dnalysis) i 2.5 , .'

I I

I

/

0.0 - '

i 0.0 0.1 0.2 0.3 0.4 0.495 Time (SEC)

Figure 4-18: Comparison of the EPRI Force Time-History For WE34 and WE35 From Test 917 With the FIXFM3

~

Predicted Force Time-History l

l 1

1

. I l

l l

l SECTION 5 METHOD OF STRESS EVALUATION l l

5.1 INTRODUCTION

The method used to combine the primary loads to evaluate the adequacy of the piping system is described in this section.

5.2 PRIMARY STRESS EVALUATION In order to perform a primary stress evaluation in accordance with the rules of the Code, definitions of stress combinations are required for the normal, upset, emergency, and faulted plant conditions as defined in Section 3. Tables 2-1 and 2-2 illustrate the allowable stress inten-sities for the appropriate combinations as discussed in Section 2.2.

Table 2-3 defines all pertinent terms.

5.2.1 DESIGN CONDITIONS The piping minimum wall thickness, t,, is calculated in accordance with the Code. The actual pipe minimum wall thickness meets the Code requirement.

The combined stresses due to primary loadings of' pressure and weight, calculated using applicable stress intensity factors, must not exceed the allowable limit. The resultant moment, Mg , is calculated using the following equation:

~

1/2 M M 2+ M 2+M Z 2 I= #wt Ywt wt ,

~

~_..

e 1 ,

.)

0766s:10 '

l -

l

where M ,M ,M = deadweight moment components

• t w #wt wt 5.2.2 UPSET CONDITIONS The combined stresses due to the primary loadings of pressure, weight, operating basis earthquake (OBE), and relief valve thrust, calculated using the applicable stress intensity factors, must not exceed the allowables. The resultant moment, M9 , is calculatea as shown below.

For seismic and relief valve thrust loading:

M I=' M*

IM w tjl+k*0BE 2 g 2

• SOT N+M #wt

+ M

1. 0BE 2g 2 YSOT U

(.

- -2 2 2 1/2 1/2

. M . M M.

z Z lZwt; 0BE SOT g b

where M ,M ,M = deadweight moment components z

• wt #wt wt

,M ,M = OBE moment components M* OBE Y OBE

• OBE M ,M ,M = relief line operation moment components

• SOT S0T SOT U U U 0766s:10 . .

' 5.2.3 EMERGENCY CONDITIONS The combined stresses due to primary loadings of pressure, weight, and safety valve thrust, using applicable stress intensification factors, must not exceed the allowable limits. The magnitude of the resultar.t moment, M j , is calculated from the moment components as shown below:

i" X SOT wt YSOT #wt Z

S0T wt E E E where M ,M ,M = deadweight moment components Ywt z

\t wt M ,M ,M = safety line operation moment components

• S0T SOT S0T E E E 5.2.4 FAULTED CONDITIONS The combined stresses due to the primary loadings of pressure, weight, safe shutdown earthquake (SSE), and S0T p , using applicable stress intensification factors, must not exceed the allowable limits. The magnitude of the resultant moment, M j , is calculated from the three moment components as shown below:

f 2 , g 2 1/2 + l 2 M g M I= M*S0T SSE wt F

6 2 2 1/2 2

+ M + M + M YSOT p YSSE #wt

-44~

0766s:10-26

2 2 1/2 '2 1/2

+ +

+

[M# M M

1. wt

( ,

S0T F

SSE where M ,M ,M = deadweight moment components Z

• W t #wt wt

,M = SSE moment components M*SSE,M#SSE #

SSE l M ,M ,M = maximum of S0Tg and SOT m ment components Z E

• SOT SOT p SOT p F

For the safety and relief piping, the faulted condition load conbination of pressure, welght, and valve thrust is considered as given in Tables 2-1 and 2-2 and defined in Table 2-3. The pipe break loads (MS/FWPB or LOCA) can be ignored for the PSARY system. These loads have very little impact on the pressurizer safety and relief system when compared to the loading conditions discussed in this report.

5.3 SECONDARY STRESS EVALUATION The conbined stresses due to all thermal loadings, using applicable stress intensification factors, must not exceed the allowable limit of SA for thermal only or (S h + SA ) f r thermal, pressure, and wei gh t. For the resultant moment loading, M j, thermal moments are conbined as shown below:

2 2

+ MZ -M Z 1/2 M

I= M* MAX - M* MIN g#MAX , g#MIN h2 MAX MIN /

,M # = maximum thermal moment considering all thermal cases M* MAX,M # MAX MAX including normal operation M ,M ,M = minimum thermal moment considering all thermal cases x 7 MIN # MIN MIN including normal operation 0766s:10-27

SECTION 6 RESULTS 6.1 EVALUATION PRIOR TO EPRI TEST PROGRAM i

Tne Beaver Valley Unit 1 safety and relief valve discharge piping system had received a very detailed thermal hydraulic and structural dynamic evaluation to ensure the operability and structural integrity of the l system. This structural evaluation, including the thermal hydraulic analysis, was based on the criteria and methods that were current prior to the availability of the data from the EPRI Test Program. The thermal hydraulic forcing functions were generated assuming simultaneous opening of either the safety valves or the relief valves, since they represent the worst applicable loading conditions for the piping and supports for this specific layout. These forcing functions were then used as input to the structural evaluation in which the primary and secondary stresses were determined. The methods used and the loadings considered are consistent with Sections 2.0, 3.0, 4.0 and 5.0 of this report, respec-tively. Results of this extensive analysis and evaluation have demon-strated that the PSARV piping met all the applicable design limits for the various loading cases. In addition, the acceptability of the valve nozzles and equipment nozzles was assured for the applied loads.

6.2 EVALUATION SUBSEQUENT TO EPRI TEST PROGRAM The Beaver Valley Unit 1 pressurizer safety and relief valve discharge piping system has received a detailed thermal hydraulic analysis and structural evaluation to ensure the operability and structural integrity of the system. The methods used and the loadings considered are consistent with Sections 2.0, 3.0, 4.0 and 5.0 of this report.

0766s:10 1

' 6.2.1 THERMAL HYDRAULIC RESULTS The thermal hydraulic analysis used computer programs which have been shown to match the results of the EPRI Test Program (Section 4.5.2).

Hydraulic forcing functioas were generated assuming the simultaneous opening of either the safety valves or the relief valves since these represent the worst applicable loading cases for the piping and supports  !

I of this specific layout.

Table 6-1 shows the maximum forces on each straight run of pipe for the simultaneous ooening of all three safety valves while Table 6-2 shows the maximum forces for the simultaneous opening of all three relief valves. To account for uncertainties in the valve flow capacities due to tolerances and deviations, a conservative factor of over 1.20 was included in the maximum rated valve mass flow rate for these cases.

This resul ts in conservative forcing functions.

! For the safety valves opening case, hot loop seals were assumed to exist upstream of the valves since the piping has been insulated to eliminate cold loop seals which can induce severe hydraulic forces on the pipng sys tem. The loop seal temperature distribution for this case was presumed to be consistent with the distribution in EPRI test 917. That is, the loop seal temperature at the valve inlet was about 300*F, and approximately eight feet upstream, the loop seal liquid temperature was near the system saturation temperature of 655'F. Based upon engineering judgement, significant flashing of hot water near the valve occurred for test 917, thus reducing the downstream loads significantly.

Based on analytical work and tests to date, all acoustic pressures in the upstream piping calculated or observed prior to and during safety valve hot or cold loop seal discharge are below the maximum permissable pressure. The piping between the pressurizer nozzle and the inlet of tne safety valves is 6-inch schedule 160. The calculated maximum upstream pressure for this size of piping is below the maximum per-missable pressure. A similar evaluation of this inlet piping pheno-menon, applicable for temperatures below 300*F, was conducted and the results are documented in a report entitled " Review of Pressurizer

~

0766s:10-29

Safety Valve Performance as Observed in the EPRI Safety and Relief Valve Test Program", WCAP-10105, dated June 1982.

6.2.2 STRUCTURAL RESULTS Stress summaries for the valve discharge loading cases considered are provided in Tables 6-3 through 6-15. Sketches of the structural models are shown in Figures 6-1 and 6-6.

For purposes of providing stress summaries, the system was broken up into the following three sets of sections: (Figure 6-7)

Section 1: Piping between the pressurizer and the safety valve cutlet nozzles (upstream of valves).

Section 2: Piping between the pressurizer and the relief valve outlet nozzles (upstream of valves).

Section 3: Piping between the safety and relief valve outlet nozzles and the pressurizer relief tank (seismically designed downstream portion).

The results of this extensive analysis and evaluation demonstrated that the piping meets the applicable design Ifmits for the various loading cases. Support load combinations were consistent with piping combina-tions. Applicable allowables for the supports were obtained from the ASME Code, Section NF. In addition, the acceptability of the valve nozzles and equipment nozzles was assured for the applied loads.

6.3

SUMMARY

OF RESULTS AND CONCLUSIONS The thermal hydraulic analysis and structural evaluation of the Beaver Valley Unit 1 pressurizer safety and relief valve discharge piping system have been completed. In summary the operability and structural integrity of the system have been ensured for all applicable loadings and load combinations including all pertinent safety and relief valve discharge cases.

0766s:10 i -

TABLE 6-1 HYDRAULIC FORCES - SAFETY LINE I

Pipe Type Force (LBF) Pipe Type Force (LBF) 1 130 23 130 2 180 24 180 3 3800 25 3700 4 3800 26 3800 5 3800 27 3800 6 2800 28 3500 7 3100 29 3800 8 4800 30 11000 9 3200 31 4800 10 4000 32 130 11 + 12 + 13 44000 33 180 14 9000 34 3000 15 11000 35 3800 16 20000 36 3900 17 12000 37 3900 18 19000 38 3700 19 5200 39 12000 20 11000 40 5400 21 10000 The pipe type numbers correspond to that shown on Figure 4-1.

0766s:10-31

.. . - - . - _ _ _ . - - . - - , _ = _ . - _ .

]

TABLE 6-2 HYDRAULIC FORCES - RELIEF LINE 1

Pipe Type Force (LBF) Pipe Type Force (LBF) 1 70 20 7000 2 250 21 9000 3 60 22 2400 4 220 23 5300

5 340 24 4600 6+7+8 440 26 330 9 + 10 + 11 850 27 390 12 + 13 + 14 3000 28 600

.l 15 + 16 2700 29 630 l 17 6000 30

• 31 + 32 1950 18 7000 33 330 19 12000 34 + 35 + 36 3200 The pipe type nunbers correspond to that shown on Figure 4-2.

l 0766s:10-32 - . , - .

l

. 1 1

• TABLE 6-3 ]

PRIMARY STRESS StM4ARY - SECTION 1 Piping System: Pressurizer Safety Line Conbination 1 - N Node Maximum Allowable Point Piping Component Stress (ksi) Stress (ksi) 5350 Straight run 4.2 15.9 5380 Bu tt wel d 5.7 15 .9 5330 Elbow 4.3 15.9 f

See Tables 2-1 through 2-3 for load conbinations and definitions.

0766s:10-33

TABLE 6-4 PRIMARY STRESS

SUMMARY

- SECTION 1 Piping System: Pressurizer Safety Line Combination 2 - N + OBE + SOT g Node Maximum All owabl e Point Piping Component Stress (ksi) Stress (ksi) 5350 Straight run 4.6 19.1 53 80 Bu tt wel d 6.5 19.1 5370 Elbow 4.8 19.1 i

See Tables 2-1 through 2-3 for load combinations and definitions.

- ~

0766s:10-34

- TABLE 6-5 PRIMARY STRESS SIM4ARY - SECTION 1 Piping System: Pressurizer Safety Line i

Conbination 3 - N + SOT g Node Maximum All owabl e Point Piping Component Stress (ksi) Stress (ksi) 7280 Straight run 7.9 28.6 4

5380 Butt wel d 12.0 28.6 5370 Elbow 8.9 28.6 See Tables 2-1 through 2-3 for load conbinations and definitions.

0766s:10-35

, TABLE 6-6 PRIMARY STRESS

SUMMARY

- SECTION 1 Piping System: Pressurizer Safety Line Combinations 4 and 5 - N + SSE + SOT F

Node Maximum Allowable Point Piping Component Stress (ksi) Stress (ksi) 7280 Straight run 7.9 38.2 5380 Butt weld 12.0 38.2 5370 Elbow 8.9 38.2 See Tables 2-1 through 2-3 for load combinations and definitions.

0766s:10 - _ _ _ _ _ _ _

' ~

. \

TABLE 6-7  !

PRIMARY STRESS SUlHARY - SECTION 2 Piping System: Pressurizer Relief Line Conbination 1 - N Node Maximum Allowable Point Piping Component Stress (ksi) Stress (ksi) 3175 Straight run 7.5 15.9 4160 Bu tt wel d 9.1 15 .9 1060 Elbow 4.8 15.9 1180 Reducer 6.2 15.9 4000 Tee 6.9 15.9 i

i See Tables 2-1 through 2-3 for load combinations and definitions.

l l

0766s:10-37

TABLE 6-8 PRIMARY STRESS SUM 4ARY - SECTION 2 Piping System: Pressurizer Relief Line Combination 2 - N + OBE + SOTg Node Maximum All owabl e Point Piping Component Stress (ksi) Stress (ksi) 3175 Straight run 11 . 4 19.1 3220 Bu tt wel d 17.2 19.1 3110 Elbow 8.0 19.1 1170 Reducer 10.3 19.1 4000 Tee 11.2 19.1 See Tables 2-1 through 2-3 for load combinations and definitions. '

0766s:10-39

4

. TABLE 6-9 PRIMARY STRESS SUMARY - SECTION 2 Piping System: Pressurizer Relief Line Conbination 3 - N + S0Tg Node Maximum Allowabl e Point Piping Component Stress (ksi) Stress (ksi) 3175 Straight run 8.5 28.6 4160 Butt wel d 10.6 28.6 1040 Elbow 5.6 28.6 1180 Peducer 7.8 28.6 4000 Tee 8.8 28.6 See Tables 2-1 through 2-3 for load conbinations and definitions.

i 0766s:10-39 l

TABLE 6-10 PRIMARY STRESS

SUMMARY

- SECTION 2 Piping System: Pressurizer Relief Line Combinations 4 and 5 - N + SSE + S0Te l '

Node Maximum Allowable Point Piping Component Stress (ksi) Stress (ksi) 3175 Straight run 11.6 38.2 3220 Butt weld 17.8 38.2 l

1 3110 Elbow 8.4 38.2 1170 Reducer 10.7 38.2 4000 Tee 11.4 38.2 See Tables 2-1 through 2-3 for load combinations and definitions.

t 0766s:10 I -

- TABLE 6-11 PRIMARY STRESS StM4ARY - SECTION 3 Piping System: Pressurizer Safety and Relief Line Downcomer Conbination 1 - N Node Maximum Allowable Point Piping Component Stress (ksi) Stress (ksi) 1997 Straight run 6.3 15.9 1410 Butt weld 6.3 15.9 1340 Elbow 8.3 15.9 1440 Reducer 5.7 15.9 3280 Branch 9.6 15.9 See Tables 2-1 through 2-3 for load combinations and definitions.

t 0766s:10-41

- TABLE 6-12 i

PRIMARY STRESS SUtNARY - SECTION 3 Piping System: Pressurizer Safety and Relief Line Downcomer Co41 nation 4 - N + SOT'J Node Maximum All owabl e Point Piping Component Stress (ksi) Stress (ksi) 1530 Straight run 16.0 33.7 1540 Butt wel d 16.2 33.7 1540 Elbow 22.5 33.7 1440 Reducer 10.5 33.7 3280 Branch 20.6 33.7 See Tables 2-1 through 2-3 for load cosinations and definitions.

~ ~

0766s:10-42

.- .- = . _ _ _ _ _ _ _ _ _ _ _ = - - _ - - _ _ _ -

. TABLE 6-13 PRIMARY STRESS SUMARY - SECTION 3 Piping System: Pressurizer Safety and Relief Line Downcomer Conbination 3 - N + OBE + SOT g Node Maximum Allowabl e Point Piping Component Stress (ksi) Stress (ksi) 1530 Straight run 16.0 28.6 1540 Butt wel d 16.2 28.6 1540 Elbow 22.5 28.6 1440 Reducer 10.5 28.6 3280 Branch 22.2 28.6 ,

See Tables 2-1 through 2-3 for load conbinations and definitions.

1 0766s:10-43 - _ . - _ _ -

. TABLE 6-14 l

PRIMARY STRESS SUMARY - SECTION 3 l

Piping System: Pressurizer Safety and Relief Line Downcomer Co61 nation 2 - N + SOT g Node Maximum Allowable Point Piping Component Stress (ksi) Stress (ksi) 1530 Straight run 22.5 33.7 i

1540 Butt wel d 22.6 33.7 1540 Elbow 32.7 33.7 1440 Reducer 16.3 33.7 7000 Branch 29.1 33.7 See Tables 2-1 through 2-3 for load cosinations and definitions.

0766s:10 44

TABLE 6-15 PRIMARY STRESS SIM4ARY - SECTION 3 Piping System: Pressurizer Safety and Relief Line Downcomer Combinations 5 and 6 - N + SSE + SOT Node Maximum All owable Point Piping Component Stress (ksi) Stress (ksi) 1530 Straight run 22.5 38.2 1540 Butt wel d 22.6 38.2 1540 Elbow 32.7 38.2 1440 Reducer 16.3 38.2 7000 Branch 29.2 38.2 See Tables 2-1 through 2-3 for load conbinations and definitions.

0766s:10-45

, LEGEND. .

VARIABLE X SPRING HANGER fl./ RIGID SUFPORT 1030 g U

<> 53E0 J 6360 7360 1 SNi'BBER / l o 0210 I

! U l1000 SAFETY VALVE SUPPORT R ! N C ?.* O D E L STIF F NESS ( l OR r.* A T RI X l o 8210 l , 5230 o I

d' Q REDUCER TEE i

l l I 8110 BRANCH o 7220 ONNECTION t>

PRESSURIZER UPPER LATERAL SUPPORTS 8030 4 V

N OR M VALVE < s j

l j B310 PRESSURtZER I NOZZLE l PRESSURIZER

/ g I

I O 8020

(

\

PRESSURIZER \ PRESSURIZER SUPPORT SUPPORT SKIRT N

-- 8000

( PRESSURIZER Figure 6-1. Pressurizer Model ,

1 i

e

. c k E' e' e

. g, , ,,

y e

[. ' ~ _~ -

\T

\/-

2E

g  :

.- I 1

-(.

{  ;. 3

• • .E. E S e V i i- . ;

.. ./

E, r

e 57 e

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~

E/  :

• 1 2  ;

'; mame ; ,

• 3 i E y * '

'5 0

-. . p z e.

[*

. $=

i v ~

o e : < r e I ~

!-f- c -

9 e e  :

[ -

g g

c

= / s J 9

:

E

• # et W

L Q C - .-

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SECTION 2 *

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UNMARKED: SECTION 3 k

Section Definition for . Stress Summary Figure 6-7.

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e SECTION 7 MODIFICATIONS 7.1 ELEVATION OF If>OP SEAL TEMPERATURE It was shown during the EPRI testing that the piping loads developed f rom hot loop seals are of a smaller magnitude than those with colder temperature profiles. The piping between the pressurizer and the valve inlet flange will be covered with an insulated enclosure to reduce the piping loads. The boxes will be constructed in such a way as to allow heat from the non-flooded portions of the seal piping or from an adjacent. exposed section of the pressurizer to maintain the required temperature. The temperature of the loop seal piping just upstream of inlet flange will be elevated to a minimum of 3100 F. This J is slightly higher than the temperature observed during the EPRI Tenting Program.

To provide added reliability, the design will not employ any active heating or control devices. The verification of the proper func-tioning of the seal enclosures will be performed by testing.

7.2 PIPE SUPPORT MODIFICATION Due to the previous analysis described in Section 6.1 and the increased seal temperature, the support load increase has been small and as such the need for modifications has been minor. The snubber of one support will be repositioned to within the movement envelope specified by the vendor. In addition, to the above various anchor bolts will undergo proof testing to higher magnitudes than used during their original installation. The testing will allow the use of I higher allowable design bolt loads. The f actor of safety based on ultimate loads will not be lower than 4.0.

7.3 CONCLUSION

The installation of the above modification will render the safety and relief valve piping in full compliance with all the analysis criteria described in this report.

- - _ - - _ _ _ . _ _ _ . - - _ _ - . _ - _ - _ _ _ - _ . _ . . . . .J