ML16256A199

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Revision 309 to Final Safety Analysis Report, Chapter 3, Design of Structures, Components Equipment and Systems, Appendix 3.9E - Asymmetric Loads
ML16256A199
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WSES-FSAR-UNIT-3 3.9E-1 Revision 307 (07/13)

APPENDIX 3.9E ASYMMETRIC LOADS (DRN 03-2056, R14)This appendix was written with respect to main c oolant loop breaks (MCLBs), which were subsequently eliminated from consideration of mechanical (dynamic) effects via leak-before-break (LBB) arguments.

Elimination of MCLBs by LBB and replacement with branch line pipe breaks (BLPBs) is discussed in Section 3.6.2. LBB methodology is discussed in Section 3.6.3. (EC-8458, R307)This appendix has been retained for the historical record because most of the Design loads that are retained in the RCS component specifications for extended power uprate to 3716 MWt and the Replacement Steam Generators are due to the postulated MCLBs. For extended power uprate to 3716

MWt and the Replacement Steam Generators, branc h line pipe breaks (BLPBs) were postulated, and their effects on the RCS, its supports and interfacing components were compared to those from MCLBs.

Most loads due to BLPBs are bounded by loads due to MCLBs. Where new bounding component loads were generated by BLPBs under power uprate conditions , the component was re-evaluated with respect to the criteria of Section 3.9.3 to demonstrate continued acceptability. (DRN 03-2056, R14; EC-8458, R307)

WSES-FSAR-UNIT-33.9E-

21.0INTRODUCTION

In February 1979, the NRC, in Question 110.1, required LP&L to expand FSAR Subsection 3.9.1.5 tomore specifically address the consideration of asymmetric loads on reactor coolant system components and supports which could result from postulated reactor coolant pipe breaks within cavities located inside the containment. Enclosure 1, attached to the question, described the information required.LP&L had tried unsuccessfully, on several past occasions, to obtain clarification of this particular question.The clarification was warranted because of the following:a) The probability of the type of pipe failures required to be assumed for theseanalyses is very low;b)The nature of the analyses required to perform the stated assessment can vary in sophisticationto a significant degree so that the time and expenditures required for the performance of saidassessment can also vary quite substantially.c)LP&L has already modified the plant original design to severely limit the area of the breaks thatcould be postulated to occur in both the reactor and steam generator/pump compartment and has increased the load carrying capacity of the fuel spacer grids and guide tubes; andd)an analysis of the response of the primary system to break at locations specified in Enclosure 1was performed and reported in FSAR Appendix 5.4A, and no formal feedback has been obtained to date form the Commission for that submittal.On August 13, 1980, LP&L personnel and LP&L consultants met with the NRC MEB staff inBethesda to present LP&L's intended approach in responding to Question 110.1. At this meeting, LP&L indicated that several of the components listed in Enclosure 1 to Question 110.1 had already been assessed by either specific consideration of asymmetric loads or in the design of the component/structure. LP&L further indicated that it intended, to the extend deemed technically feasible, to demonstrate the adequacy of the remaining components/structures by comparison type analyses with other similar plants which had been previously analyzed.On October 1, 1980, at the draft SER Review Meeting between the NRC MEB Staff, MEB Consultants,LP&L Staff and LP&L's Consultants, LP&L's Consultants presented the basis and the results of an evaluation of the various components/structures listed in Enclosure 1 to Question 110.1. At theconclusion of the meeting and as indicated in the "Action Items" of the meeting minutes (letter from LP&L to R.J. Bosnak, dated October 24, 1980), LP&L was to provide a written report of the evaluation.

Accordingly, such a report containing the information presented at the meeting and responses to the MEBStaff/Consultants questions was submitted to the NRC on November 14, 1980.

WSES-FSAR-UNIT-33.9E-3Initial MEB Staff review of the report generated a set of informal questions and requests for additionalinformation. Responses to the questions and requests were discussed at a meeting with the NRC in Bethesda on January 22, 1981. At this meeting the NRC MEB Staff recommended that LP&L respond to the asymmetric load questions in a manner and format consistent with NUREG-0609, a copy of which was made available to LP&L Consultants (Summary of meeting on LOCA Asymmetric Load Analysis, Docket 50-382, dated February 14, 1981).Accordingly, the subsequent sections which present the basis, analytical methods and results of theevaluation of the capability of the various components/structures to withstand asymmetric loads are organized in the same way as NUREG-0609.Replies to questions raised by the NRC Staff and their Consultants, and additional information theyrequested to enable them to conduct expeditious review, are incorporated either in the subsequentsections or in attachments at the end of the report.Table 1.1 summarizes the status of the assessment of each component/structure requested in Question110.1. Details of the assessment are presented in the following sections.

2.0BACKGROUND

During a postulated loss-of-coolant accident, thermodynamic and hydrodynamic induced loads occurthrough the primary system. When the loss-of-coolant accident is in the form of a circumferential pipe rupture at the inlet nozzle of the reactor pressure vessel, a decompression of the reactor pressure vessel occurs within a short period of time. Decompression waves which originate at the postulated break travel around the inlet plenum and propagate downward along the downcomer annulus. The finite time required by the decompression disturbances to travel about the vessel causes a transient pressure differential field t be created across the core support barrel (CSB) and the vessel inner surface. This field imposes a transient asymmetric loading on the core support barrel as well as the vessel itself. Since the postulated pipe break is located within the biological shield wall, the blow down fluid flashing into the reactor cavity also causes a transient pressurization acting externally on the vessel. This external pressurization is also asymmetric. The internal asymmetric loading (IAL) and external asymmetric loading (EAL) act in the same direction for breaks occurring in the cold leg piping. For breaks in the hot legs, the internal asymmetric load is virtually absent in the horizontal direction, hence the two loads are additive in the vertical direction only. These loadings are transmitted to the reactor vessel support system. The resultant reaction forces at the support interfaces must be considered in the evaluation of the adequacy of the support system together with the thrust load resulting from the break and other operating loads. Normal operating loads, together with postulated seismic loads, as well as the EAL and IAL for both cold leg and hot leg breaks at the vessel nozzle had been previously analyzed in FSAR Appendix 5.4A with regard to vessel support.Breaks outside the reactor cavity can result in IAL imposed on the reactor pressure vessel and internals,and in EAL on the reactor coolant pump and steam generators. For the breaks outside the cavity, the adequacy of the primary system supports has been assessed for those breaks at appropriate primarysystem locations listed in Appendix 3.6A, as part of the design requirement.

WSES-FSAR-UNIT-33.9E-43.0DEVELOPMENT OF LOADING FUNCTIONS3.1DETERMINATION OF POSTULATED BREAK SIZESThe evaluation of the primary system for LOCA loads requires definition and development of the loadingtransients associated with the postulated pipe ruptures. In turn these transients depend on the postulated break locations, and the break conditions (i.e., break area and opening time), which depend on the primary system configuration and the presence of restraints to limit the size of the break. The Waterford 3 primary system has been provided with system restraints specifically designed to limit the size of the breaks in the main piping and such that there will be no unacceptable deformation in the RCS piping.The Waterford 3 primary system piping was subjected to the a detailed stress survey. Implementation ofthe stress survey criteria indicated in FSAR Subsection 3.6.2.1.1.1, which criteria are consistent with thecriteria delineated in Reference 3, results in the postulation of a pipe break location at the terminal ends of each leg of the RCS main loop piping and at two of the three intermediate elbows in each of the pump suction legs. Implementation of the detailed finite element stress analysis criteria of Subsection3.6.2.1.1.1 results in the postulation of a longitudinal type pipe break at each elbow break location,oriented within 90 degrees circumferentially from the crotch of the elbow.The design basis pipe break sizes are as presented in Reference 3, except as discussed in Subsection3.6.2.1.1. The set of pipe stops and system restraints used in Waterford 3 are within a range ofparameters for which these break sizes were confirmed using Reference 3 methodology.The break opening times and the break sizes are also the same as given in CENPD 168 since they resultfrom the detailed analyses presented in that document for a system which is identical to Waterford 3.The circumferential break opening times and break sizes for Waterford summarized in Table 3.1.Longitudinal breaks are reported in FSAR Table 5.4A-3.Figure 17 and 22 and Table 4-1 of CENPD 168 (Reference 3) describe the Waterford 3 RCS restraintsand pipe stops.The time varying forcing functions used and the time varying responses of the RCS main loop piping forWaterford are presented on Figures 15, 16 and 27 through 30 of CENPD 168 (Reference 3).Each of the postulated breaks produces transient loading conditions on the primary system. These loadscan be characterized into four components:a)Subcooled blowdown loads - dynamic hydraulic forces within the primary system due torapid subcooled depressurization of the system.

WSES-FSAR-UNIT-33.9E-5b)Cavity pressure loads - resulting from pressurization of the space in the reactorcavity or in the steam generator compartment around the components.c)Strain Energy Loads - due to the release of strain energy where the rupture occurs.

d)Jet Impingement and Jet Thrust Loads - which act at the break location.

Each of these load components have been addressed in the Waterford 3 asymmetric load assessment.

3.2INTERNAL SUBCOOLED BLOWDOWN LOADSA thermohydraulic analysis was performed in accordance with the procedures and models developed inReference 4 for the controlling breaks of those described in Subsection 3.1 of this report. The thermohydraulic analyses for this evaluation used an approach which anteceded the presently approved CE approach embodied in their Topical Report CENPD-252-P as referenced in NUREG-0609. The approach used in the Waterford 3 analyses, which are explained in Appendix 5.4A, produces subcooled loads which are larger than those which would be produced by using the computer program CEFLASH-4B described in the topical above. The methodology of Reference 4 utilized the computer program WATERHAMMER and CEFLASH-4 for the analysis of cavity breaks only.In the methodology of Reference 4, both programs were used since WATERHAMMER predictedasymmetric pressure differentials during the initial subcooled portion of the blowdown that were larger than the corresponding differentials predicted by CEFLASH-4.WATERHAMMER results were thus employed in the subcooled regime and joined to CEFLASH-4 resultswhich were then used for subsequent pressure difference distribution calculations by CE.Reference 4 provides comparisons of this analytical approach with experiments and also the models usedfor WATERHAMMER and CEFLASH-4 to represent both hot leg and cold leg breaks.The pressure and flow parameters resulting from the application of the above thermohydraulic modelswere used to compute three-dimensional time history forcing functions acting on the reactor vessel and the internals, using the break sizes and opening times described in Subsection 3.1The loading histories thus derived were combined with the cavity pressure loads strain-release loads andthrust loads, and applied to the structural model of the primary system.3.3CAVITY PRESSURE LOADSAnalyses to calculate the asymmetric subcompartment pressures were performed by Ebasco using massand energy release data provided by Combustion Engineering.The CEFLASH-4A computer program was used to calculate the mass and energy release rates from thepostulated reactor coolant system pipe ruptures. This program is described in Subsection 6.2.1.1.4 of CESSAR (Reference 5). The reactor coolant system nodalization scheme utilized to compute the mass and energy releases is given in FSAR Figure 6.2-20.

WSES-FSAR-UNIT-33.9E-6FSAR Subsection 6.2.1.2.3 describes the assumptions made in order to maximize the short-termblowdown rates to the the subcompartments.The same section also describes the assumptions made in analyzing the pressure response of the reactorcavity and steam generator subcompartments.The analyses of these subcompartments utilized the RELAP-3 Mod 68 computer code (Reference 6),modified to allow the use of a multiplier in the Moody choked flow correlation of 0.6. The RELAP-3 Mod68 code was the best code available at the time of this analysis, and predated the RELAP-4 Mod 5 currently utilized at Ebasco for such analyses. The coefficient of 0.6 was used for physical junctions.Further description of the modeling assumptions used in the analyses are provided in FSAR Subsection 6.2.1.2.3.The asymmetric pressure loads calculated by these analyses were converted to forces time historiesacross the reactor vessel and steam generators, which combined with the asymmetric internal subcooleddecompression loads, strain energy release loads and thrust loads, were applied to the structural model ofthe Reactor Coolant System. The asymmetric pressure time histories were also applied to the biological shield wall and a the steam generator subcompartment walls as equivalent static loads using peak differential pressures calculated in the different regions of the pertinent subcompartment models.4.0STRUCTURAL ANALYSIS4.1PRIMARY SYSTEM ANALYTICAL MODELThe major components of the Reactor Coolant System are designed to withstand the forces associatedwith the design basis pipe breaks discussed in Section 3.1 of this report in combination with the safe shutdown earthquake and normal operating conditions. The forces associated with the postulated pipe breaks include pipe thrust forces and tension release forces at the break location as developed in CENPD 168 (Reference 3), resultant subcompartment differential pressurization forces and hydraulic forces acting on the reactor internals and the reactor pressure vessel.The model utilized for the dynamic analysis is a lumped parameter model discussed in CENPD 168(Reference 3) which includes details of the reactor vessel and supports, major connected piping and components, and the reactor internals. This model is also described in FSAR Appendix 5.4A.Appendix 5.4A also describes the approach taken to reduce the detailed model of the RCS and reactorinternals to the condensed model used to evaluate the overall system response to the LOCA loads.4.1.1REACTOR AND STEAM GENERATORThe reactor vessel representation is shown on Figure 5.4A-9 of FSAR Appendix 5.4A. The internal andexternal asymmetric forces are applied at the indicated nodes. The steam generator vessel representation is shown in the flexibility analysis model of Figure 5.4A-7 of FSAR Appendix 5.4A.

WSES-FSAR-UNIT-33.9E-74.1.2PRIMARY COOLANT PIPINGThe model of the primary piping is also shown on Figure 5.4A-7 of FSAR Appendix 5.4A.

4.1.3COMPONENT SUPPORTS The Waterford 3 component supports are modeled as springs with gaps. The stiffnesses of the springsfor the pump and steam generator supports are defined in Table 4-1 of Reference 3 along with the hot gapallowed. The vessel support stiffnesses and gaps are as follows.In the vertical direction there is no gap between the vessel and supporting ring girder except a maximumhot gap of 0.015 inches at the hold-down studs is provided. The support stiffness is 2.255 x 10 7 lb/inupward and 1.087 x 10 8 lb/in in the downward vertical direction.Horizontally, a minimum stiffness of 2.864 x 10 7lb/in exists in the tangential direction. The nominalhorizontal gap between the reactor vessel and the cavity wall is less than 0.035 inches in the tangential direction.A detailed description of the reactor vessel steam generator and pumps support and restraint system isgiven in FSAR Subsection 3.8.3.1.5. FSAR Figure 3.8-34 illustrates the reactor vessel support system.

Figures 3.8-35 and 3.8-36 illustrate the steam generator supports and Figure 3.8-37 shows the reactor coolant pumps supports. Figures 3.8-38 and 3.8-39 show details of the pipe restraints designed to limitthe break sizes.4.1.4INTERNALS The model of the internals is shown on Figure 5.4A-8, of FSAR Appendix 5.4A.4.1.5FUELThe fuel model is shown on Figure 5.4A-8 of FSAR Appendix 5.4A. Additional modeling of the fuel for thepurposes of its analysis is addressed by the Response to Question 231.2.4.1.6CONTROL ELEMENT DRIVE MECHANISMS (CEDMs)

CEDMs were not specifically modeled for Waterford. For a discussion of the CEDM analysis please referto Subsection 5.1.8 of this report.4.1.7PIPING ATTACHED TO THE RCSTwo different models have been used for the analysis of RCS attached piping in Waterford 3. One modelis used for a dynamic non-linear analysis of one of the lines. One line only was analyzed in this fashion since the computer execution cost is very high. The model that was utilized is shown on Figure 4.1. The computer program utilized in the analysis is PLAST which is described in detail in FSAR Appendix 3.6B.

The program allows consideration of geometric (i.e., gaps) as well as material (i.e., plasticity) non-linearities in both the piping and the supports/restraints placed on the pipe.

WSES-FSAR-UNIT-33.9E-8The analysis performed with PLAST as evidenced in the results discussed in Subsection 5.1.7 of thisreport did not reveal any material non-linearities and hence it was in effect an elastic, non-linear analysis.The remaining lines were analyzed using the computer program PIPESTRESS2010 described in FSARSubsection 3.9.1.2.1, expanded to permit computation of the maximum linear elastic response of the line, support/restraint system by the generalized response analysis method. This method with consideration of the higher modes, is chosen instead of true time history for cost saving purposes even though it is knownto produce conservative bounding answers.The line analyzed by non-linear analysis in true time history has also been analyzed by the generalizedresponse method so that direct comparison is possible between the two analyses.4.2BIOLOGICAL SHIELD AND SECONDARY SHIELD WALLSThe description of the biological or primary shield wall is provided in FSAR Subsection 3.8.3.1.1. Thedescription of the secondary shield wall is provided in FSAR Subsection 3.8.3.1.2. FSAR Subsection 3.8.3.4.1.2 describes the models and computer programs used to design both of these walls.5.0RESULTS OF ASSESSMENTResults of the assessment of the Waterford 3 RCS System and its support to withstand the loads resultingfrom the postulated design basis loads of Section 3.1 of this report are given in the following subsections.5.1REACTOR VESSEL AND STEAM GENERATOR5.1.1REACTOR VESSEL The acceptance criteria for the reactor vessel are given in FSAR Subsection 3.9.1.4.1.1. Analysesreported in FSAR Appendix 5.4A indicated that when the LOCA resulting stresses are combined withnormal operating and seismic stresses, the vessel meets the acceptance criteria.5.1.2STEAM GENERATOR AND REACTOR COOLANT PUMPSThe steam generator has also been demonstrated to meet the acceptance criteria under the design basisloads resulting from the breaks postulated in Section 3.1 of this report. The pumps have not been specifically analyzed except as part of the reactor coolant piping analysis. Refer to Subsection 5.1.7 of this report.Maximum reactor vessel displacements have been imposed upon a flexibility analyses model such asdescribed in Figure 5.4A-7 of FSAR Appendix 5.4A in order to account for the load on the steam generator supports as a result of vessel motion.

WSES-FSAR-UNIT-33.9E-95.1.3REACTOR VESSEL SUPPORTSThe analysis reported in FSAR Appendix 5.4A has demonstrated the adequacy of the reactor vesselsupports to withstand the asymmetric loads resulting from the 350 in.

2 cold leg vessel inlet, 100 in.

2 hotleg vessel outlet, and 600 in.

2 hot leg steam generator inlet guillotine breaks. The response of the vesselsupports for other breaks needed no evaluation since it is clear that those breaks would result in lower loads. The absence of external asymmetric loads for those breaks, postulated to occur outside the reactor vessel cavity, is what produces the lower overall loads since (a) thrust loads remain the same or are less for cold leg and hot leg breaks respectively outside the reactor cavity as they are for the cold leg and hot leg breaks inside the cavity, and (b) internal asymmetric loads are not significantly affected by the size of the break area alone but rather depend on the combination of the break area and the break opening time. Sensitivity studies performed for instance, for cold leg breaks, have shown that a full areaguillotine break (1414 in.

2) requiring approximately 30 msec to fully open, a 576 in.

2 guillotine breakrequiring 18 msec to open, and a 188 in.

2 guillotine break requiring 8 msec to open, result in internalasymmetric loads differing only by a few percent. Refer to Table 5.8 for the results of such sensitivity studies. Figure 5.1 shows the behavior of the break area developed as a function of time following an assumed instantaneously occurring severance of the cold leg pipe at the vessel inlet nozzle. This figure shows that to each break area there corresponds an opening time.Table 3.1 reports the opening time required to achieve the circumferential breaks postulated in the RCS ofWaterford 3 which are also tabulated in FSAR Table 6.2-1.FSAR Table 3.8-36 indicates the margins between the allowable stress or load on the reactor vesselsupport and the actual stress or load.FSAR Table 3.8-22 specifies the load combinations which were employed in the assessment.

5.1.4BIOLOGICAL (PRIMARY) SHIELD WALL The adequacy of the biological shield is insured by design since the wall has been designed to withstandloads in excess of those transmitted by the reactor vessel supports as computed in FSAR Appendix 5.4Ain combination with the asymmetric pressure loads across the wall as described in FSAR Subsection 6.2.1.2, and other normal operating and seismic loads.A comparison between the calculated design loads for this wall and the ultimate capacity of the wall isshown in FSAR Table 3.8-32/ FSAR Table 3.8-21 lists the loading combinations employed in thisassessment of the secondary shield wall.5.1.5ASSESSMENT OF STEAM GENERATOR AND PUMP COMPARTMENT WALL(SECONDARY)The adequacy of the compartment wall surrounding the reactor coolant pumps and steam generator isalso assured by design. Refer to FSAR Subsections 3.8.3.3.1 and 3.8.3.3.2 for the structural design and Subsection 6.2.1.2 for the subcompartment analyses providing the asymmetric loads across the wall for the postulated breaks (see FSAR Table 6.2-1 for the breaks postulated in the compartments). FSAR Table 3.8-32 indicates the margins of safety for this wall.

WSES-FSAR-UNIT-33.9E-105.1.6ASSESSMENT OF STEAM GENERATOR, REACTOR COOLANT PUMPSUPPORTSThe asymmetric pressure histories reported for the breaks postulated in FSAR Subsection 6.2.1.2 havebeen used to determine the asymmetric loads across the steam generators and reactor coolant pumps.

These components have been designed to accept these loads as stated in Subsection 5.1.2 of this report.

The adequacy of the steam generator and reactor coolant pump supports is also assured by design.Refer to FSAR Subsections 3.6.2.3 and FSAR Appendix 3.6A for the design of those supports andrestraints. FSAR Subsection 3.8.3.4.1.2, Items b3), b5), and b6) describe the basis and models used for the design. FSAR Tables 3.8-33, 3.8-34, 3.8-35, and 3.8-31, illustrate the margins between allowable and maximum calculated stresses for the pump supports, steam generator upper and lower supports and reactor coolant pipe stops respectively.5.1.7ASSESSMENT OF PRIMARY COOLANT PIPINGThe adequacy of the reactor coolant piping under LOCA conditions had been verified when piping rupturerestraints limiting the break areas were backfitted into the original design (refer to FSAR Subsection3.6.2.3 and FSAR Appendix 3.6A). The Waterford 3 design falls within the bounds of the primary system design of CENPD 168 (Reference 3), hence the RCS piping is adequate to withstand the LOCA loads by reference to CENPD 168. Moreover, the Waterford 3 design being very similar to that of St. Lucie 1 (Reference 7), and both having seismic loads which are an order of magnitude less than the allowable loads, and the St. Lucie 1 RCS piping having been demonstrated (Reference 7) to be adequate by plant specific analysis for combined LOCA and seismic loads, then it is judged that the Waterford 3 RCS piping should also be adequate since its expected stresses are lower than those of St. Lucie 1 for the following reasons: (a) design basis break areas in Waterford 3 are smaller than break areas of St. Lucie 1 at corresponding locations, hence asymmetric effects are smaller, and (b) conservatively estimated vesselmotions resulting from asymmetric loads due to pipe breaks within the cavity in Waterford 3 aresignificantly smaller than the corresponding vessel motions in St. Lucie 1 (refer to Figures 5.2, 5.3 and 5.4). These two results stem from the significant difference in the forces across the vessel generated by the pressures in the reactor cavities. Figure 5.5 shows the horizontal and vertical forces acting across the vessel for various plants with reactor cavities which are not necessarily very similar. The force computed for the Waterford 3 break area of 350 in 2 in the cold leg from pressures reported in FSAR Subsection6.2.1.2 is seen to be considerably less than that for the generic plant full break, and to essentially fall in line with the trend of lateral force vs. break area shown in Figure 5.5.Table 5.1 compares the seismic moments at selected points in the RCS main piping for St. Lucie 1 andWaterford 3.5.1.8ASSESSMENT OF CEDMS Control rod insertion for break areas exceeding 0.5 square feet is not required due to NSSS design.Hence, it is only necessary to demonstrate that the CEDMs retain their integrity.

WSES-FSAR-UNIT-33.9E-11The CEDMs in Waterford 3 are identical to those of the CE generic plant (Reference 1) and St. Lucie 1(Reference 7), with the exception that in the latter two plants, all nozzles are of equal length; whereas Waterford 3 CEDM nozzles have unequal length resulting in all CEDMs tops being at the same elevation.

This difference causes the CEDMs differing in length from those analyzed for the CE generic plant and St.

Lucie 1 (referred to hereinafter as the short nozzle CEDMs) to respond to lower frequencies on a mode bymode comparison than those previously analyzed. Table 5.2 lists the frequencies of the first seven modes for both the shortest and longest Waterford 3 CEDMs.In Waterford 3 the conservatively estimated amplitudes of the Reactor Vessel head motions (whichprovide the excitation to the CEDMs) are at most one-half of those of the generic CE plant for the cold leg break and the amplitudes of the motions for the hot leg break are another 30 percent lower. This is evident from Table 5.4A-1 of FSAR Appendix 5.4A wherein peak forces on the supports from the 100 in 2outlet nozzle guillotine break are 72 percent of those resulting from the 350 in 2 guillotine break at the inletnozzle. Since, as stated in FSAR Subsection 5.4A.6 of that Appendix, the horizontal effect of all loads on the system causes the vessel to traverse the horizontal support gaps, impact the supports and remain incontact with the support, oscillating about a deflected, close gap position, it follows that the horizontaldisplacements from cold and hot leg breaks are simply proportional to the loads after the initial deflection.

Because of the orientation of the supports with respect to the break, the initial deflection to close the gaps for a hot leg break is the same as for a cold leg break. Subsequent displacements for the hot leg break, however will be about 30 percent smaller since the computed loads from FSAR Appendix 5.4A are about30 percent smaller.For the cold leg break, the dominant frequency of the driving force is between 9-12 Hz for Waterford 3.This is sufficiently close to possible modal resonances for any length CEDMs to be concerned with amplification of motion. For the generic plant and for St. Lucie 1, the dominant frequency of the drivingmotion is about 17 Hz. On a purely elastic basis, long-term excitation of their CEDMs (all short) could beexpected to result in two-fold amplification of motion as measured at the center of the mass. Results of an elasto-plastic analysis for the St. Lucie 1 CEDM nozzle (Reference 7) revealed instead an amplification of approximately 25 percent. This is attributed to plasticity in the nozzle as well as short time application of the input motions. A similar behavior can be expected for Waterford 3, except that a higheramplification would be predicted on an elastic basis (three and 1/2 times that of St. Lucie 1 for any lengthnozzle). The expected amplification from an elasto-plastic analysis would thus be of the order of 85 percent. Since the amplitudes of the driving displacements in Waterford 3 are half of those of St. Lucie 1, the maximum moment that can be expected for Waterford 3 is less than that computed for St. Lucie 1 (which was 173,000 in-lbs) and is expected to be approximately 130,000 in-lbs.*

WSES-FSAR-UNIT-33.9E-12The ultimate moment that the Waterford 3 nozzle can carry is approximately 356,m000 in-lbs (same as St.Lucie 1) as shown in Figure 5.6 and has been computed using the approach of Gerber (Reference 2) with minimum ASME property values for the nozzle material. To assure that the pressure boundary is not violated, the criterion used is from the ASME Code, Appendix F, that the integrity of the pressure boundary is assured if the applied loads do not exceed 70 percent of the plastic instability load. Since the maximum LOCA moment for cold leg breaks in Waterford 3 is well below this 70 percent criterion (249,000 in-lbs), it can be stated that the Waterford 3 CEDMs are adequate for cold leg breaks. The maximum seismic moment that the nozzle sees is 104,000 in-lbs by previous actual plant seismic analysis. The absolute addition of the LOCA and seismic moments would thus be less than 70 percent criterion.For hot leg breaks, the dominant frequency of the driving force is about 16-17 Hz, hence it is not near anyresonance with Waterford 3 CEDMs and further it would excite higher modes than the cold leg break.

Since the amplitude of the driving motions for the hot leg break are even less than those for the cold leg break, the adequacy of the CEDMs for this break is assured.5.1.9RCS ATTACHED PIPING ASSESSMENTFor the ECCS lines, an assessment of the adequacy of the lines has been initially made by comparing therouting of the Waterford 3 lines with that of the ECCS line which had been analyzed in St. Lucie 1. That analysis had shown that some plasticity might occur in the vicinity of the first elbow after the nozzle. For the RCS attached line which is shown in FSAR Figures 3.6A-25 and 3.6A-26 for instance (one of the six lines attached to the RCS for Waterford), the first elbow is that one directly below the intersection of line 1RC42-2RL2 and line 1RC14-45RL2. That initial analysis had been predicated on the assumptions that one can conservatively estimate the bending moment at that location by use of the simple equation where a is the displacement at the nozzle, M is the maximum bending moment, L is the length (the length of the pipe section between the nozzle and the elbow), E is Young's Modulus, and I is the moment of inertia.

Referring to the same example of FSAR Figures 3.6A-25 and 3.6A-26, the nozzle displacements are applied at the intersection of 1RC42-2RL2 and the centerline of 1RC14-45RL2, the length L is measured from that point downward to the midpoint of the first elbow, which is then assumed to be fixed and restrained against rotations by the rest of the piping system via line 1RC14-45RL2. Inherent in this simplistic approach was the modeling of a dynamic event (i.e., time dependent displacement or accelerations applied at the nozzle resulting in inertia forces which induce primary stresses) by a staticcantilever subject to a displacement .M = 3 EI/L 2_______________________*based on Dynamic Amplification Factor = 1 where f is the driving frequency and f n the modalfrequency.

2 n f f-1 WSES-FSAR-UNIT-33.9E-13The loads resulting from the simplistic model application therefore were taken as primary loads eventhough as static anchor displacement they could be construed to be secondary loads. The reason for originally employing the simplistic model were two-fold: (a) the exorbitant cost of performing potentially elasto-plastic nonlinear analyses for each of the six lines attached to the RCS piping and (b) the reasonably good agreement between the moments computed in the line by this method vs. the sophisticated analytical method employed in St. Lucie 1.For instance, the maximum bending stress computed by the above formula for St. Lucie 1 is 22,880 psiwhich compares very well with the 19, 280 psi computed by time history nonlinear elasto-plastic analysis of that line. That analysis had also shown that everywhere else in the line, stresses in combination with seismic stresses were within allowable limits for St. Lucie 1.In Waterford 3, the maximum bending moments had originally been computed on the basis that nozzledisplacements would be half or less than the St. Lucie 1 nozzle displacements (since vessel displacements are one-half or less). The maximum bending moments were computed for two RCS attached lines; one a 12 inch and one a 14 inch nominal line, are 3.2 x 10 6 in-lbs and 4.1 x 10 6 in-lbs,respectively and they occurred a the first elbow near the nozzle at the RCS pipe.Because of the first elbow plasticity, as computed in this manner, consideration of simultaneous seismicloads had been given as follows: the maximum bending moment due to LOCA and that due to seismic had been added absolutely and their sum has been compared to 70 percent of the ultimate capacity of the elbow to carry moment without collapse. The calculated moments are compared with 70 percent of themaximum bending moment carrying capacity of those elbows computed by the methodology of Gerber(Reference 2), which are 5.34 x 10 6 in-lbs and 5.81 x 10 6 in-lbs, respectively for the 12 inch and the 14 inch pipes scaled by the B 2Index in the ASME Code, where B 2 = 1.9026 for the 12 inch line and B 2 = 2.04609 for the 14 inch line.The simpler method of computing the capability of the elbow to sustain an applied moment had beencompared to a finite element analysis method employed in Reference 1 and was found to given excellentagreement. For instance, the simple method predicted maximum collapse moment for the elbow of an ECCS line in St. Lucie 1 to be 5.94 x 10 6in-lbs whereas the finite element method of Reference 1 for avirtually identical elbow computed it to be 5.5 x 10 6 in-lbs.For the two lines examined, the total bending moment (LOCA and seismic) was less in one instance andslightly higher in the other instance than 70 percent of the collapse moment. Hence, it could be concludedthat the ECCS lines would retain their integrity and function during a simultaneous LOCA and seismic event. This conclusion is predicated on the similar conclusion derived in the CE analyses of the lines attached to the RCS for the generic plant and for specific plants as well (Reference 1). In this analysis CE has computed that even at 70 percent of the ultimate moments, plastic strains are of the order of 2percent. At that strain level, CE concludes that functionability is not impaired. Refer to Subsections 4.5.3and 4.9.3 of Reference 1 for a description of the methods employed in CE's analyses, and for the results. provides additional details on the comparative analyses used in this evaluation.

WSES-FSAR-UNIT-33.9E-14Subsequent questions raised during the draft SER review meeting (between MEB Staff, MEB Consultants,LP&L's Staff and LP&L's Consultants) in New York on October 1, 1980, as to whether the seismic supports of the RCS attached lines would be able to withstand the reaction forces resulting from LOCA-induced motions, caused departure from the simple analysis which had been employed to determine the bending moment in the pipe. An elasto-plastic analysis has been performed on the basis of bounding Waterford 3 specific vessel displacements derived from FSAR Appendix 5.4A, which are assumed to adequately represent the ECCS nozzle displacements. The analysis concludes that the analyzed line and its supports/restraints are capable of withstanding loads resulting from the combined postulated LOCA and seismic event. This analysis has been carried out for one of the RCS attached lines only, because of the large cost involved.The analysis utilizes the PLAST Code which is a direct integration piping dynamic analysis program whichallows the consideration of geometric (i.e., gaps) as well as material (i.e., plasticity) nonlinearities in thesupports/restraints and pipe itself. A description of this code is provided in FSAR Appendix 3.6B. The input motions at the nozzle utilized in the full dynamic analysis of the Waterford line are shown in Figure 5.7 where X and Z represent the time histories of the nozzle deflections in the two orthogonal horizontaldirections and Y is the time history of the vertical nozzle deflection. Displacement time histories havebeen utilized instead of acceleration time histories since the former were available and the PLAST code can accept either as input.The output of the program provides information regarding the maximum reaction forces due to the inertialLOCA motion alone at the various restraints, and a also the maximum moments in the different points of the piping system. The maxima are obtained by searching through the entire time histories of the forces and moments calculated by the program.The piping model utilized for the dynamic analysis is shown on Figure 4.1. the maximum moment(resultant) in the Waterford 3 piping system is 7.8 x 10 5 in-lbs occurs at the first elbow (closest to thenozzle to the primary piping).The maximum reaction forces at various restraints is reported in Table 5.3.

The maximum bending moment in the piping calculated by the actual dynamic analysis are considerablybelow 70 percent of the ultimate moment carrying capability. Strains are elastic, and so that the piping experiences extremely small deformations and retains its cross-sectional area, thus ensuring functionality of the pipe.Table 5.3 lists the reaction loads computed for Waterford 3 for LOCA motions, together with a descriptionof the type of support/restraints employed for the restraints. Seismic loads and thermal loads are also listed. The overall loads computed by summing thermal and the LOCA and seismic loads combined inSRSS fashion is compared to the capacity of the support/restraint to failure.From Table 5.3 it can be concluded that the supports/restraints are capable of withstanding loads resultingfrom a combination of postulated LOCA and seismic events.

WSES-FSAR-UNIT-33.9E-15This analysis conducted for the 1RC14-45RL2 line also provides confirmation of the conservatism of theassessment of the first elbow bending moments by use of the simple formula previously used. For this line the bending moment computed by elasto-plastic detailed analysis is 7.8 x 10 5 in-lbs which can becompared with a calculated 4.1 x 10 6in-lbs determined by the simple formula.The remainder of the lines attached to the RCS have been analyzed using a linear elastic analysisdescribed in Subsection 4.1.7 of this report.The method employed is not a true time history method. Rather it is a generalized response methodutilizing modal analysis. In this method the maxima of the modal amplitudes are summed at each point in the line, regardless of the time at which they occur and of their sign by the SRSS method or absolute method depending on the spacing of the modes. The stresses in the line and reactions at the support restraints are then computed for the three orthogonal excitations by the SRSS method. This method is guaranteed to provide a conservative bound on the value of the stresses in the lines and the loads in the supports/restraints, but is used since it can be run reasonably economically in the computer, and hence permits ready conservative assessment of all of the lines attached to the RCS.As a measure of the conservatism of the method, line 1RC14-45RL2, which has been analyzed by thedetailed non-linear analysis utilizing PLAST, has also been analyzed by the generalized response method.

The results of the two analyses are shown in Table 5.4 for the restraint support reactions and Table 5.5 forthe moments at the first elbow (region of highest stress).Table 5.4 lists two outputs for the generalized response analysis. The first column assumes that thesnubbers in the line are not present to simulate the behavior predicted by the non-linear analysis wherein the gaps in the snubbers were shown not to close. The last column includes the effect of the snubbers,which can be modeled only as linear springs.Except in the immediate vicinity of the first restraint/snubber the two results are not too significantlydifferent. As expected, the linear generalized response method severely overpredicts both the stresses in the line and the reactions at the supports. Table 5.6 shows the results obtained for the restraints/supports of the remaining lines by the linear generalized response and also repeats those results for the 1RC14-45RL2 line. Only three more lines are shown since the other two are virtually identical to two of those shown. Table 5.7 compares the stresses in the lines at the high stress points. Results for the remaining lines are very comparable to those obtained for the 1RC14-45RL2 line. FSAR Figures 3.6A-8 through3.6A-10 and 3.6A-14 through 3.6A-18 show portions of the isometrics for the lines which have beenanalyzed and indicate the position of the supports/restraints. This means that detailed non-linear time history analysis of the other lines can be expected to provide results similar to those obtained for the 1RC14-45RL2 line shown in Table 5.3.The above, coupled with the fact that the vessel motions used are a conservative bound of thoseexpected, confirms that the lines attached to the RCS and their supports/restraints can adequately withstand the postulated LOCA.>

WSES-FSAR-UNIT-33.9E-165.1.10ASSESSMENT OF REACTOR INTERNALSMotions of the core barrel, the reactor vessel, as well as the relative motion between core barrel andvessel and hydraulic loads, are the important parameters in assessing the adequacy of the reactor internals. The relative motion between core barrel and vessel represents the new element in the assessment since it was not not considered in the analyses reported in FSAR Subsection 3.9.2 which had utilized very conservative hydrodynamic asymmetric loads, but had held the reactor vessel fixed.

Previously the asymmetric loads were not part of the design loads.A comparison of the vessel motions (Figures 5.2 and 5.3) resulting from the most significant of thepostulated breaks: namely, the break at the vessel inlet nozzle, obtained for Waterford 3 to those computed for similar plants (i.e., the generic CE plant (Calvert Cliffs, see Reference 1) and the St. Lucie 1plant (Reference 7) shows that the Waterford 3 vessel deflections are significantly smaller. This is due in large part to the fact that the break area in Waterford 3 is significantly smaller (i.e., limited by design) than the corresponding areas in the other two plants. Another significant difference besides the smallermagnitude of the vessel deflections in Waterford 3 is that the period of oscillation is longer.

Corresponding accelerations are thus much lower in Waterford 3 than the generic plant or St. Lucie 1.The vessel deflections of Waterford 3 are further compared with those calculated for a plant subjected to a cold leg nozzle inlet break area of 188 in

2. The comparison shows similar displacement time histories inmagnitude and frequency for the Waterford 3 and the plant with 188 in 2 break area. The largerdisplacement in the plant with the smaller break area resulted from the larger gaps between the reactor vessel nozzle pads and the supports. The break area at the reactor vessel outlet nozzle for Waterford 3 (100 in 2) is slightly smaller than the corresponding break for the generic CE plant (135 in 2). It can beexpected therefore that at worst the motions of the reactor vessel resulting from that hot leg break will be similar in both amplitude and periodicity for both plants but that the Waterford 3 amplitudes will besomewhat lower. For the generic CE plant (Reference 1) vessel motions resulting from the hot leg break at the reactor vessel have been computed to be less than 1/3 of those resulting from the cold leg break at the vessel. Since corresponding cold leg break vessel motions for Waterford 3 are essentially 1/2 of thegeneric plant motions, the cold leg break inside the cavity in Waterford 3 will produce vessel motions which are approximately 30 percent larger than the motions produced by the hot leg break. This isconfirmed by the arguments presented in Subsection 5.1.8 wherein a 30 percent reduction of the hot leginduced motions vs. the cold leg break induced motion is justified by comparing the computed peak loads on the supports.As indicated above, the vessel motions for Waterford 3 are at most one half of the corresponding motionsof the generic plant (see Figures 5.2 and 5.3) and have about the same frequency content initially but subsequently they exhibit a longer period. The internal asymmetric hydraulic forces across the barrel, are essentially unaffected by the break size WSES-FSAR-UNIT-33.9E-17and can thus be considered nearly the same in both plants. Sensitivity studies have shown that internalasymmetric hydraulic loads only decrease about one percent when the break is reduced from full area opening in 30 msec. to an area of 188 in 2 opening in 8 msec. Table 5.8 summarizes these results. Therelative motion between the vessel and the core barrel for Waterford 3 will be slightly smaller than that for the generic plant on the basis of slightly smaller internal hydraulic loads and the absolute values of core barrel and in-phase vessel motions for Waterford 3 being about half that for the generic plant. The in-phase nature of the barrel and vessel motions is confirmed by Figure 5.8 applicable to the plant with a break area of 188 in 2.It can therefore be expected that reaction forces on the Waterford 3 internals will be less than thecorresponding forces for a cold leg break in the CE generic plant; whereas, they will be nearly the same for hot leg breaks since the break areas in both plant are approximately identical (100 square inchesinside the cavity for Waterford 3 and 135 square inches for the generic plant). The hot leg break outside the cavity is larger (600 in 2), however, it gives rise to no cavity pressure. Hence, a significant load, whichwould contribute to the motion of the vessel if the break had been inside the cavity, is absent.Since the internals had been analyzed for a full hot leg break with a fixed vessel, and since the vesselmotions in the absence of cavity pressure are expected to be 10 percent less than corresponding motions when cavity pressures are present, the determining hot leg break would be the hot leg break within the cavity. In the generic plant analysis it has been found that the cold leg break in the cavity results in more severe loading on the internals.Table 5.9 compares the internals for Waterford 3 with those of the generic plant. Table 5.10 providesadditional comparisons. As can be seen, the Waterford 3 internals, with one exception unimportant for this evaluation, are capable of withstanding higher loads than the generic plant design. The additional margin has been computed by considering the relative capacity of each element to accommodate bending(both meridional and circumferential) hoop stresses, local and global shear, and buckling; thus assuming the minimum capacity irrespective of the loading condition.A comparison of the Waterford 3 internals with the generic plant internals indicates that generally whererelatively low margins to design values had been computed for the generic plant, the Waterford 3 internalshave at least 11 percent capacity, with much larger margins in most of the internals components. These added margins, coupled with the fact that loads are expected to be lower as a result of the limited break areas, assure the acceptability of Waterford 3 internals under the combined LOCA and seismic loads.The fuel alignment plate is the only component whose capacity in Waterford 3 is less than that for thegeneric plant. Stresses in that plate are not critical for the generic plant, it is expected that similar conclusions can also be drawn for Waterford 3.

WSES-FSAR-UNIT-33.9E-185.1.11ASSESSMENT OF FUEL ASSEMBLIESAs indicated in Table 1.1, at the draft SER review meeting with the MEB Staff/Consultants on October 1,1980, and at the meetings on January 22, 1981 in Bethesda with the MEB and CSB Staff, the adequacy of Waterford 3 fuel assemblies will be provided in LP&L's response to FSAR Question 231.2, as Question 231.2 specifically addresses the ability of fuel assemblies to withstand combined seismic and LOCA mechanical loads. However, as indicated at the October 1, 1980 meeting, and subsequently in the NRC minutes of the January 22, 1981 meetings, the Waterford 3 fuel assembly evaluation for LOCA will be limited to the cold leg vessel inlet break, since all prior analyses of fuel assemblies (generic plant, St.

Lucie 1) have indicated that the cold leg break at the vessel inlet nozzle produces maximum LOCA loads on the assemblies. Attachment 2 provides the basis for arriving at the above conclusion.

6.0CONCLUSION

SBased on the results of the evaluation contained in this report, it is concluded that all the reactor systemcomponents/structures cited in Question 110.1 (except fuel assemblies which will be addressed under FSAR Question 231.2), can withstand the combination of loads due to a postulated, but highly unlikely,design basis LOCA plus seismic event and that Waterford 3 can operate without undue risk to the healthand safety of the public. An evaluation using more sophisticated techniques than contained in this report would only quantify the safety margins that already exist based on the evaluation contained herein.

Hence, from a value-impact standpoint it is believed that further analysis using more sophisticated techniques is not warranted.At the conclusion of the draft SER Review Meeting and a subsequent meeting in Bethesda, the MEB Staffprovided oral acceptance of our proposal to assess the adequacy of fuel assemblies subjected to LOCA loads resulting from the controlling break; i.e., a postulated break at the cold leg vessel inlet nozzle. The basis for selecting this break location is contained in the report. The Waterford 3 fuel assembly assessment will also be made on this basis.At the same time the analyses of the internal hydraulics, the vessel motions, etc. required to provide thebasic information needed to establish the fuel alignment plate and core support plant motions for this"worst" break, will provide information regarding the asymmetric loads across the internals, the actual vessel motion, etc. Thus the fuel analysis will verify the validity of the technical bases used in this report.

WSES-FSAR-UNIT-33.9E-1

97.0REFERENCES

1) Calvert Cliffs Nuclear Power Plant Units No. 1 and 2, Docket Nos. 50-317 and 50-318, Reactor Coolant System Asymmetric Loads Evaluation Program - Final Report.2) Gerber, T.L., "Plastic Deformation of Piping due to Pipe Whip Loading," ASME Paper 74-NE-1, 1974.3) Combustion Engineering Topical Report, CENPD-168, "Design Basis Pipe Breaks," July 1975.4)CENPD-42, "Topical Report on "Dynamic Analysis of Reactor Vessel Internals", December 1973.5)CESSAR, Combustion Engineering Standard Safety Analysis Report, Combustion Engineering, Incorporated, NRC Docket No. STN-50-470.6) Rettig, W.H. et al, "RELAP A Computer Program for Reactor Blowdown Analysis," Idaho Nuclear Corporation, IN-1321, June 1970.7) Florida Power & Light, Docket No. 50-335, "Reactor Coolant System Asymmetric LOCALoad Evaluation," July 1980.

WSES-FSAR-UNIT-3TABLE 1.1 (1 OF 2)ASSESSMENT OF STRUCTURES/COMPONENTS OF QUESTION 110.1 Component/StructureAssessmentStatus Evaluation Basis References CommentsReactor Pressure VesselCompletePlant Specific AnalysesFSAR App 5.4ASteam GeneratorsComplete DesignCE Design Reports Reactor Coolant PumpsComplete DesignCE Design Reports Reactor Vessel SupportsCompletePlant Specific AnalysesFSAR App 5.4A Steam Generator SupportsComplete DesignFSAR Sect 3.6.2.3and App. 3.6AReactor Coolant Pump SupportsComplete DesignFSAR Sect 3.6.2.3and App. 3.6ABiological Shield WallComplete DesignFSAR Sect 3.8.3.3.1, 3.8.3.3.2 and 6.2.1.2Steam Generator, R C Pump Compartment WallComplete DesignFSAR Sect 3.8.3.3.1, 3.8.3.3.2 and 6.2.1.2RCS Main PipingCompletePlant Specific Analyses andFSAR Sect 3.6.2.3Waterford had been analyzed for all postulatedand Reference to other PlantApp. 3.6Abreaks w/o consideration of asymmetric Analyses(Reference 1)loads and for a 350 in 2 cold leg and 100 in 2 hot leg guillotine break at the reactorvessel nozzle for asymmetric loads also (see text).RCS Attached PipingCompleteDetailed non-linear analysisReference 1 andGeneric Plant Analyses were for both cold(ECCS, etc.)of one line and detailedReference 7and hot leg guillotine breakslinear analysis of remaining(See Text) lines WSES-FSAR-UNIT-3TABLE 1.1 (2 OF 2)ASSESSMENT OF STRUCTURES/COMPONENTS OF QUESTION 110.1 Component/StructureAssessmentStatus Evaluation Basis References CommentsRCS Attached Piping SupportsCompleteDetailed non-linear analysis(See Text)and Restraintsof one line and detailedlinear analysis of remaining linesCEDMsCompleteComparison to previouslyReference 1Generic Plant Analyses were for bothanalyzed CEDM'sand Reference 7cold and hot leg guillotine breaks(See Text)Reactor InternalsCompleteComparison to analyses ofReference 1 (See Text)Generic Plant Analyses were forsimilar plantsboth cold and hot leg guillotinebreaksFuelWill be(See Text)Prior analyses done for similar CEaddressed inplants (with 14 x 14 fuel however) response tohave indicated that the cold leg Q 231.2vessel inlet break is the determiningbreak for fuel assessment. Comparative analyses indicate that there should be no problem with coolability of the core.

WSES-FSAR-UNIT-3TABLE 3.1DESIGN BASIS BREAK OPENING TIMES, SIZES AND LOCATIONSFOR RCS POSTULATED BREAKSMinimum Opening TimeReactor Cavity (from CENPD-168, Reference 3) 100 in 2 hot leg guillotine break20 msec. 350 in 2 discharge leg guillotine 6 msec.Steam Generator Compartment 600 in 2 hot leg guillotine14 msec. 430 in 2 suction leg guillotine11 msec. 592 in 2 suction leg guillotine17 msec. 480 in 2 discharge leg guillotine28 msec.

WSES-FSAR-UNIT-3TABLE 5.1COMPARISON OF SEISMIC MOMENTS (IN-KIPS)SSE FOR ST. LUCIE 1 AND WATERFORD 3 ATPOINTS IN THE RCS MAIN INTACT LOOPST. LUCIE 1WATERFORD 3RCP Discharge Nozzle 5910 6899RCP Suction Nozzle 7256 6932RV Inlet Nozzle 5272 7038RV Outlet Nozzle 2535 8400 WSES-FSAR-UNIT-3TABLE 5.2MODAL FREQUENCIES (HZ) OF WATERFORD 3 CEDMsMode No.ShortestLongest 1 3.4 2,7 2 5.1 4.9 3 11.5 10.9 4 13.5 11.8 5 13.8 13.6 6 23.8 15.0 7 43.3 37.1 WSES-FSAR-UNIT-3TABLE 5.3WATERFORD 3 ECCS LINE REACTION FORCES ON SUPPORTS/RESTRAINTS (KIPS)ThermalSupport/Restraint LOCA SSE(Normal)Total*Capacity(Refer to Figure 6 for location of Supports/Restraints)RCSR-92 019.2 019.291.0(Snubber)RCRR-94 11.0/-4.717.8 .92921.826.0(Strut)RCRR-292 2.1/-1.516.4 9.38225.929.0(Strut)RCRR-100 1.3/-1.123.1 6.00829.135.3(Strut)RCRR-183 2.0/-1.9 5.0 .212 5.627.0(Strut)RCRR-184 031.5 031.591.0(Snubber)RCSR-186 027.3 027.392.0(Snubber)RCRR-293 1.1/-0.912.2 3.29215.527.0*Total = Thermal + (LOCA 2 + SSE 2)1/2 WSES-FSAR-UNIT-3TABLE 5.4COMPARISON BETWEEN RESULTSOR ANALYSIS FOR 1RC14-45RL2 OBTAINED BY NONLINEAR TIME HISTORYAND LINEAR ELASTIC GENERALIZED RESPONSE ANALYSISPROGRAMPIPESTRESS 201050 MODES 7 LEFTOUT FORCE*PLAST 2267 (MAX.)(GENERALIZED RESPONSE METHOD) NODE POINT NONLINEAR ELEASTIC SNUBBERS DELETED SNUBBERS ENGAGED Fx (lb) Fy (lb) Fz(lb) fX (lb) fY (lb) Fz (lb) Fx (lb) Fy(lb) Fz (lb) 71101328134 35015 8002110.311191 10528 104293.6 784616464 36080 1001252.4 10135 131958.417933 39118 16051005.8 6025 6197 20 226 2367 2641 235 476.8 2028 5504 2905 415.4 1884 2638 321 279 2248 2807 2909 272.3 3195 4667 2812 264 1052 1845 2817 264.5 429 522 2917 315.6 1518

2794 2818 170.9 1081 317 2820 35.2 59.8 109.1 54 328 150 67 70 112 34/3541292.3 204.9 1150.5 7660 2374 6411 3831 3067 4393 8 (v.s.) 122.3 481

458 6 (snub) 39072 130 (snub) 16 (snub) 2217 4759 2350 (snub) 2421 1398 2805 (snub) 1031 1786 2819 (snub)

1211 2009 (snub) 3901NOTE:*Modes up to 96 Hz and Leftout Force to account for other modes.

WSES-FSAR-UNIT-3TABLE 5.5MAXIMUM BENDING MOMENTS IN 1RC14-45RL2 COMPUTER BY NONLINEARELASTIC ANALYSIS AND GENERALIZED RESPONSE ELASTIC ANALYSISGeneralized Response Line NodeForce (lb) Moment (Ft-lb) Stress (psi) (Calc #) Point Name Fx Fy Fz Mx My Mz Moment Pressure Total 1RC14-45RL2 377115645076979167915240981029422759 1132634085 1Elbow 477115645076979189515189651542525589 1132636916K = 1 x 108 lb/inAt Point #1 559904872131999 69643 664715425 9594 1132620920(Snub. Deleted) 1Elbow500059904872131999 142803 123672375219457 1132630783

_________________________________________________________________________________________________________________________

____________________________________________________Non-Linear Analysis Line Node Moment (Ft-lb) Stress (psi) (Calc #) Point Name Mx My Mz 1RC14-45RL2 365067 703-1686 1Elbow 4 -64383 -189 1301 5 -47917 -34334-3594 1Elbow 5000 -50125 3858 3347 WSES-FSAR-UNIT-3TABLE 5.6 (1 of 4)MAXIMUM REACTION FORCES AT SUPPORTS/RESTRAINTSOF RCS ATTACHED PIPING 1RC14-45RL2POINT------------FORCES IN POUNDS------------

NO. Fx Fy Fz RESTRAINT 7 0. 0.28931.VAR SUPPORT 8 0. 486. 0.RESTRAINT 80010170. 0. 0.RESTRAINT 10 0.17596. 0.

RESTRAINT 10014911. 0. 0.RESTRAINT 13 0.19062. 0.RESTRAINT 1605 0. 4981 0.RESTRAINT 20 0. 2046 0.RESTRAINT 235 0. 5554 0.

RESTRAINT 2905 0. 4537. 0.RESTRAINT 321 0. 2566 0.RESTRAINT 2909 0. 4537 0.

RESTRAINT 2812 0. 0. 954.RESTRAINT 2817 0. 0. 461.RESTRAINT 2917 1531. 0. 0.RESTRAINT 2818 0. 167. 0.RESTRAINT 2820 34. 0. 0.

RESTRAINT 2820 0. 44. 0.RESTRAINT 2820 0. 0. 93.ROT.RESTRAINT 2820ROT.RESTRAINT 2820ROT.RESTRAINT 2820 SNUBBER 610695. 0. 0.SNUBBER 130 0. 0.22192.SNUBBER 16 5175. 0.11107.

SNUBBER 2350 4944. 0. 2856.SNUBBER 2805 2143. 0. 3710.SNUBBER 2809 1204. 0. 0.SNUBBER 2009 0. 0. 1941.

WSES-FSAR-UNIT-3TABLE 5.6 (2 of 4)MAXIMUM REACTION FORCES AT SUPPORTS/RESTRAINTSOF RCS ATTACHED PIPING 1RC12-40RL2BPOINT------------FORCES IN POUNDS------------

NO. Fx Fy Fz ANCHOR 4095 990. 2516. 880.ANCHOR 430 442. 605. 330.RESTRAINT 2601 1604. 0. 2378.RESTRAINT 2600 0. 9599. 0.

RESTRAINT 320 0. 7978. 0.RESTRAINT 3800 0. 8727. 234.RESTRAINT 3801 8526. 0. 0.

SNUBBER 4000 0.17464. 0.RESTRAINT 4001 4482. 0. 0.RESTRAINT 4101 0. 0.37917.VAR SUPPORT 4300 0. 179. 0.RESTRAINT 610 0. 8454. 0.

RESTRAINT 6301 6157. 0. 4153.RESTRAINT 6700 9571. 0. 2875.RESTRAINT 7000 786. 0. 4169.

RESTRAINT 5900 1988. 0. 0.RESTRAINT 770 0.12825. 0.RESTRAINT 3841 0. 0. 2973.RESTRAINT 3945 1094. 0. 3180.RESTRAINT 3961 0. 2074. 0.

RESTRAINT 397 858. 0. 2141.RESTRAINT 4010 628. 0. 293.RESTRAINT 4020 0. 2200. 0.RESTRAINT 4045 827. 0. 331.RESTRAINT 4046 0. 2026. 0.

VAR SUPPORT 414 0. 6. 0.RESTRAINT 415 120. 0. 300.SNUBBER 4150 0. 629. 0.

RESTRAINT 4165 0. 677. 0.RESTRAINT 4180 840. 0. 336.RESTRAINT 4221 0. 620. 0.RESTRAINT 4220 392. 0. 309.RESTRAINT 4240 0. 319. 0.

WSES-FSAR-UNIT-3TABLE 5.6 (3 of 4)MAXIMUM REACTION FORCES AT SUPPORTS/RESTRAINTSOF RCS ATTACHED PIPING 1RC12-38RL1BPOINT------------FORCES IN POUNDS------------

NO. Fx Fy Fz ANCHOR 5143 1383. 2230. 1213.ANCHOR 5431 6551. 3765. 6466.VAR SUPPORT 192 0. 22. 0.RESTRAINT 19415241. 0. 0.

SNUBBER 195 0. 8843. 0.RESTRAINT 197 0. 0.14542.RESTRAINT 1995 0. 9114. 0.RESTRAINT 200 8894. 0. 0.RESTRAINT 2015 0. 9757. 0.

RESTRAINT 206 0.16443. 0.RESTRAINT 5025 2424. 0. 0.RESTRAINT 5075. 0. 2619. 0.

RESTRAINT 5075 1069. 0. 2486.RESTRAINT 5114 0. 2407. 0.RESTRAINT 511 886. 0. 2067.SNUBBER 51622572. 0. 0.RESTRAINT 532 0. 0.34011.

SNUBBER 517112549. 0. 0.RESTRAINT 517 0.10675. 0.RESTRAINT 5188 0.21046. 0.SNUBBER 519 0. 0.43548.RESTRAINT 522 0.21847. 0.

RESTRAINT 5361 0.27771. 0.

WSES-FSAR-UNIT-3TABLE 5.6 (4 of 4)MAXIMUM REACTION FORCES AT SUPPORTS/RESTRAINTSOF RCS ATTACHED PIPING 1RC12-39RL2APOINT------------FORCES IN POUNDS------------

NO. Fx Fy Fz ANCHOR 7405 2609. 112. 2946.ANCHOR 905. 334. 341. 137.RESTRAINT 903 4095. 0. 1740.SNUBBER 1306 0. 0. 6084.

SNUBBER 1501 0.10073. 0.RESTRAINT 1903 9466. 0. 0.VAR SUPPORT 1901 0. 109. 0.

RESTRAINT 2103 0.13407. 0.SNUBBER 2101 0. 0.18462.RESTRAINT 2406 0.19308. 0.RESTRAINT 260123145. 0. 0.RESTRAINT 5101. 0. 6846. 0.

RESTRAINT 5401 3215. 0. 402.VAR SUPPORT 5903 0. 10. 0.RESTRAINT 5901 0. 0. 5826.

SNUBBER 63 0. 2498. 0.SNUBBER 6403 4091. 0. 0.VAR SUPPORT 6401 0. 20. 0.RESTRAINT 6705 0. 1386. 0.RESTRAINT 7201 0. 1072. 0.

SNUBBER 7496 0. 0. 1381.RESTRAINT 78 0. 1131. 0.RESTRAINT 8005 810. 0. 0.RESTRAINT 84 175. 0. 371.RESTRAINT 8701 0. 284. 0.NOTE:1RC14-44RL1 is similar to 1RC14-45RL21RC12-40RC1A is similar to 1RC12-40RL2B WSES-FSAR-UNIT-3TABLE 5.7MAXIMUM STRESSES IN RCS ATTACHED LINES RESULTINGFROM LOCA MOTION AND INTERNAL PRESSURE Line(Calc #)NodePointNamePressureStress (psi)TotalStress (psi) Line(Calc #)NodePoint NamePressureStress (psi) Total Stress (psi)1RC12-40RL2B 46 11419 360241RC12-39RL2A 29 11419 34878 1Elbow 1Elbow 47 11419 41784 30 11419 26756 44 11419 31459 1Elbow 3 22759 34085 45 11419 266761RC14-45RL2 1Elbow 48Tangent 6885 30713 4 25589 369161RC12-38RL1B 5301Tangent 6885 46418 5 9594 20920(Snub.Deleted) 1Elbow 5301 11419 427185000 19457 30783 1Elbow 5302 11419 53527 3 19925 31251 529 11419 266841RL14-45RL2 1Elbow 1Elbow 4 33448 44775 530 11419 27249 5 13236 24563 527 11419 26797(Snub.Engaged) 1Elbow 1Elbow1000 21921 33247 528 11419 26249 WSES-FSAR-UNIT-3TABLE 5.8A COMPARISON ON THE RESULTS OF THE COLD LEG BUILLOTINEWITH DIFFERENT BREAK OPENING TIMES AND ASSUMPTIONSBreak Opening TimeBreak AreaMaximum Pressure Differences (milliseconds) (in

2) Across the CSB(psi)___________________________________________________________________________________ 1 5761120 18 576 372 8 188 37430 1414 378 WSES-FSAR-UNIT-3TABLE 5.9MECHANICAL SYSTEMS AND COMPONENTS - COMPARISON OF STRUCTURAL DESIGN PARAMETERS Parameters Generic Plant Waterford 3Added Capacity of LP&L Internals (Percent)Structural Rmean, in. 75 1/4 75 1/4Hoop 20Bending+ 40Upper CSB t, in. 2 1/2 3Shear* 17 L, in. 135 5/8 138 7/16Buckling 72

Rmean, in. 74 7/8 75 1/4Hoop 42

.Bending+ 66Middle CSB t, in. 1 3/4 2 1/2Shear* 21 L, in. 144 3/4 160 1/8Buckling 133 Rmean, in. 74 5/8 75 1/4Hoop 32Bending+ 75Lower CSB t, in. 2 1/4 3Shear

  • 51 L, in. 38 33 5/8Buckling 133Lower cylinder ID, in. 141 141Hoop No increaseCore cylinder OD, in. 145 145Binding+ (No increase meridion) 29 AxiallyShear* 11Support cylinder L, in. 42 37 5/8Buckling No increaseStructure supported CSB CSBCircumferential Bending 30 Lower Lower Flange FlangeFuel Alignment Plate No change (not important)(CSB OAL, in. 328 1/2 347Core shroud supportCore Support Plate No increase UGS Hoop 27 Rmean, in. 72 5/8 72 1/2Bending+ 56Cylinder t, in. 2 2 1/2Shear* 24 L, in. 24 24Buckling 98Beams in. 24 x 1 1/2 24 x 1 1/2 No ChangePlate on Top of Beams t, in. 4 3 1/2Shear -12.5 (This plate is not important ShroudsThe two are the same butBending -3.0 since beams are weaker)the support distance is +30shorter for Waterford 3Legend:CSB - Core Support BarrellUGS - Upper Guide Structure

+ - Minimum of axial or meridional bending

  • - Minimum of overall or local shear WSES-FSAR-UNIT-3TABLE 5.10COMPARISON OF OPERATING CONDITIONS AND DIMENSIONS FOR THE WATERFORD 3*AND GENERIC PLANT AFFECTING INTERNAL ASYMMETRIC LOAD ANALYSISOperating ConditionWaterford 3Generic PlantSt. Lucie 1Normal Operating Presssure (psi) 2250 2250 2250Normal Operating Temperature H.L. (F) 611 598 592Normal Operating Temperature C.L. (F) 553 548 543Total Flowrate (lb/hr)148 x 10 6130 x 10 6130 x 10 6DimensionsHot Leg I.D. (in.) 42 42 42Cold Leg I.D. (in.) 30 30 30Reactor Vessel I.D. (in.) 172 172 172RV Nozzle Support Interface Gap (in.) .035 .125 .215* Other dimensions as given in Table 5.9 WSES-FSAR-UNIT-33.9E-40ATTACHMENT 1 (Appendix 3.9E)SAFETY INJECTION LINE STRESSES DUE TO LOCA + SEISMIC LOADING WSES-FSAR-UNIT-33.9E-41WATERFORD 3SAFETY INJECTION LINE STRESSES DUE TO MOTION OF REACTORThe stresses developed in the Waterford SES 3 Safety Injection Lines are estimated by comparison with aprevious analysis. Table 1-1 provides such a comparison.

WSES-FSAR-UNIT-33.9E-42TABLE 1-1STRESSES IN ECCS LINESPlantSt. Lucie 1 Waterford 3 Line1-B-11RC12-40RL2B1RC14-45RL2 O.D.12.750 inches14.000 inches I.D.10.126 inches10.270 inches11.636 inches Pressure2235 psi2235 psi2235 psi Pressure10,860 psi11,490 psi13,240 psi\

Maximum Dynamic2/3 inch1/3 inch1/3 inchDisplacementLength of First 10.73 ft.7.00 ft.7.07 ft.Appropriate SectionBending (LOCA only)19,280 psi (1)26,830 (2)28,890 psi(1)22,800 psi (2)Equation 9 (3)45,000 psi52,530 psi72,350 psiSince the maximum stresses calculated are at or above the 3 Sm limit of 48,000 psi, it is necessary to demonstrate theintegrity/functionality of these lines. To do this, the maximum benidng moment which can be sustained by the pipe is computedand compared to bending moment actually developed in the lines. (The maximum sustainable bending moment is taken to be 70 percent of Gerber's (Reference 2) result; the actual bending moment is computed by the simplified mehtod as explained inSubsection 5.1.9 of Appendix 3.9E). Table 1-2 summarizes such a comparison.

(1) from PLAST detailed analysis (2) Calculated from Simplified Method for Computation of Pipe Stress due to Specified Displacements M = 3EI/L 2 (3) = B 1 x Press + B 2 x Bend; B 1 = 1.000, B 2 = 1.817 (St. Lucie 1), B 2 = 1.9026 (Waterford 3 - 12 inch pipe)

B 2 = 2.04609 (Waterford 3 - 14 inch pipe)

WSES-FSAR-UNIT-33.9E-43TABLE 1-2MAXIMUM LOAD CARRYING CAPABILITY OF FIRST ELBOW IN ECCSPIPE VS COMPUTED MOMENTSPlantSt. Lucie 1 Waterford 3 Line1-B-11RC12-40RL2B1RC14-45RL2

t/R o0.206 0.1940.169*0.3090.2920.253*/R o0.0484 0.0458 0.0362 3 n+3)2 (40 N n I o93.3 x 10 3 (psi)4 ()RR o n i n 33++165.9158.8189.9(*/)R o n0.6950.6910.671 M max+5.94 x 10 6 in-lb5.34 x 10 6 in-lb5.81 x 10 6 in-lb70% Mmax4.15 x 10 6 in-lb3.74 x 10 6 in-lb4.07 x 10 6 in-lb MBending2.8 x 10 6 in-lb3.2 x 10 6 in-lb4.1 x 10 6 in-lb MSeismic6.,5 x 10 5in-lb3.98 x 10 5 in-lb2.76 x 10 5in-lb+Computed by the method of Gerber (Reference 2) for straight pipes, and scaled down by the B 2 factor in equation 3

.

WSES-FSAR-UNIT-33.9.E-44ATTACHMENT 2 (APPENDIX 3.9E)JUSTIFICATION OF COLD LEG INLET BREAK AS THE DETERMININGBREAK FOR FUEL WSES-FSAR-UNIT-33.9.E-45JUSTIFICATION OF COLD LEG INLET BREAKAS THE DETERMINING BREAK FOR FUELAnalyses performed for the generic CE plant (Reference 1), as well as specific plants like Fort Calhoun,have indicated that the determining break insofar as fuel analysis is concerned is a cold leg guillotine break at the reactor vessel inlet nozzle. This conclusion is based on the fact that the response of the fuel for these plants has been analyzed for both a full area guillotine break at the reactor vessel inlet nozzle and a limited area (135 in

2) guillotine break at the reactor vessel outlet nozzle, and that the former breakhas been found to be limiting break.Another result of these analysis has been that the beam-column effect, due to concurrent lateral and axialloading is also more pronounced for inlet break, but that this effect does not significantly increase the maximum fuel bundle stresses. (A dynamic beam-column analysis was performed for the plants of Reference 1 to determine any aditional bending stresses and stability of the fuel assembly due to concurrent lateral and axial loading, and that analysis had also shown that the beam-column effects are more sensitive to lateral bending moments than to axial forces.)The lateral bending moments on the fuel assembly due to the input excitation of the core support plate,fuel alignment plate, and core shroud displacement time histories have been found to be significantly larger for the reactor vessel inlet breaks than for the reactor vessel outlet break.These analyses have been performed for a 14 x 14 fuel and therefore detailed stress results have limitedapplicability to Waterford 3. However, some of the conclusions reached for these plants have a direct bearng on Waterford 3 fuel assemblies also:a)The determining load condition on fuel is the lateral bending moments due to the excitation of thecore support plate, fuel alignment plate and core shroud lateral displacement time histories.

These were largest in the generic plant (Reference 1) for the inlet break; however, the inlet break in the plant is a full area break whereas in Waterford it is a limited area break (350 in 2). In boththe generic plant and in Waterford, the outlet break area is limited. For the generic plant this break is 135 in 2 while in Waterford 3 that break area is 100 in 2.Figures 2-1, 2-2 and 2-3 compare the lateral motions of the generic plant vessel for cold and hotleg breaks. Significant displacements only occur in the x-direction (parallel to the hot legdirection) for hot leg break, so only that direction is shown.The motions of the vessel, together with the internal asymmetric hydraulic load determine thelateral motions of the core support plate, fuel alignment plate, andcore shroud. The vesselmotions of the inlet break are clearly much larger than those for a hot leg break for the generic CE

plant.

WSES-FSAR-UNIT-33.9.E-46For Waterford 3, the vessel motions (lateral) are essentially half of those of the generic plant (seeFigures 2-4 and 2-5 for an inlet break 350 in 2). The motions of the Waterford vessel resultingfrom a hot leg outlet break of 100 in 2 would be very similar to those of the generic plant (135 in 2).However, the amplitudes of these motions would still be approximately 30 to 40 percent less than the corresponding motions for the 350 in 2 cold leg inlet guillotine break.In addition, the hot leg break would produce essentially no asymmetric internal hydraulic loadsacross the core barrel. Hence, it can be concluded that the inlet break is the determining breakfor establishing the largest lateral bending moments on the fuel bundle for Waterford also.b)Axial loads can be higher for the hot leg breaks, however, the generic plant analysis has indicatedthat they are not as significant as lateral bending loads in determining fuel stability and bending stresses. This, coupled with the fact that the beam-column analyses performed for the generic plant indicated negligible additional effects on the fuel bundle stresses, over those predicted from the lateral loading analysis, is indicative of the cold leg break being the determining break for the fuel assessment.c)For axial loads, the hot leg breaks would provide the dominant loads. The Waterford 3 plant hastwo postulated hot leg breaks, a 100 in 2 break at the vessel outlet and a 600 in 2 break at the S.G.inlet. (Lateral Vessel motions, resulting from either breaks, would be considerably smaller than those resulting from the 350 in 2 cold leg vessel inlet break. The 600 in 2 break would produceabout 90 percent of the vessel motions produced by the 100 in 2 break.) These two breaks requiredifferent times to open. In any case, Waterford 3 has been analyzed for the pure axial loads resulting from a full area hot leg break (the so-called core bounce analysis), and the fuel was shown to be adequate for tha tbreak which produced axial loads that are larger than either the 100

in 2 or 600 in 2 hot leg breaks.It is therefore concluded that the cold leg vessel inlet break is the determining break fordemonstrating the adequacy of Waterford 3 fuel assemblies.

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