ML17207A876

RCS Asymmetric LOCA Evaluation.

ML17207A876
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Site:	Saint Lucie
Issue date:	03/03/1980
From:	FLORIDA POWER & LIGHT CO.
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REACTOR COOLANT SYSTEM ASYMMETRIC LOCA LOAD EVALUATION ST. LUCIE UNIT 1 DOCKET NO. 50-335

)farch 3, 1980

SUI&fARY, In May 1975 the NRC Staff was informed by a pressurized water reactor licensee that loads resulting from a hypothetical rupture of the reactor coolant cold leg pipe in the immediate vicinity of the reactor pressure vessel (RPV) may have been underestimated.

In November 1975 the Staff agreed that these loads should be considered and evaluated on,a generic basis.

I Florida Power & Light's response to the Staff's letter of November 26, 1975 indicated that the support system design incorporated the reaction forces associated with the large arbitrary reactor coolant pipe ruptures, and that further, it had been shown to acceptably accommodate the additional loads associated with differential pxessures within the reactor cavity as

,shown in Appendix 3H of the Final Safety Analysis Report.

The Staff requested that further internal asymmetric load (IAL) evaluations be conducted. FP&Ls letter of February 9, 1976 documents the Company's commitments to evaluate the reactor vessel support capability for the limit-ing break, a commitment which is restated in Supplement 2 to the Unit 1 Safety Evaluation Report (SER), dated Harch 1, 1976.

In September 1977 FP&L transmitted to the NRC a report assessing the margin in design of the vessel supports when the internal asymmetric loads are added to all previous loads. The report concluded that the supports would adequately withstand all the loadings. However, since the analysis did not account for gaps between the vessel and the core barrel, and also the vessel and the support structures, an analysis was initiated at the same time to account for these effects.

The Staff's letter of February 16, 1978 requested that the evaulations conducted to date be expanded in scope to include an assessment of the reactor pressure vessel, fuel assemblies and internals, control element assemblies, primary coolant piping and attached ECCS piping, all primary system supports, and the biological and secondary shield walls for a spectrum of breaks in the primary system. FP&Ls March 1978 response stated that its August 1977 report was fully responsive to the Staff's SER requirement, that the St. Lucie 1 design was acceptable and that the large instantaneous pipe breaks being postulated weie overly conservative. The response went on to say that FP&L would pursue additional analyses once the Staff approved the analytical methods used in the August 1977 report. This reply notwith-standing, FP&L being sympathetic with the Staff's desire to assess any potential risk to public health and safety from postulated events, expanded the analysis referred to above, to also assess the additional items identified by the Staff.

This report discusses the results of this expanded analysis. The combina-tion of thrust, external, and internal asymmetric loads resulting from the inlet pipe circumferential break present the largest load to the vessel supports among those that would ensue from any of the design basis breaks listed in Appendix 3EI of the PSAR.

The results confirm that the vessel supports will adequately withstand all the loads resulting from the postulated circumferential break in the vessel inlet pipe. The cold leg guillotine break in the cavity is the beak which results in the largest loading of the vessel supports. There-fore the vessel supports are clearly adequate for all other break locations.

This reaffirms the conclusions of the August 1977 report.

Results also show that all supports for the primary system are adequate for all break locations, that the stresses in the intact primary piping arid attached lines are sufficiently low to ensure performance of intended func-tions, and that the biological shield wall performs its intended function.

The secondary shield wall is designed for postulated primary system ruptures within the steam generator subcompartments.

The analyses of the adequacy of fuel assemblies internals, and control element assemblies is in progress. Results are expected in July of 1980.

The Staff, at a meeting in January 1980, further requested that seismic loads be separately identified. All results presented herein, as well as the August 1977 xeport, include the SRSS combination of LOCA and seismic loads, consistent with the requirement of HUREG-0484. Por those combinations design seismic loads have been used and are hereby attached for use by the Staff. In all cases, design seismic loads are considerably higher than calculated peak seismic loads.

Qualitatively, the small displacements observed for the vessel and core barrel for the worst break analyzed, strongly suggests that the analyses now in progress of the fuel/internals and CED>fs will indicate acceptable results.

It must also be noted that since the submittal of the August 1977 report to the Staff, additional work has been reported, to support PP&Ls contention that the types of instantaneous pipe breaks being postulated by the Staff are excessively conservative.

l. 0 INTRODUCTION During a postulated loss of coolant accident 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 over a short period of time. Decompression waves originated.

at the postulated break travel around the inlet plenum and propa-gate downward along the downcomer annulus. The finite time required by the decompression disturbances to travel about the vessel causes a transient pxessure differential field to 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 blowdown fluid flash-ing into the reactor cavity also causes a transient pressurization acting on the vessel. This external pressurization is also asymmetric.

The internal asymmetric loading (IAL) and the external asymmetric loading 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 suppoxt system together with the thrust load resulting from the break, other operating loads, and postulated seismic loads. The seismic loads and normal operating loads, as well as the EAL have been previously analyzed in Appendix 3H of the Final Safety Analysis Report.

Breaks outside 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 is assessed for full breaks at appropriate prima~ system locations. The cavity breaks are determining breaks for the assess- .

ment of the adequacy of piping attached to the primary system piping.

The circumferential pipe rupture at the inlet nozzle of the reactor pressure vessel is determined to be the design basis break for the evaluation of the vessel support adequacy. A break at the outlet nozzle would not produce a horizontal asymmetric pressure loading to the vessel. Consistent with the Final Safety Analysis Report, a 4.0 sq. ft. cold leg guillotine break at the inlet nozzle is chosen for the analyses of the vessel support adequacy.

2. 0 METHOD OP ANALYSIS 2.1 Reactor Vessel Su orts The. adequacy of the reactor vessel supports is evaluated by determining the loads acting on the pximary system which result from a postulated break at the inlet cold leg nozzle; the response of the primary system to the application of these loads; and the reaction forces generated by this=response at the reactor vessel supports. The loads acting on the primary system consist of normal plus seismic loads, the thrust load, external asymmetric loads, and internal asymmetric loads. The latter three are combined in true time history fashion, added to the normal loads reactions, then the xesul-tant reaction loads at the supports are combined with design reaction loads resulting from the postulated seismic (SSE) events by SRSS techniques, to obtain the overall reaction load at each of the supports. Design seismic loads are provided for each primary system support in the three orthogonal direc-.

tions in Table 1. It should be emphasized that computed peak seismic loads are in general substantially less than the design seismic loads; thus providing an element of conservatism in this analysis. Table 2 gives a sample comparison of calculated and design seismic loads at representative locations.

The following subsections describe the methodology employed to evaluate each of the thrust, external asymmetric and internal asymmetric loads. Inherent in the eyaulation of these loads is the detemination of the time. required to open up the break to the area being analyzed.

2.1.1 Break Opening Time and Thrust Loads The St. Lucie plant primary coolant piping'in the vicinity of the vessel is restrained from unlimited motion following complete severance in the portion within the cavity by restraints in the primary shield wall penetrations and wire ropes around the reactor coolant pumps. This restraining system has been previous-ly described in the FSAR, Following an arbitrarily assumed instantaneous severance of the pipe at the nozzle, the two ends of the broken pipe separate under the action of the thrust imposed by the instantaneous tension release followed by the blowdown of the escap-ing fluid, and form a combined break area which varies with time as given in the following equation:

mR.SX(T) 2 .

s'v. mg(R.+R ) t (v) = + 2R j m-( Bn 2g) f

<5 where Ri and R are the inner and outer pipe radii, t 0

is the pipe thz.ckness, x is the axial sepanation of the two ends which varie's with time 7, and g = cos (2) 2R.

i wherein y is the radial separation of the two broken ends which also varies with time.

This equation is solved in iterative fashion together with the equation for the combined tension release and blowdown force, given below (v) 2 V (T)

F(x) P dl (v)A +

p p dl (x)A g (3) to yield the correct forcing function and break area as a function of time. In equation (3), P and pd are the pressure and fluid density in the 3xscharge leg, respectively, A is the cross-sectional area of the pipe, V the bloBdown velocity, and A (7) is defined in (1) above.

The motion of the piping system under the application of the force given by (3), is computed by modelling the leg, the pump, and the cross over leg with 'ischarge an el@to-plastic finite element computer program, PLAST considering the steam generator and the vessel to remain motionless.

Results of the analyses indicate that- at least 18 msec.

are necessary for the pipe ends to separate the overall area of 4. 0 sq. ft. ref er md to in the FSAR. This analysis also indicates that as a result of plastic rotation at the pump, it is possible for the pipe ends to separate further, to a maximum area of 7.78 sq. ft.

The time required for this area to be achieved, however, would be in excess of 25 msec. The longer time required for opening the larger break insures that the IAL result-ing from the two breaks are virtually identical. The larger break does however result in a larger external

horizontal asymmetric load (external vertical asym-metric loads are virtually identical for the two breaks). Since the 4.0 sq. ft. break had been one the design basis breaks in the FSAR, all analyses

'f used that break area. However, consideration is given to whether the system is capable of accommodat-ing the larger break. As discussed in the subsequent section, the system is in fact adequate for the largest of the breaks.

2.1.2 External Asymmetric Pressure Loads (Reactor Cavity)

The reactor subcompartment analysis for St. Lucie Unit 81 had beenperfoimed for stipulated LOCA conditions including a 4.0 sq. ft. cold leg guillotine break, and the results had been submitted to the NRC. in the FSAR and approved by the NRC during the course of the operating license review. The results for the 4.0 sq.

ft. cold leg guillotine break, as reported in Reference 2, have been directly used in'the present study. This results. in conservatism of the analysis since the cavity response had been, predicated on a break opening time of 10 msec, whereas 18 msec. is needed to achieve this size break. The peak external asymmetric forces across the reactor vessel, that would result from the larger 7.78 sq.. ft.. break, would be approximately 40 percent larger. This is predicated on a ratio of 1.39 between peak and average energy flow to the cavity resulting from a 7.78 sq. ft. and a 4.0 sq. ft. cold leg break respectively.

In the original analysis, however, two elements of conservatism had been introduced. First, the mass and energy releases had been increased by 10 percent and second, all insulation had been assumed to reamin in place in the reactor cavity and vent areas for the purposes of volume and vent area calculations in the mathematical model. The insulation in the upper cavity reaches would be crushed against the vessel upon cavity pressurization, resulting in an increased volume of approximately 15-20 percent.

Hence, realistic modeling of the insulation behavior, coupled with removal of the 10 percent conservatism in the mass and energy release would result in a pre-dicted external asymmetric pressure load and cavity pressure load from a 7.78 sq. ft. break which is only 15 to 20 percent higher than those conservatively pre-dicted.

2.1.3 Internal Asymmetric Pressuxe Loads The model used to determine the pressure field at every point in the primary system following the postu-lated primary system breaks, from which the internal asymmetric forces on the vessel and core support barrel are deduced, is shown in Figure l.

3/

The RELAP-4thexmal hydraulic code is used to compute the thermodynamic properties in the model volumes and junctions. Results of the RELAP-4 model have been compared toresults achieved. by modelling the system with WHAM-6for the period of time duxing which the latter can be applied with confidence, which is also the period of time of interest., Figure 2 shows the model employed for WHAM-6. A similar WHAM model and assumptions in its use, had been p~viously submitted to the Staff in the August 1977 report. The results of the two models axe in good agreement, with RELAP-4 predicting a larger pressure differential across the core support barrel.

Results of the internal asymmetric loads analysis indicate that the peak forces across the core support barrel and the vessel are virtually insensitive to the break area, but extremely sensitive to beak opening times. For instance, a change in axea from 1 sq. ft.

requiring 8 msec. to open to approximately 9.81 sq. ft.

(complete double-ended area break) with an opening time.

of 36 msec., only results in a 2 to 3 percent increase in peak internal asymmetxic loads, whereas a decrease in opening time from 36 msec. to 1 msec. for the full break brings about a threefold increase in internal asymmetric load.

2.1.,4 Vessel and Primary System Structural Model A non-linear elastic time history dynamic analysis of three-dimensional mathematical model of the reactor coolant system including details of the reactor internals, pressure vessel, supports, and piping was performed for the postulated pipe break to provide reactor vessel support reaction,forces.

The structural model employed is shown in Figures 3(a) and 3(b). This model is three-dimensional and has 981 total static degrees of fxeedom and 77 mass .degrees of freedom. The reactor vessel and all internal components are mo'delled at internal and support interfaces.

The STRUDL 5/ computer code generates the condensed stiffness matrix used in the dynamic'analysis from the physical definition of the structure. Hydrodynamic effects, including both virtual mass and annular effects are accounted for in the coupling between the RPV and the CSB, and between the CSB and the core shroud. The hydrodyamic (added) mass matrix is evaluated using the ADifASS code..

The dyanmic analysis to determine the systyy response was performed using the computer code DAGS and DFORCE .

The reactor pressure vessel support system is described in the FSAR. The modelling of the steel portion of the support is identical to that described in the FSAR in Appendix 3H. The basic model of the biological shield wall is also identical. However, a more refine( analysis is employed for the latter, utilizing a NASTRAN- nonlinear solution procedure employing quadrilateral and triangular plane stress concrete cracking finite elements, instead of the STARDYNE method of solution described in Appendix 3H of the FSAR.

2.2 Reactor Coolant Pi in Connected Pi n and Other RCS Su orts 2.2.1 Steam Generator Supports Outside the reactor cavity, breaks have been assumed at appropriate locations.

The RCS supports most affected are the lower steam generator supports.

The primary system model is analyzed on an elastic basis for both hot leg and cold leg breaks, the hot leg break at the steam generator inlet being the determining event for the Steam Generator support.

This analysis is a static analysis which employs the compu(p) code 51EC-21 (Hare Island piping flexibility code) .

Both LOCA and des'ign seismic loads are included in the analysis.

2.2.2 ECCS and Other Connected Piping The analysis of the stresses generated in the ECCS lines and other lines attached to the primaxy loop involved a two step process. First, the time histories of the dis-placements are generated at each nozzle attaching said piping to the primary loop. The "worst" time history, irrespective of the location at which is occurs is

applied to the line which by configuration and other loading (normal and seismic) would result in the highest stresses. The stresses induced by LOCA motions for this particular configuration are added-to previously computed normal and seismic (SRSS) stresses. The determining break for ECCS line evalua-tion is the cold leg nozzle break in the cavity.

2.2.3 Reactor Coolant Piping The structural model for the primary system is also utilized to determine the stress conditions in the intact portion of the reactor coolant loop.

3.0 RESULTS OF THE ANALYSES 3.1 Vessel Su orts The loads calculated for each reactor vessel support by the method outlined in Section 2.1.4 are reported in Table 3 for the break chosen for the analysis; i.e., the 4.0 sq. ft. cold leg break at the inlet nozzle; for a range of reactor vessel support stiffnesses. This range covers the possible values of the overall stiffness of the individual'actor vessel supports, the real value being somewhere between the two extremes. It is not possible to quantify the stiffness value more precisely since the modelling of the boundary condition representing embedded steel in the biological shield is subject to variation.

In the support analyses however, the higher loads resulting from the use of the highest stiffness, have been utilized.

This insures again that the absolute maximum load per support is computed. In reality, lower values are expected.

The capability of the reactor vessel supports is given in Figure 4 and Table 4 respectively for the RPV support pad capability and the weakest link in the steel support/biolo-gical shield structure.

Since the capability of the supports exceed the maximum loads computed for the given break, it is concluded that the exist-ing support system is adequate for that break.

As stated in Section 2.1, it is possible that, as a consequence of the broken discharge line rotation about the pump, a larger break area could form within the cavity, up to a maximum of 7.78 sq. ft. This larger break area, requiring a proportion-ately longer time to open, has virtually no effect on thrust and internal asymmetric loads, but would increase the horizontal external asymmetric load by approximately 15-20 percent over that

used in the analysis, as explained in Section 2.1.2. The EAL represents approximately 40 percent of the overall load.

Hence, a 20 percent increase in this load would result in less than a 10 percent increase in the overall loading. From Table 4 and Figure 4, it can be seen that this increase would be accommodated by the margins existing in the support system.

It is therefore concluded that the reactor'essel supports can withstand the largest break in the cold leg piping within the cavity.

Since cold leg breaks outside the cavity do not produce EAL loads and since the'IAL is virtually unaffected by the area of the break as explained in Section 2.1.3, it is also conclu-ded that the reactor vessel supports are capable of withstanding any load resulting from postulated ruptures outside the cavity.

A detailed analysis of the reactor loads resulting from hot leg breaks within the cavity has not been performed. The masons are as follows: the stiffness of the hot leg,pipe combined with the steam generator restraining action, results in a break area within the cavity which is smaller than the cold leg break area, hence resulting EAL would be lower than calculated for the cold leg break; although the thrust force initially would be larger, the IAL would not be colinear with thrust and EAL, but would in fact be approximately orthogonal to them. The resultant horizontal loads on the vessel supports therefore, would clearly be smaller.

For instance, the reactions at reactor vessel supports, due to a hot leg break have been compared to the reactions due to a cold leg break for thrust and subcompartment pressure only.

Horizontal Hot Horizontal Cold Le Break (Ki s) Le Break (Ki s)

Cold Leg Spt 4270 3270 Hot Leg Spt 3275 Although the load on the cold leg support is more severe for a hot. leg break than for a cold leg break, when the effects of internal asymmetric loads-are added, the cold leg break will govern.

Vertical loads would be of the same'order of those experienced as a result of cold leg breaks, and the capacity of the support system to accommodate vertical loads is significantly higher than its horizontal capability. Hence clearly the reactor vessel support system is also capable of withstanding the effects of postulated hot leg breaks inside and outside the reactor cavity.

A similar conclusion had been, reached in our August 1977 report. Differences in maximum loads reported herein from those reported in the August 1977 report are two fold. The August 1977 report did not consider internal gaps or gaps between'he support pads and the support structure. The August 1977 report considered therefore that all loaded supports would be loaded simultaneously and share the load equally. The agreement of the overall loading. between the present and the August 1977 results, confirm that the approach taken in 1977 to assess the loads was not unreasonable.

3.2 Other RCS Su orts The only supports on the primary system, other than the vessel supports, are the steam generator supports. Results of the analyses of the loads imposed on these supports from both hot and cold leg breaks in the system in combination with seismic loads, indicated that none of the design loads have been exceed-ed, with exception of the loads on the four holdown bolts at the vessel end of the steam generator sliding base and the sliding base itself. The computed and design loads are shown in Table 5. Individual examination of the sliding base, the bolts, and bolt anchorages however indicates that all can acceptably withstand the applied loads. It is therefore conclu-ded that the existing supports design is adeq'uate.

3.3 Reactor Coolant Pi in Table 6 reports the elastically calculated pipe rupture- and seismic loads on intact reactor coolant piping associated with the broken loop for the worst break, which is the cold leg .

guillotine break at the vessel safe end. Examination of this table reveals that all loads fall within the allowable loads with the exception of the load at the RCP discharge nozzle, which exceed the allowable by about,l3 percent, on an elastic basis.

Since this analysis is predicated

~ on a 4.0 sq. f t. cold leg break, by the arguments presented in Section 3.1, consideration of the largest break that could occur at the vessel safe end; i.e., 7.78 sq. ft., requires that an increase in load of less than 10 percent be examined to assess the a'dequacy of the coolant piping. Such an increase'can be readily accommodated at the RCP suction and RV outlet nozzles. The RCP discharge would be more overstressed (on an elastic basis) and the RV inlet would be very slightly overstressed.

Since only the fluid retaining integrity of this coolant piping needs to be maintained during the postulated LOCA, an analysis conducted on an elasto-plastic basis would conclude that this

integrity would be maintained at those nozzles. Since the amount of overstressing calculated on an elastic basis is relatively small, a plastic analysis was not considered necessary.

During the performance of this particular'analysis it was calculated that the snubbers on the reactor coolant pumps are overstressed. These snubbe~ are not needed for these events. However their failure could affect the results.

Hence, the analysis was repeated by taking no account of the snubbers. Results are also reported in Table 6 As can be

~

clearly seen, the effect of the presence or absence of the snubbers is negligible.

3.4 ECCS and Connected Pi in The stresses computed from the analysis described in Section 2.2.2 are within 10 percent of the allowable, and hence it is concluded that the ECCS piping and other piping connected to the primary loop, is not adversely affected by the postulated event.

Table 7 compares the peak computed stresses, which include no mal and seismic loads to the allowable stresses.

The margin existing between peak stresses calculated on an elastic basis and stresses that would be allowed within an elasto-plastic analysis further indicates that this attached piping would be able to withstand the imposed loads from the 7.78 sq. ft. larger cold leg guillotine break.

3. 5 Seismic Loads

, Pursuant to the Staff's request at the January 16, 1980 meeting, Table 1 provides the design seismic loads at the various support points in the Reactor Coolant System.

3.6 Control Element Assemblies Although the analysis of the Control Element Drives response to postulated LOCA events is in progress, but will not be comple-ted until July 1980, i't is germane to point out that the CEAs are not needed for breaks 'in the RCS which exceed 0.5 sq. ft.

The assumption of a complete guillotine will result in breaks larger than 0.5 sq. ft.

4.0 CONCLUSION

Even though some analyses have not yet been completed, results obtained to date demonstrate that the existing design has signifi-cant capabi/jtyl)o accommodate the postulated events. Additional informatio~ which has become available since the August 1977 report, and which reinforces our contention, stated in that report, demonstrates that such events are of an acceptably low probability and cannot happen in the manner postulated for this analysis. The foregoing reaffirms our conclusion that the design of St. Lucie l

Unit is acceptable.

5.0 REFERENCES

"PLAST An Elasto-Plastic Computer Program for Stress Analysis of 3-D Piping Systems and Components Subject to Dynamic Forces", submitted to the NRC as ETR-1001 Ebasco Topical Report.

2/ St. Lucie Unit No. 1 FSAR, Docket No. 50-335, Amendment 44.*

3/"RELAP 4 A Computer Program for Transient Thermal Hydraulic Analysis of Nuclear Reactors and Related, Systems", User's Manual, ANCR-NUREG-1335.

4/ Fabic, S., Computex Program WHAM for Calculating Pressure, Velocity and Force Transients in Liquid Filled Piping Network", Kaiser Engineering Report No. 67-49-R, November 1967.

"ICES STRUDL II The Structural Design Language Engineers User's Manual",

MIT Press, Cambridge, Massachusetts, 1968.

6/"ADMASS A Computer Code for Fluid Structure Interaction Using the Finite Element Technique", Ebasco Services, Incorporated, 1979.

"DAGS CENPD 168, Revision 1 Design Basis Pipe Breaks", September 1976.

"DFORCE Design Basis Pipe Breaks", September 1976.

"MSC/NASTRAN User's Manual", McNeal Schwandler Corporation, Los Angeles, California.

10/ "WCAP 9570 Class 3 Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Thru Wall Crack", by Palusamy, S. S., and A. J. Hartmann, October 1979.

Ayers, D. J. and T. J. Griesbach, '-'Opening and E~tension of Circumferential Cracks in a Pipe Subjected to Dynamic Loads", Fifth International Conference of Structural Mechanics in Reactor Technology, Berlin, Germany, 1979.

Griffin, J. H., "MEC-21 A Piping Flexibility Analysis Program", TID-4500 (31st edition), LA-2924, UC-38, July 14, 1964.

TABLE 1 ST. LUCIE 1 NORMAL AND SEISMIC SUPPORT LOADS (X106LB.)

IIOAMAL OPEIIATING ODE SEISMIC DBE SEISSIIC I

THEGMAL + AV1 CONDITION DEAD DEAD LOAD VIEIGHT V(EIGHT SY SY tV H1 0 .028 .005 .002 .G44 .011 .005 1.288 f N V$

V1 .GGG 1.155 .032 .335 .04G .064 .670 .092 NV1 4195 i.360 HZ .091 1.226 .001 +.355 2.452 .003 +.710

.264- -.528 ivt V2 .G64 .72G .017 .253 .380 035 .507 .761 I $ $ ACIOII V$ $ $ ($

$ Vt(OIIIA(ACIIO(N NV2 a.195 k.215

.079 1.139 .019 .270 2.278 .038 .540 $ ($

(

.060 .349 ~120 .6SS N V3 .741 .367 .G23 .743 .006 .50G .746 NV3 x.195  %.215 $ ($

YI Z11 0 0 0 0 S.195 0 0 S.390 .v Z12 0 0 0 0 4197 0 0 S.394 Y$ h Y1 .300 0 .016 .019 0 .033 .039 0 Y2 .300 .315 .057 .155 .060 .1'14 211 .121 Y3 1.300 1 1.009 .086 .417 .071 .173 .835 .143 Y4 .300 .320 .054 .152 058 .108 .305 .116 YI Vf x 0 0 0 0 0 0 0 0

$ (fAN C(N(AAION NY a.376 S.300 LOW(N$V(YOA1$

Table 2 Comparison of Peak Calculated and Design Seismic (DBE) Loads at Representative Locations Design (Kips) Calculated (Kips)

Horizontal Vertical Horizontal Vertical Cold Leg Spt 2,455 . 1,268 522.6 354. 6 Hot Leg Spt l,293 762 515.0 429.4

Table 3 St. Lucie Unit fi'1 RV SUPPORT IfAX ABS REACTIONS (KIPS) - LOCA + SEISMIC (SRSS) 4 FT CLG BREAK AT NOZZLE lA OR 2A LOCATIONS RV SPPT STIFFNESS VALUES

='4.62 x 6 6-K 10 lb/in K = 77.54 x 10 lb/in 6 = 75.83 x 10 6 K ~ 59. 71 x 10 lb/in K lb/in 8'1A SPPT Vertical 1397 2317 Horizontal 1502 1587 ulB SPPT Vertical 2800 2251 Horizontal 5331 5473-Hot Leg SPPT Vertical 3458 3048 Horizontal 7493 7777 For LOCA + Nop reactions, add these values to the vertical results:

~21A SPPT 710~ K 81B SPPT 726. K Hot Leg SPPT 1157 K

+

For break at nozzle 1B or 2B, the loads on the cold leg supports would be reversed

Table 4 St, Lucie 1 Reactor Pressure Vessel Support Capacity Steel support structure horizontal 8400 kips*

(concrete is limiting)

Steel support structure vertical downward 12000 kips Reactor Cavity Wall horizontal ".13000 kips*~

Reactor Cavity Wall vertical not limiting Reactor Support Pads horizontal See Figure 4 Reactor Support Pads vertical See Figure 4 Load on individual girder

• < Allowable resultant asymmetric mechanical load transmitted along girders to concrete, based on rebar mean axial stress'being within yield.

TABLE 5 STEAM GENERATOR LOWER SUPPORT CALCULATED AN D DESI G N LOADS IREFER TO TABLE 1 FOR SYMBOLS) CL GUILL NO. 1 HL GUILL NO. 2 DESIGN LOAD SUPPORT + DBE IRSS) + DBE IRSS) LOCA+ DBE Z11 727 42 3,600 Z12 806 -1,868" FRONT Y1 -406.9 -2487.7 -1,770 SIDE Y2 -756.5 -1176.4 -1I737 BACK Y3 -605.0 +1249.0 ,691 SIDE Y4 -300.1 -1175.9 -1,734 X-STOP -5194.9 5,648 78 278 ,301 Z1 194 40 -1,574 Z2 301 40 1,800 IK AND FT KIPS) 9 g STEAM GENERATOR/

SLIDING BASE SUPPORT SKIRT IH L GUILL) RSS SLIDING BASE INTERFACE LOCA LOCA 5 DBE DESIGN LOADS

.Fx 5205 5205 5653 Fy -3582 3582 -2,471.0 6.8 11.0

'x Fz 418.3 32.0 24.0 Mz M609 4614 1003 "A NEGATIVE SIGN MEANS TENSION

Table 6 St. Lucie Unit IIl Reactor Coolant System Reactor Pressure Vessel and Reactor Coolant Pump Nozzle Loads Due to a 4 ft Reactor Vessel 1A Inlet Nozzle- Guillotine Break PIPE RUPTURE RSS MOMENT (In-Ki s)

RCP Snubber RCP Snubber Seismic Moment Allowable Moment Nozzle Not Ac~tin (In-Ki s) (In-Ki s)

RCP Discharge 109,300 109,600 5, 910 96,810 RCP Suction 50,500 54,550 7,256 78,965 RV Inlet 71,750 71,910 5,272 78,965 RV Outlet 50,150 50,170 2,535 279,340

Table 7 St. Lucie Unit No. 1 Connected Piping Stresses Calculated vs. Allowable 4.0 sq. ft. CLG Inlet Break Design Point Calculated Stress (Refe r to Fi ure 5) (E u. 10 ASME) Allowable Stress 39,070 48,600 75,152* 48,600 3 75,030* 48,600 X

5 41,835 48,600 6 43,475 48,600 47,690 48,600 33,430 48,600 20,171 48,600 Functionability and integrity are assured if Level B (upset conditions) limits of the ASME Boiler and Pressure Vessel Code,Section III, Division 1 are not exceeded. Functionability is important at points 5 and 6 where the valve is. At points 2 and 3, these limits are exceeded. However, Level D (faulted limits) are not exceeded at these two points. Level D limits are used to demonstrate that integrity is maintained. Equation (9) at those two points would yield 45,043 psi. and 44,479 psi respectively with an allowable of 48,600 psi.

FIGURE 1 RELAP4 510DEL FOR ST. LUCIE PR)MARY COOLANT SYSTEM 3')

1$ 13 14 58 15 16 5 6 8 32 0 04 59 30 11 60 83 96 7 7 12 33 62 57 4 3 62 63 4 8 6

3 95 e oo 23 24 9 10 11 1Z 31 13 14 15 16 17 19 20 54 41 42 4$ 45 46 47 48 21 22 25 23 24 51 QS3, 26 59 50 85 37 73 0>>

76 7S 51 49 48 47 61 63 (A1) 35 40 80 81 52 53 (81) Q69 71 70 43 42 41 55 S6 S7 Qoo Q72

'(a2) 74 73 98 VAt.VB 46 45 (82)

FIGURE 2 ST. LUCIE 1 RCS- V/HAM/6 MODEL FOR IAL pss Qss Qss 11 13 107 01 02 87 106 53 86 74 10 12 81 14 072 58 eo 85 108 71 104 103 102 82 83~ 84 89

>>5 75 109 76 110 77 70 69 0 80 2 49 95 62 51 57 61 g 114 8P 81 82 78 0,0, 0, Qs 0 96 2'00 98 Qss 101 99 047 68 1 18 20 19 67 66 Q1 17 13 '4 79 15 12 26 28 46 24 18 19 16 31 25 27 0

42 35 22 44, 0 77 46 38 23 39 32 Qss 40 33 0

41 43 Qss oo 52 0~

45 76 I0 55 0

49 41 (A1) (81)

P14 53 75 P1 P10 58 60 42 74 62 = 56

~~X 69 61 4 73 38 o 120 40 69 72 71 70 Pso 0 BRGAK P5 P18 (A2) (82)

>>9 67 118 21

~83 0 122 0~

FIGUPESA ST. LUCIE 1 REDUCED MODEL OF REACTOR INTERPeALS 38

'A 18 UGS :A Z'2. 16 H

X - DI R IS // TO OUTLET NOZZLES'"

28I 10 '9

.CSB 26'4,', CORE SHROUD FUEL' 22',

20, A

13 H LEGEND A ~ AXIALGAP H = HORIZONTAL GAP

// = PRELOADED COUPLING

'SEE FIGURE 3b FOR DETAILS OF REACTOR VESSELS AND PIPING

= GAP COUPLING

= COLINEAR CONNECTOR

v 1

'FIGURE 38 ST LUCIE1 REDUCED MODEL OF REACTOR COOLANT SYSTEM o 9918 REACTOR VESSEL TO S.G. NO. 1' TO S.G. NO. 2 r r r 9999 HOT LEG 9909 INTERNALS

'SEE FIGURE 3a FOR DETAILS OF INTERNALS) 9905 o 9914 r o 9913 0o LUMPED MASS POINT OF APPLIED FORCE r

~

r

.FIGURE 4 ST. LUCIE 1 .-:

REACTOR PRESSURE VESSEL SUPPORT PAD CAPABILITY I

hC

,='4.0 SO. FT. CLG

'COMPUTED MAX LOADS i (LOCA+ SEISMIC+ Mop)

K- I QA (HOT LEG SUPPORT 10 QB! VNBROl<EN COLD LEG SUPPORT HOT LEG SUPPORT BROI<EN COLD LEG SUPPORT QA

~l 0

COLD LEG SUPPORTS Z'I 0'

4 0:

~I

p. 1; 2. 3'. 4: 5I 6: 7., 8, 9,10 ..11 12;13:14,15,-16 Rq x 10 (K)

VERTICAL LOAD.

FIXED 4

5,' '-

37

"-~

x 35 36 DISP LACEMENTS SPECIFIED 30 29 42 28 43 8

30 27 45 27 10 31 45 11 FlGURE 5': 'l2 FIXED 24-SAFETY INJECTlQN 23 22 LINE 1-8-1 (PREVIOUS 972) 21 t

FLGRIDA PORFB 8 LIGHT OOMPANY 13 14 20 15 ST LUCIE NO. 1 18 17 ~O

'< 's~

I