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| number = ML17207A876
| number = ML17207A876
| issue date = 03/03/1980
| issue date = 03/03/1980
| title = RCS Asymmetric LOCA Evaluation.
| title = RCS Asymmetric LOCA Evaluation
| author name =  
| author name =  
| author affiliation = FLORIDA POWER & LIGHT CO.
| author affiliation = FLORIDA POWER & LIGHT CO.
Line 17: Line 17:


=Text=
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{{#Wiki_filter:REACTOR COOLANT SYSTEM ASYMMETRIC LOCA LOAD EVALUATION ST. LUCIE UNIT 1 DOCKET NO. 50-335
{{#Wiki_filter:REACTOR COOLANT SYSTEM ASYMMETRIC LOCA LOAD EVALUATION ST.
          )farch 3, 1980
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.
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.
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
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.
,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.
The Staff requested that further internal asymmetric load (IAL) evaluations be conducted.
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.
FP&Ls letter of February 9,
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.
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.
This report discusses the results of this expanded analysis.
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.
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.
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.
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 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 analyses of the adequacy of fuel assemblies internals, and control element assemblies is in progress.
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.
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.
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.
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.
: 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
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.
: vessel, a decompression of the reactor pressure vessel occurs over a short period of time.
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.
Decompression waves originated.
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.
at the postulated break travel around the inlet plenum and propa-gate downward along the downcomer annulus.
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- .
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.
ment of the adequacy of piping attached to the primary system piping.
This field imposes a transient asymmetric loading on the core-support-barrel as well as the vessel itself.
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.
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.
: 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-.
This external pressurization is also asymmetric.
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 internal asymmetric loading (IAL) and the external asymmetric loading act in the same direction for breaks occurring in the cold leg piping.
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.
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.
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:
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   .
mR.SX(T) 2 Bn mg(R.+R ) t (v) =
s'v.          mg(R.+R ) t (v) =           + 2R j m-( Bn        2g) f
<5
              <5 where Ri and     R are the inner and outer pipe radii, t 0
+ 2R j m-( s'v. 2g) f where Ri and R
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.
are the inner and outer pipe radii, t 0
i wherein y is the radial separation of the two broken ends which also varies with time.
is the pipe thz.ckness, x is the axial sepanation of the two ends which varie's with time 7, and g = cos 2R.i (2) 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)
This equation is solved in iterative fashion together with the equation for the combined tension release and blowdown force, given below V
F(x)   P dl (v)A +
(T) 2 F(x)
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.
P (v)A
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.
+ p (x)A (v) dl p
Results of the analyses indicate that- at least 18 msec.
dl g
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.
(3) to yield the correct forcing function and break area as a function of time.
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
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
'ischarge leg, the pump, and the cross over leg with 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. refer 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
horizontal asymmetric load (external vertical asym-metric loads are virtually identical for the two breaks).
                                                              '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.
Since the 4.0 sq. ft. break had been one
'f the design basis breaks in the FSAR, all analyses 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)
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.
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.
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.
The results for the 4.0 sq.
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.
ft. cold leg guillotine break, as reported in Reference 2, have been directly used in'the present study.
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.
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.
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
3/
: deduced, is shown in Figure l.
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.
The RELAP-4thexmal hydraulic code is used to compute 3/
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.
the thermodynamic properties in the model volumes and junctions.
requiring 8 msec. to open to approximately 9.81 sq. ft.
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.
(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.
of 36 msec.,
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.
only results in a 2 to 3 percent increase in peak internal asymmetxic loads, whereas a decrease in opening time from 36 msec.
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.
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 STRUDL computer code generates the condensed 5/
The dyanmic analysis       to determine the systyy response was performed using       the computer code DAGS and DFORCE .
stiffness matrix used in the dynamic'analysis from the physical definition of the structure.
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.
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.
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 hydrodyamic (added) mass matrix is evaluated using the ADifASS code..
The   RCS supports most affected are the lower steam generator supports.
The dyanmic analysis to determine the systyy response was performed using the computer code DAGS and DFORCE.
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.
The reactor pressure vessel support system is described in the FSAR.
This analysis is a static analysis which employs the compu(p) code 51EC-21 (Hare Island     piping flexibility code)     .
The modelling of the steel portion of the support is identical to that described in the FSAR in Appendix 3H.
Both   LOCA and   des'ign seismic loads are included   in the analysis.
The basic model of the biological shield wall is also identical.
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
: 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.
applied to the line which by configuration and other loading (normal and seismic) would result in the highest stresses.
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.
The stresses induced by LOCA motions for this particular configuration are added-to previously computed normal and seismic (SRSS) stresses.
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.
The determining break for ECCS line evalua-tion is the cold leg nozzle break in the cavity.
In the support analyses however, the higher loads resulting from the use of the highest stiffness, have been utilized.
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.
This insures again that the absolute maximum load per support is computed. In reality, lower values are expected.
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.
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.
This range covers the possible values of the overall stiffness of the individual'actor vessel
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.
: supports, the real value being somewhere between the two extremes.
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
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.
used in the analysis, as explained in Section 2.1.2.
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.
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.
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.
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.
A detailed analysis of the reactor loads resulting from hot leg breaks within the cavity has not been performed.
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.
The masons are as follows:
Horizontal Hot     Horizontal Cold Le Break (Ki s)   Le Break (Ki s)
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
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.
: 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.
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.
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 Le Break (Ki s)
Horizontal Cold Le Break (Ki s)
Cold Leg Spt Hot Leg Spt 4270 3270 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.
A similar conclusion had been, reached in our August 1977 report.
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.
Differences in maximum loads reported herein from those reported in the August 1977 report are two fold.
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 .
The August 1977 report did not consider internal gaps or gaps between'he support pads and the support structure.
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.
The August 1977 report considered therefore that all loaded supports would be loaded simultaneously and share the load equally.
Since this analysis is predicated
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.
                          ~                  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.
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. ft. 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
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.
integrity would be maintained at those nozzles.
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.
Since the amount of overstressing calculated on an elastic basis is relatively small, a plastic analysis was not considered necessary.
Hence, the analysis was repeated by taking no account of the snubbers. Results are also reported in Table 6 As can be
During the performance of this particular'analysis it was calculated that the snubbers on the reactor coolant pumps are overstressed.
                                                            ~
These snubbe~
clearly seen, the effect of the presence or absence of the snubbers   is negligible.
are not needed for these events.
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.
However their failure could affect the results.
Table   7 compares the peak computed stresses, which     include no mal   and seismic loads to the allowable stresses.
: Hence, the analysis was repeated by taking no account of the snubbers.
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.
Results are also reported in Table 6 ~
: 3. 5   Seismic Loads
As can be clearly seen, the effect of the presence or absence of the snubbers is negligible.
    , 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.
3.4 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.
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.
Table 7 compares the peak computed stresses, which include no mal and seismic loads to the allowable stresses.
The assumption   of a complete guillotine   will result   in breaks larger than 0.5 sq. ft.
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==
==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 Unit l is acceptable.


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
==5.0 REFERENCES==
Unit is acceptable.
"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.
==5.0       REFERENCES==
St. Lucie Unit No.
 
1 FSAR, Docket No. 50-335, Amendment 44.*
"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/
2/ St. Lucie Unit     No. 1 FSAR, Docket No. 50-335, Amendment 44.*
"RELAP 4 A Computer Program for Transient Thermal Hydraulic Analysis 3/
3/"RELAP 4 A Computer Program for Transient Thermal Hydraulic Analysis of Nuclear Reactors and Related, Systems", User's Manual, ANCR-NUREG-1335.
of Nuclear Reactors and Related, Systems",
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.
User's Manual, ANCR-NUREG-1335.
"ICES     STRUDL   II The Structural Design   Language Engineers User's Manual",
Fabic, S.,
MIT Press,     Cambridge, Massachusetts,   1968.
Computex Program WHAM for Calculating Pressure, Velocity 4/
6/"ADMASS     A Computer Code   for Fluid Structure Interaction Using the Finite Element Technique", Ebasco Services,       Incorporated, 1979.
and Force Transients in Liquid Filled Piping Network", Kaiser Engineering Report No. 67-49-R, November 1967.
"DAGS     CENPD 168,   Revision 1 Design   Basis Pipe Breaks", September 1976.
"ICES STRUDL II The Structural Design Language Engineers User's Manual",
"DFORCE       Design   Basis Pipe Breaks", September 1976.
MIT Press, Cambridge, Massachusetts, 1968.
"MSC/NASTRAN       User's Manual", McNeal Schwandler Corporation, Los Angeles, California.
"ADMASS A Computer Code for Fluid Structure Interaction Using the Finite 6/
 
Element Technique",
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.
Ebasco Services, Incorporated, 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.
"DAGS CENPD 168, Revision 1 Design Basis Pipe Breaks",
Griffin, J.       H., "MEC-21   A Piping Flexibility Analysis Program", TID-4500 (31st edition),   LA-2924, UC-38, July 14, 1964.
September 1976.
"DFORCE Design Basis Pipe Breaks",
September 1976.
"MSC/NASTRAN User's Manual", McNeal Schwandler Corporation, Los Angeles, California.
"WCAP 9570 Class 3 Mechanistic Fracture Evaluation of Reactor Coolant 10/
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.)
TABLE 1 ST. LUCIE 1 NORMAL AND SEISMIC SUPPORT LOADS (X106LB.)
IIOAMAL OPEIIATING           ODE SEISMIC                DBE SEISSIIC I
CONDITION LOAD H1 V1 NV1 0
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$
.GGG 4195
V1        .GGG      1.155    .032    .335    .04G    .064      .670    .092 NV1      4195        i.360 HZ                  .091  1.226     .001   +.355   2.452     .003     +.710
.028 1.155 i.360 IIOAMAL OPEIIATING THEGMAL+
                                                .264-                      -.528         ivt V2        .G64        .72G  .017    .253    .380    035      .507      .761   I                     $ $ ACIOII V$ $ $ ($
DEAD DEAD VIEIGHT V(EIGHT
                                                                                                          $ Vt(OIIIA(ACIIO(N NV2      a.195      k.215
.005
                        .079 1.139   .019     .270   2.278   .038       .540                   $ ($
.032 ODE SEISMIC SY
.002
.335
.G44
.04G
.011
.064 DBE SEISSIIC SY
.005
.670 1.288
.092 I
f N
AV1 V$
tV HZ V2 NV2
.G64 a.195
.091
.72G k.215 1.226
.017
.001
.253
+.355
.264-
.380 2.452 035
.003
.507
+.710
-.528
.761 ivt I
$$ACIOIIV$$$($
$Vt(OIIIA(ACIIO(N V3 NV3 Z11 Z12 Y1 Y2 Y3 Y4 x
NY x.195 0
0
.300
.300 1.300 1
.300 0
a.376
.741
%.215
.367 0
0 0
.315 1.009
.320 0
S.300 0
0
.016
.057
.086
.054 0
.079 1.139
.019
.060
.G23 0
0
.019
.155
.417
.152 0
.270
.349
.743 S.195 4197 0
.060
.071 058 0
2.278
.006 0
0
.033
.1'14
.173
.108 0
.038
~120
.50G 0
0
.039 211
.835
.305 0
.540
.6SS
.746 S.390 S.394 0
.121
.143
.116 0
Y$
$($
YI
.v h
Vf
$(fANC(N(AAION LOW(N$V(YOA1$
YI
(
(
                                      .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
N
                                                                                                      $ (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)
Table 2
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
Comparison of Peak Calculated and Design Seismic (DBE) Loads at Representative Locations Design (Kips)
Horizontal Vertical Calculated (Kips)
Horizontal Vertical Cold Leg Spt Hot Leg Spt l,293 762 2,455 1,268 522.6 515.0 354. 6 429.4


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
Table 3
                        ='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:
St. Lucie Unit fi'1 RV SUPPORT IfAX ABS REACTIONS (KIPS) - LOCA + SEISMIC (SRSS) 4 FT CLG BREAK AT NOZZLE lA OR 2A LOCATIONS 8'1A SPPT RV SPPT STIFFNESS VALUES K ='4.62 x 10 lb/in 6
            ~21A   SPPT     710~ K 81B   SPPT     726. K Hot Leg   SPPT       1157   K
K
+
~ 59. 71 x 10 lb/in 6
For break at nozzle     1B or 2B, the loads on the cold leg supports would be reversed
K = 77.54 x 10 lb/in 6-K = 75.83 x 10 lb/in 6
Vertical Horizontal 1397 1502 2317 1587 ulB SPPT Vertical Horizontal 2800 5331 2251 5473-Hot Leg SPPT Vertical Horizontal 3458 7493 3048 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*
Table 4
(concrete   is limiting)
St, Lucie 1 Reactor Pressure Vessel Support Capacity Steel support structure horizontal 8400 kips*
Steel support structure     vertical downward           12000 kips Reactor Cavity Wall         horizontal                 ".13000 kips*~
(concrete is limiting)
Reactor Cavity Wall          vertical                  not limiting Reactor Support Pads        horizontal                See Figure   4 Reactor Support Pads        vertical                  See Figure   4 Load on individual girder
Steel support structure vertical downward 12000 kips Reactor Cavity Wall Reactor Cavity Wall Reactor Support Pads Reactor Support Pads horizontal vertical horizontal vertical
*< Allowable resultant asymmetric mechanical load transmitted along girders to concrete, based on rebar mean axial stress'being within yield.
".13000 kips*~
not limiting See Figure 4
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/
TABLE 5 STEAM GENERATOR LOWER SUPPORT CALCULATED AND DESI G N LOADS IREFER TO TABLE 1 FOR SYMBOLS)
SLIDING BASE SUPPORT SKIRT         IH L GUILL)         RSS     SLIDING BASE INTERFACE                        LOCA          LOCA 5 DBE   DESIGN LOADS
SUPPORT Z11 Z12 FRONT Y1 SIDE Y2 BACKY3 SIDE Y4 X-STOP Z1 Z2 CL GUILLNO. 1
          .Fx                          5205              5205            5653 Fy                        -3582            3582        -2,471.0 6.8            11.0
+ DBE IRSS) 727 806
          'x Fz 418.3            32.0 24.0 Mz                        M609                4614            1003 "A NEGATIVE SIGN MEANS TENSION
-406.9
-756.5
-605.0
-300.1 78 194 301 HL GUILLNO. 2
+ DBE IRSS) 42
-2487.7
-1176.4
+1249.0
-1175.9
-5194.9 278 40 40 DESIGN LOAD LOCA+ DBE 3,600
-1,868"
-1,770
-1I737
,691
-1,734 5,648
,301
-1,574 1,800 IK AND FT KIPS) 9 g STEAM GENERATOR/
SLIDING BASE SUPPORT SKIRT INTERFACE
.Fx Fy Fz'x Mz IH L GUILL)
LOCA 5205
-3582 M609 RSS LOCA 5 DBE 5205 3582 6.8 418.3 4614 SLIDING BASE DESIGN LOADS 5653
-2,471.0 11.0 32.0 24.0 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)
Table 6
RCP Snubber         RCP Snubber         Seismic Moment Allowable  Moment Nozzle                                          Not Ac~tin           (In-Ki s)     (In-Ki s)
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  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
Nozzle RCP Discharge RCP Suction RV Inlet RV Outlet RCP Snubber 109,300 50,500 71,750 50,150 RCP Snubber Not Ac~tin 109,600 54,550 71,910 50,170 Seismic Moment (In-Ki s) 5, 910 7,256 5,272 2,535 Allowable Moment (In-Ki s) 96,810 78,965 78,965 279,340


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
Table 7
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.
St. Lucie Unit No.
1 Connected Piping Stresses Calculated vs. Allowable 4.0 sq. ft.
CLG Inlet Break Design Point (Refe r to Fi ure 5)
Calculated Stress (E u.
10 ASME)
Allowable Stress 3
X 5
6 39,070 75,152*
75,030*
41,835 43,475 47,690 33,430 20,171 48,600 48,600 48,600 48,600 48,600 48,600 48,600 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')
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
32 5
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>>
6 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
04 59 11 60 30 57 3
                  '(a2)                                             74              73 98 VAt.VB             46               45 (82)
1$
13 14 58 83 96 15 16 8
12 33 7
7 4
62 63 4
8 62 e oo 9
10 6
3 95 11 23 24 1Z 13 14 15 31 16 17 54 19 20 4$
41 42 45 46 21 22 25 51 23 47 24 48
: QS3, 26 59 73 50 0>>
37 85 51 76 7S 49 48 47 61 52 (A1) 80 81 53 63 35 (81) 71 70 43 42 41 40 Q69 55 S6 S7 Qoo
'(a2) 98 VAt.VB 74 73 46 45 (82)
Q72


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
FIGURE 2 ST. LUCIE 1 RCS-V/HAM/6 MODEL FOR IAL pss Qss Qss 107 74 106 072 108 71 104 103 10 11 13 12 102 81 14 82 02 01 87 53 86 58 eo 85 83~ 84 89 109 110 70 69
      >>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
>>5 75 76 77 g
42 35 22 44, 0    77     46 38 23 39 32 Qss 40 33 0
114 8P 81 78 82 1
41 43 Qss oo 52 0~
Q1 17 0
45 76 I0          55 0
18 19 20 13 '4 26 18 28 27 0,0, 0, 80 047 79 46 Qs 2
49 41 (A1)             (81)
0 101 68 15 12 19 16 31 49 51 57 95 62 61 96 2'00 99 67 66 24 25 98 Qss 0
P14 53       75 P1                        P10 58         60       42 74         62           = 56
0 35 42 22 43 0
                                                                                                                                                          ~~X 69          61 4    73    38 o               120 40 69 72 71 70 Pso 0                         BRGAK P5         P18 (A2)           (82)
38 44, 77 46 23 0~
                                                  >>9            67 118                  21
39 32 Qss 33 40 (A1)
                                                                        ~83 0                                  122 0~
(81) 41 Qss 45 52oo 53 58 60 42 69 61 4 I
76 0 75 74 73 38 0
41 55 49 62
=
56 P1 P14 P10 ~~X o
>>90 40 72 120 69 71 67 118 ~83 0
70 Pso 21 122 0~
BRGAK P5 (A2)
P18 (82)


FIGUPESA ST. LUCIE             1 REDUCED MODEL OF REACTOR INTERPeALS 38
FIGUPESA ST. LUCIE 1 REDUCED MODEL OF REACTOR INTERPeALS UGS 38
                                                            'A 18 UGS                                :A Z'2.                           16 H
'A 18
X - DI R IS // TO OUTLET   NOZZLES'"
:A Z'2.
28I                     10                 '9
16 H
                                                          .CSB 26'4,',                       CORE SHROUD FUEL' 22',
X - DIR IS // TO OUTLET NOZZLES'"
20, A
28I 10
13                         H LEGEND A ~ AXIALGAP H = HORIZONTAL GAP
'9 26'4,',
      // = PRELOADED COUPLING
22',
                              'SEE             FIGURE 3b FOR DETAILS OF REACTOR VESSELS AND PIPING
FUEL' CORE SHROUD
        = GAP COUPLING
.CSB 20, A
        = COLINEAR CONNECTOR
13 H
LEGEND A ~ AXIALGAP H = HORIZONTALGAP
// = PRELOADED COUPLING
= GAP COUPLING
= COLINEAR CONNECTOR
'SEE FIGURE 3b FOR DETAILS OF REACTOR VESSELS AND PIPING


v 1
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
'FIGURE 38 ST LUCIE1 REDUCED MODEL OF REACTOR COOLANT SYSTEM TO S.G. NO. 1' rr r
                  'SEE   FIGURE 3a FOR DETAILS OF INTERNALS) 9905 o 9914 r o 9913            0o  LUMPED MASS POINT OF APPLIED FORCE r
o 9918 9999 REACTOR VESSEL HOT LEG TO S.G. NO. 2 9909 INTERNALS
                              ~
'SEE FIGURE 3a FOR DETAILSOF INTERNALS) 9905 r
r
r
~ r o
9914 o
9913 o
LUMPED MASS 0 POINT OF APPLIED FORCE


                              .FIGURE 4         ST. LUCIE 1   .-:
.FIGURE 4 ST. LUCIE 1 REACTOR PRESSURE VESSEL SUPPORT PAD CAPABILITY I
REACTOR PRESSURE VESSEL SUPPORT PAD CAPABILITY I
hC K-10 HOT LEG SUPPORT
hC
,='4.0 SO. FT. CLG
                                                                          ,='4.0 SO. FT. CLG
'COMPUTED MAX LOADS i (LOCA+SEISMIC+ Mop)
                                                                          'COMPUTED MAX LOADS i (LOCA+ SEISMIC+ Mop)
I QA (HOT LEG SUPPORT QB!VNBROl<EN COLD LEG SUPPORT BROI<EN COLD LEG SUPPORT QA
K-                                                                I QA (HOT LEG SUPPORT 10 QB! VNBROl<EN COLD LEG SUPPORT HOT LEG SUPPORT BROI<EN COLD LEG SUPPORT QA
~l 0
~l 0
COLD LEG SUPPORTS Z'I 0'
Z'I0' 4
4 0:
0:
~I
~I COLD LEG SUPPORTS p.
: p. 1;   2. 3'. 4: 5I     6:   7.,   8,   9,10 ..11     12;13:14,15,-16             Rq x 10 (K)
1; 2.
VERTICAL LOAD.
3'.
4:
5I 6:
7.,
8, 9,10..11 12;13:14,15,-16 Rq x 10 (K)
VERTICALLOAD.


FIXED 4
FIXED "-~
5,'                                                       '-
4 5,'
37
'- 37 x
                                                                                                      "-~
35 36 DISP LACEMENTS SPECIFIED 30 29 42 43 28 8
x                                                        35                 36 DISP LACEMENTS SPECIFIED 30 29 42                   28 43 8
30 27 27 45 10 31 45 11
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
'l2 FlGURE 5':
FLGRIDA PORFB 8 LIGHT OOMPANY               13 14 20 15 ST LUCIE NO. 1                                     18 17 ~O
24-FIXED SAFETY INJECTlQN LINE 1-8-1 (PREVIOUS 972)
FLGRIDA PORFB 8 LIGHT OOMPANY ST LUCIE NO. 1 13 14 15 17 ~O 21 20 18 23 22 t


  '< 's~
'< 's~
I}}
I}}

Latest revision as of 16:21, 8 January 2025

RCS Asymmetric LOCA Evaluation
ML17207A876
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Site: Saint Lucie NextEra Energy icon.png
Issue date: 03/03/1980
From:
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Text

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 Bn mg(R.+R ) t (v) =

<5

+ 2R j m-( s'v. 2g) f 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 2R.i (2) 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

(T) 2 F(x)

P (v)A

+ p (x)A (v) dl p

dl 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

'ischarge leg, the pump, and the cross over leg with 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. refer 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

'f the design basis breaks in the FSAR, all analyses 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.

The RELAP-4thexmal hydraulic code is used to compute 3/

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 computer code generates the condensed 5/

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 Le Break (Ki s)

Horizontal Cold Le Break (Ki s)

Cold Leg Spt Hot Leg Spt 4270 3270 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. ft. 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 Unit l 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.

St. Lucie Unit No.

1 FSAR, Docket No. 50-335, Amendment 44.*

2/

"RELAP 4 A Computer Program for Transient Thermal Hydraulic Analysis 3/

of Nuclear Reactors and Related, Systems",

User's Manual, ANCR-NUREG-1335.

Fabic, S.,

Computex Program WHAM for Calculating Pressure, Velocity 4/

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.

"ADMASS A Computer Code for Fluid Structure Interaction Using the Finite 6/

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.

"WCAP 9570 Class 3 Mechanistic Fracture Evaluation of Reactor Coolant 10/

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.)

CONDITION LOAD H1 V1 NV1 0

.GGG 4195

.028 1.155 i.360 IIOAMAL OPEIIATING THEGMAL+

DEAD DEAD VIEIGHT V(EIGHT

.005

.032 ODE SEISMIC SY

.002

.335

.G44

.04G

.011

.064 DBE SEISSIIC SY

.005

.670 1.288

.092 I

f N

AV1 V$

tV HZ V2 NV2

.G64 a.195

.091

.72G k.215 1.226

.017

.001

.253

+.355

.264-

.380 2.452 035

.003

.507

+.710

-.528

.761 ivt I

$$ACIOIIV$$$($

$Vt(OIIIA(ACIIO(N V3 NV3 Z11 Z12 Y1 Y2 Y3 Y4 x

NY x.195 0

0

.300

.300 1.300 1

.300 0

a.376

.741

%.215

.367 0

0 0

.315 1.009

.320 0

S.300 0

0

.016

.057

.086

.054 0

.079 1.139

.019

.060

.G23 0

0

.019

.155

.417

.152 0

.270

.349

.743 S.195 4197 0

.060

.071 058 0

2.278

.006 0

0

.033

.1'14

.173

.108 0

.038

~120

.50G 0

0

.039 211

.835

.305 0

.540

.6SS

.746 S.390 S.394 0

.121

.143

.116 0

Y$

$($

YI

.v h

Vf

$(fANC(N(AAION LOW(N$V(YOA1$

YI

(

$($

N

Table 2

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

Horizontal Vertical Calculated (Kips)

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

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 8'1A SPPT RV SPPT STIFFNESS VALUES K ='4.62 x 10 lb/in 6

K

~ 59. 71 x 10 lb/in 6

K = 77.54 x 10 lb/in 6-K = 75.83 x 10 lb/in 6

Vertical Horizontal 1397 1502 2317 1587 ulB SPPT Vertical Horizontal 2800 5331 2251 5473-Hot Leg SPPT Vertical Horizontal 3458 7493 3048 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 Reactor Cavity Wall Reactor Support Pads Reactor Support Pads horizontal vertical horizontal vertical

".13000 kips*~

not limiting See Figure 4

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 AND DESI G N LOADS IREFER TO TABLE 1 FOR SYMBOLS)

SUPPORT Z11 Z12 FRONT Y1 SIDE Y2 BACKY3 SIDE Y4 X-STOP Z1 Z2 CL GUILLNO. 1

+ DBE IRSS) 727 806

-406.9

-756.5

-605.0

-300.1 78 194 301 HL GUILLNO. 2

+ DBE IRSS) 42

-2487.7

-1176.4

+1249.0

-1175.9

-5194.9 278 40 40 DESIGN LOAD LOCA+ DBE 3,600

-1,868"

-1,770

-1I737

,691

-1,734 5,648

,301

-1,574 1,800 IK AND FT KIPS) 9 g STEAM GENERATOR/

SLIDING BASE SUPPORT SKIRT INTERFACE

.Fx Fy Fz'x Mz IH L GUILL)

LOCA 5205

-3582 M609 RSS LOCA 5 DBE 5205 3582 6.8 418.3 4614 SLIDING BASE DESIGN LOADS 5653

-2,471.0 11.0 32.0 24.0 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)

Nozzle RCP Discharge RCP Suction RV Inlet RV Outlet RCP Snubber 109,300 50,500 71,750 50,150 RCP Snubber Not Ac~tin 109,600 54,550 71,910 50,170 Seismic Moment (In-Ki s) 5, 910 7,256 5,272 2,535 Allowable Moment (In-Ki s) 96,810 78,965 78,965 279,340

Table 7

St. Lucie Unit No.

1 Connected Piping Stresses Calculated vs. Allowable 4.0 sq. ft.

CLG Inlet Break Design Point (Refe r to Fi ure 5)

Calculated Stress (E u.

10 ASME)

Allowable Stress 3

X 5

6 39,070 75,152*

75,030*

41,835 43,475 47,690 33,430 20,171 48,600 48,600 48,600 48,600 48,600 48,600 48,600 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')

32 5

6 0

04 59 11 60 30 57 3

1$

13 14 58 83 96 15 16 8

12 33 7

7 4

62 63 4

8 62 e oo 9

10 6

3 95 11 23 24 1Z 13 14 15 31 16 17 54 19 20 4$

41 42 45 46 21 22 25 51 23 47 24 48

QS3, 26 59 73 50 0>>

37 85 51 76 7S 49 48 47 61 52 (A1) 80 81 53 63 35 (81) 71 70 43 42 41 40 Q69 55 S6 S7 Qoo

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

Q72

FIGURE 2 ST. LUCIE 1 RCS-V/HAM/6 MODEL FOR IAL pss Qss Qss 107 74 106 072 108 71 104 103 10 11 13 12 102 81 14 82 02 01 87 53 86 58 eo 85 83~ 84 89 109 110 70 69

>>5 75 76 77 g

114 8P 81 78 82 1

Q1 17 0

18 19 20 13 '4 26 18 28 27 0,0, 0, 80 047 79 46 Qs 2

0 101 68 15 12 19 16 31 49 51 57 95 62 61 96 2'00 99 67 66 24 25 98 Qss 0

0 35 42 22 43 0

38 44, 77 46 23 0~

39 32 Qss 33 40 (A1)

(81) 41 Qss 45 52oo 53 58 60 42 69 61 4 I

76 0 75 74 73 38 0

41 55 49 62

=

56 P1 P14 P10 ~~X o

>>90 40 72 120 69 71 67 118 ~83 0

70 Pso 21 122 0~

BRGAK P5 (A2)

P18 (82)

FIGUPESA ST. LUCIE 1 REDUCED MODEL OF REACTOR INTERPeALS UGS 38

'A 18

A Z'2.

16 H

X - DIR IS // TO OUTLET NOZZLES'"

28I 10

'9 26'4,',

22',

FUEL' CORE SHROUD

.CSB 20, A

13 H

LEGEND A ~ AXIALGAP H = HORIZONTALGAP

// = PRELOADED COUPLING

= GAP COUPLING

= COLINEAR CONNECTOR

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

v 1

'FIGURE 38 ST LUCIE1 REDUCED MODEL OF REACTOR COOLANT SYSTEM TO S.G. NO. 1' rr r

o 9918 9999 REACTOR VESSEL HOT LEG TO S.G. NO. 2 9909 INTERNALS

'SEE FIGURE 3a FOR DETAILSOF INTERNALS) 9905 r

r

~ r o

9914 o

9913 o

LUMPED MASS 0 POINT OF APPLIED FORCE

.FIGURE 4 ST. LUCIE 1 REACTOR PRESSURE VESSEL SUPPORT PAD CAPABILITY I

hC K-10 HOT LEG SUPPORT

,='4.0 SO. FT. CLG

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

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

~l 0

Z'I0' 4

0:

~I COLD LEG SUPPORTS p.

1; 2.

3'.

4:

5I 6:

7.,

8, 9,10..11 12;13:14,15,-16 Rq x 10 (K)

VERTICALLOAD.

FIXED "-~

4 5,'

'- 37 x

35 36 DISP LACEMENTS SPECIFIED 30 29 42 43 28 8

30 27 27 45 10 31 45 11

'l2 FlGURE 5':

24-FIXED SAFETY INJECTlQN LINE 1-8-1 (PREVIOUS 972)

FLGRIDA PORFB 8 LIGHT OOMPANY ST LUCIE NO. 1 13 14 15 17 ~O 21 20 18 23 22 t

'< 's~

I

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