ML20237C409
| ML20237C409 | |
| Person / Time | |
|---|---|
| Site: | Byron, Braidwood, 05000000 |
| Issue date: | 11/30/1987 |
| From: | Alper N, Augustine D, Ditommaso S, Pujadas S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19302D195 | List: |
| References | |
| WCAP-11387, WCAP-11387-R02, WCAP-11387-R2, NUDOCS 8712210304 | |
| Download: ML20237C409 (150) | |
Text
l WESTINGHOUSE CLASS 3 WCAP-11387 Rev. 2 4
BYRON /BRAIDWOOD T REDUCTION FINAL HOT LICENSING REPORT NOVEMBER 1987 S. N. DiTommaso D. B. Augustine N. Alper a S. A. Pujadas i
- WESTINGHOUSE ELECTRIC CORPORATION Power Systems
' P. D. Box 355
' Pittsburgh, Pennsylvania 15230 8712210304 871204 gDR ADOCK 050 4 I
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- The authors wish to acknowledge the contribution of the following individuals:
E. Manz
. E. White C. Thompson M. Emery M. Wengerd J. Bass M. Weaver i G. Smith T. Miller G. Whiteman
- D. Paulsen R. Osterrieder
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- M. Kachmar a -
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WESTINGHOUSE PROPRIETARY CLASS 2 TABLE OF CONTENTS Page 4 Executive Summary i 1.0 Program Description j 1.1 Definition of. Goals I l 1.2 Applicable Criteria 1 1.3 Scope 2 2.0 Determination of Reduced Temperature Operating Conditions 2.1 Design Power Capability Parameters 3 1 2.2 Control and Protection System Setpoints 3 2.3 Design Transients 4 1
3.0 Safety Evaluations '
3.1 Loss of Coolant Accident (LOCA) 6 3.2 LOCA Hydraulic Forces 7 3.3 Subcompartment Mass / Energy Release 9 3.4 Non-LOCA Transients 13 3.5 Systems and Components Evaluation 44 3.6 Technical Specification Impact 60 3.7 Final Safety Analysis Report (FSAR) Impact 60 3.8 Operating Window Concept 61 4.0 LicensingCriteriaReview 4.1 10 CFR 50.59 (Unreviewed Safety Question) 63 4.2 10 CFR 50.92 (Significant Hazard Determination for 63 Issuance of Amendment) 4.3 10 CFR 50.36 (Technical Specifications) 63 4.4 10 CFR 50.71 (FSAR Update) 64
.. 5.0 Conclusion 65
,". Appendix A - FSAR Mark-ups for Large and Small Break LOCA Appendix B - Technical Specification Mark ups 0181v:1o-021287
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EXECUTIVE
SUMMARY
In order to reduce.the' potential for the initiation and propagation of primary water stress corrosion cracking (PWSCC) in the Model D steam generator tubes, Connonwealth Edison is taking the action to operate Byron 1 & 2 and Braidwood 1 & 2 with a reduced primary system vessel outlet temperature (Thot) "I' 600*F. This constitutes a reduction of 18.4*F from the original design Thot of 618.4*F. Details of the safety analyses of the Nuclear Steam Supply System (NSSS) and turbine generator which justify plant operation at 100% power at this temperature condition are provided in this report.
The evaluation and analysis effort was divided into fourteen distinct tasks:
- 1. Feasibility of Maintaining Current Zero Load Temperature
- 2. Power Capability Parameters 3.. LOCA Evaluation
- 4. , Control System Setpoints
- 5. NSSS Design Transients
- 6. Non-LOCA Evaluation
- 7. LOCA Hydraulic Forcing Functions
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- 8. Pressurizer Study
- 9. Safety Evaluation of Components
- 10. Operating Window Concept
- 11. Licensing Support
- 12. Turbine Limits
- 13. Containment Nass/ Energy Release 14.ProjectCoordination 1
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l i The evaluations demonstrate that no unreviewed safety question is involved.
Reanalysis and evaluation of specifically impacted FSAR accidents has l.' supported this conclusion by showing that the probability of occurrence, possibility of new accidents or margin of safety in any Technical l- Specifications basis has not changed from the original design. Options selected by Commonwealth Edison which incorporate such operating margin as 10%
! steam generator tube plugging, 5% safety injection flow reduction, K(z) curve modification and increased FAH and Fg has been factored into the LOCA analysis. The K(z) curve modification necessitates a change to the Technical Specifications. Further analyses of the non-LOCA transients and a concurrent l
change to core limits is required before Fg and FAH can be revised in the Technical Specifications.
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1.0 PROGRAM DESCRIPTION l
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1.1 Definition of Goals In order to reduce the potential for the initiation and propagation of primary water stress corrosion cracking'(PWSCC) in the Nodel D steam generator tubes, i
Commonwealth Edison is providi'ng the necessary analysis, documentation and licensing effort to support operation at reduced temperatures for all Byron and Braidwood nuclear units. The effort performed will allow for 100% thermal l power operation in the range of hot leg temperatures between 600'F and 618.4*F. This temperature range is known as the " operating window". It should be noted that Byron Unit 1, which is the only unit of the four units
. presently in commercial service, is currently limited by the high pressure turbine to a hot leg temperature approximately 6*F higher than 600'F at 100%
thermal power. The reason for evaluating a minimum hot leg temperature as low as 600'F is because high pressure turbine modifications which would permit full thermal power operation at 600'F are possible and are being considered by
. Commonwealth Edison. j The efforts performed for this program demonstrate the capability of the Byron and Braidwood units to operate at the reduced temperatures on a permanent basis up to the end-of-life of the units. In addition, in order to achieve ,
maximum operating flexibility, the necessary justification to operate over a l range of primary temperatures (the operating window concept referred to above) is provided. It is Commonwealth Edison's intention to select the desired operating temperature within the range on a cycle-to-cycle basis.
i 1.2 Applicable Criteria The quality assurance requirements defined in WCAP-9245, revision 8, interim I
change A, Commonwealth Edison addendum number 1 (March 1985) apply to this program. Equipment reviews and evaluations have been performed in accordance
, l with Westinghouse and industry codes, standards and regulatory requirements applicable to the Byron and Braidwood units per the NSSS contracts and associated change notices as of'the date of the Thot reduction proposal (August 1985).
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J For a description of the applicable licensing criteria refer to Section 4.0.
1.3 Scope , ,
Evaluations and/or analyses have been perforsed to assess the impact of reduced temperature operation on the following systems and components:
l Steam generators Reactor vessel Reactor vessel internals Control rod drive mechanism .
Reactor coolant pumps-Pressurizer Reactor coolant system piping and primary component supports i Balance of plant and auxiliary systems ,
Valves j j
! Turbine-Generator l
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The following safety analyses have been performed to address the effects of )
reduced temperatures:
Large Break FSAR Loss-of-Coolant Accident (LB LOCA)
LOCA hydraulic forces Small Break FSAR LOCA (SB LOCA)
LOCA mass / energy releases to containment subcompartment ~
Non-LOCA FSAR Chapter 15 transients 1
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i 2.0 DETERMINATION OF REDUCED TEMPERATURE
. OPERATING CONDITIONS 2.1 Design Power Capability Parameters Design power capability parameters which were used as a basis for the evaluations / analyses of the effects of reduced temperature operation are defined in Table 2.1-1. Column 1 lists the original design parameters with vessel outlet temperature .(Thot)at618.4*F,ThermalDesignFlow(TDF)and 100% power. Column 2 defines the parameters at the reduced Thot of 600'F with TDF and 100% power. This column represents the most limiting set of parameters which was developed during the Thot reduction program for evaluation. It should be noted, however, that where higher primary {
temperatures were conservative for a given evaluation or analysis, those l analyses which were performed at the original design temperatures (column 1) were not changed, and are to be considered bounding for the purposes of this program. This was done in order to demonstrate acceptable results for the range of operating temperatures bounded by columns 1 and 2.
.The steam generator performance parameters which were used for the LB LOCA analyses incorporate 10% steam generator tube plugging and reduced TDF (reduced by 2% for conservatism). This was done so that LOCA analyses need not be repeated in the event of steam generator tube plugging (up to 10%) or low measured reactor coolent flow. The LOCA analyses were the only safety j analyses performed for this program which specifically incorporate steam generator tube plugging margin. (Similarly, increases in Fg and FAH were addrassed only in the LOCA analyses.) The steam generator parameters used as input to the LOCA analyses are tabulated in section 3.1 of this report.
2.2 Control and Protection System Setpoints l A detailed accounting of the effects of reduced temperature operation on
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- non-safety related control system setpoints has been performed. Changes to these setpoints have been identified to maintain plant response consistent with the original design. ,
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The evaluations performed in Sections 3.1 and 3.4 have verified that
,- modification of protection system setpoints is not necessary to operate within analyzed space at reduced temperatures. ,
2.3 Design Transients Revised NSSS design transients have been generated to account for reduced primary side operating temperatures. These transients have formed the bases against which the system and component safety evaluations in Section 3.5 have been conducted.
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5 e TABLE 2.1-1 COMPARISON OF ORIGINAL AND REVISED DESIGN POWER CAPABILITY REDUCTIM PARAMETERS FOR BYRON AND BRAIDWOOD THOT CURRENT REDUCED DESIGN TEMPERATURES PARAMETERS:
(1) (2) 100 100 NSSSPower(%)
3425 3425 (NWT) '
1 Thermal Design Flow (GPM) 94400 94400 RCSPressure.(psia) 2250 2250 RCSTemperatures('F):
621.7 603.5
' Core Outlet 618.4 600.0 Vessel Outlet Core Average 591.8 572.2 Vessel Average 588.4 569.1 558.4 538.2 Vessel / Core Inlet 558.1 537.9 Steam Generator Outlet Steam Generator:
Steam Temperature ('F) 543.3 522.1 Steam Pressure (psia) 990 827 6 15.13 15.03 SteamFlow(10 lb/hrTotal) 440 440 FeedTemperature(*F)
.00005 .00005
- Approx. Fouling Factor
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- ZeroLoadTemperature('F) 557 557 0181v:1D-021287 5 .
i 3.0' SAFETY EVALUATIONS Chapter 3.0 summarizes the results of all of the NSSS related safety ;
evaluations performed for the Byron /Braidwood T
, hot reduction program. The l purpose of these evaluations is to assess the impact of reducing primary side temperatures on the safety and operation of all four Byron /Braidwood units. 1 For the evaluation of the effects of reduced temperatures on the small and large break LOCA analysis, advantage is taken of state-of-the-art computer f
codes which requires reanalysis of the FSAR cases in their entirety. This !
l' generates greater margin in terms of peak cladding temperature and, along with i several conservative assumptions factored into the analysis, increased operating flexibility.
3.1 Loss-Of-Coolant Accident (LOCA) , j l
['. Both small break and large break LOCA conditions have been reanalyzed utilizing the NOTRUMP and BASH computer codes, respectively. The following
. parameters have been revised as indicated in the new analysis:
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Original Revised TotalPeakingFactor(Fn) 2.32 2.40 Enthalpy Rise Peaking FHetor (FAH) 1.55 1.62 Steam Generator Tube Plugging (uniform) 0% 10%
Safety Injection Flow -
5% reduction ,
K(z) curve -
eliminate third '
line segment Results have indicated increased margin to peak cladding temperature (PCT)
I limits. The reanalysis is presented in Appendix A of this report as marked-up sections of Chapters 6.2.1.5 and 15.6.4 of the Byron /Braidwood FSAR. This is ;
done in order to present the information in a familiar, standard format which will assist in identification of sections affected by reanalysis. In this manner, ease of review and concurrence with the new methodology can be i
. obtained.
It should be noted that the increases in F and g FAH were addressed only in the LOCA analyses. Revisions of these parameters in the Technical Specifications cannot be made until appropriate changes to core limits are
. made.
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l 3.2 Loss of Coolant Accident (LOCA) Hydraulic Forces
.- 3.2.1 Introduction
. The purpose of the LOCA hydraulic forces safety evaluation is to demonstrate that the original LOCA hydraulic forces supplied for the dynamic analysis are j still conservative and applicable for all four Byron /Braidwood units with i respect to the reduction of RCS primary fluid temperatures. To demonstrate i
that the original LOCA hydraulic forces are still conservative, the expected increase in the magnitude of the peak forces due to a temperature reduction !
will be accounted for by the margin that is available through a reduction in the assumed break area.
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Previous analysis of the 144 in2 reactor pressure vessel inlet (RPVI) nozzle j
break for Byron 1 concluded that an offset of increased forces due to l increased water density can be achieved based on break size reduction. A l
2 reduction of the RPVI break from 144 in to ( ]a,c in was justified 2
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based on a detailed study of as-built physical restrictions to pipe 2
, dislocation in the event of a break. A break size of [ ]a,c in was l determined to be sufficient to offset the forces due to increased density. In order to verify that the most severe conditions were analyzed, an additional evaluation was performed for Byron 1&2 and Braidwood 1 & 2. The analysis performed on the Byron /Braidwood units, as described below, takes advantage of the leak-before-break methodology.
The following assumptions were made in performing the LOCA hydraulic forces evaluation for Byron 1&2 and Braidwood 1 & 2: 1) There is enough margin in utilizing the leak-before-break methodology to account for the increase in peak forces, so a complete dynamic analysis will not have to be performed. 2)
Only one break location will be analyzed to show the effects of decreasing the primary temperature. The break location chosen was the accumulator branch line.
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3.2.2 Method of Analysis I
. The analysis to determine the effects of lowering RCS temperatures involved a
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break of the accumulator line at reduced temperatures in order to establish a I 2
sensitivity to the 144 in RPVI nozzle break at original operating temperatures.
The comouter codes used to evaluate the postulated LOCA were MULTIFLEX (Version 1.0), LATFORC, and FORCE 2.(1) MULTIFLEX calculates the thermal hydraulics of the RCS due to a postulated LOCA. LATFORC takes output from MULTIFLEX and calculates the horizontal forces on the core barrel and reactor j pressure vessel. FORCE 2 takes output from MULTIFLEX and calculates vertical forces on the reactor internals.
1 The following will be used as acceptance criteria to verify that the original LOCA hydreulic forces are still applicable to RCS reduced temperature
. operation: The peak forces resulting from the accumulator line break at )
reduced temperature operation must be lower than the peak forces produced by a )
. RPVI nozzle break at current operating temperatures.
1 3.2.3 Results j Results for the LOCA hydraulic forces show the calculated peak forces for the accumulator line break at reduced temperatures are ( )%a,c lower than the peak LOCA hydraulic forces produced by a 144 in2 RPVI nozzle break at current operating temperatures.
In conclusion, on the basis of this evaluation, the use of the leak-before-break methodology compensates for the increased peak forces from reducing RCS primary temperatures. Finally, from the evaluations performed for all four units, it is concluded that the original Dynamic Analysis (2) for Byron 1 & 2 and Braidwood 1 & 2 is acceptable for the new reduced
,' temperature operating conditions from the standpoint that the peak forces are not increased.
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3.2.4 References
, 1. "MULTIFLEX, A FORTRAN-IV Computer Program for Analyzing Thermal-Hydraulic-Structure System Dynamics," WCAP-8708-PA-VI, September 1977, Westinghouse Proprietary.
- 2. " Dynamic Analysis of Reactor Pressure Vessel for Postulated Loss-of-Coolant Accidents - Byron /Braidwood Power Stations," WCAP-8939, August 1977, Westinghouse Proprietary.
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3.3 Subcompartment Mass / Energy Release l 3.3.1 Introduction i I
This evaluation concerns subcompartment. analysis, which determines the I differential pressure response for the reactor cavity, upper pressurizer
- cubicle, loop compartments and steamline pipe chase. Since the subcompartment )
pressures are strongly dependent on the mass and energy release rates, an evaluation of the postulated pipe breaks has been performed.
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Short term mass and energy release rates tend to increase when operating temperatures are lowered. The critical flow tnrough the pipe break increases for subcooled blowdown when temperatures decrease (i.e., subcooling increases). Since the proposed operating loop temperatures for the T hot reduction program are lowered, an evaluation of the short term mass and energy release rates was performed. The operating parameters listed in Table 2.1-1 for reduced temperatures have been used.
The current subcompartment analysis design basis pipe breaks, which are analyzed in FSAR Section 6.2.1.2, are as follows: 1) reactor cavity -
2 150 in2 reactor vesssi inlet nozzle (144 in break plus conservatism), 2) upper pressurizer cubicle - double ended spray line rupture, 3) loop
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- l. compartments - double-ended hot leg and double-ended cold leg ruptures and 4)
I steamline pipe chase - double-ended main steam line rupture. These analyses
. have been reevaluated based on the use of leak-before-break for Byron 1&2 and Braidwood 1&2 for all reactor coolant loop piping.
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i 3.3.2 Reactor Cavity
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. For the reactor cavity, the postulated reactor vessel inlet break area is l eliminated based on use of leak-before-break methods. The reduction in l temperature will not have any effect on the reactor cavity design limits since these were based on a significantly larger break area than is now postulated.
3.3.3 Upoer Pressurizer Cubicle For the upper pressurizer cubicle, the postulated double ended spray line l rupture area is reduced to 74% of the break flow area assumed in the design basis analysis because of actual pipe flow area. The design basis analysis l was based on the following nominal flow areas, not acco.unting for pipe wall l 2
thickness: 1) 6-inch spray suction line (28.27 in ) and 2) 4-inch spray line connection to the pressurizer (12.57 in2 ). The actual break area i expected is based on the 6-inch schedule 160 and 4-inch schedule 160 pipe that
. is installed in the plant. This consideration reduces the break areas to i 21.15 in2and 9.28 in 2for the 6-inch and 4-inch pipe sections, respectively. The estimated increase in subcooled break flow due to lowering the loop temperatures would be { Ja,c %. The net reduction in peak mass and energy release rates is to approximately 78% of the design basis analysis value. ;
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3.3.4 Loop Compartments
. For the loop compartments, the double-ended Hot Leg and Double-ended Cold Leg postulated ruptures can be eliminated by the consideration of
. leak-before-break. The limiting break size for subcompartment design will then be the largest auxiliary line connected to the reactor coolant loop piping. The largest pipes are a 14 inch diameter pressurizer surge line for the hot leg and a 10 inch diameter Safety Injection line for the cold leg.
For the hot leg, a four loop plant analysis for a plant similar in operating temperature and loop layout to Byron 1&2 and Braidwood 1&2 has shown that a
[ Ja,c% increase in peak mass and energy release rate would be expected for a 22*F reduction in hot leg temperature. The net effect of the reduction.in hot leg temperature and the elimination of the double-ended hot leg break is a significant reduction in peak mass and energy relese due to the significant reduction in maximum break area. l 1
For the cold leg, an estimated [ ]a,c% increase in peak mass and energy release rate would be expected due to the lowering of the vessel inlet temperature. The net effect of the reduction in temperature and the
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elimination of the double-ended cold leg break is a significant reduction in peak mass and energy release due to the significant reduction in maximum break area.
3.3.5 Steamline Pipe Chase The double-ended main steamline rupture data in FSAR Section 6.2.1.2 will remain limiting when the changes to system parameters for T hot reduction are considered. The FSAR results for the steamline pipe chase remain unchanged.
See Section 3.4.3.15 of the report for further information.
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3.3.6 Conclusion
+ The design basis mass and energy release rates were reviewed based on the leak-before-break methodology that is part of the design basis for Byron 1&2
- i. Braidwood 1&2. This methodology eliminates large loop pipe breaks from consideration for subcompartment mass and energy release. The large pipe breaks considered previously in the subcompartment design are more limiting I than the auxiliary line breaks remaining when leak-before-break is considered even with a significant reduction in loop temperatures.
l The postulated reactor vessel inlet break is eliminated based on i leak-before-break methodology. The postulated double-ended spray line break area is reduced by considering the actual 4-inch and 6-inch schedule 160 pipe flow areas. . The postulated primary coolant loop breaks in the loop compartments are eliminated based on leak-before-break and the remaining auxiliary line breaks are of significantly less area.
In all cases, the justifiable reduction in postulated break area more then
,~. offsets the effect of increased subcooled blowdown. Since the mass and energy release rates are the driving force for the subcompartment pressure cransient, a reduction in peak release rates leads directly to lower subcompartment peak pressures. This indicates that the FSAR design basis analysis is still
- limiting when considering changes in initial plant conditions due to T hot reduction in conjunction with reductions in postulated break areas. For the postulated rupture of the main steamline, the FSAR design basis analysis envelopes the results for operation at reduced T hot. Thus, operation at reduced T hot is justified based on an evaluation of subcompartment analyses for postulated primary and secondary pipe ruptures inside containment.
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- 3.4 Non-LOCA Transients 1
- 3.4.1. Introduction .
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i This section summarizes the impact of a primary system temperature reduction l on the Byron and Braidwood 'FSAR Chapter 15 non-LOCA Accident Analyses. The f l
purpose of the proposed temperature reduction is to minimize steam generator.
tube corrosion and, consequently, the hot leg temperature may be as low as 600*F. With this reduction in temperature, new design operating parameters have been developed. Therefore, the impact of the new design conditions on the design basis accident analyses must be addressed. This evaluation will also permit. operation with a nominal average temperature between the current design (588.4*F) and the new temperature reduction design (569.1'F), to be specified on a cycle by cycle basis.
For the majority of the non-LOCA transients, an evaluation of the impact of the T hot reduction has been made. This is presented as Section 3.4.3. The ,
events explicitly reanalyzed for the Thot reduction are presented in Section j 3.4.4. All evaluations and analyses employ conservative assumptions l consistent with those used in the FSAR II) unless explictly noted for each reanalysis. This evaluation assumed an FAH of 1.55 and an Fg of 2.32.
The purpose of this evaluation is to show that either the current analysis of record (FSAR) is bounding er, if reanalysis is required, that all applicable j
acceptance criteria for that event are satisfied.
1 The parameters modified by the Thot reduction that have an impact on the non-LOCA transients are: 1 A) A 19.3'F reduction in nominal RCS average temperature.
. B) A corresponding 20.2'F reduction in the nominal core inlet temperature.
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C) A 163 psi reduction in nominal secondary steam pressure.
D)
A 5% Narrow Range Span (NRS) reduction in nominal steam generator water level. ,
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i 3.4.2. Code Description A11'the events reanalyzed in Section 3.4.4 used the LOFTRAN computer code.
LOFTRAN (2) is a digital computer code, developed to simulate transient-behavior in a multi-loop pressurized water reactor system. The program simulates the neutron kinetics,, thermal-hydraulic conditions, pressurizer, steam generators, reactor coolant pumps, and control and protection system 3
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operation. The secondary side of each steam generator utilizes a homogeneous saturated mixture for the thermal transients.
l 3.4.3. Non-LOCA Transients Evaluated The T het reduction will lower the initial average temperature assumption in all full power transients by 19.3'F, and there is no restriction on reactor L- operation. The Improved Thermal Design Procedure (ITDP) methodology (3) is utilized in the analysis of Condition 11 and Condition III DNB events for Byron /Braidwood. With no change in RCS pressure, NSSS power, or flow and a reduction in temperature, any DNBR values calculated for the Thot reduction l
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will not be less than those generated for the FSAR. This forms the basis for the evaluations of all DNB transients. Furthermore, all transients or cases analyzedatno-loadconditions(hotzeropower)arenotimpactedbytheT hot i reduction, since design conditions at hot zero power are not changed.
The only concern presented by the Thot reduction is the possibility of pressurizer overfill for Condition 11 events. Water relief through the pressurizer is unacceptable for.any Condition 11 transient since this would !
violate the following specific criterion for events of moderate frequency, as stated in the NRC Standard Review "An Planincident (4): of moderate frequency should not generate a more serious plant condition without other faults occurring independently." The 19.3*F decrease in primary temperature '
increases the density of the coolant. For long term heatup transients, a greater potential for filling the pressurizer exists, due to the greater
". initial RCS mass.
A calculation was performed to generate the Overtemperature Delta-T (OTDT) trip setpoint lines applicable to the conditions of the T hot reduction and the lowering in the reference average temperature used in the OTDT and I 14 .
[ -_- - - _ _ - _o m v.io-o ins 7
.- OverpowerDelta-T-(OPDT)setpointequations. The validity of the steam generator safety valve line of the core thermal limits was checked as well as the applicability of the current OTDT and OPDT coefficients defined in the Technical Specifications. The results of the calculation show that the core limits are still protected and the current setpoint coefficients do not need to be changed. The results of this calculation are presented in Figure 3.4-1.
This evaluation assumes that the reference average temperatures.used in the OTDT and OPDT setpoint equations are rescaled to the Thot reduction nominal average temperature. It also assumes recalibration of the NIS excore l
detectors to compensate for the increase in coolant density at a lower l
temperature. These two items should be performed each time the cycle average temperature is changed. The effects of the Thot reduction were also l
evaluated for any potential impact on fuel temperature calculations. It was reduction is bounded by the current FSAR concluded that the Thot calculations.
3.4.3.1. Feedwater System Malfunction This accident is described in Section 15.1.2 of the FSAR. The hot zero power (HZP) case is unaffected because the no load temperature remains unchanged.
For the full power case, this is a DNB transient analyzed with the ITDP methodology. Since the initial steam generator water level will be lower for the T hot reduction, this will delay the actuation of the high-high water level signal. However, with respect to Figure 15.1-2 of the current '
Byron /Braidwood FSAR,"the DNBR value is relatively constant before reactor trip and well above the limit value. Thus, delaying the high-high signal which generates a turbine trip, and consequently reactor trip, will not significantly impact the results. Since the initial temperature is lower than the current analysis, the minimum DNBR calculated would remain above the limit
. value. This transient is a cooldown event and there would be no potential for filling the pressurizer. Therefore, the Thot reduction will have no adverse impact on the event. The conclusions presented in the FSAR remain valid.
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q 3.4.3.2. Excessive Increase in Secondary Steam Flow This event is described in Section 15.1.3 of the FSAR. This is a Condition II
- DNB cooldown transient analyzed with ITDP and is similar to a Feedwater System Malfunction (cooldown). Based on the discussion in the previous section (3.4.3.1) of the reduction in nominal temperature, there will be no adverse impact on this event. The conclusions presented in the FSAR remain valid. l 3.4.3.3 Inadvertent Opening of a Steam Generator Relief or Safety Valve L I This accident is described in Section 15.1.4 of the FSAR. This transient is explicitly analyzed at hot zero power conditions and, therefore, is not impacted by the TMt reduction. The conclusions presented in the FSAR remain valid.
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3.4.3.4 Steam System Piping Failure
- This accident is described in Section 15.1.5 of the FSAR. As reported in the FSAR, the most limiting steamline break case is analyzed at hot zero power.
Since the noload conditions are not changed, the analysis is not impacted by the T bt reduction. The conclusions presented in the FSAR remain valid.
3.4.3.5 Turbine Trip This accident is described in Section 15.2.3 in the FSAR. This is a Condition II DNB event analyzed with ITDP. As detailed in Section 3.4.3, the Thot reduction has no adverse impact on DNB. The initial increase in pressurizer water volume terminates with reactor trip. The times of reactor trip for the four cases analyzed are not adversely affected by the Thot reduction since the reactor trip is generated by a high pressurizer pressure signal.
. Therefore, there is no potential for filling the pressurizer. The conclusions presented in the FSAR and the Overpressure Protection Report remain valid.
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I 3.4.3.6 Decrease in Reactor Coolant System Flow Rate These accidents are described in Section 15.3 of the FSAR and include a)
Partial loss of flow, b) Complete loss of flow, c) RCP shaft seizure, and d)
RCP shaft break.
These transients are analyzed similarly and can be ' evaluated together. The loss of flow transients are of short duration (ten seconds),
trip on low flow and undervoltage (which are not impacted by the Thot reduction), and are essentially DNB transients. The initial conditions of each event define the severity of the results, and reactor trip occurs within a few seconds, terminating the transient before pressurizer. level significantly increases. Since the T hot reduction will lower the inlet l
temperature, less limiting DNBR values would be calculated for any reanalysis. Thus, the results shown in the FSAR are limiting and the conclusions remain valid. l l
3.4.3.7 Uncontrolled RCCA Bank Withdrawal From Suberit8cel ,
This accident is described in Section 15.4.1 of the FSAR. This transient is analyzed at hot zero power and is not impacted. The conclusions presented in the FSAR remain valid. ,
3.4.3.8 Uncontrolled RCCA Bank Withdrawal at Power l
This accident is described in Section 15.4.2 of the FSAR. This is a Conditio 11 DNB event analyzed with ITDP. As detailed in Section 3.4.3, the Thot "
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reduction has no adverse impact on DNB. With respect to pressurizer overfill, l l
previous analyses demonstrated that the part power (60% Rated Thermal Power) l l case was limiting. It was concluded that pressurizer overfill was not a concern for all insertion rates at 60% RTP. Thus, the conclusions presented in the FSAR remain valid. ,
. . 4 3.4.3.9 RCCA Misoperation This set of accidents is described in Section 15.4.3. of the FSAR. These transients are analyzed with the ITDP methodology with minimum DNBR as one of the acceptance criteria. The initial assumptions for these events are sisicio-0212:7 U s
E limiting at nominal full power conditions and, as previously noted, DNB events benefit from a reduction in nominal temperature. Furthermore, for these events, peaking factors impact the analysis of these transients. The limiting RCCA misoperation transient is the statically misaligned RCCA (Section 15.4.3.2.a.3.) To address the effects of the T reduction on peaking hot factors, the most reactive rod was reanalyzed for the misaligned RCCA event, and the peaking factors remain bounded by the maximum allowable value. The DNB design basis is met, and the conclusions presented in the FSAR remain valid. I 3.4.3.10 startup of an Inactive RCP at an Incorrect Temperature l
This accident is described in Section 15.4.4 of the FSAR. This transient is analyzed for Byron /Braidwood starting from a suberitical mode with loop stop valves closed in one loop and is not impacted by the Thot reduction. The i . conclusions p, resented in the FSAR remain valid.
6 3.4.3.11 Chemical and Volume Control System Malfunction (Boron Dilution)
This accident is described in Section 15.4.6 of the FSAR. The only case i l potentially affected by the Thot reduction is the Dilution During Full Power !
Operation. This event is analyzed to determine the time available for operator action before a return to criticality. Significant time exists for the operator to terminate the dilution before the pressurizer fills. Only the changes in reactivity insertion due to the changes in temperature need to be ,
reduction. An evaluation has been performed at addressed for the Thot 1 beginning-of-life (BOL) to address the changes in the initial boron
' concentration for criticality and the constant boron worth as a function of the decrease in temperature. The small change (10 ppm) in the initial bcron concentration is still bounded by the conservative assumptions used in the ,
FSAR analysis (1600 ppm initial concentration). There was no change in constant boron worth. Therefore, the results provided in the FSAR are unchanged and the conclusions presented remain valid. .
O f '
18 ,
oisir.10-ortzs7
_ _ _ _ _ _ _ = _ _ _ _ _ - _ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ - _ _ _ _ _
i 3.4.3.12 RCCA Ejection The het zero These accidents are described in Section 15.4.8 of the FSAR.
~
power cases are not impacted by the Thot reduction. The full power cases f benefit from a reduction in initial temperature making the fuel rod thermal transients less severe.
The severity of the results are defined by initial conditions and the ejected rod. parameters assumed (which are not changed).
The cases presented in the FSAR are still bounding and the conclusions remain valid.
3.4.3.13 Increase in Reactor Coolant Inventory The limiting These accidents are described in Section 15.5 of the FSAR.
transient is the Inadvertent ECCS Actuation and is analyzed at nominal full The transient is analyzed with the ITDP methodology with power conditions.
minimum DNSR as the acceptance criterion. Nuclear power decreases immediatei due to boron injection, but steam flow does not decrease until the turbine throttle valves are fully open. The mismatch between turbine load and nuclear
~
power causes pressurizer pressure and water vclume to decrease (a cooldow ]
l transient). As previously noted, DNB events benefit from a reduction in nominal temperature. Thus, the conclusions presented in the FSAR remain valid.
3.4.3.14 Inadvertent Opening of a Pressurizer Safety or Relief Valve The transient is l This event is described in Section 15.6.1 of the FSAR.
~
analyzed with the ITDP methodology with minimum DNBR as the acceptance f
criteria and the initial assumptions for this event are limiting at nominal ]
j full power conditions. As previously noted, DNB events benefit from a #
The increase in pressurizer water volume is reduction in nominal temperature.
terminated by reactor trip on low pressurizer pressure which would not be Thus, the conclusions presented in the FSAR
. changed by the Thet reduction. 1
"~ remain valid.
19 ot81v 1D-o21287
.- 3.4.3.15 Steamline Break Mass & Eneroy Releases
.- This section will discuss the impact of the Thot reduction on the steamline break mass and energy releases inside containment and the superheated niass and energy releases outside containment.
The mass and energy releases fbr the inside containment analyses are discussed '
in Chapter 6 of the FSAR. For the inside containment analysis, blowdown data was generated at 0, 30, 70, and 100% power. The 0% power analysis will not change since the Thet reduction does not affect conditions at het zero power. For the at power analyses, the initial primary temperatures and secondary steam pressures will be lower than those used in the FSAR analyses.
The assumptions on the initial conditions are taken to maximize the mass and total energy released. Since the mass blowdown rate is dependent on steam
- pressure and steam pressure will be reduced, the initi6l mass blowdown rate will be lower. With the nominal steam generator mass reduced for the Thot reduction, the total mass released through the break will be lower than calculated in the current analysis. This combined with lower temperatures
~
- will result in a lower integrated mass and a lower integrated energy released into containment. Thus, the blowdown data used in the FSAR remains bounding.
The superheated mass and energy release analysis outside containment was performed to address equipment qualification issues. The study performed for Byron /Braidwood analyzed 70% and 100% power cases. For the outside containment analyses, initial conditions are assumed in order to obtain the
[
Ja,e With a reduction of initial temperature and steam pressure,
[ la c Since the low-low steam i generator water level trip signal is not being changed, [ l 1
3"'C Safety injection and steamline isolation l Ja,e Therefore, l l ', on low steamline pressure [
[
]"'C are expected if all these conditions l Ja,c, the current were considered. [
mass and energy releases do not have to be revised and remain bounding, j c181mto-0212s7 20 .
L 3.4.4 Non-LOCA Transiegg Rsanalyzed .
The only concern presented by the Thot reduction is the possibility ef pr.essurizer overfill for Condition II events. The reduction in nominal
-average temperature at nominal conditions-is 19.3'F. This decrease in primary temperature increases the density of the coolant.and, during any heatup transient, increases the potential for filling..
The limiting transients for pressurizer overfill are the Loss of All Non-Emergency AC Power and the Loss of Normal Feedwater. For' comparison 3
purposes, the pressurizer fills at 1865 ft .
Since changes are made to the. initial steam generator water level and nominal
' steam pressure is significantly decreased, the Feedline Break event may be impacted. Although these two changes are somewhat competing, the final acceptance criteria might not be met. Therefore, the Feedline Break transient, both with and without offsite power, is also reanalyzed.
3.4.4.1 Loss of Non-Emergency AC Power (Less of Offsite Power)
~
This accident is described in Section 15.2.6 of the FSAR.
Nethods To ensure that the pressurizer does not overfill, the direction of conservatism in the initial conditions was examined. When this transient is analyzed for the FSAR, the conservatism in the initial cenditions were taken to maximize stored energy and reduce decay heat removal. For the T hot Since reduction, a sensitivity was performed to maximize pressurizer insurge.
a maximum water mass in the primary system, given a constant volume, is I
desired, the' uncertainties and errors on the initial average temperature were
~ subtracted from the nominal value. The pressurizer pressure control system was assumed to be available as well. The transient was also reanalyzed consistent with the FSAR analysis in that initial power is at 102% of the EngineeringSafetyFeatures(ESF) designating. The assumptions made in this analysis are detailed below:
o 21 als1v:10-o21287 J
l l
l l
- a. The plant is operating at 102% of ESF Rated Thermal Power (3579 MWt).
- b. An initial average temperature minus uncertainties is assumed.
- c. The initial pressurizer water level is assumed to be at the nominal setpoint plus uncertainties.
- d. Pressurizer Power Operated Relief Valves (PORVs) are assumed operable to maximize pressurizer water volume,
- e. The maximum pressurizer spray flow rate is assumed to maxicize pressurizer water volume,
- f. Initial steam generator water level is 61% of narrow range span.
- g. An auxiliary feedwater flowrate of 612 gpm is delivered to four steam i
j generators.
3
- h. An auxiliary feedwater line purge volume of 50 ft was assumed, f
)
All remaining assumptions are consistent with the analysis presented in the I
. FSAR. Power is assumed to be lost to the reactor coolant pumps (RCP's) at the time of rod motion. The reanalysis used LOFTRAN (2) to obtain the plant transient following a loss of offsite power.
Results The transient response of the RCS following a loss of AC power is shown in Figures 3.4-2 and 3.4-3. The calculated sequence of events is listed in Table 3.4-1. The plot of pressurizer water volume clearly shows that the pressurizer does not fill.
i Conclusions The natural circulation capability of the RCS is sufficient to remove decay Thus, a loss heat following a RCP coastdown to prevent fuel or clad damage.
of AC power does not adversely affect the core, the RCS, or the steam system l ,' and the auxiliary feedwater capability is sufficient to preclude water relief through the pressurizer relief or safety valves. Therefore, since the pressurizer does not fill, the conclusions presented in the FSAR remain valid.
I' 22 c1:1v.10-o212s7 l
i 3.4.4.2 Loss of Normal Feedwater Flow
^
This accident is described in Section 15.2.7 of the FSAR. This transient is essentially the same as the Loss of AC Power described in the previous section, the difference being the lack of a loss of offsite power coincident with reactor trip. ,
Nethods Consistent with the Loss of AC Power reanalysis, the uncertainty and errors on the initial average temperature are subtracted from the nominal value. The assumptions made in this analysis are detailed below:
- a. The plant is operating at 102% of ESF Rated Thermal Power (3579 MWt).
- b. An initial average temperature minus uncertainties is assumed.
- c. The initial pressurizer water level is assumed to be at the nominal setpoint plus uncertainties. j I d. Pressurizer Power Operated Relief Valves (PORVs) are assumed operable
- to maximize pressurizer water volume.
- e. The maximum pressurizer spray flow rate is assumed to maximize pressurizer water volume.
- f. Initial steam generator water level is 61% of narrow range span.
- g. An auxiliary feedwater flowrate of 612 gpm is delivered to four j l
intact steam generators.
3
- h. An auxiliary feedwater line purge volume of 50 ft was assumed. .
All remaining assumptions are consistent with the analysis presented in the FSAR. The reanalysis used LOFTRAN (2) to obtain the plant transient following a loss of normal feedwater flow.
. Results The transient response of the RCS following a loss of normal feedwater flow is
~
shown in Figures 3.4-4 and 3.4-5. The calculated sequence of events is listed in Table 3.4-2. The plot of pressurizer water volume clearly shows that the pressurizer does not fill.
c1 1<.10-o212s7 23
i v
Conclusions ]
k 2 The reanalysis shows.that a loss of normal feedwater 'does not adversely affect the core, the RCS, or the steam system and the auxiliary feedwater system.
capability is sufficient to preclude water relief through the pressurizer j
relief or safety valves. Therefore, since the pressurizer does not fill, the
- conclusions presented in the FSAR remain valid.
3.4.4.3 Feedwater System Pipe Break-This accident is described in Section 15.2.8 of the FSAR and is a Condition ,
event. This transient is analyzed both with and without offsite power available, the difference being the loss of offsite power coincident with reactor trip.
Methods The assumptions made in this analysis are detailed below:
- a. The plant is operating at 102% of ESF Rated Thermal Power (3579 MWt).
- b. An initial average temperature plus uncertainties is assumed.
- c. The initial pressurizer water level is assumed to be at the nominal setpoint plus uncertainties.
- e. Initial steam generator water level is 61% of narrow range span plus 5% HRS for the faulted loop and minus 5 % for the intact loops.
f.
Reactor trip is assumed to be actuated when the water level reaches 13% NRS in the faulted loop.
- g. An auxiliary feedwater flowrate of 459 gpm is delivered to three l
i ~ intact steam generators. 3
- h. An auxiliary feedwater line purge volume of 50 ft is assumed. .
All remaining assumptions are consistent with the analysis presente? in the FSAR. The reanalysis used LOFTRAN III to obtain the plant transient following a feedwater line rupture. For the case without offsite power available, power is assumed to be lost to the RCPs at the time of rod motion.
24 E-----_d1 Ply:1D-o212s7 .
4 1
Results )
i l
- j The transient response of the RCS following a feedwater line rupture with i offsite power available is shown in Figures 3.4-6 through 3.4-9. The j calculated sequence of events is listed in Table 3.4-3. The case without {
offsite power available is shown in Figures 3.4-10 through 3.4-13, and the sequence of events is given as Table 3.4-4. For both cases, the plots of faulted and intact loop temperatures show that there is no bulk boiling in the hot leg. Therefore, the core remains covered and in a coolable geometry.
Conclusions 1
The reanalysis shows that for a postulated feedwater line rupture, the assumed auxiliary feedwater system capacity is adequate to remove decay haat, prevent overpressurization of the RCS, and prevent the uncovering of the et.e.
Therefore, since the core remains covered, the conclusions presented in the FSAR remain valid.
3.4.5 Steam Generator Sensitivities The Byron and Braidwood Units have different Model D steam generators. Byron and Braidwood Unit I have D4 steam generators while Byron and Braidwood Unit 2 have D5 steam generators. Previous analytical work has always had to address the difference in the steam generators. The significant difference between the D4 and D5 steam generators is the level measuring system (Narrow Range
" Span- NRS). A D4 has a NRS of 233 inches while a D5 steam generator has a more compact NRS (120 inches).
Calculations were performed using the nominal conditions indicative of the T reduction program, and these calculations concluded that for all l hot applications, the D4 steam generator was limiting. Thus, all analyses for the
. T reduction were based on the D4 model. Hence, this evaluation bounds
, hot all four Byron and Braidwood units.
25
3.4.6 Conclusion
- Based on the analyses and evaluations presented in this section, there are no
- outstanding issues for the FSAR Chapter 15 non-LOCA events that would prohibit Byron 1 & 2 and Braidwood 1 & 2 from operating under the conditions defined by the T hot reduction program. As previously discussed, the impact of a reduction in temperature was not as limiting as the current design conditions for many FSAR transients. The only impacted events were heatup transients ,
where the possibility of filling the pressurizer for a Condition 11 event existed. The analyses presented concluded that pressurizer overfill is not a concern for these events. This evaluation also concluded that no changes to reactor trip setpoints specified in the Technical Specifications are required. lq 3.4.7 References !
- 1. Byron /Braidwood Final Safety Analysis Report, Amendment 47, April 1986.
q
- 2. Burnett, T.W.T., et al., "LOFTRAN Code Description,' WCAP-7907-P-A
, l (Proprietary), WCAP-7907-A (Non-Proprietary),
f April 1984. i
- 3. Chelemer, H., et al., " Improved Thermal Design Procedure," WCAP-8557-P r
(Proprietary), WCAP-8568 (Non-Proprietary),
July 1975, i
- 4. NUREG-0800, " Standard Review Plant for the Review of Safety Analysis Reports for Nuclear Power Plants - LWR Edition", Revision 1, July 1981.
l
)
O 1
TABLE 3.4-1 TIME SE00ENCE OF EVENTS FOR A LOSS OF NON-EMERGENCY AC POWER EVENT TIME (SEC51 Main feedwater flow stops 10 Low-low steam generator water 52 level trip setpoint reached Rods begin to drop 54 Reactor coolant pumps 54 begin to coastdown
- One motor driven auxiliary 113 feedwater pump starts and supplies four steam generators
- 310 Peak water level in the pressurizer occurs 1 Cold auxiliary feedwater is 259 l delivered to the steam generators ,
l Core decay heat decreases down to the 325 ,
l level of auxiliary feedwater heat removal capability Qe l -
e 27 otsiv:1o-o212:7 k-- -_-_- - - -
I TABLE 3.4-2 l
~
TIME SE00ENCE OF EVENTS FOR A LOSS OF NORMAL FEEDWATER FLOW J
l EVENT TIME (SECS) l 1
Main feedwater flow' stops 10 l Low-low steam generator water 52 level trip setpoint reached y Rods begin to drop 54 Peak water level in the 113 ,
pressurizer occurs !
One motor driven auxiliary 113 feedwater pump starts and supplies four steam generators Cold auxiliary fee'dwater is 259
'- delivered to'the steam generators 1 Core decay heat plus pump heat 600 decreases down to the level of auxiliary feedwater heat removal capability i
l l
f '-
l l
28 i, 0181v:iD-021287
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i.
TABLE 3.4-3 i
. TIME SEQUENCE OF EVENTS FOR A FEEDWATER LINE RUPTURE WITH OFFSITE POWER AVAILABLE i
EVENT TIME (SECS) ,
Main feedwater line break occurs 10 Low-low steam generator water 31 level trip setpoint reached in the faulted steam generator Rods begin to drop 33 l l
One motor driven auxiliary 92 feedwater pump starts and supplies three intact steam generators !
- Cold auxiliary feedwater is 238 delivered to the intact steam generators
- 1 Low steamline pressure setpoint 420 1 reached in the faulted loop All main steamline isolation 428 valves close l 1
Steam generator safety valve setpoint 982 l reached in the intact loops , l
- Pressurizer water relief begins 1744 Core decay heat plus pump heat 5000 decreases down to the level of auxiliary feedwater heat removal capability 1
= l l
l e
cisiv.to-0212s7 29 ;
i i
i L
TABLE 3.4-4 )
TIME SEQUENCE OF EVENTS FOR A .
f l FEEDWATER LINE RUPTURE WITHOUT OFFSITE POWER AVAILABLE l
EVENT TIME (SECS) 1 Main feedwater line break occurs 10 i
Low-low steam generator water 31 level trip setpoint reached in '
the faulted steam generator Rods begin to drop 33 Reactor coolant pumps 54 I
. l begin to coastdown i
^
- One motor driven auxiliary 92 feedwater pump starts and supplies three intact steam generators Cold auxiliary feedwater is 238 delivered to the intact steam generators ;
i Low steamline pressure setpoint 508 reached in the faulted loop All main steamline isolation 516 valves close Steam generator safety valve setpoint 1378 reached in the intact loops Core decay heat decreases down to the 2000 level of auxiliary feedwater heat removal capability 0181v:1D-021287 30 ,
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. for a Feedwater Pipe Rupture with Offsite Power 37 0181v:1D-020187
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1 0181v:1D-D20187 42
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3.5 Systems and Components Evaluation 3.5.1 General . .
Evaluations were performed for the following systems and components to determine the impact of the reduced temperature conditions on them:
Steam generators Reactor vessel Reactor vessel internals Control rod drive mechanisms Reactor coolant pump motors
- Pressurizer RCS piping and supports Balance of plant and auxiliary systems Valves
. Turb.ine generator The primary emphasis of these evaluations was placed on confirming the structural integrity of the systems and components at the reduced temperature i
conditions.
l 3.5.2 Steam Generators I
a) Structural Considerations A comparative structural evaluation of the critical components of the steam generator was used to address safe operation at the reduced T hot c ndition.
The analysis was performed according to the requirements of the ASME B&PV l
Code,Section III, Subsection NB. The critical components considered were the tubesheet and shell junction, the divider plate, the feedwater nozzle, and the
) ,-
steam generator tubes. It was assumed that the primary side pressure remains l
essentially unchanged at 2250 psia while the reactor vessel outlet temperature l ,
is reduced by 18.4*F to 600.0*F, resulting in a reduction in steam generator I
secondary side steam pressure.
44 ,
0181v:1D-c21287
The evaluation was based on the results of previous structural analyses of the Model D4 and D5 steam generators. These results were scaled by a secondary side pressure ratio which was used to determine the total stresses (pressure and thermal) for the tubesheet, tubesheet to shell junction and divider plate.
For the tubesheet evaluation two locations, the tubesheet centers and the tubesheat and shell junction, were considered. The results of the evaluation show that the maximum stress intensities expected for those locations remain below the allowable stress intensities for all conditions analyzed in the NSSS design transients. The fatigue usage factors are also acceptable.
For the steam generator divider plates, stress intensity ranges are close to the 35 , limit. Thus, the divider plate fatigue evaluations were performed using elastic plastic analysis as stipulated in NB-3228.4(c). The resulting
. fatigue usage factors in the critical areas, i.e., the divider plate fillet and drain hole, are acceptable.
In considering the feedwater nozzle and flow restrictor assemblies, only the feedwater nozzle was evaluated since it was found to have the largest fatigue usage factor in the previous evaluations. In an analysis of the critical transients, it was found that all steam temperatures, saturation pressures and feedwater pressures for the reduced Thot were equal or less than the corresponding values for the original Thot. Therefore, the fatigue usage !
factor calculated previously is considered applicable for the reduced Thet condition.
In evaluating the steam generator tubes, it was determined that the maximum stresses for the critical loading conditions for the reduced temperatures were j enveloped by those for the normal case. In both cases, the fatigue usage l factors are negligible. In addition, the qualitative evaluation of preheater l l
tube vibration for the flow condition at 100% power shows that vibration potential for reduced Thot was no more severe than that for the original
- T hot condition. l In conclusion, the critical components of the steam generators are found to i
satisfy the requirements of the ASME B&PV Code,Section III, Subsection NB for the reduced T hot condition.
- mumu.umstm9 @ _j
b) Tube Corrosion Considerations The T hot reduction program has been evaluated with regard to steam generator
. degradation using a corrusion algorithm which provides a quantitative indication of steam generator corrosion susceptibility. Types of steam generator degradation considered are denting, OD and ID initiated. stress corrosion cracking, and pitting. The results of this study indicated that operation at reduced temperatures could lead to a signifiesnt decrer.se in corrosion concern for the steam generators. The operating temperatures wnich were used in the algorithm are as follows:
Original T Target T Vessel Outlet 618.4 F 600.0 F Steam Temperature 543.3 F 522.1 F
- Additional information also considered in the algorithm are individual plant characteristics such as materials of construction for the steam generator and other secondary side components, type of cooling water, and various other design considerations.
l The results for each plant (Byron 1 & 2, Braidwood 1 & 2) can be seen in Figures 3.5-1 through 3.5-4 which are plots of Cumulative Chemistry Operating Experience Factor (CCOEF) versus time. The CCOEF is a measure of the relative corrosion susceptibility of the steam generators. The higher CCOEF values for Byron 1 and Braidwood 1 versus those for Byron 2 and Braidwood 2 result from the differences in materials of construction and design between the Model 04 and Model D5 steam generators. The Byron steam generators show higher CCOEF values than their comparative Braidwood steam generators due to differences in cooling water type and tubesheet joints.
For each steam generator it can be seen that operation at the target
.' temperatures will lead to a significant reduction in CCOEF value at a given time. This reduction could be translated to an increase in operating life for l
- the steem generators. For Byron 1, the original operating temperatures have a CCOEF of 300, reached after approximately seven years. When operating under the reduced temperatures this same CCOEF is not reached until approximately 12 years. This indicates that a significant increase in the life of the steam sistrio-ortza7 46 ,
generator could be realized if operated at th'e lower temperatures when considering the steam generator degradation mechanisms previously mentioned.
Similar scenarios could be constructed for the other three plants as all
~
showed improvement with the reduced temperatures. It should be noted that the results reported here assume that the steam generators are being run with secondary side chemistry that is at least as gocd as that specified by EPRI.
For ID stress corrosion cracking (ID SCC), results indicate the reduction in operating temperatures could also decrease the concern for this form of degradation. The algorithm, which uses an activation energy of 33 kcal/mol for ID SCC was used to study each plant and for each case the results indicated an approximately 33% reduction in propensity for ID SCC.
From the above discussion, it can be concluded that the T hot reduction program is a significant contributor to improved steam generator operability.
. 3.5.3 Reactor Vessel Stress analyses, stress evaluations or stress report reviews were performed on all of the various locations in the reactor vessels which were analyzed in the reactor vessel stress reports, to determine the effects of the Thot reduction parameters and transients. Based on this work, .it is concluded that the operation of the reactor vessels throughout their 40 ' year design life ;
under the conditions of the T hot reduction program is acceptable. All of 1 the stress intensity and usage factor limits of the 1971 edition of Section III of the ASME Boiler and Pressure Vessel Code with Addenda through Summer 1973, to which the Byron 1 & 2 and Braidwood 1 & 2 reactor vessels were designed, are met with incorporation of the Thot reduction program.
The following analyses, evaluations, and reviews were performed:
(1) Finite element stress analyses were performed for the control rod drive mechanism housings and the bottom instrumentation tubes. These analyses show that there is no increase in the maximum ranges of primary plus secondary stress intensity and fatigue usage factors reported previously in the reactor vessel stress reports.
0141v:1D-021287 47
(2) Stress evaluations using a combination of text book calculations, conservative assumptions, and values from previous analyses were performed on the outlet nozzles, inlet nozzles, and main closures of the reactor
- vessels. These evaluations made use of simplified elastic plastic analysis to result in fatigue usage factors less than 1.0 at reduced temperature conditions.-
(3) The stress reports were reviewed to estimate the effect of the That reduction on the vessel shell and the core support guides. These reviews demonstrated that the maximum fatigue usage factors remain less than 1.0. ,
It should be noted that a reactor vessel stress report addendum will be issued to document all of the analyses and evaluations discussed above.
The conclusion of the efforts described above show that the operation of the {
~
Byron 1 & 2 and Braidwood 1 & 2 reactor vessels throughout their 40 year I design life under the conditions of the Thot reduction program is acceptable. .
)
3.5.4 Reactor Vessel Internals In order to address the effects on the reactor vessel internals of lower operating temperatures and higher fluid density due to the Thot reduction, the following analyses were performed.
System thermal hydraulic analyses included: a) evaluation of the effects on core bypass flow, b) pressure drop distribution in the reactor vessel, c) component hydraulic lift forces and, d) pressure relief hole velocities.
Conclusions of these analyses indicate that the effects of the Thot reduction program on the Byron 1 & 2 and Braidwood 1 & 2 units result in acceptable values of core bypass flow, pressure drops, component lift forces and relief hole velocities.
,- i A control rod drop time evaluation indicates that the impact of primary temperature reduction on margin between actual drop time and the technical specification limit for Byron 1 & 2 and Braidwood 1 & 2 with the 17 X 17 Optimized Fuel Assembly (OFA) and full-length hafnium Rod Cluster Control Assembly (RCCA)isinsignificant.
o181v:1D-021287 4B
Component thermal stress analysis was performed to evaluate the impact of
,. revised system design transients due to the Thot reduction. The analysis results show that the fatigue usage factors at the critical locations of the reactor internals are within the ASME Code allowables.
Flow induced vibrational analysis of the reactor internals was performed to evaluate the impact of. higher fluid density due to the T hot reduction.
Results indicate that the stresses and vibrational amplitudes remain small and there is no adverse impact on the strue.tural integrity of the internals due to l flow induced vibrations during the T hot reduction mode of operation. -
An evaluation was performed for the Reactor Vessel Level Measuring System i Upper Probe Housing based on the reduced temperature conditions. All stresses and fatigue usage factors are below the ASME Code allowables. Therefore, I l
there is no impact to this component caused by the temperature reduction for
~
the T hot reduction program.
~
3.5.5 Control Rod Drive Mechanisms I
e The applicable design documentation for Byron 1 & 2 and Braidwood 1 & 2 has been reviewed with respect'to the revised operating parameters and normal, upset, faulted, test and emergency condition transients. A review has been performed of the T hot reduction parameters with respect to the CRDM Generic Stress Report as well as each plant's specific Stress Report and concludes that the operability and integrity of the CRDM are not impacted by the T hot reduction parameters.
3.5.6 Reactor Coolant Pumos 4 An estimate has been made of the changes which can be expected in the Reacter l
. Coolant Pump (RCP) motors as a result of the increased load. Since all of the i meters have slightly different operating points (a result of miner variations !
, in im:eller and loop geometry) this analysis has focused on enanges from the average operating point.
o181v;1D-022EE7 49
The actual coolant temperature reduction seen at the impeller of the RCP is
, 20.2'F. This results in a load increase of approximately 145 HP for each motor. The expected changes in operating parameters resulting from this
. increased load are listed below:
- 1. Speed reduced by 0.38 rpm
- 2. Efficiency assumed unchanged
- 3. Power Factor - assumed unchanged
- 4. Current - increased by 10.7 amps
- 5. KW Input - increased by 116.6 KW
- 6. Stator Temperature Rise - increased by 1.4*C It should be noted that by design the efficiency and power factor of large -
induction motors change little around the rated operating point. Therefore, the assumption of constant values is not unreasonable.
The actual operating point of the motors after the hot loop temperature reduction is still less than the guaranteed design point of 7000 HP. None of the operating parameters has increased (or decreased) beyond allowable limits. The motors are acceptable for use without modification.
It has also been concluded that for the T hot reduction program no change will be required to the existing RCP hardware or spare parts intended for use. l l
3.5.7 Pressurizer The pressurizer components were evaluated for reduced Thot perating l conditions. Initial concerns centered on the acceptability of fatigue usage at components with limited margin available, including the upper shell (spray impingement), the spray nozzle and the surge nozzle. The reason for this concern is the increased change between the pressurizer temperature and that
~
- of the pressurizer spray, which is at cold leg temperature.
Reported fatigue usage at the spray nonle and spray impingement area of the upper shell was reduced from previously reported values, as a result of an I update to the spray flowrate as indicated in the equipment specification o181 v:1D-022687 50
r during the boron concentration equalization event. A more realistic value is assumed for the flow rate which provides some benefit for the transient, consistent with the value used for more recent equipment specifications.
Impact to reported fatigue usage of other pressurizer components was found to be small or negligible, due to the existing conservatism inherent in the simplification process of enveloping groups of transient events with a single event of greater severity.
I No impact to the satisfaction of applicable ASME Code criteria has resulted reduction program.
from the modifications associated with the Thot 3.5.8 Piping and Supports The structural evaluation of the primary loop piping, the primary equipment l nozzles, the primary equipment supports, and the auxiliary line piping have
- been reconciled to the existing analysis. The fatigue evaluation of the primary loop piping and the auxiliary line piping has been performed and
~
- reconciled. The fatigue usage factors for the loop piping were modified slightly upward to accommodate the revised thermal transients. This modification to the usage factors has no effect on the conclusions originally reached in the analysis that indicates acceptable results.
3.5.9 Auxiliary Systems A review of the following auxiliary systems was conducted to determine th's impact of the reduced primary side temperatures and resulting secondary side ;
i parameters.
3.5.9.1 Chemical Volume Control System (CVCS) l The regenerative and excess letdown heat exchangers are designed to function .
~
- adequately at temperatures < 560*F. Since T eold is reduced by '
approximately 20*F to 538.2*F as part of the Thot reduction program, this change provides additional margin to the maximum allowable temperature.
l oisir.to-orirs7 51
l 1 l
3.5.9.2 Safety Injection System (SIS)
- The Safety Injection System is designed to provide appropriate flow from the Refueling Water Storage Tank or Accumulator regardless of RCS temperatures, so the T hot reduction has no impact on the SIS.
3.5.9.3 Residual Heat Removal System (RHRS)
The Residual Heat Removal System initiates at 350'F, regardless of full power f i
RCS temperatures. Therefore, it is not impacted by the temperature change, 3.5.9.4 Auxiliary
- Systems Tanks, Heat Exchancers, Pumos and Valves 1
Auxiliary Systems Tanks The Auxiliary System Tanks were designed in ace.ordance with the design transients specified in Systems Standard Design Criteria (SSDC) 1.3, Revision 2. These transients were evaluated to support operation at the c nditions. It was determined that all the design transients reduced T Mt l utilized for the current auxiliary systems tanks design are either unchanged or still bounding for the reduced Tbt c nditions. The term still bounding means that the transient temperature changes are less severe at reduced l
temperatures. t
\
Auxiliary Systems Heat Exchancers l
The Auxiliary System Heat Exchangers were designed in accordance with the 1 These transients were l design transients specified in SSDC 1.3, Revision 2.
evaluated and modified as necessary to support operation at the reduced temperatures. It was determined, with one exception, that all the design
~
tramients utili2ed for the current heat exchanger designs are either unchang;d Sr still bounding for the reduced temperature conditions.
The one transient that is impacted is considered in the design of the Letdown Heat Exchanger and Letdown Reheat Heat Exchanger.
It is the transient for the For the Charging Flow Shut Off and Re-Initiated (with Continued Letdown).
52 .
0181c1D-c21287
reduced temperature conditions the initial and finsi temperature is reduced by 10*F. This transient was reviewed and its effect on the fatigue analysis of both heat exchangers was judged to be insignificant.
,' Auxiliary Systems Pumps and Valves The Auxiliary System Pumps and Valves were procured in accordance with the design transients specified in SSDC 1.3, Revision 2 fI) These transients
)
were evaluated to support operation at the reduced temperature conditions. It I was determined, with one exception, that the transients remain the same or are enveloped by the previous transients used in the procurement of the auxiliary systems pumps and valves.
The exception in the transient for Cherging Flow Shut Off and Re-initiated (with Continut3 Letdown), for which the initial and final temperatures of the transient are reduced by 10*F. The minor change to this transient has an
. insignificant effect on the auxiliary pumps and valves. )
3.5.9.5 Auxiliary Systems Summary 1
The revised NSSS design transients which affect the LVCS, SIS and RHRS were reviewed and judged to be either less severe or not impacted as a result of reducing temperatures. The revised NSSS design transients which apply to the other auxiliary components (tanks, heat exchangers, pumps and valves) were reviewed and determined to be either less severe or to have negligible impact I on those components. On this basis and the bases described above, the T hot reduction will not have any adverse effects on the auxiliary systems.
3.5.10 Pressurizer and Steam Generator Valve Capacities An evalution was performed to review the minimum required flow capacities for the pressurizer safety valves, pressurizer PORV's, pressurizer spray valves
- and steam generator safety valves, for the reduced temperature conditions shown in Table 2.1-1. The purposes of this evaluation are to determine if the flow capacities assumed in the existing FSAR non-LOCA transients analyses must 0181v
- 1D-112387 53 !
i
change in order to accomodate acceptable analysis results at the revised conditions, and, if the assumed flow capacities must be increased, are the
, ,o currently installed capacities sufficient to allow the increase.
. The results of this review demonstrated that it was not necessary to assume higher valve capacities for the non-LOCA transients; where these capacities currently appear in the FSAR analyses , the same values are still valid.
Furthermore, using the current FSAR values for the valve capacities does not j result in more severe transient results than previously reported. This re-confirms that in every case listed above, the installed val a capacities are more than sufficient to meet the requirements of those valves.
l 3.5.11 Turbine-Generator The Turbine-Generator (TG) unit is a tandem-compound six-flow unit with a 40-inch last row blade. The unit is composed of one double flow high pressure turbine element and three double flow low pressure elements, plus a moisture separator-reheater with two stages of reheat. The unit was designed for l '. operation with steam conditions corresponding to a steam generator thermal output of 3425 megawatt thermal (MWT), 990 psia, and 543.3*F. An evaluation ,
~
was performed of the mechanical adequacy of the high pressure (HP) turbine, l low pressure (LP) turbine and moisture separator reheater under the revised operating conditions.
The revised steam generator outlet conditions will increase the steam volumetric flow to the high pressure turbine by approximately [ Ja,c% at I 3425 MWT. The turbine was not designed to be able to pass that volume of flow; as a result the electrical load generating capability of the TG unit will be reduced by [ ]"'C%. Modifications to the HP turbine to optimize operation with the revised steam generator conditions will allow the unit to regain [ Ja c% of the lost capability, while having the flexibility of allowing the unit to return to operation under the 990 psia original steam
.' generator conditions with an electrical load loss of approximately [ ]a,cy from the original calculations.
0181v:1D-112387 54 i
The most important factor to consider when evaluating the effect on the LP turbine is the location of the saturation line. The probability of having a turbine missile is increased if a disc is moved from the superheated region of the Mollier diagram to the wet region. From a review of the LP expansion
, lines from a T hot f 618*F to 600*F there was no change in the operating )
environment (superheated or wet steam) of the respective discs. Therefore there is no increase in the probability of a turbine missile.
Operating the unit under the reduced temperature steam conditions would have very little effect on the moisture separator-reheater. From a mechanical j adequacy standpoint, the change in operating conditions would have a negligible effect. 1 In summary, the reduced temperature conditions will not affect the safety considerations of the turbine generator or moisture separator reheaters.
- 3.5.12 Systems / Components Evaluations Summary 1
The evaluations presented above demonstrate that there is no safety or operational concern with operation of Byron 1 & 2 and Braidwood 1 & 2, i incorporating the design parameters shown in Table-2.1-1. In general, the
~
systems and components listed in Section 3.5 were reviewed to determine the effects of revised NSSS Design Transients and revised design power capability parameters. Emphasis was placed on determining the impact on fatigue usage factors (calculated from stress analyses) of the revised temperatures.
3.5.13 References l
- 1. Systems Standard Design Criteria (SSDC) 1.3, Rev. 2, dated April 15, 1974.
0181v;1 D-112387 55 I I
l
\
l
\
BYRON i T-44DT fit! DUCTION PMOORAN 1
i ao : : : :
- e 350- -
300- .. ,
i 390- -
1 g
o 200- -
.. Dristnal T 8 : Torget T 150- -
l 100- - ..
t 30- . ..
C l l l '! l 'I : :
o a 4 e e .to la 14 is se 20 TIME! (Years)
Figure 3.5-1 Byron 1 Cumulative Chemistry Operating Experence Factor vs. Time i
I I
0181v:10-020187 55 I
i
.=
BYRON 2 T HOT REDUCTION PROGRAN 200- l l l l l l l ! l 175- - X -
$50-- --
125-- --
. g g 200-- --
Origine! T g ; Terget T 75-- --
50-- --
25-- --
0 l l l l l l l l l 0 2 4 e a 10 12 14 is se 20 TIME (Year e)
Figure 3.5-2 Byron 2 Cumulative Chemistry Operating Experence Factor vs. Time 1
0181v;10-020187 l
l
l BRAIDWOOD i T-HOT REDUCTION PROGRAN 350 l l l l l l l l l 300-- --
250-- -.
. 200-- .-
t o Origine! T y : Target T
$50-- --
300-- ..
50-- --
0 l l l l l l l l l 0 2 4 5 e 10 12 id as se 20 TIME (Years)
Fioure 3.5-3 Braidwood 1 Cumulative Chemistry Operating Experence Factor vs. Time 0181v:1D-020187 58
BRAIDWOOD E T-44DT REDUCTION PROGRAN l
15e 125--
100--
Original T
. 75- -
- Terget T 8
50--
25-- .
l l l l l l 0 l l 10 12 14 15 18 20 P 2 4 5 5 t
t TIME (Yesce)
Figure 3.5-4 Braidwood 2 Cumulative Chettetry Operating Experence Factor vs. Time 0181v:1D-020187 59
3.6 Technical Specification 'mpact l One of the primary goals of the T hot reduction program has been to achieve as low a vessel outlet temperature as possible while minimizing the impact on the Technical Specifications. The initial task of the program involved a
~
study which determined the feasibility of developing steam dump setpoints for '
the 100% power parameters with T The study concluded hot reduced to 600"F.
that a vessel outlet temperature of 600'F could be achieved without affecting the original design no load temperature of 557*F, which was identified as a parameter indicative of significant impact on safety related setpoints and analyses. The evaluations and reanalyses described in previous sections have verified that no change in the Technical Specifications is necessary to implement operation at a vessel outlet temperature of 600*F. The current Technical Specification limits provide proper assurance that the plant will operate within the bounds of the existing accident analyses even V,1en in the reduced power and temperature mode, tiowever, as a result of options factored into the revised small break LOCA analysis, changes to the Technical Specifications are necessary. These are identified in Appendix B.
A statistical setpoint study performed previcusly for Byron Units 1 and 2
. provided increased margin in total allowance to various Technical Specification related instrumentation setpoints. Based on an evaluation of the sensitivities of the study for the reduced temperature parameters, it has been concluded that the setpoint allowances accounted for in the statistical evaluation remain va' lid.
3.7 Final Safety Analysis Report (FSAR) Impact The effects of reduced temperature and power have been evaluated for impact on the current systems, components and accident analyses as identified in Section 1.3 of this report. Appropriate changes to the FSAR will be identified and ,
docketed as an amendment.
/
0181v:10-020187 60 . %
3.8 Operating Window Concept in the past, nuclear power plant design and supporting de:ign basis analyses j have been focused towards permitting one single point of operation of a power
," plant at 100% rated thermal power. The intent of this section is to describe .
one of the principal goals of the T hot reduction program, which is to l eva'luate the capability and submit the necessary documentation to support !
operation of a power plant for a locus of operating points which permits variation in reactor coolant system temperatures.
Operation in future cycles, after the completion of this program, will have j the flexibility to operate at 100% power anywhere within an operating window )
bounded by a T f 618.4*F and 600*F, as determined by other plant l hot conditions such as percent of steam generator tubes plugged. It should be noted that, since steam generator tube plugging reduces heat transfer area, steam generator outlet pressure decreases as tubes are plugged, if primary
- side temperatures are held constant. At a certain minimum value of steam pressure, further pressure decreases result in the inability of the turbine to pass sufficient steam flow to maintain 100% thermal power output from the NSSS. By increasing T hot incrementally within the range of temperatures specified above, compensation can be made for reduced steam pressure to allow the unit to achieve 100% thermal power operation.
Detailed analyses performed in the analysis and evaluation sections of this report form the bases for the operating window. This involved reviewing all scope items at both the original and reduced Thot conditions and verifying that the temperatures and conditions in between these values are bounded by such analyses. One example of how this was accomplished is in the area of component stress analysis. The equipment designers were informed at the initiation of the program that the purpose of the program was to justify operation of the Byron /Braidwood units within the range of temperatures noted above. Since they had already performed the necessary work to analyze
- equipment stresses the original design temperatures, the bulk of the effort was spent in doing the same for the reduced temperatures. The NSSS design
' transients used as input for the equipment specifications were reviewed and modified as necessary. For nany of the transients, both the reduced temperature transient and the original transient were supplied in order that D181v:1D-021267 61
the stress analyst could determine which transient was more severe. For other transients, no change was made to the original curve since it was determined that the original transient was the bounding transient for all temperatures in the operating window. Where reduced temperatures caused the transients to
. become more severe in their impact, the new transients were supplied. It should be noted that, in general, reduced temperatures caused more severe transients on the cold leg, due to an increased zero load to full load temperature swing (see Table 2.1-1).
Determination of the target T hot at which the plant will operate will be made on a cycle-to-cycle basis prior to each plant start up. Implementation of this operating window is desired in order to achieve an economic benefit as well as operating flexibility. These benefits include increased net electrical generation and improved plant capacity factor as a result of being able to achieve 100% power at a wider range of operating conditions.
1 l
l l
l 0181v:1D-021287 62
4.0 LICENSING CRITERIA REVIEW 4.1 10 CFR 50.59 (Unreviewed Safety Question)
Changes made to an existing nuclear power plant must be evaluated under this regulation for impact on technical specifications, involvement of an unreviewed safety question and impact on procedures and analyses as described in the Final Safety Analysis Report (FSAR).
The evaluations presented in Section 3.0 demonstrate that no unreviewed c.;afety question is involved. Reanalysis and evaluation of specifi u ,1y impacted FSAR accidents has supported this conclusion by showing that the probability of )
occurrence, possibility of new accidents or margin of safety in any Technical Specifi ations basis has not changed from the original design. Plant 3 operation at reduced temperatures is not a test or experiment. However, as a result of options intended to provide operating margin that were factored into the LOCA reanalysis, Technical Specification changes are necessary. These are identified in Appendix B.
4.2 10 CFR 50.92 (Significant Hazard Determination for Issuance of Amendment) )
The criteria in this regulation must be considered when a proposed change to an existing plant also involves an amendment to the operating license. An evaluation of significant hazard considerations accompanies the application for amendment submittal.
4.3 10 CFR 50.36 (Technical Specifications)
This regulation defines the type of information to be included in Technical ,
l Specifications at the time of application for an operating license. The necessary amendments to Technical Specifications for this program are done in accordance with 10 CFR 50.90 (See Appendix B). i i
1 o181v:1D-021287 63
4.4 10 CFR 50.71 (FSAR Update) l
,. Faragraph (e) of this regulation provides guidelines for periodic updates of the FSAR. -The four Bryon/Braidwood units share a common FSAR and periodic
.' amendments will continue to be routinely provided as modifications accumulate. Revised FSAR sections, as necessary, will be provided at the conclusion of the program specifically addressing changes from a reduced T Changes to the FSAR associated with the LOCA reanalysis are supplied hot.
in Appendix A in order to ease review of the modifications.
1
)
l l
l
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0181v:10-021287 64
_s
5.0 CONCLUSION
The proceeding report addressed the efforts pei fermed to justify plant operation at 100% power for the Byron and Braidwood units. Additionally,
.' operation within a range of primary side temperatures corresponding to vessel average temperatures of 569.1*F to 588.4*F was also justified. Safety analyses or evaluations were performed in the following areas:
i LOCA and non-LOCA safety analyses Mass and energy releases NSSS systems and components 1 NSSS/ BOP interface The results demonstrate that the safety of these units is not compromised by I operating the units at reduce temperatures within the range described above.
- 1 l
l 0181v:1o-021287 65
)
'4 ,
APPENDIX A FSAR MARK-UPS FOR LARGE AND SMALL BREAK LOCA 4
0181v:1D-021287
.} >.
~ - }
m/s-rsnR
- 6. 2.1. 5 Minimum containment Pressure analysis for Performance capability studies of Mroency core M11mm svstem -
The containment backpressure used for the limiting omsk cp = 0.6, DECIs break
- for the Eccs analysis presented in section 15.6.5 is I presented in Figure 6.2-24. The containment backpressure is calculated using the methods and assumptions described in Appendix A of Reference 12. Input parameters, including the containment initial conditions; net free containment volume; passive sink materials, thicknesses, and surface areas; and starting time and member of containment cooling systems used in the analysis, are described below.
i 6.2.1.5.1 Mass and Enerov Release Lata The mass # energy releases to the containment during the blowdown and reflood portions of the limiting kreak transient are pre-sented in Tables 6.2-51 and 6.2-52. The mass and energy releaset from the broken Acop accumulator are given in Table 6.2-53.
The mathematical models which calculate the mass and energy releases to the containment are described in subsection 15.6.5.
Since the requirements of Appendix K of 10 crR 50 are very
- specific in regard to the modeling of the RCs during blowdown and the models used are in conformance with Appendiz K, no alterations to those models have been made in regard to the mass and energy releases. A break spectrum analysis is performed (see references in subsection 15.6.5) that considers various break i sizes, break locations, and Noody discharge coefficeints for the double-ended cold leg guil1otines, which do affect the mass and energy released to the containment. This effect is considered for each case analyzed. During refill, the mass and energy released to the containment is assumed to be sero, drich I minimizes the containment pressure. During reflood, the effect of steam water mixing between the safety injection water and the steam flowing through the RCS intact loops reduces the available energy released to the containment vapor space and therefore tends to minimise containment pressure.
- 6. 2.1. 5. 2 Initial containment Internal conditions The following initial values were used in the analysis:
containment pressure 14.7 psia containment temperature 90* F
., 5tWST temperature 35' F Outside temperature -108 ,F Relative humidity 98.85 O
. J..
2/3-75AR 1
.- i The containment initial conditions of 908 F and 14.7 psia are l
- representative 1y low values anticipated during normal f ull-power operation. , f
)
6.2.1.5.3 containment volume The volume used in the analysis is 2.809 x 106 ft8 6.2.1.5.4 Active seat sinks The containment spray system and the containment atmosphere re-circulation system fan coolers operate to remove beat from the containment.
Pertinent data for these systems, which were used in the analyses, are presented in Table 6.2-54.
A curve of the temperature versus beat load is provided in Figure 6.2-25 for the estimated capacity of one fan cooler.
The sump temperature was not used in the analysis because the maximum peak cladding temperature occurs prior to initiation of the recirculation phase for the containment spray system. In addition, beat transfer between the sump water and the containment vapor space was not considered in the analysis.
6.2.1.5.5 steam-Water Mixina Water spillage rates fra the broken loop accumulator are deterr.ined as part of the core reflooding calculation and are included in the containment (Coco) code calculation model. *
- 6. 2.1. 5. 6 Passive Beat sinks The passive beat sinks used in the analysis and their thermophysical properties are given in table 6.2-55.
i
- 6. 2.1. 5.7 Beat Transfer to Passive Beat sinks The condensing beat transfer coefficients use$ for beat transfer to the steel containment structures is given in Figure 6.2-261or the limiting break. The containment temperature transient for the limiting break is shown in Figure 6.2-27.
~
- 6. 2.1. 5. 8 Eher Parameters No ot.ber parameters have a substantial effect on the minimum
- containment pressure analysis.
I
a B/B-FSAR TABLE 6.2 - 51
- DECLG BIEWDOWN MASS AND ENERGY REYMSES f CD= 0.~ 6 )
TIME MASS FLOW ENERGY TIDW (SEC) fLB/SEC) (BTU'S/SEC)Y10-6 0.0 9,521 5.329 0.05 65,659 36.590 0.20 66,477 37.030 1.0 63,301 35.474 2.0 55,090 31.349 3.0 43,246 24.984
. 4.0 36,607 21.598 5.0 31,284 18.997 6.0 29,5$4 18.092 7.0 28,046
- 17.382 8.0 24,920 16.013 9.0 21,563 14.383
. 10.0 18,787 12.811 12.0 15,023 10.396 14.0 11,781 8.506
. 16.0 8,253 6.504 18.0 5,446 4.794 20.0 3,484 3.163 22.0 1,900 1.247 24.0 5,756 2.590 26.0 5,440 2.027 28.0 5,210 1.477 30.0 164 0.0167 6.2-161
.t B/B-PSAR
.- TABLE G.2 - 52
' . DIGC REFLOOD MASS AND ENERGY REftASES (CD=0.6)
TIME MASS FIDW ENERGY FIDW fSEC) (LB/SEC) fBTU'S/SEC) 45.917 0.0 0.0 46.542 0,,02345 30.0 46.742 0.02355 30.5 l 46.942 0.02366 30.6 1
47.342 0.02838 36.8 56.198 1392.55 204,220.0 70.254 262.33 192,646.0 87.554 282.28 191,277.0 106.854 293.68 186,600.0 127.554 302.5 181,020.0
' 149.554 310.41 174,950.0
, 250.004 219.388 123,650.0 l
1 '
l 6.2-162 9
.i',, -
5/B-FSAR
~
- TABLE 6.2-53 3ROKEN EDOP INJECTION SPILL TO CONTAINMENT (CD=0.6)
THE SacKEN LOOP $8tJECT30N SPILL DUR!nG BLOWDOWN 25
( c2751.252 2h# % Y !r T 164029.635 4J 99.620 1.010 2544.826 151722.525 59.620 i I
2.010 '2380.4 4 141946.018 M .4FD 3.010 2245.612 133M 3.399 59.620 l 6.010 2131.746 127094.707 59.620 S.010 2033.828 121256.821 99.620 6.010 1947.9f4 116137.606 59.620 7.,010 1871.60u 111585.157 59.620 8.010 1803.066 107498.916 99.620 9.C 10 1741.217 103811.328 99.620 10.010 1685.141 '8004M.128 59.620 11.010 1634.080 97423.870 59.620 12.010 1587.275 94633.315 99.620 13.010 1544.112 92059.966 90.620
. 14.010 1504.1 % 09677.777 99.620 15.010 tot 7.016 47463.521 59.620 to.010 1432.548 45410.877 $9.620 17.010 1400.541 83500.238 99.620
. 18.010 1370.575 81713.693 99.620
+ 19.010 1342.579 40044.564 $9.620 20.010 1316.419 78484.916 59.620 21.010 1291.806 77017.444 99.620 22.010 12M.673 ?$6:1.271 59.620 23.010 1266.835 74336.274 99.620
. 24.010 1225.912 73088.903 59.620 25.010 1205.902 71895.463 59.620 26.010 11M.894 70762.637 59.620 27.010 1228.397 MM3.439 55.453 28.010 1212.587 67342.395 55.536 29.010 1997.468 e6365.093 55.422
~
6.2-163
~
o
B/B-FSAR TABLE 6.2-54
' ACTIVE REAT SINK DATA POR MINIMUM POST-LOCA CONTAINMENT PRESSURE I Containment Spray System Parameters A. Maximum spray system flow, total 8118 gym B. Fastest post-LOCA initiation of spray system assuming offsite power loss at start of LOCA 35 sec l II containment Atmosphere Recirculation Fan Coolers e
A. Maximum number of fan coolers operating 4
- B. Fastest post-LOCA initiation assuming offsite power loss at start of LOCA 15 see C. Performance data See Figure 6.2-25 for fan cooler temperature versus heat load curve.
6.2-164
b B/B-FSAR
~
TABLE 6.2-55 l PASSIVE HEAT SINr. DATA FOR MINIMUM ,
l POST-LOCA CONTAINMENT PRESSURE I
A. Heat Sink Description f Slab Slab Material Surface 2
Number Description Material Thickness (FT) Area (FT )
1 Structural Steel Carbon Steel 0.02083 202007.6 Structural Steel Carbon Steel 0.25583 305.56 2
' Carbon Steel 0.1325 84.028 3 Structural Steel Carbon Steel 0.19792 285.
4 Structural Steel Carbon Steel 0.20833 B50.
5 Structural Steel Carbon Steel n.23958 745.
~
6 Structural Steel carbon Steel 0.125 1055.
7 Structural Steel Carbon Steel 0.1 04 852.94 8 Structural Steel Carbon Steel 0.04167 40138.
9 Structural Steel Carbon Steel 0.025 42665.9 10 Structural Steel Carbon Steel 0.16667 506.5 11 Structural Steel carbon Steel 0.1875 125.4 12 Structural Steel concrete 1.0 96562.9 13 Internal Concrete Concrete 1.0 14063.5 14 Internal Concrete Contaiment Floor Concrete 0. 5 15 Steel 0.03362 788.7 Containment Floor Containment Floor Concrete 0.5 i
~
16-19 Foundation'and Sump Concrete 0.5 Foundation and Sump Steel 0.02292 1762.1
'- Foundation and Sung Concrete 0.5
. 6.2-165
. }
I .. .
S/8-TSAR TABLE 6.2-55 (Cont'd)
Slab Slab Material Surface I
Number Description Material Thickness (FT) Area (FT )
Foundation and Sump Concrete 0.5 Foundation and Sump Steel 0.01563 9652.5 Foundation and Suw Concrete 0.5 Foundation and Sump Concrete 0.5 Foundation and Sute Steel 0.04899 22371.
Foundation and Sump Concrete 0.5 Foundation and Sump Concrete 0.5 Foundation and Sump Steel 0.15276 2878.7 Foundation and Sug Concrete 0.5 20 containment Wall Concrete 0.5 Steel 0.020B3 110355.
02ntainment Wall Containment 211 Concrete 4.0 B. Thermophysical Properties Themal Density specific Heat Conductivity 1b/ft 3 Btu /1b F Btu /hr-ft F
~
145 0.156 0.92 Concrete 490 0.12 27.0 Carbon Steel 6.2-166
,b-50.
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D. 50. 100. 150. 200. 250. ..
TIME (SECOND5)
) >.
BYRON /BRAIDWOOD STATIONS
- FINAL SAFETY ANALYSIS REPORT Figurs E.2-24 Containrnent Presue DECLG (Cp = 0,6) .
NOMTA/AL DOYON JV)
280 270 -
~
260 -
250 -
240 - j j
230 -
g 220 -
1 w 210 -
e D
Q 200 5
g 190 -
N 180 -
ONE FAN COOLER E ESTIMATED CAPACITY 2 ENTERING WATER TEMP. = 45'F g 170 -
E
$ 160 -
E w 150 -
140 -
130 -
120 -
110 -
100 90 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 HEAT LOAD (MBTU/HR) l BYRON /BRAIDWOOD STATIONS J
- FINAL SAFETY ANALYSIS REPORT l Figure 6.2-25 One Fan Cooler Estimated Capacity l
J
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f
- 2o N .lE*04-4 i E
$ B00.- $
_5 u>
g 600.-
v .
i U
C a 400.-
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$ TIME (SECONDS)
. SYRON/BRAIDWOOD STATIONS FINAL SAFETY A88ALY388 AEPORT
., Fipsre 8.2-26 '
Hest Trarsfer Coefficients verws Time DECLG (C = 0.6)
N6MTA/AL (UK 1 d.
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u 250.-
b y
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E O 1 y 100..
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TIME ISECONDS) e e
SYMMMIDWOOD STATIONS
. FINAL SAFETY ANALYS15 REPORT YW S.2 27 Containment Tempses versus Time DECLG (Cp = 0.6)
\
M W DAL YAAY f D 2:3 *f)
B/B-FSAR 15.6.5 Loss of coolant Accidents Resulting From a Spectrum of Fostd{ ate 5 Piping BreaEs'Within the Reactor Coofant Pressure Boundary 15.6.5.1 Identification of Causes and Frequency Classification A LOCA is the result of a pipe rupture of the RCS pressure boundary. For the analyses reported here, a major pipe break (large break) is defined as a rupture with a total cross-sec-tional area equal to or greater than 1.0 square foot (ft2),
This event is considered an ANS Condition IV event, a limiting
- fault, in that it is not expected to occur during the lifetime of the plant but is postulated as a conservative design basis (see Section 15.0.1).
A minor pipe break (small break), as considered in this sec-tion, is defined as a rupture of the reactor coolant pressure boundary (see Section 5.2) with a total cross-sectional area less than 1.0 f t2 in which the normally operating chargina system flow is not sufficient to sustain pressurizer level and pressure. This is considered a Condition III event, in that it is an inf requent f ault which may occur during the life of the plant.
The Acceptance Criteria for the LOCA is described in 10CFR50.46 (2) as follows:
- a. The calculated peak f uel element clad temperature is below the requirement of 2200er.
- b. The amount of fuel element cladding that reacts chemically with water or steam does not exceed 1%
of the total amount of Zircaloy in the reactor.
15.6-13
.s .
B/B-FSAR
('
The clad temperature transient is terminated at a c.
- time when the core geometry is still amenable to cooling. The localized cladding oxidation limits of 17% are not exceeded during or after quenching.
- d. The core remains amenable to cooling during and after the break.
- e. The core temperature is reduced and decay heat is removed for an extended period of time, as required by the long lived radioactivity remaining in the core.
These criteria were established to provide signficant margin in Emergency Core Cooling System (ECCS) performance following a LOCA. Reference (3) presents a recent study in regard to the probability of occurrence of RCS pipe ruptures.
In all cases, small breaks (less than 1.0 f t2) yield results with more margin to the Acceptance Criteria limits than large breaks.
15.6.5.2 Sequence of Events an6 ,Syst3ms Operations a Should a major break occur, depressurization of the RCS results LJ in a pressure decrease in the pressurizer. The reactor trip
, signal subsequently occurs when the pressurizer low pressure trip setpoint is reached. A safety injection signal is gener-ated when the appropriate setpoint is reached. These counter-measures will limit the consequences of the accident in two ways:
- a. Reactor trip and borated water injection comple-ment void formation in causing rapid reduction of power to a residual level corresponding to fission product decay heat. However, no credit is taken in the LOCA analysis for boron content of the injection water . In addition, the insertion of control rods to shut down the reactor is neglected in the.large break analysis.
- b. Injection of borated water provides for heat transfer from the core and prevents excessive clad temperatures.
b 15.6-14
4
.I... .
- c. s -
B/B-FSAR Description of Large Break LOCA Transient The sequence of events following a large break LOCA are pre-
. sented in Figure 15.6-4.
Before the break occurs, the unit is in an equilibrium condi-tion,.i.e., the heat generated in the core is being removed via the secondary system. During blowdown, heat from fission pro-duct decay, hot internals and the vessel continues to be trans-ferred to the reactor coolant. At the beginning of the blow-down phase, the entire RCS contains subcooled liquid which transf ers heat from the core by forced convection with some fully developed nucleate boiling. Thereafter, the core heat transfer is based on local conditions with transition boiling and forced convection to steam as the major heat transfer mechanisms.
The heat transfer between the RCS and the secondary system may be in either direction depending on the relative temperatures.
In the case of continued heat addition to the secondary, secon-dary system pressure increases and the main steam safety valves may actuate to limit the pressure. Makeup water to the secon-dary side is automatically provided by the auxiliary feedwater system. The safety injection signal actuates a feedwater iso-lation signal which isolates normal feedwater flow by closing the main feedwater isolation valves and also initiates emer-gency feedwater flow by starting the auxiliary feedwater pumps. The secondary flow aids in the reduction of RCS pres-sure.
When the RCS depressurizes to 600 psia, the accumulators begin to inject borated water into the reactor coolant loops. Since the loss oi off site power is assumed, the reactor coolant pumps are assumed to trip at the inception of the accident. The effects of pump coastdown are included in the blowdown analysis.
The blowdown phase of the transient ends when the RCS pressure (initially assumed at 2280 psia) falls to a value approaching that of the containment atmosphere. Prior to or at the end of the blowdown, the mechanisms that are responsible for the bypassing of emergency core cooling water injected into the RCS are calculated not to be effective. At this time (called end-of-bypass) refill of the reactor vessel lower plenum begins. Refill is complete when emergency core cooling water has filled the lower plenum of the reactor vessel which is bounded by the bottom of the fuel rods (called bottom of core recovery time).
The reflood phase of the transient is defined as the time period lasting from the end-of-refill until the reactor vessel
[
15.6-15
(" . ,
B/B-FSAR has been filled with water to theFrom extent-that thestage the later. core-tempera-of blow-
- /(.A ture rise has been terminated.
down and then the beginning-of-reflood, the safety injection
- accumulator tanks rapidly discharge borated cooling water into
~~ the RCS, contributing to the filling of the reactor vessel downcomer. The downcomer water elevation head provides the driving force required for the reflooding of the reactor core.
The low head and high head safety injection pumps aid in the filling of the downcomer and subsequently supply water to main-tain a full downcomer and complete the reflooding process.
Continued operation of the ECCS pumps supplies water during long term cooling. Core temperatures have been reduced to long term steady state levels Afterassociated with the water dissipation level of resi-of the refueling dual heat generation.
l water storage tank reaches a minimum allowable value, coolant for long term cooling of the core is obtained by switching to the cold leg recirculation. phase of operation in which spilled borated water is drawn from the engineered safety features sumps by the low head safety injection (residual heat removal)
The Containment spray pumps and returned to the RCS cold' legs.
System continues to operate to further reduce Containment pres-sure. Approximately 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after initiation of the LOCA the ECCS is realigned to supply water to the RCS hot legs in order to control the boric acid concentration in the reactor vessel.
Description of Small Break LOCA Transient
~
lh
- As contrasted with the large break, the blowdownThus, phaseforofthe the small break occurs over a longer time period.
small break LOCA there are only three characteristic stages, ~
i.e., a gradual blowdown in which the decrease in water level is checked, core recovery, and long term recirculation.
15.6.5.3 Core and System Performance 15.6.5.3.1 Mathematical Model The requirements of an acceptable ECCS evaluation model are presented in Appendix K of 10CFR50 (2).
! Q uauu vu LesCewere v=1uo uivu z w o un. vi iuei rua mocea.s investigated in NUREG-0 onsidera-posed E?'the ese fuel rod models tions of the possi from us for by available has been demonstrated to be CHMENT ISB).
improvements in t evaluation mode ayerryijQe, he auer*Tsed,to perform this analysis is deeme y -^a--r"st?-e, 2-f terrptr'Ir.
( Also, datte ausa~o ist.)
(
15.6-16
B/B-FSAR
'Large Break LOCA Evaluation Model
([
The analysis of a large break LOCA Transient is divided .into
' three phases: (1) blowdown, (2) refill, and (3) reflood.
There are three distinct transients analyzed in each phase, including the thermal-hydraulic transient in the RCS, the pressure and temperature transient within the containment, and the fuel and clad temperature transient of the hottest f uel rod in the core.
Based on these considerations, a system of inter-related computer codes h_as been developed for the analysis of the LOCA.
y N SE R"T ~
G. .m ; y u vu wi ...Iw.. ..ym. wI w. i~-CA .u.1 1 .i.
bas
~
m.
ology is given in Reference (4). This document de se mputer the ma phenomena modeled, the interfaces among the pliance codes, an e features of the codes which ensure D, COCO, The SATAN-VI, with the Acce nee Criteria. analysis are and LOCTA-IV co . which are used in the gh (8); code modifi -
described in deta'il References (5) th through (11), (25), (20 .z. C eferences (
cations are specified the core heat transfer (2 8) . These codes are use o ass geometry and to determine if t ont ore remains amenable to the blowdown, refill, and cooling throughout .no subs The SAT computer code reflood phases of the raulic transien n the RCS during
' analyzes the thermal- LOOD computer code is o to calculate blowdown and the ing the refill and reflood ph a of the
. e this transient COCO computer code is used to caleu e the accident. T the Contains pressure transient during all three phase d LOCA ysis. Similarly, the LOCTA-IV computer code is L tg e6mpute the thermal transient of the hottest fuel rod dur
_.a thrse rk----
SATAN-VI is used to calculate the RCS pressure, enthalpy, density, and the mass and energy flow rates in the RCS, as well as steam generator energy transfer between the primary and secondary sytems as a function of time during the blowdown SATAN-VI also calculates the accumulator phase of the LOCA. water mass and internal pressure and the pipe break mass and energy flow rates that are assumed to be vented to theAt the end of the blow Containment during blowdown. Also at the these data are transferred to the WREFLOOD code.
end-of-blowdown, the mass and energy release rates during blow-down tion of are transferred topressure the Containment the COCO code during response for usethis in first the determina-Additional SATAN-VI output data from the phase of the LOCA.end-of-blowdown, including the core inlet flow rate and
?
enthalpy, the core pressure, and the core power decay transient, are input to the LOCTA IV code.
- LocBART
/
15.6-17
.~
INSERT - 1 The description of the various aspects of the IDCA analysis methodology is given in References 4, 10, 15, 20, and 21. These documents describe the major phenomena modeled, the interfaces among the computer ,codw, and the features of the codes which ensure compliance'with the Acceptance Criteria. The SATAN-VI (Reference 5) , WREFIDOD (Reference 6) , COCO (Reference 7), BART (Reference 20), BASH (Reference 21) and IDCBART (Reference 8,21) codes are used to assess the core heat transfer geometry.
and to determine if the core remains amenable to cooling throughout and subsequent to the blowdown, refill and reflood phases of the IDCA.
G I
i e
1 l
I
1 B/B-FSAR
( ]
.- -< LInse n k.
th allhydraulic model to betbraineMe core flood [ng rat [
.' (i.e., he coolant rate at which coolant enters the bottom of uench
- ore), pressure and temperature, and th the LOCA.
i front heigh uring the refill and reflood phase addition to
' #REFLOOD also culates the mass and energye Since mass flow rate to the Containment th gh the break. oding rate and the
- he Containment depen upons the core on of the Containment Local core pressure, whi a fu codes are interactively
> backpressure, the WREFLOOD a o the LOCTA-IV code in that Linked.- WREFLOOD is also lia from LOOD are used by LOCTA-I V i
- hermal-hydraulic parame e fuel tempera e. LOCTA-IV,is used in its calculation of is of the LOCA trans t to calculate the
- hroughout the anaure and metal-water react of the hottest f uel clad temp r od in the e.
1978 The 1 e break analysis was performed with the Februa 9 on of the evaluation model which,oincludes . a ens modificat
'1', l', 19,
.rlfrrrtr' in *r'rrrrrrr
' The analysis in this section was performed with the upper head fluid temperature equal to the reactor coolant system cold leg fluid temperature, achieved by increasing the upper head
- cooling flow (Reference . ggg4 3]
'/ Enall Break LOCA Evaluation Model __ .
LTesert Us m .m ==eti usw.a u us w=
u.w o au w=
--- F .u ty.i.
= m u-is extension of the FLASH-4 code (12) developed The WFLAS at the Westin >use Be'ttis Atomic Power Laboratory. e RCS.
program p its a detailed spatial representation o by flowpaths.
The RCS is noda ed into volumes interconnect e intact loops the broken loop is eled explicitly with The transi behavicr of the ,
Lumped into a second .
e gover g conservation equations system is determined fr d through the system. A of mass, energy and momentum given in Reference (13).
detailed description of WFLAS analysis inv es, among other things, The use of WFLASH in the reactor core a heated control the representationsociated bubble rise mode permit a volume with the re height calculation. The mult ode dcapa- spatial transient si e program enables an explicit and deta bility of In part .laralit repre es aation of various system components.
proper calculation of the behavior of the loop en J
rina a loss of coolant transient.
Clad thermal analyses are performed with the LOCTA-IV code
- which uses the RCS pressure, fuel rod power
- (Reference 8)
(
l 15.6-18 ll
INSERT - 2
- ~
WREFLOOD, using input from the SATAN-VI code, calculates the time to bottom of core recovery (BOC), RCS conditions at BOC and mass and energy
,- release from the break during the reflood phase of the LOCA. Since the mass flow rate to the containment depends upon the core flooding rate and the local core pressure, which is a function of the containment backpressure, the WREFLOOD and COCO codes are interactively linked. The BOC conditions calculated by WREFLOOD and the containment pressure transient calculated by COCO are used as input to the BASH code. Data from both the SATAN-VI code and the WREFLOOD code out to BOC are input to the LOCBART code which calculates core average conditions at BOC for use by the BASH code.
BASH provides a more realistic thermal-hydraulic simulation of the reactor core and RCS during the reflood phase of a large break LOCA.
Instantaneous values of the accumulator conditions and safety injection flow at the time of completion of lower plenum refill are provided to BASH by WREFLOOD. Figure 15.6-5 illustrates how BASH has been substituted for WREFLOOD in calculating transient values of core inlet flow, enthalpy, and pressure for the detailed fuel rod model, LOCBART. A more detailed description of the BASH code is available in Reference 21. The BASH code provides a much more sophisticated treatment of steam /Awater more flow phenomens dynamic
- in the reactor coolant system during core reflood.
interaction between core thermal-hydraulics and system behavior is expected, and experiments have shown this behavior. The BART code has
- been coupled with a loop model to form the BASH code and BART provides the entrainment rate for a given flooding rate. The loop model determines the loop flows and pressure drops in response to the calculated core exit flow
~
determined by BART. The updated inlet flow is used by BART to calculate a new entrainment rate to be fed back to the loop code. This process of transferring data between BART, the loop code and back to BART forms the calculational process for analyzing the reflood transient. This coupling of the BART code with a loop code produces a more dynamic flooding transient, which reflects the close coupling between core thermal-hydraulics and loop behavior.
The cladding heat-up transient is calculated A bymore LOCBART which detailed is a description of combination of the LOCTA code with BART. During reflood, the the LOCBART code can be found in Reference 8,21.
LOCBART code provides a significant improvement in the prediction of fuel rod behavior. In LOCBART BART the empirical FLECHT correlation has been employs rigorous mechanistic models to replaced by the BART code.
generate heat transfer coefficients appropriate to the actual flow and heat transfer regimes experienced by the fuel rods.
. INSERT - 3 Modeling features necessary to account for the reactor barrel-baffle
.- region and the reactor fuel assembly thimbles were included in this analysis as presented in References 15, 20,and 21. The impact of a no l
single failure assumption for the ECCS was examined by re-analyzing the most limiting bleak with maximum ECCS flows as required by Reference 22. l l
l l
l
l i
)
Insert A
~
Small Break LDCA Evaluation Model For loss-of-coolaggl'ccidents 9
computerduecodeto small breaks is used less than to calculate the1 square foot, the NOTRUMP5 transient depressurization of the RCS as well as to describe the mass and enthalpy of flow through the break. The NOTRUMP computer code is a state-of-the-art one-dimensional general network code consisting of a number of advanced features. Among these features are the calculation of thermal non-equilibrium in all fluid volumes, flow regine-dependent drift flux calculations with counter-current flooding limitations, mixture level tracking logic in multiple-stacked fluid The NO1 RUMP nodes and regine-dependent heat transfer correlations.
small break LOCA emergency core cooling system (ECCS) evaluation model was developed to determine the RCS response to design basis small break LOCAs and to address the NRC concerns expressed in NUREG-0611,
" Generic Evaluation of Feedwater Transients and Small Break Loss-of-Coolant Accidents in Westinghouse-Designed Operating Plants".
In NOTRUMP, the RCS is nodalized into volumes interconnected by flowpaths. The broken loop is modelled explicitly, with the intact loops lumped into a second loop. The transient behavior of the system is determined from the governing conservation equations of mass, energy, and momentum applied throughout the system. A detailed description of the NOTRUMP code is provided in References 11 and 12.
The use of NOTRUMP in the analysis involves, among other things, the representation of the reactor core as heated control volumes with an associated bubble rise model to permit a transient mixture height calculation. The multinode capability of the program enables an explicit and detailed spatial representation of various system components. In particular, it enables a proper calculation of the behavior of the loop seal during a loss-of-coolant accident.
l I .
~
. . ~
J- . .
l B/B-FSAR history, steam flow past the uncovered part of the cor9 and
- (-
A mixture height history from the WFLASH hydraulic calculations as input.
Figure 15.6-48 presents the hot rod power shape utilized to This power perform the small break analysis presented here. shape wa tion of power versus core height and also local power is maximized in the upper regions of the reactor core (10 ft. to 12 ft.). This power shape is skewed to the top of the core l
with the peak loc'al power occurring at the 4414r 10 0 f t. core l
elevation.
This is limiting for the small break analysis because of the core uncovery process for small breaks. As the core uncovers, the cladding in the upper elevation of the core heats up and is The cladding sensitive to the local power at that elevation.
temperatures in the lower elevation of the The core, peak cladbelow the two temperature phase mixture height, remains low.
occurs above 10 feet.
Schematic representations of the computer code interfaces are given in Figures 15.6-5 and 15.6-6.
- The small break analysis was performed with the Westinghouse ECCS Small Break Evaluation Model using the NOTRUMP code, approved for this
- C_'unebytheNuclearRegulatoryCommissioninMay1985(referto Re!arance 12).
15.6.5.3.2 Input Parameters and Initial Conditions Table 15.6-2 lists Laportant input parameters and initial con-ditions used _in the analysis.
7 C[ntert 4;)pre;cated in thi; ;;; tion x;; p :f::::' cith
~'c . -..aaysi r vessel upper head temperature equal to the RCS cold n g
ture. The effect of using the cold leg temperatur(24).
re tem the reac vessel upper head is described in Reference upflow In addition thethodology analysis in this section utilized 0).
barrel-ba f fle described in Reference s that are input The bases used to a et is the numerical valrvatively determined have been co rences 21, 22, and 23).
I parameters to the ana from sensitivity studies efer to R endix K regarding specific ,
In addition, the requiremen of ing models which provida a nodel features were met by sel he analysis. The assump-i signficant overall conservatonditions the reactor and tions made pertain to the time that the LOCA associated safety syst equipment at t king factors, the
>ccurs and include re, ch items as the core and the performance of e ECCS. Decay nservatively
. :ontainment 1est generate preshroughout the transient is also calculated 7, En a rdance with the methodology presented in referene g (Cp = 0.6) was repeated ass gg worst break for this plant
_ ri ;1: f ri1=: (=ri== ::f: gn :d ' .
15.6-19
i l
INSERT ,1 A range of reactor operating temperatures were analyzed in order to justify operation at 1004 power between 600 to 622.3 Degrees-F in the RCS hot legs and 535.6 to 560.1 in the RCS cold legs respectively. A full spectrum was done at the reduced operating temperatures and the most limiting break was repeated at the higher operating temperatures.
Additionally, the most limiting of the RCS operating conditions was reanalyzed assuming no single failure (maximum safeguards) in accordance with the methodology presented in references 21 a:nd 22.
The basis used to select the numerical values that are input parameters to the analysis have been conservatively determined from sensitivity studies (refer to Paferences 16, 17, 18, and 21). In addition, the requirements of Appendix K regarding specific model features were met by selecting models which provide a significant overall conservatism in the analysis.
The assumptions made pertain to the conditions of the reactor and associated safety system equipment at the time that the IDCA was hypothesized to occur and include such assumptions as the core peaking
- factors, core power shape, the containment pressure, and the performance of the ECCS. Decay heat generated throughout the LOCA transient is also conservatively calculated.
A specific requirement for LOCA analysis (10CFR50, Appendix K) is that a range of power distributions shall be studied and the distribution resulting in the most severe calculated consequences shall be used in the analysis. The analysis was performed with a chopped cosine power distribution. Reference (21) addendum 1 has presented generic studies demonstrating that the chopped cosine power shape results in the most severe calculated consequences.
B/B-FSAR.
'( ' 15.6.5.3.3 Results Large Break Results 11 10 14 IDCA sensitivity studies, Based on the /esu[ts of the limiting large break was found (Ref erences X, K, and to be the double ended cold leg luillotine (DBCLG) . Therefore, only the DECLG break isCalculations considered in were the performed large breakfor BCCS a range performance analysis. The results of these of Moody break discharge coefficients.
calculations are summarised in Tables 15.6-1 and 15.6-3.
{
The mass and energy release data for the break resulting in the highest calculated peak clad temperature are presented in .
l Section 6.2.1.5.
Figures 15.6-7 through 15.6-33 present the parameters of For all principal interest from the large break ECCS analyses.
cases analysed transients of the following parameters are presented: )
l
- a. Bot spot clad temperature. (Figures 15.6-7, 15.6-7ag 15.6-22, 15.6-28)
'6h (Figures 15.6-8,
- b. Coolantfpressure in the reactor core. l 15.6-8a3 15.6-23, 15.6-29 j
- c. Water level in the core and downcomer during reflood.
(Figures 15.6-9 , 15. 6-9 a 15.6-24, 15.6-30) l l
'th s rate. (Figures 15.6-10, 15.6-10aA j d.
Core 15.6-25,refloodin$1) 15.6-
- e. Thermal ower during blowdown. (Figures 15.6-11, !
15.6-11 15.6-26, 15.6-32) ,b1 I 4
l
- f. Containment Pressure (Figures 15.6-12, 15. 6-12 a$ 15.6-27, 15.6-33)
The Containment pressure transient resulting from a LOCA is presented in Section 6.2.1.5.
For the limiting break analyzed, the following additional transient parameters nre presented:
(Figure
- a. Core flow during blowdown (inlet and outlet) .
15.6-13)
- b. Core heat transfer coefficients. (Figure 15.6-14) .
- c. Hot spot fluid temperature. (Figure 15.6-15) 15.6-20
r e
B/B-FSAR (Figure Is d. Mass released to Containment during blowdown. l 15.6-16) ,
- e. Energy released to Containment during blowdown.
(Figure 15.6-17)
- f. Fluid quality in the hot assembly during blowdown. '
(Figure 15.6-18)
- g. Mass velocity during blowdown. (Figure 15.6-20)
- h. Accumulator water flow rate during blowdown. (Figure 15.6-19)
- 1. Pumped safety injection water flow rate during reflood. (Figure 15.6-21) 0 ([oseer
" 5L/ . _
mo.. mum vi.u sw y s 6ute caAcuAstea ror a large brea of
' ll ich is less than the Acceptance Criteria er reaction
- t200*F of 50.46. The maximum local met ent limit of 174 as
.s 5.974 which ell below the embri e metal-water reaction is required by 10CFR50. . he tota s compared with the 14 less than 0.3% for all breand the temperature transient is
- criterion of 10CFR50. when the core ge is still amenable
- erminated at a a result, the core temperatu 11 continue to
- o coolin he ability to remove decay heat generate the fuel
<!ro x~~
- '- d A n
.arind af time will be provided.
Small Break Results
- -[Inseek b) sne ca Aculatea peak clad tempergat
~ calculated
- .. vowu y&wv4vusAy, resul om a small Basedbreak onLOCA is less thane LOCA sensi-the results for a large br ng small break was the tivity studies (Ref ere eter rupture of the RCS found to be less than a 10ange of saa k analyses are cold leg. Therefor The ablishes the limiting brea .
nd presented e
wh ese analyses are summarized in Tables .
Figures 15.6-34 through 15.6-47 present the principal For parameters of interest for the small break ECCS analyses.
all cases analyzed the following transient parameters are presented:
- a. RCS pressure. (Figures 15.6-34,15.6-41,)15.6-42 15.4 - 42 o, ls.6- 42k o
- b. Core mixture height. (Figures 15.6-35, 1s.6-43, 15.6-44r, 15 L
- Ms.,15.L- Hb) f 15.6-21
t l l
. INSERT - 5 The maximum cladding temperature calculated for a hypothetical double ended severance of the RCS cold leg piping was 1871 F which is less than the Acceptance criteria limit of 2200 F0 specified by 10CFR50.46. This result was calculated assuming a discharge coefficient of 0.6 for the break and with the RCS operating at a nominal hot leg temperature of 622.3 0F. Analysis performed assuming the RCS to be operating with a reduced hot leg temperature of 6000 F were found to be less limiting than the results obtained when the RCS was assumed to be operating with a hot leg temperature of 622.3 F. Additionally, the analysis with a nominal 0
hot leg temperature of 622.3 F was repeated assuming no single failure within the ECCS or ESF and the results were less limiting than the results calculated assuming a single failure in the ECCS. j The maximum calculated local metal-water was 2.414% which is well below tha embrittlement limit.of 17% specified in 10CFR50.46. The total core wide metal-water reactions is less than 0.3% for all breaks, as compared with the 1% criterion of 10CFR50.46 and in all cases the cladding
~
temperature transient was terminated at a time Ethen the core geometry was still amenable to cooling. As a result, the core temperature will continue to drop and the ability to remove decay heat generated in the
- fuel for an extended period of time will be provided.
These results provide assurance that operation with the RCS hot leg temperature in the range of 600 to 622.3 F can be accomplished within the requirements of 10CFR50.46 and Appendix K to 10CFR50.46.
)
i l
1 f
Insert B Small Break Results This section presents the results of a spectrum of small break sizes analyzed for the Byron /Braidwood Units. As noted previously, the calculated peak clad temperature resulting fromBased a smallonbreak LOCA of the results is less than that calculated for a large break.
LOCA sensitivity studies (Refarence 16 and 13) the limiting small break was found to be less than a 10-inch diameter rupture of the RCS cold leg. The worst break size (small break) is a 3-inch diameter break in the cold leg. This limiting break size was also analyzed for the nominal Thot condition to show that the higher temperatureThe time sequence o results in afor the results less all severe transient.
the breaks analyzed is shown in Tables 15.6-1 and 15.6-4.
During the earlier part of the small break transient, the effect of the break flow is not strong enough to overcome the flow maintained by the reactor coolant pumps through the core as they are coasting down following reactor trip. Therefore, upward flow through the core is maintained. The resultant heat transfer cools the fuel rods and cladding to very near the coolant temperature as long as the coreThis effect is eviden remains covered by a two-phase mixture.
accompanying figures.
.- B/B-FSAR
/ '
- c. Hot spot clad temperature. (Fi 15.6-46 3
$.4 -46 a,,15'.6 - 44 b)gure 15.6-36, 15.6-4 5,
- d. Core power after reactor trip. (Figure 15.6-37)
- e. Pumped sarety injection. (Figure 15.6-47)
For the limiting break analyzed, the following additional transient parameters are presented:
- a. Core steam 110w rate. (Figure 15.6-38)
- b. Core heat transfer coe fficient. (Figure 15.6-39)
- c. Hot spot fluid temperature. (Figure 15.6-48)
/b30 4 The maximum calculatedp [' eak clad temperature for all small breaks analyzed is 17.. 40cF. These results are well below all Acceptance Criteria limits of 10CFR50.46 and in all cases
- are not limiting when compared to the results presented for large breaks.
15.6.5.4 Radiological Consequences of a Postulated Loss-of-Coolant Accident (See Subsection 6.2.1.3) k* The results of analyses presented in this section demonstrate that the amounts of radioactivity released to the environment in the event of a loss-of-coolant accident do not result in doses which exceed the guideline values specified in a 10 CFR 100.
Two analyses are perfor d : .) a realistic analysis, and (2) an analysis based on Re g l at Guide 1.4, Revision 2. The parameters used for eaca f se analyses are listed in Table 15.6-9. In addition, an evaluation of the offsite dose result-ing from purging the containment for hydrogen control, and an evaluation of the offsite doses resulting from recirculation loop leakage are presented.
Fission Product Release to the Containment In order to evaluate the radiological consequences, the of fsite doses are based on the conservative fission product release given in Regulatory Guide 1.4.
I Thus, a total of 100% of the noble gas core inventory and 25%
of the core iodine inventory is assumed to be immediately available for leakage from the primary containment. Of the halogen activity available for release, it is further assumed that 91% is in elemental form, 4% in methyl form and 5% in i particulate form. The total core noble gas and iodine inven-tories are given in Table 15.0-8, while the activity in the o
containment atmosphere immediately following the LOCA and available for leakage is shown in Table 15.6-10.
15.6-22
B/B-PSAR i
- 15.
5.7 REFERENCES
- 1. T. W. T. Burnett et al. , "LOFTRAN Code Description,"
WCAP-7907, June 1972. Also supplementary information in letter from T. M. Anderson, NS-TMA-1802, May 26,1978 and NS-TMA-1824, J une 16, 1978.
- 2. " Acceptance. Criteria for Emergency Core Cooling Systems for Light Water Cooled Nuclear Power Reactors," 10CFR50.46 and Appendix K of 1CCFR50. Federal Register, Volume 39, Number 3, January 4, 1)74.
- 3. " Reactor Safety Study - An Assessment of Accident Risks in U. S. Commercial Nuclear Power Plants," WASH-1400, NUREG-75/014, October 1975.
- 4. Bordelon, F. M., Massie, H. W. and Zordan T. A.,
" Westinghouse ECCS Evaluation Model - Summary," WCAP-8339, July 1974.
~
- 5. Bordelon, F. M. , et al. , " SATAN-VI Program: Comprehensive SpaceTime Dependent Analysis of Loss of Coolant," WCAP-8302 (Proprietary) and WCAP-8306 (Non-Proprietary), June 1974.
- 6. Kelly, R. D., et al. , " Calculational Model for Core
, tE) Refloodine Af ter a toss of Cooiant Accident (WRErLOOo C od e ) ," WCAP-8170 (Proprietary) and WCAP-8171 (Non-Proprietary) , June 1974.
- 7. Bordelon, F. M. and Murphy, E. T. , " Containment Pressure Analysis Code (00CO) ," WCAP-8327 (Proprietary) and WCAP-8326 (Non-Proprietary) , June 1974.
- 8. Bordelon, F. M. , et al. , "LOCTA-IV Program: Loss of Coolant Transient Analysis," WCAP-8301 (Proprietary) and
- WCAP-8305 (Non-Proprietary) , June 1974. ,
- 9. Bordelon, F. M., et al., " Westinghouse ECCS Evaluation Model - Supplementary Information," WCAP-8471 (Proprietary) and WCAP-8472 (Non-Proprietary) , April 1975.
1120 -P- A sq
- 10. " Westinghouse ECCSjEvaluation Model - 0 ; f t - . .,,,
Version," WCAP&69t (Proprietary) and WCAP6 Q2,l-8 (Non-Pr oprie tary) , IL.- ':: 10 ? T.
Februsry 1982.
t
- 11. 1:tt:: "r cr-92t, f:tre 2:nu::y 23, 1075, C. rich:1 ding::.
(Z;; t ic.; h;;;; ) .; 0. L. ~lmeeollv GLC).
[
h 15.6-27 i l
l
I l
B/B-FSAR 1
.* 11. Meyer, P. E., NJIRIMP A Nodal Transient Small Bred and General Uetwerk Code , WCAP-10079-P-A, August 1985. (Proprietary)
{
- 12. Ime, H. , h=ld , S. D. , Tauche, W. D., Scir.rarz, W. R. , Westimhouse l
small Bred ECCS Evaluation Model Usim the N7tRDMP code, {
WCAP-10054-P-A, August 1985. (Proprietary) {
l E., Willis, J. M.,_
- 23. R w a-ht, S. D., Osterrieder, R. A., Wills, M. l Westiuscuse Small Break 10CA ECCS Evaluation Model Generic Study With Yhe H7IPUMP Cb5e, WCAP-11145-P-A, June 1986. (Proprietary) 4 l lf. Letter NS-CE-1672, dated February 10, 1978, C. Eicheldinger
- (Westinghouse) to J. F. Stolz (NRC).
. : mil 2, n. C., Theapseu, C . :: . , c. .l . , ";:e s ;.ing hames ergency Core Cooling System Evaluation Model for
- A zing Large LOCA's During Operation with One Los," Out
- of S ice for Plants Without Loop Isolation Val WCAP-91 February 1978.
" Westinghouse ECCS Ev ation Model,
.7. Eicheldinger, .,
Proprietary February 1978 Ve i on ," WCAP-9 2 20-P-Version), WCAP-922 -A (Non-Prop etary Version), Februar.r
- {' } 1978.
.8. Letter from T. M. Anders Westinghouse Electric Cc,rporation to John S z of t Nuclear Regulatory Cxunission, letter mber NS-TMA- 1, November 1, 197a .
l
- 19. Letter from to R.. Anderson L. Tedesco of Westingho e Electric of the Nuclea Regulatory Cor por atio 11, 1978.
Commiss n, letter number NS-TMA-2014, Decem
- .9. er from T. M. Anderson of Westinghouse Elect,riu orporation to R. L. Tedesco of the201', Nuclear Regulatory Occcric; 11, IMO.
N -[ 7 j , } - -"-1:: ':0 T".*.
" Westinghouse Eriergency l
.JHT. Johnson, W. J. and Thompson, C. N ,
/F Core Cooling System Evaluation Moul - Modified October 1975 Version," WCAP-9168 (Proprietary) and WCAP-9169 (Non-Pr opr i et a r y) , September 1977.
l
- f. " Westinghouse ECCS Evaluation Model Sensitivity Studies,"
WCAP-8341 (Proprietary) and WCAP-8342 (Non-Propr ietar y) ,
/6 July 1974. .
15.6-28
- - - - - - - - - _ _ _ _ _ _ _ _ _ _ . _ a
B/8-FSAR
.- 17 {
- M. Salvatori, R., " Westinghouse ECCS - Plan Sensitivity
(- Studies, WCAP8340 (Proprietary) and WCAP-8356 (Non-Proprietary) , July 1974.
is l
- 75. Johnson, W. J., Massie, E. W. and Thompson,M., C. Sensitivity
" Westinghouse ECCS-Four Loop Plant (17x17)
Studies", WCAP-8565-P-A (Proprietary) and NCAP-8566-A (Non-Proprietary), July 1975.
19 l
f4. Letter from T. M. Anderson of of Westinghouse Electric the Nuclear Regulatory Corporation to John Sto3r.
Commission, letter number NS-TMA-2030, January 1979.
-+ f::: T. M. ?.;f::::: ef " :tir;hre Emet-i-4&r : r ": ar Corpor to Dr. Denwood F. Ross, Jr. of the itegulatory ion, letter number NS- 54, December, 1980.
& better from T. M. on of We ouse Electric
. Denwood F. Ross, Jr. e Nuclear Corporation llegul Commission, letter number NS-TMA-
. . . . . sons Westinghouse Electric Corporation l N )df. Lett from E. P. Rahe t, to Robert L. Tedesco of the U.S. Nuclear Regulatory Commis-
- sion, letter number NS-EPR-2538, December 22, 1981.
@ : : . :.= fa :. r. re.; ;f T;;ti.c.;;;; ::::tri: c:r;::: tie-t 2: r: ?- Mille e f the " _ E Murle er Pr;;1_ __ _ , C - -' r -
cirn, 1:tt:: ni ':: M ErrZ-25??, r;c d :: ", I?"2.
J A Cmf ater Code for the Best f20. %9, N. Y. , " % RT - A test;nte. Analysis \
) Rev.1 0;t.A AJ/eds. 1-3 (Propr;etary), Arch ,l%4-I
'9 b
] ZI . hbad;,T.N., et al., "The 198I Version of the Wes+ing ouse Eccs Evaluction hodel Usin3 the BRSH Code " wcA kev 2 wit.h a)<Ienda. C Proprietary), Aups+, lib 6.
I i
l 15.6-29 i l
B/B-FSAR TABLE 15.6-1 (Sheet 1 of 5)
TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN REACTOR COOLANT INVENT _OlR TIME EVENT (SEC)
ACCIDENT Inadvertent opening Safety valve opens fully 0.0 of a pressurizer safety valve Low Pressurizer Pressure 41.7
' reactor trip setpoint reached Minimum DNBR occurs 43.7 Rods begin to drop 42.6 Large break LOCA
- 1. Start 0.0 DECLG Cp = 0.8 (TM Goo *F) Reactor trip signal O.5 07 4 +f A 4Per Safety injection signal /.52 Accumulator injection begins ll.5 W End-of-bypass g 4,04 3M M )
p,g J End-of-blowdown
~
h 15.6-30
~u
B/B-FSAR TABLE 15.6-1 (Cont'd)
TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN REACTOR COOLANT INVENTORY TIME (SEC)
ACCIDENT EVENT Purp injection begins R[,.ff,96TS l Botton of aore recovery gf.(,6-99M Accumulator empty ff,(/ f 49%
0.0
- 2. DECLG Cp . 0.5 5.dar t Meristum ; .,
- (..
s fr;:: ds Reactor Arip signal -
g,5/;t, (>,M-O 14
- l> W f.)
o T*
~
(-
Safety injection signal /. 72, M
{.'
Accumulator injection begins /f.8 M ,
q.
)/ .// M F.t.t{-of-bypass c, >
End-of-bloufown Kg.// Nh*3
'g 4
(
Pump itdatitt,sn begins 2 f,/7 2k J3 Il Bottom of core recovery 9F /.[4 M B d <
[. 3 M h l 4
k, Accoh.ulator empty
~
%.l e, .Ja
's \
h[ Q
- t. . ,.
, )
1-
'y M.6-31 1 1 '
~
=:L L t
B/B-FSAR TABLE 15.6-1 (Cont'd)
TIME ACCIDENT EVENT (see)
- 0. Y
- 3. DECLG C a Orfr S tart D.O bbm Reactor trip signal O. Ff., W Safety injection signal 2 .12. 4.44 Accumulator injection begins /y,f -Hv6-
~
End-of-bypass ,pg 26,98
. End-of-blowdown g pgM Pump injection begins gg jg 36d3
~
Bottom of core recovery %.fd41,30 Accumulator empty 70,*f/6h49 0'
- 4. WY WART ~ 0.6 Q T g*
- O A 3 ) R Oe:in& T/2rf SZCspAL O'*
gpptry rm gT M M HS l' H pp p 17.Qt74W W lW
- W (ND -of -BYNS 6Jp-of- 8Lo*U 30'W pgup tv,rrerseu8 & # 2" W'VE
~
g g a y a f c o d ifA % ' f e'/
AggpasA22 SW7Y SA Sl 15.6-31a
s/s-rsAR TABLE 15.6-1 (cont'd)
(
TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN REACTOR COOLANT INVENTORY TIME (SEC)
EVENT ACCIDENT 6 0.0 h#. DECLG CD=0.f Start
~
Reactor . trip signal 0 62 &
Aanmum 4v+
Safety injection signal /,66 Selc3uard5)
~ Accumulator injection begins /4. 8 4A+
End-of-bypass 3 0.5 B 2 2 . 0 --
e End-of-blowdown 3c.5S 9(rr*-
Pump injection begins 2.6.54 e654-Bottom of core recovery 44.87 6 Accumulator empty lo 4 .l f N Small break LOCA 7- 0.0 Start
- 1. 4 inch (7g= 6oo.o*F) Reactor trip signal 15.7 era-
' Sefety Iniec%n Si3nal 2 +.7 Top of core uncovered 2654 770J l 15.6-32 u_____-_______-__ -
c B/B-FSAR
TABLE 15.6-1 (Cont'd)
TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE DECREASE IN REACTOR COOLANT INVENTORY
' TIME (SEC)_
EVENT ACCIDENT Accumulator injection begins hl$ 4444-Peak clad temperature occurs 2982. 17::.:
Top of core covered f2.06 4 M G-0.0
' 3 Star t
- 2. finch (T;p Loo.o*F) Reactor trip signal 7.I 464-12.6 6dety Inject;.s Signal *ee-Top of core uncovered 105 4 444 Accumulator injection begins 2.09l
( B72 W+
Peak clad temperature occurs 2(o47 -GGG-
' Top of core covered 0.0 Start
- 3. / inch (T : 600.o*F) 4.I M et Reactor trip signal Sdet/ Injection Si9nal 6.3 Top of core unccvered 727 M 15.6-33
P B/B-FSAR TABLE 15.6-1 (Cont'd)
TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN REACTOR COOLAMT INVENTORY TIME (8eC)
EVENT ACCIDENT Accumulator injection begins 933 yMME Peak clad temperature occurs 777 220.25-
/39';E ?f ?-
Top of core covered j
. i 0.0
- 4. 6 inch Start Reactor trip signal 2.3
. (THOT= 600.0*F)
Safety Injection signal 3.2 l Top of core uncovered 311 l
, l Accumulator injection begins 348 Peak clad temperature occurs 412 l Top of core covered 427 0.0
- 5. 3 inch Start
{
Reactor trip signal 8.7 l (THOT= 622.3*F) 18.8 Safety Injection signal Top of core uncovered 1130 I
Accumulator injection begins 2127 Peak clad temperature occurs 1773 Top of core covered 2394 15.6-34
l l
l B/B-FSAR TABLE 15.6- 2
(
INPUT PARAMETERS USED IN THE ECCS ANALYSIS Licensed core power (a), (MWt) 3411 Peak linear power, includes 1024 factor (kW/ft) +ihM- I3.32.
Total peaking factor, Ff w 2.40
- ,=
Axial peaking factor, Fg
/ * +816 Power shape Chopped cosine Large break See Figure 15.6-48 Small break Optimized 17x17 Fuel assembly array 950
- Accumulator . water volume, nominal (ft 3/ accumulator)
Accumulator tank volume, nominal (ft /3 accumulator) 1350 600 Accumulator gas pressure, minimum (psia)
See Figures 15.6-21 t Safety injection pumped flow and 15.6-47 See Sec. 6.2 Containment parameters Initial loop flow (1b/sec)(redumi and ne+d Twt) 9992- 98f 3 I i vesse1 inlet temperature (OF)(etduceA ed naeat Tw.r) e F35.d 660 I >
Vessel outlet temperature (DF)(yedwd ad naa0nal Thot) fif 9 600.O hl*N 3 2280 Average Reactor coolant pressure (psia)
Steam pressure (psia)(redued and neWr4l Tg) A90- 776 h Y77 Steam generator tube plugging level (t) # /0 (a) Two percent is added to this power to account for calorimetric error.
15.6-35
i i
l
' I B/B - FSAR l TABLE 15.6 - 3 ,
I LARGE BREAK LOCA RESULTS FUEL CLADDING DATA REDUCED THOT REDUCED THOT REDUCED THOT 0
(600 F) (6000F) (6000F)
C D=0.8 C D=0.6 C D=0.4 DECLG DECLG DECLG RESULTS FOR N LOOP Peak clad temperature (DF) 1628.78 1754.002 1516.67 {
Peak Clad temperature 6.0 6.25 7.0 location (PT)
Local Zr/H 2 O reaction, 0.831 1.714 0.652 naximum (%)
Local Zr/H2 O location (FT) 7.0 6.0 7.0 Total Zr/H 2O reaction (%) <0.3 <0.3 <0.3 Hot Rod burst time (Sec) NB 1 49.16 NB 1 Hot Rod burst location (PT) N/A 6.0 N/A 15.6-36 1 Burst was not calculated to occur for this case. NB "No Burst".
'I .
B/B - FSAR e TABLE 15.6 - 3 (CONT'D)
, LARGE BREAK LOCA RESULTS FUEL CLADDING DATA NOMINAL THOT NOMINAL THOT (622.30F) (622.30F)
MAX ECCS C D=0.6 C D=0.6 DECLG DECLG RESULTS FOR N LOOP Peak clad temperature (OF) 1870.67 1815.96 Peak Clad temperature 6.25 6.25 location (PT)
Local Zr/H 2O reaction, 2.414 1.8114 maximum (%)
Local Zr/H2O location (FT) 5.75 7.0 I
Total Zr/H 2O reaction (%) <0.3 <0.3
. Hot Rod burst time (Sec) 45.2 45.31 Hot Rod burst location (PT) 5.75 5.75 15.6-36a
~
t F t 0 3 0 3 A A
=
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6 f 3 xr 6 2 2 0 N N 7 2 I 8 1 0 1 <
H2 1' 6 3 4
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n 8 2 2 0 N N I 0 1 0 1 <
1 6 1 t
9 0 3 0 3 A A ic 2 / /
R F n 5 2 2 0 N N G
- I 1 1 0 1 <
0 4 D 4 1 G 0 i 0 l
r 6 u =
W s 4 0 1 0 3 A A U
C hs tc 5 2 0
/
N
/
N n 0 2 1 1 <
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1 4
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- U 6 F 5 S 1 A 6 0 5 0 3 A A 1
hc 1 / /
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E m 2 D
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u u i t
i t
i t e a t t c a ca m c a a a co i o u e t l 2
= e e r r y y l L l m O 02 = a
- m O s b b 2 2 h
f t H H I h .
li t
d a
d a /r /r Z
/r Z d d s l l Z o o
- e c c l l l r r R k k a a a t t a a c c t b b o e e o n I l
l P P l I ' l n
Hs'm h
MT3LL MFLEB
.l ESIADW l 95 WC#tt l
- "" I sa tuL Tts e t ano, an attut eDD. g CE tulATES ICT EDC G& MCtWT 000. A C BL E LAnt. h.t.s. Et ulte.T 000 TEpFIBAtual, kKu&I.
AC NOT AL50 A15t4.180 CALCEATis fletutual.(LOCTA CDR1 TEWERATutt SET)i A g 6.g.g.
i, IC1 ASEt4LY. M elAll vtLKlft. a D 41TT. ;
entuunt '
' CDRI FL90CIE BAlt 3a.11 War BNfmALPt Sec1TIgul I
A " "IE l CALtiLattl att PL9001C INICt!C. Eit. SC5 CBC1Tiens palm SC5 Egeltlets NFLug SAttuLifti RCE. O kt.
151 A11t4L1 AT 80Catt m PLL'ID EECITIWt$
i I 8 3 accLattaTOR. St Flow, IIL11. Dratt trit 45E fEas gar 1 Alm ol1 pat 15unt
- 3870 Car 1A!Mui
=
CALDI.ittl trf!Lt. FLagets BAft alt IILit. gutast ALLIA5! GAft Ftt> ttl BuRIE StrigDD.
stFL90D . SALCILATES EDITAlWEN1 i
" M EBCC) strico/ coco l CALCutatts touTAtorwT pa(15utI (CDCD pr.T) l 1' i BYRDN/SRAIDWOOD STATIONS FINAL SAFETY ANALYStB REPORT Figure 15.6-5. 1 Code inter' ace Description for l Lay Break Model !
l a
t l
N L
0 O T C .
R T A
U CORE PRESSURE, CORE M
FLOW, MIXTURE LEVEL, P AND FUEL ROD POWER HISTORY 3
0 < 11ME < CORE COVERED v
1 i
BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figure 15.6 6, Code Interface Description for Small Break Model i
. - . ~ . - -
-]
1 i
i j
1 (i
l l
)
- .25E+04 a
$ .225E+04 w
$ .2E+04 E
~ !.* j
.175E+04
',/ h. )
g
'15E+04
[V (
N--'m~
' .125E 04 g
- +
o .l I .1E+04 750.
h 500.
C a 250.
"a u g 1 D. 20, 40. 60. 80. 100. 120. 140. 160. j TIME (SEC.)
l SYRDN/BRAIDWDDD STATIDNS
.* FINAL SAFETY ANALY8tg agPORT Figure 15.6-7 Peak Clad Temperature-DECLG ,
(Cp = 0.6) Minimum Safeguercis N6Mfi2hL*TileY(C2A3 *fN
r f
- .25E 04 u
$ .225E 04 U
e .2E+04 8 ^
~
.175E+04
/ *
' w
.15E 04 '
8
" .125E 04 l %'y/
G 1
=. .1E 04 1
- 5 750.
. h 500.
c o 250.
5 u g 13 . 20. 40. 60. 80. 100. 120. 140. 160.
TIME ISEC.)
SYRON/8RAIDWOOD STATIONS FINAL SAFETY ANALYSTS REPORT Figure 15.6 74 Paek Clad Temperature-DECLG {
(CD = 0.6) Maximum Safeguards N6MTL!AL 'YNOTON S
( .- .
4
- .25E 04 i, s
$ .225E=24 u
E .2E 24
.175E 94 15E*D4
" .125E 84 -
- V E .lE 94 c.
6 758.
e, e see.
.- a g 250.
d e 1
15 . 20. 40. 60. 80. 100. 120. 140. 160.
TIME fBCC.I I
i BYRON /SRAIDWDOD STATIDNS
- I I
I FINAL SAFETY ANALYSIS REPDRT j I
~
FIGURE 15.6-78 I I I l PEAK CLAD TEMPERATURE I
I I l REDUCED THDT 4600.0*F) i oEcLS - c..o.. !
E .
ssee E
Esees 1- \ s.
\ %
,e= '
N see L
N k % % .
's as s as se er s is it s se er s as av s os or s 184 8 kcl BYRON /SRAIDWOOD STATIONS t FINAL SAFETY ANALYsit REPORT Figure 15.6 8
~
~
Core Pressure DECLG (Cp = 0.6) Minimum Safeguards N6Mf2)Al Tint ~ (C'at.3*FN
4 ssee a
5 me 1- N s.
\ N. 4
. g N u see N
A e as s 7s is 13 s is 37 s se sa s as 37 s se er s ilE IEC3 SYRON/BRAlOWOOO STATIONS FINAL SAFETY ANALY888 REPORT Figure 15.6 Core Pressure DECLG
. (Cp = 0.6) Maximum SI
_ N8 prs /Ax We r N22. 3 *f O
. s rsee
- i s j Esees l
1see r T
,.ee N- _
\ %
see N
\ * % ,
's as s 7s is it s is it s se er s as 37 s se Ins tse secci I
i BYRON /BRAIDWDDD STATIONS 1 l I
." i FINAL SAFETY ANALYSIS REPORT '
- l 1
FIGURE 15.6-8B I
~-
i l I CORE PRESSURE TRANSIENT I i i I REDUCED THOT (600 0*F)
~ l DECLS - C.=0.6 ._,
I.
e e
l 22.
20- bu DWA"UNEDYO'EEYM'Y'YO'. NAN -
n, E18.
$ 16.
d 14 12.
10 * '
g' hgknW'*%ANNWJ4%%O7E' c.W,y/ h
[,$'
, 6.
4 D. 50. 100. 150.
. 200. 250.
TIME (SEC)
DOWNCOMER LEVEL (-----) , CORE LEVEL (** ***)
BYRON /BRAIDWOOD STATIDNS FINAL SAFETY ANALY8tt REPORT Figure 15.64 -
Reflood Transtent-Core & Dowf$mtrer Water Levels DECLG (Cp *D.6)
AnpryA heYbWbN
s.
22.
y w= =:,w.mmm . . =.v:. . . ..: e
- V b 18.
v>
d 1s.
5
_J 14.
12.
9 a +
3g' .
- 8. f 4- $ WJf*4%WN'%$W4TSshwW Yh
- 6. 7, I
4 B. 50. 100. 150, 200. 250.
l TIME (SEC)
DOWNCOMER 12 VEL (----) , CORE IZVEL (*****)
I SYRDN/BRAIDWOOD STATIONS j FINAL BAFETY ANALY8tl REPORT t Figure 15.6 9 %
Reflood Transient-Core & Downoomer water Levels DECLG (Cp = 0.8) Maximum safspuerds
- !J9pryAL TROT k.M 2 'E) i 1
L ,
t - i 22.
N' %%YA%%%W'M/A%WMMWWMVMVNMVN/#AMMMt#fM
_ 20. .
4
)
b 18. / -
w Y d 16, l
5
_a 14 1
- 12. !
g f[ 4% y /M A# ""# $Dd%%gggg
- 6. : 5 C
l !
4 D. 50. 100. 150. 200. 250.
TIME ISEC)
N N (-----),
CORE LEVEL (eeeee)
- 7-i
, I I BYRDN/BRAIDWOOD STATIONS i I I l FINAL SAFETY ANALYSIS REPORT
--- H a
l I FIGURE 15.6-9B i l l I REFLOOD TRANSIENT - CORE 8* DOWN WATER LEVELS (FT)
I l i I MEDUCED THQT (600.0*F) I i DECLB - Ce,=0.6 a u - ,.
I
\
r 1
- 10. .
{
.6 9.. -
8.
U
- .s M S 7- 2 b
56.
_ E z 5. 4 5 d 2 4 8 r
5 o m 5 5.
.2
- 2.
1.
O D. 25. 50, 75. 100. 125. 150, 175. 200. 225. 250.
71ME ISEC)
BYRON /BRAIDWOOD STATIONS f FINAL SAFffY ANALYSit AEPORT Figure 18.6-10 Reflood Transient Core ledet Velochy DECLG (C = 0.6) Minimum safegueds o
10.
9.
8.
u 7. .
N ,
s s. '
.z. 5.
I 4.
U5 o
. I v>
5.
- 2. -
1.
0 B. 25. 50. 75. 100., 125. 150, 175. 200. 225, 250.
TIME ISEC1 I
I i
BYRON /8RAIDWOOD STATIONS
~
FINAL SAFETY ANALYSTS REPORT
~
Figure 15.61@
t Reflood Transient Core inlet Velocity DECLG (Cp = 0.6) Maximum Esfoggerds
& @ M fs./AL W O T* f ( 9 3 3
- 0 =
1
L -
l
- 10. .8 9.
8.
U
.6 M U
U 7. b M Z h 6.
.a
'd z 5. 3 j
' g E d- 5 8 5- .2 5
2.
. 1.
B.
0 50. 75. 103. 125. 150. 175. 200. 225. 250.
B. 25.
TlHE ISECl I
I BYRON /BRAIDWDQD STATIONS I
I I e i, FINAL SAFETY ANALYSIS REPDRT FIGURE 15.6-109 ~l 1
I 1
i I l REFLDOD TRANSIENT CORE INLET VELDCITY l l i l M NtptttmL THOT (600.0*F) I I DECLS - C =0.6
f l
I e
s I
C i 4
4 A
N_
's as s 7s as sa s is it s se sa s as at s as sa s fitt (kcl l
BYRDN/BRAIDWODD STATIONS
) FINAL SAFETY ANALYS18 REPORT f
Figure 16.6-11 Core Power Transient DECLG (Cp= 0.6) Minimum Safeguards
- We+cruAL -mrun.2*FJ 1
~
I e
e o
a E
(
's as s as se ses is 17 s se an e as af s se sa s SHE IKCI SYRON/SRAIDWOOD STATIONS f
FINAL SAFETY ANALYSIS REPORT Figure 15.6110~
' Core Power Trans'ent-DECLG (Cp = 0.6) Maximum Si NGM1"NAL mot fW3Y e
N eu w wa-- e -
- .'~- . - . .
l 53 1
e e
a l
a
(
8
'S 38 87 s es ys a y, , , ,, , ,, ,, ,
y 1
l l
I
~
h BYRON /BRAIDWOOD STATIONS
~ I i l I FINAL SAFETY ANALYSIS REPORT 1
l FIGURE 15.6-119 I 1 I l CORE POWER TRANSIENT I
I i I REDUCED THOT (600.0=F) 1 I DECLS - C.=0.6
I 50.-
G G
s 25.-
W
@ 20.-
0 E
is.-
5 E 10.-
E E
o s..
B .O . 50. 100. ,150. 200. 250.
TIME iSECONDS1 SYRONltRAIDWOOD STATIONS FINAL SAFETY ANALYSl8 REPORT Figure 15.612 .
- Containment Pressure DECLG (Cp= 0.6) Minimum $stepueds Mpf])fL*~f5VfUY Y S
e 50.-
o m
.n 2 d.'
w R20..
so u
0:
- n. 35.-
z LJ
.E_se.-
E e.-
~
z o
u 5.-
150. 200. 250.
, 00. 50. 100.
. TIME (SECONDS)
/
I /
/
9 't /
/. , '
I /~ g e
)_
, ! 1,'
/ g 1
, r
-/ 1 fii / -
l','/
I ' p p
i .
,I SYRON/8RAIDWOOD STATIONS *
((
FINAL SAFETY A*tstYsis REPORT__]
. r e j I r j .
Figure 16M)4 l 7i '
'f ' Containment Pressure DECLG
.r' '
/
>j (Cp = 0.6) Maximum Safeguertis
, l
& f M TJJA L f o W l M '] b -
L l
./
a
' i
- t 5i
[ '
y
~ ~ ~ ~ ~ ~
i __3 } ~ . . ~
g- g ,
y* .v, m J.N., @ .. Mr W> . ,, - -
%^u stg a y.-i- ;>
,x ,8,3; ,: y,
. +
<u . . , -
\\.<g.
.y.. .
g_Q,v 1 .e.y . . .
u . .
s,". -r-y>, ; , [%. ,; ;;a'y 4
,N ** I;} P .,(:l 'c/-
- g. f .
4 .,,
. . , ,/ :e,>; 'y .
..,y q ,- } ty
- . v.
, 1 7 \ 4 h3 .
y\et o
., 'i
- . . 'p,
,a 'I ch s 3, u,
- s. . ,.
4, s r t e.
- s. 'n 1: t ,1 t .
' N ;' ' r A j- w
}' . I
- 6,* };- 'ff V' Vg ,
f.'.,' p, y '_' ,
. ti. , ,. ; n f
1
,. - ' y b '50.- 3t i ,
't
( ,
'f. ~) r ,I r
' C.
- ,- ll -s 8h i - ev.s 25. < s i i. , , o T.
m s as .p p' *,
%Qq .,h in m .y.
g- N' 1
[40.J..[A;\
rn u2 <t .
'N ~
u1 r
.' t,.4 ;t '. / t' '
/
<x ... ' ,{ '
g J
,Ns O , 15j Js LL
' it,5 % ' , e. -s p ~N i, ;; x ,
'n ,, <' .
,2 '.c we
, s s -s
, a,.
y i
/,
Q '
j; 3' I *;
a oe 'f;
- n
- 11 1 tc 2 .
cp 4 .w ,. .
( i.
, ,. ,i.w tj, ., ..,, '
, , -f
.< }
. .x
, ,. y ,,,, .
J t.
U *0 c. , ,, 50. .,.100s 150. 200, 250.
L, ,
9 <
TIME ISECONDS) s
'x -
, r;: ..
's i
c 4
t,
~ f ; .. < t- 6 -
t, n ,
>r ii y
. L.: ni b t tt' i % '
-}j[
( i .r
(%
y ) ^. ,
ha h '
F t' ,
A ; .
- n. i..-
I-
/f .,
o, . .
it; {:
E f, e r s
- t F SYRON/BRAIDWOOD STATIONS I
I r
l
~
n 3
m- ,
l, FINAL SAFETY ANALYSIS REPDRT
~
/!
,b9 i F2GURE 15.6-129 l
/ .j g
- CONTAINMENT PRESSURE I
g REDUCED THOT (600'.0*F) 3 DECLg - c =0.6 ,
t e t r
. . . . . 1 M "
V.see E
S_
esse W
l,,,, A L
f , h -
j ve%g Dees
.aees
-esse h
'" o as s rs se it s is it s ss as s as er s se sa s FIT. EeCCI BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYSl8 REPORT
- Figure 15.6-13 Core Flow (Top and Bottom)
DECLG (Cp= 0.6) O pf JJSMTLIAL %nT" /?.2s.1 *f I _
, t . .
- u. 50. p e
- 9 45.
W I
- 4e.
O fi --
G 55.
8 m 50. p'
@ 25. l 1 AA
. 15.
\ Af ? l\ h
.g py i 2
^
E 10.
E
- y 5-B. 20. 40. 60. 80. 100. 120. 140. 160.
TIME ISEC.1 SYRON/BRAIDWOOD STATIONS FINAL RAFETY ANALYS18 REPORT Figure 15.514 j Core Heat Trarsfer Coefficient DECLG (CD = 0.6) t Ahxwat "Ar tw s *r2 1
+ - .25E 04 a
S .225E 04 y
w
$ .2E 04
~ .175E 04 S
w .15E 04 N '
@ .125E+04 c.
g .it.c~.
( s. -
v O
5 750.
a
~ 500. IU L l u 3j
~
250.
E
~
D. 20. 40. 60. 80. 100. 120. 140. 160.
TIME ISEC.1 SYMON/BRAIDWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figurs 15.6-15 Fluki Temperature - DECLG (Cp= 0.6)
ACM27)k %.s7- (L "..T *Y)
I 1
1
I
- )
vesst.s l esset.s \
l seest.s n
g deset s
(
e-.s \8 seest's
\ ,
N tesst.s
\
N % r %
Q s
- '" * 's as a rs se sa s is sr. s se se s ss at s as sa s w e aste 1 .
l.
,. BYRON /SRAIDWOOD STATIONS FINAL 4AFETY ANALYSIS REPORT
~ Figure 15.6-16 treak Flow Rete DECLG (Cp = 0.6) AtgN24W M l N 0 k r u / L ~11te r t w . 2 V ) :
1
- 8555c*8 es550 0 g _.. <
r asset *e \ ,
N isost.e
\
Is08t*9 y%
sseet.? N s
- " " s as s 7s is er s is or s se er s ss sr s as sa s 114 ts[Cl BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYSit REPORT Figure 15.617 Brook Energy Released to Containment 4 U DECLG (Co = 0.6) MgTgp(M SifCJ NdimlAL THer- (412.1 of)
~
1.6
~
~
b 1.4 E
E 1.2 1- iv g )
S .6 w
c5
.e C 1 . i l
_a a w-g .2 d 160.
- 40. 60. 90. 100. 120. 140.
E. 20.
-, TIME ISEC.)
BYRON /BRAIDWOOD ETATIONS FINAL SAFETY ANALYSTS REPORT Figure 15.618 Fluid Quanty-DECLG
- g (Co
- 0.6)
A0J)T)JM M R z!3 "YJ
i I
.~
i esse j b# \
so,e W
R c' awe I.,,,
seer 1 e ea s 7s se se s is d it s se as s a er s as sa s 134 BECs I
1 i
BYRON /BRAIDWOOD STATIONS
. FINAL SAFETY ANALYSTS REPORT Figure IL S 19 Accumulator Flow (Blowdown)
DE CLG (Cp = 0.6) M,,,,... .Su..ne,o N6Pru4L net cn2. 3 *FL .
i ee
- 50.
d 9 4s.
N
[ 4 0 '.
S l 2 ss. !
50.
25.
20.
C
.-. 1 5 .
O d
10.
. ~n n~ n em n e h o
}
f I Ii V\jW 3 /H Eg J
- 80. 100. 120. 140. 160.
D. 20. 40. 60.
TIME (SEC.)
BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALY8tl REPORT Figure 15.6 20
- Mass Velocity-DECLG ,
(Co aD.6) hl0M7WAL ~7y~&T (d 22..?
4
4 10.-
c.-
a B.-
7.-
U g 6.-
5 L 5.-
N 4.-
$a
" 5.-
2.-
- 1. -
2.
- 2. 25. 50. 75. 100. 125. 150. 175. 200. 225. 250.
TIME ISECOND5)
I )
j BYRON /BRAIDWOOD STAT 1DNS I t
l l i FINAL SAFETY ANALYSIS REPORT _ _;
I I
FIGLAE 15.6-21 I I I l PUMPED ECCS FLOW (REFLOOD)
I i
~
l I i NDNINAL THOT (622.3*F) I l DECLS - C,=0.6 i
t
.~
u.
.25E 04
$ .225E 04 u
EE .2E*04 s
~ .175E 04
.15E+04 #
A a
yg l'
o
" . 125E 04 'j N
\-/ N k .1E+04 N- m c.
5-750.
h 500.
C o 250.
3 o e D. 20. 40. 60. 80. 100. 120. 140. 160.
TIME ISEC.)
{
l I
i 1
BYRDN/BRAIDWOOD STATIDNS
, FINAL SAFETY ANALygts REPORT Fleure 15.6-22 Paek Ciad Temperature-DECLG (Cp = 0.8) j ffNQ9 'Titr (Ut *FO l l
i
- aus. =
1 Eseen.
flat.
b i
.. \ -
N N
see.
N N
- s. a. a. s. e. nn. u. u. n. u. se. m. **. m.
sw (estl SYRON/SRAIDWOOD STATIONS FINAL SAFETY ANALYSTS REPORT Figure 15.623 Cort Ptsenne DECLG (Cp
- 0A) kD'd & 7007[d6'? ,
)
I l
25.
C 1MW m;:: : :,n :.: . : :::.:
g 20. y 3
w t, is .
a 30- , y, :g% :. ; =:, w =,7;. ,:.-.=
- 5. f D. 50. 100. 150. 200. 250.
TIME (SEC)
DoWNCOMER LEVEL (-----), CORE LEVEL (*****)
SYRON/BRAIDWOOD STATIONS FINAL SAFETY ANALY8tB REPORT r Figure 16.6 24 Reflood Transient-Core & Downoorner Water Levels DECLG (Cp = 0.8) 0(mfd W5T (@ W k- -- -
s.
10.
.6 9.
- 8. -
_ u u 7. .6 M R D E 6. b z
.5b* 5
> 4 x
h o
O o
y s. 5 2.
}. j D. 25. 50. 75. 100. 125. 150. 175. 200. 225. 250.
TIME ISEC1
. BYRON /8RAIDWOOD STATIONS FINAL SAFETY ANALY&ts REPORT
- Fipwe 15.8-25 Reflood Transient Core inlet Velocky DECLG (C D
- E8) 1 SfDo911) tee 7" fi9!'YJ
i e
/
3.3 l
- .8 I
8
.3
( -
S. 82. 82. Id. 84. 88. AS. 37. 74 36.
- E. A. d. 4.
n,s eset.
t I SYRON/BRAIDWOOD STATIONS FINAL SAFETY ANALYSTS REPORT Figure 15.6 26 Core Power Transient DECLG (Co = 0.8L g f(M*b 7EeTOMYJ --
,. 50.-
G B 25.-
S w
!E 20. -
N u
' 35.-
E W
z E 20.-
E O
O 5.-
0 .O . 25. 50. 75. 100. 125. 150. 175. 200. 225. 250.
- TIME ISECONDS)
(
SYRON/BRAIDWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figure 15.6-27 i Containment Premore h a D ls* % % rJ 4
I
- c l
t
- .25E 84 a
w
.225E.84 E .2E.84 ,
8 1
.175E 54 35E*B4 _- _'
8
- .325E 04 y [
/
E .it e4 I\ >
750.
o 588.
E p 250.
d e
- 50. 75. 120. 125. 350. 17E. 200. 225.
- B. 25.
TIME ISEC.I I
l l
I BYRDN/BRAIDWDDD STATIONS FINAL SAFETr ANALYSTS REPORT
- Figure 15.6-28 Peak Cled Temperature DECLG (Cp = 0.4)
$[DfdAD ~7ddr UM*fb ,,
l t
7598 E
Esese 1588 "
r
. N N
. N i
%~ %
5 5 le Il M E5 Se 55 de 85 sig tite BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYS48 REPORT Figure 15.6-29 Core Prosaure DECLG (Cp = 0.4)
Af00Lfu 'Tye7- r&M $
I J
S 25.
g w w: = ::::::::
a 20. ,
G d
D 15.
_s k t '
10.
74 p y yc- ~ : -. . . .
.s 5.
f' 0 100. 150. 200. 550.
~
- 2. 00.
TIME (SEC)
DOWNCOMER LEVEL (-----), CORE LEVEL (*****)
BYRON /8RAIDWOOD STATIONS FINAL SAFETY ANALYSTS REPORT Figure 15.6 30
- Reflood Transient-Core & Downzmer Water twels DECLG (CD= 0.4) f6Docfd Tsai Geo*FJ
4
- 10. ,g S.
~
- 6. G h 7.
- 5 2 6. g
~
s k i s. .a ,
3 o
d' w
E 5.
- .2 2.
1.
E
- t. 25. 50. 75. 100. 125. 150. 175. 202. 225. 250.
TIME ISECl BYRON /BRAIDWOOD STATIONS blNAL SAFED ANALYST 8 REPORT Figure 15.6-31 Reflood Transient Core Ir.!et Velocity DECLG (Cp = 0.4)
N U dfD ~)pfr (fde*f.)
f
. . . = . . . . - - -
s
\
A
, I ___.
k
. TiteI ISCCI BYRON /8RAIDWOOD STATIONS FINAL SAFITY ANALYSIS REPORT Figure 15.6-32 Core Power Transient i
DECLG (Cp = 0.4)
~
kfDD.CfD ~IMd7 (609*V]
' ~ ~ '
ws..
50.-
E
- 5
- c. 25.-
W 8 2e.- -
0 E
,_ is.. -
5 m
- 2. 20.
5 5
u F,-
l *0 . 50. 300. 150. 200. 250.
TIME lSECONDS)
BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figurs 15.6-33 Containment Prwisure DECLG (Cp = 0.4) kN_Yifff) EnT fbW 1
l 1
1 BYRON /BRAIDWOOD (CAE/CBE/CCE/CDE) -NOTRUMP SMALL BREAK 3 IN - REDUCED THOT - 17X17 OFA 275 PSIG
.24E*04
.22E*04
.2[*04 8 _
$ .lBE*24 -
E_
- y. LEE *D4 i 8 4 --
E.uEas4
~
'h.12E*04 T\ N e
m .1E*84 BBC.
60 6. ~
- 2. 580. .lE*B4 .tSE*B4 .2E*B4 .25E*B4 .5E.84 TIME ISEEt
,. BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figure 15.6-34.
RCS Depressurization Transient (3 Inch Break)
i
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e BYRON /BRAIDWOOD [CAE/CBE/CCE/CDE) NOTRUMP SMALL BREAK 3 IN - REDUCED THOT - 17X17 OFA 275 PSIG 56.
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- BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figure 15.6 35.
Core Mixture Height (3 Irxt Break)
f BYRON /BRAlOWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figure 15.6-36.
Clad Temperature Transient (3 Inch Break)
BYRON /BRAIDWOOD (CAE/CBE/CCE/CDE) SB LOCTA
- 3. INCH BREAK - B75' BACKFILL FUEL CLAD AVG. TEMP. HOT ROD 5023.
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Core Power After Reactor Trip
t .
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TIME isECi l '# BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYSTS REPORT t Figure 15.6-41. RCS Depressurization Transient (4 Inch Break)
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FINAL SAFETY ANALYSIE REPORT Figure 15.6-42. RCS Depressurization Transient (6 inch Break) l l
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l I L___________________ ..
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, TIME ISEEl BYRON /BRAIDWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figure 15.6 43.
Core Mixture Height (4 Incil Break)
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Core Mixture Height (6 inch Breek)
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- Ficfre 15.6-44a Core Mixture 11eight (2 Incti BEeak)
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i r SYRON/BRAIDWOOD STATIONS FINAL SAFETY ANALYSl8 REPORT Figure 15.6-45. Clad Temperature Transient (4 Dxt Break) BYRON /BRAIDWOOD iCAE/CBE/CCE/CDE) SB LOCTA 4 INCH PREAK - 275 BACKFILL FUEL CLAD AVG. TEMP. HOT ROD t see2. C
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- TIME (SECl O
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i e r BYRON / BRA!DWOOD STATIONS FINAL SAFETY ANALYSIS REPORT Figure 15.6-46. Clad Temperature Transient (6 inch Break) BYRON / BRA]DWOOD (CAE/CBE/CCE/CDE) SB LOCTA . 6 INCH BREAK - 275 BACKFILL FUEL CLAD AVG. TEMP.HDT ROD t 5000. E
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f Figure 25.6-46a Clad Temperature Tnmsient (2 Int Break) BYRON /BRAIDuGOD (CAE/CBE/CCE/CDE) SB LOCTA
. 2 INCH BREAK - 275 BACKFILL FUEL CL AD AVG. TEMP. HOT ROD t
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4 Y l 1 Figure 15.6-46b Clad Temperature Transient l (3 Iru:t1 Break - Tm. DSS .3
- F)
BYRON /BRAIDWOOD ICAE/CBE/CCE/CDE)-SB LOCTA
- , 3 IMCH Hi-THOT - 275 BACKFILL FUEL
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c APPENDIX B TECHNICAL SPECIFICATION MARK-UPS o Y 0181v:1D-021287
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