ML20116A859
| ML20116A859 | |
| Person / Time | |
|---|---|
| Site: | Wolf Creek  |
| Issue date: | 10/31/1992 |
| From: | WOLF CREEK NUCLEAR OPERATING CORP. |
| To: | |
| Shared Package | |
| ML20116A856 | List: |
| References | |
| NUDOCS 9210300187 | |
| Download: ML20116A859 (399) | |
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'k SAFETY EVALUATION and Y
Cycle 7 Technical Specification Change Report ' for the f Wolf Creek Generating Stationi C October 1992 - Wolf Creek Nuclear Operating Corporation l -. '
,9210300187 921028' PDR ADOCK 050004B2 -
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l 1 Executive Summary The Wolf Creek Nuclear Operating Corporation has prepared this Safety Evaluation for approval of the Technical Specification changes discussed herein. These Technical Specification changes have been prepared to support the Cycla 7 operation of the Wolf Creek Generating Station (WCGS). The Wolf Creek Generating Station is currently operating with both Standard and VANTAGE SH fuel assemblies. Commencing in Cycle 7, the reload fuel will be upgraded to incorporate the Intermediate Flow Mixer (IFM) grid feature of the VANTAGE SH fuel assembly. In addition, the safety evaluation will also support the following additional changes: Increase in FAH/Fo
. Positive Moderator Temperature Coefficient
. OTAT/OPAT Setpoint Changes
. Decrease in Allowed RCS Thermal Design Flowrate
. Main Steam Safety Valve Setpoint Tolerance Increase
. Shutdown Margin increase in Mooe 5
. Core Operating Limit Report WCGS plans to refuel and operate with the changes noted above. This report summarizes the evaluations that were performed to confirm the acceptable uso of these items. The report contains evaluations related to the following:
. Design Features
. Nuclear Analysis
. Thermal and Hydraulic Design
. Non-LOCA Analysis
. LOCA Analysis L
in addition to the Technical Specification changes noted above, the following conservative assumptions were incorporated into the analysis: 4 increased Rated Thermal Power
. Analysis Over a Range of Temperatures
. 10% Steam Generator Tube Plugging
. Thimble Plug Removal i
I
Executive Summary (continued) The Wolf Creek Nuclear Operating Corporation has undertaken an extensive analysis effort to support the proposed power uprating of the Wolf Creek Generating Station from 3411 MWth to 3565 MWth rated thermal power. This analysis effort remains in process and will be completed later with the submittal to the NRO of the Power Rerate Report. The Power Rerzte Report, upon approval of the NRC, will allow the power increase and the operation over a tempera ture range at a time scheduled to be during CyJe 7 oparation. This plan was outlined during a meeting with the NRC on April 10, 1991. Although the analysis has conservatively incorporated the increase in power assumption and has assumed a range of temperatures, the approval for operation with these two items is not being requested at this time. This will be requested later in the Power Rerate Report. The proposed Technical Specification changes have been evaluated and determined to be acceptable with respect to ali applicable acceptance criteria and that tt:e margin of safety has not been reduced.
Table of Contents Section Title Pace 1.0 Introduction 1-1 1.1 Program Description 1-2 1.2 Purpose 1-8 1.3 Proposed Technical Specification Changes 1-8 1.4 Safety Evaluation 1-14 1.5 Conclusions 1-18 1.6 References 1-20 i 20 Design Features 2-1 ; 2.1 Intermediate Flow Mixer (IFM) Grid 2-1 l 2.2 Fuel Rod Performance 2-1 2.3 Seismic /LOCA Impact on Fuel Assemblies 2-1 2.4 References 2-7 3.0 Nuclear Analysis 3-1 3.1 Introduction and Summary 3-1 3.2 Increase in FaH Limit 3-1 3.3 Increase in Fo Limit 3-2 3.4 Methodology 3-2 3.5 References 3-3 i 4.0 Thermal and Hydraulic Design 4-1 4.1 Introduction and Summary 4-1 ; 4.2 Methodology 4-1 4.3 - Hydraulic Compatibility 4-9 4.4 Core Bypass Flow 4-9 4.5 Effects of Fuel Rod Bow on DNBR 4-10 4.6 Fuel Temperature Analysis _4-11 4.7 Transition Core Effects 4-11 4.8 Conclusion 4-12 4.9 References 4-17 .i ; i
Section Title Pace 5.0 Non-LOCA Accident Analysis 5-1 5.1 Increase in Heat Removal by the Secondary System 5 5.2. Decrease in heat Removal by the Secondary System 5-58 5.3 Decrease in Reactor Coolant System Flowrate 5-117 5.4 Reactivity and Power Distribution Anomalies 5-138 5.5 increase in Reactor Coolant inventory 5-190 5.6 Decrease in Reactor Coolant Inventory 5-218 5.7 Radiological Consequences for Analyzed Accidents 5-250 5.6 References 5-275 6.0 LOCA and LOCA-Related Evaluations 6-1 6.1 Small Break LOCA 62 6.2 Large Break LOCA 6-23 6.3 References 6-66 Appendix A Technical Specification Changes / Markups Appendix B Core Operating Limits Report l i i I ii
1.0- Introduction The Wolf Creek Nuclear Operating Corporation (WCNOC) has prepared this Safety Evaluation for approval of the Technical Specification changes discussed herein. These Technical Specification changes have been prepared to support the Cycle 7 operation of the Wolf Creek Generating Station (WCGS). The Wolf Creek Generating Station is currently operating with both Standard and VANTAGE SH fuel assemblies. Commencing in Cycle 7, the reload fuel will be upgraded to incorporate the Intermediate Flow Mixer (IFM) grid feature of the VANTAGE SH fuel assembly. The IFM grid is described in the ' VANTAGE 5 Reference Core Report," WCAP 10444-P-A [1]. The application to the VANTAGE SH fuel assembly is addressed in Addendum 2-A [2]. Associated with the IFM grid is the necessary use of the WRB 2 critical heat flux (CHF) correlation; therefore this change has also been made, in addition, the safety evaluation will also support the following additional changes: Increase in FAH/Fo
. Positive Moderator Temperature Coefficient
. OTAT/OPAT Setpoint Changes
. Decrease in Allowed RCS Thermal Design Flowrate
. Main Steam Safety Valve Setpoint Tolerance increase
. Shuidown Margin increase in Mode 5
. Core Operating Limit Report WCGS plans to refuel and operate with the changes noted above. Thir report summarizes the safety evaluations that were performed to confirm the acceptable use of these items. Sections 2.0 through 6.0 of the Safety Evaluation provide a discussion concerning the Design Features of the Vantage SH with IFM fuel design, the Nuclear, Thermal-Hydraulic, Non-LOCA Accident, and LOCA Accident Analysis.
The Safety Evaluation utilizes the reload methodology discussed in References 3,4,5, and 6, as well as those listed in section 6.5. Consistent with this methodology, parameters were chosen to maximize the applicability of the analysis for future cycles. In addition to the Technical Specification changes noted above, the following cor'servative assumptions were made in the analyses:
. Increased Rated Thermal Power
. Analysis Over a Range of Temperatures
. 10% Steam Generator Tube Plugging
. Thimble Plug Removal 1-1
- 1.1 Program Description . The Wolf Creek Nuclear Operating Corporation has undertaken an extensive analysis effort to support the proposed power 9 prating of the Wolf Creek Generating Station l from 3411 MWth to 3565 MWth rated thermal power. This analysis effort remains in- l process and will be completed later with the submittal to the NRC of the Power Rerate Report. The Power Rorate Report, upon approval of the NRC, will allow the power j increase cod the operation over a temperature range at a time scheduled to be during l Cycle 7 operation [7]. l in conjunction with the analysis effort associated with the Power Rerate, a number of l other changes were incorporated into the analysis. These changes have been highlighted in the text above and will now be discuesed in more detail. l l
1.1.1 VANTAGE SH with IFM Fuel Design The Wolf Creek Generating Station is currently operating with both Standard and VANTAGE SH fuel assemblies. Commencing in Cycle 7, the reload fuel will be upgraded to incorporate the intermediate Flow Mixer (IFM) grid feature of the VANTAGE SH fuel assembly. A number of Technical Specification changes are related to the upgrade to the IFM design. These changes include:
. Revised Reactor Core Safety Limits
. Revised Critical Heat Flux Correlation
. Revised Burnup versus Enrichment Curve Revised Reactor Core Safety Limits The Reactor Core Safety Limit Curve needs to be revised due D the change in bypass flow associated with the upgrade to IFMs and the removal of Thimble Plugs. These two proposed modifications increese the bypass flow design value from 5.8% to 8.4% core bypass flow. As the Reactor Core Safety Limit Curve is directly related to core flow, thb curve is being revised. The curve is also being transferred to the Core Operating Limit Report (COLR) discussed later.
Revised Critical Heat Flux Correlation The critical heat flux (CHF) correlation is being revised, in the Technical Specification Basis, from the current WRB-1 correlation to the WPC2 correlation for use with IFMs, The discussion describing the change to the WRB-2 correlation is also being moved to the COLR. The analysis has been performed with the WRB-2 correlation except for those analyses which fall outside the range of the WRB-2 correlation. In those cases where the WRB-2 CHF correlation is not applicable, the W-3 CHF correlation continues to be used. 1-2
~ In conjunction with the change in CHF correlation and re analysis, a new correlation thermal design limit for the departure from nucleate boiling ratio (DNBR) was chosen, , This limit, since it is different from that used for the previous analysis, his been revised and included in the COLR. , l Revised Burnuo versus Enrichment Curve j Wolf Creek Generating Station's fifth reload core (cycle 6) introduced the Westinghouse Vantage 5 Hybrid Fuel (V5H) as a mix with Westinghouse Standard Fuel Assemblies (STD). The V5H fuel design utilizes Zircaloy mid-span grids contrasted to the inconel grids in the STO design. The STD design allowed initial ondchments up to 4.5 w/o based on the criticalCy analysis. To support the storage of V5H fuel in the Spent Fuel Pool, criticality analyses were pmformed by WCNOC. The analyses and evaluations performed to support the V5H fuel concluded that spent fuel criticality limits are maintained when storing V5H fuel to a maximum initial enrichment of 4.45 w/o, pruvided that the fuel burnup meets the burnup limits defined in the proposed Technical Specification change. Criticality analysis for the V5H was performed with the CASMO and PDQ code packages. CASMO is a multigroup two-dimensional, transport theory code used for burnup calculations on PWR and BWR fuel assemblies. The CASMO code has been extensively benchmarked to comparisons with experiments where isotopic fuel compositions have been determined. 1.1.2 Increased Peaking Factors (FaH /F o) The Cycle 7 reload design has assumed increased peaking factors, FaH/Fo. The full power FaH Peaking factor design limit will increase from the current value of 1,55 to 1.65. The maximum Fc peaking factor limit will increase from the current value of 2.32 to 2.50. The K(Z) envelopes will be modified. These changes will permit more flexibility in developing fuel management schemes (i.e., longer fuel cycles, improvement of fuel economy and neutron utilization, and vessel fluence reduction). 1.1.3 Positive Moderator Temperature Coefficient A positive moderator temperature coefficient (PMTC) has been assumed in the design of the WCGS Cycle 7 core. A PMTC of +6 pcm/0F from 0% power to 70% power, and then decreasing linearly to O pcm/0F at 100% power, hr.s been assumed in the analysis. The resulting boron concentration increcse aue to the PMTC has been included in the re-analysis. This change will also permit more flexibility in developing - fuel management schemes. 1-3
.i 1.1.4 OTATIOPAT Setpoint To allow'for increased operating margin and flexibility,.the OTAT/OPAT setpoints have been increased. The increased _setpoints have been as=umed in the accident analysis.
The OTAT setpoint has been changed to modify the values of r3, r3, and ra. The OPAT setpoint has been changed to modify the values of K4 , ri, r3, and r4 The current and ievised setpoints for the OTAT and OPAT equations are as follows: , Parameter Current Revised Ka 1.08 1.10 ri 8 sec. 6sec. r3 0 sec. 2 sec. ],' r4 28 sec. 16 sec. These revised setpcints are illustrated in Figure 1-1. l l l.. 1-4 l L ._ _ _ _ _ _ _ _ _ _
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40 565 570 575 580 585 590 595 600 605 610 615 620 625 630 Average Temperature (DEG F) WOLF CREEK FIGURE 1-1 Illustration of Overtemperature and Overpower AT Protection 1-5
1.1.5 Decrease in Allowed RCS Thermal Design Flowrate
'he current reactor coolant system (RCS) thermal design flewrate is 375,000 gpm. The a alysis has been performed with a revised thermal design flowrate of 374,400 gpm.
1.1.6 Main Steam Safety Valve Setpoint Tolerance Increase The analyses have been performed with an increased tolerance on the main steam l safety valves (MSSVs). The current MSSV tolerance is 11%. The proposed MSSV l tolerance is 13%, at which all analyses have been performed. i I 1.1.7 Shutdown Margin increase in Mode 5 - The shutdown margin requirement for Mode 5 has been increased in order to regain margin for the Boron Dilution Event discussed on Section 5.4.6. l l 1.1.8 Core Operating Limit Report Generic Letter 88-16, dated October 4,1988 was issued 'o encourage licensees to , I prepare changes to Technical Specifications re!cted to cycle specific parameters. These Technical Specification changes will relocate cycle-specific parameter limits from Technical Specifications to the Core Operating Limits Report (COLR). The generic letter provided guidance for relocation of certain cycie-dependent core operating limits from the Technical Specifications This would allow changes to the values of core operating limits without prior approval (i.e., license amendment) by the NRC, provided an NRC-approved methodology for the parameter limit calculation is followed. The proposed technical specification changes concern the relocation of several cycle-specific operating limits from Technical Specifications to the COLR. A new definition of the_ COLR will be added to the Technical Specifications. Additionally, certain individual Technical Specifications will be amended to note that cycle-specific parameter limits are contained in the COLR. A COLR paragraph will be added to the Administrative Controls Section (which will replace the Peaking Factor Limit Report). The COLR will be required to be submitted to the NRC to allow continued trending of the cycle-specific parameters. , The proposed changes will reference the COLR for specific parameters and will ensure that cycle-specific parameters are maintained within the limits of the COLRL The cycle-specific parameter limits proposed for relocation to the COLR as part of this license amendment request are listed in Appendix B. 1-6
o 1.1.9 Additional Assumptions incorporated into the Analys!s
. ,- 4 Additional assumptions which are not directly related to Technical Specification changes were included in the accident analysis. These include:
. Increased Core Power
. Analysis Over a Range of Temperatures
. Increased Steam Generator Tube Plugging
. Thimble Plug Removal 1.1.9.1 Increased Core Power All transients which were analyzed at 100% power correspond to the uprated core i power level of 3565 MWth. This is a conservative assumption for the accident '
analysis. 1.1.9.2 Analysis Over a Range of Temperatures An analysis over a range of hot leg temperatures was performed in the accident analysis, and extended over a range of 16.8 0F. The accident analysis presented in sections 5.c d 6.0 are to support the technical specification changes discussed in section 1.0, but the temperature range analysis is being presented as it was an analysis assumption for incorporation later in the Poiver Uprate Report. 1.1.9.3 Increased Steam Generator Tube Plugging The accident analysis has been performed assuming 10% plarn total steam generator _ i tube plugging - not to exceed 10% in any single steam generator. This assumption has been made for thoce events in which the 10% tube plugging is conservative. Those events in which the assumption of 10% tube plugging is not conservative, the increase ; in heat removal events, have continued to assume 0% steam generator tube plugging. 1.1.9.4 Thimble Plug Removal Thimble plug removal affects the core pressure drop and increases the core bypass - i flow. These assumptions have been incorporated into the accident analysis and the Thermal-Hydraulic analysis discussed in section 4.0. l 1-7 L
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1.2 Purpose Since the power uprate is scheduled to be approved and implemented during the-T middle of Cycle 7, a number of the changes listed above require approval prior Cycle 7 startup. Therefore, this Technical Specification Change Package has been prepared (prior to the upcoming Power Rerate Report) to allow the NRC review and approval of the changes prior to Cycle 7 startup. The purpose of this report is to provide the basis for proposed Technical Specification changes and demonstrate that continued safe operation will be achieved with the proposed changes. For the purpose of clarification, the term "uprating"in this report refers to the increased power level, while the term . ting" refers to the power level increase as well as the other changes and analysir. ,umptions incorporated into the entire program. Also, since all analyses included in this report have been conservatively re-analyzed with a core power of 3565 MWth,100% power refers to this core power level of 3565 MWth. 1.3 Proposed Technical Specification Changes Table 1-1 presents a lis' of the proposed Technical Specification changes. The changes have been included in Appendix A. b 1-8
s. r Tatie 1-1 Summary of Technical Specification Changes for Cycle 7 Page Section Descnotion of Change Justification h Table of Contents Several changes associated with Several items moved from the T.S. to the COLR . the COLR 1.10 COLR Definition Definition of Core Operating Limits Report 1-2 required 2.1.1 Reference to Figure 21-1 changed Figure 2.1-1 Moved to the COLR 2-1 to refer to the COLR 2-2 Figure 21-1 Revised Reactor Core Safety Limits The new limits result from an increase in the Moved to the COLR reactor coolant core bypass flow due to the VANTAGE SH with IFM fuel design. Table 2.2-1 Revised TA, Z, and ,2 terms OTAT Setpoint Optimized for increased 2-4 operating margin. Table 2.2-1 Loop Design Flow Revised Thermal Design Flow incorporated into 2-4 the analysis Table 2.2-1 Revised r terms OTAT Setpoint Optimized for increased 2-7 operating margin. Table 2.2-1 Revised fj(AI) terms OTAT Setpoint Optimized for increased 2-8 operating margin. 2-9 Table 2.2-1 Revised K4, r terms OPAT Setpoint Optimized for increased ) operating margin. 1 _ - - - . .. -
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Pace Section pescription of Chance Justification 2-10 Table 2.2-1 Revised Allowable Value OPAT Setpoint Optimized for i.w ased I operating margin. B 2-1 2.1.1 Basis CHF correlation and Fag This reflects the DNB co.Te8ation used - limit moved to the COLR for Cycle 7 design and COLR changes. B 2-5 2.2.1 Reference to Figure 2.1-1 changed Ficure 2.1-1 Moved to the COLR to refer to the COLR 3/4 1-3 3.1.1.2 Shutdow.; margin increase in Mode 5 increased margin necessa . for Boron Dilution Event flux doubling concem. 3/4 1-4 3.1.1.3 PMTC of +6 pcm/0F PMTC designed into Cycle 7. Section removed i Moved to the COLR and placed in COLR Cycle specific items removed from Technical Specifications.- 3/4 1-5 4.1.1.3 MTC 300 ppm surveillance limit Limit moved to the COLR : Moved to the COLR ; 3/4 1-14 3.1.3.1 Reference to Figure 3.1-1 changed Figure 3.1-1 Moved to the COLR to refer io Specification 3.1.3.6 3/4 1-20 3.1.3.5 Shutdown Rod insertion Limit Limit moved to the COLR Moved to the COLR 3/4 1-21 3.1.3.6 Control Rod Insertion Lisoits Limit moved to the COLR ; Moved to the COLR ; 1-10
4 Pace Section - Description of Chance Ju_stification 3/4 1-22 . Figure 3.1-1 Rod Bank insertion Limit Versus Figure moved to the COLR Thermal Power, Four Loop Operation Moved to the COLR 3/4 Z ; 3.2.1 Axial Flux Limit Specification refers to the COLR Moved to the COLR 3/4 2-3 Figure 3.2-1 % Of Rated Thermal Power vs. Cyde 7 design change Flux Differertce (DI) Moved to the COLR 3/4 2-4 3.2.2 Fo increase from 2 32 to 2.5 Increase allowed due to .wanalysis. Sedion removed &nd placedin COLR Cyde specific items removed from Technical Specifications. 3/4 2-5 Figure 3.2-2 K(Z) vs. Core Height Figure changes due to increased Fo ' Moved to the COLR Figure removed and Placed in the COLR 3/4 2-6 4.2.2.2 Revised the surveillance from Fxy This change was made due to the BYEC 3/4 2-7 4.2.2.3 to Fo monitoring methods. + Section removed and placed in COLR Cyde specific items removed from Technic 31 ; Specifications. 3/4 2-8 3.2.3 Fxy increase from 1.5 to 1.65 increase allowed due to re-analysis Moved to the COLR 3/4 ^ 3 Figure 3.2-3 - RCS Tetal Flow Rste Versus R Analysis incorporates revised thermal Moved to the COLR design flow . i 1-11
J, . Pace Section Description of Chance Justification
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3/4 2-14 3.2-5 includes RCS P.awrate as a DNB Refers to the COLP for RCS Flowrate 3/4 2-15 Table 3.2-1 Parameter and associated changes since Figure was meved to the COLR , to the COLR 3/4 7-3 Table 3.7-2 Steam Line Safety Valve Lift Setting Analysis incorporates setting tolerance cha'ge Tolerance Change ' t 3/4 9-1 3.4.9.1 Boron Concentration Limit moved to the COLR t Moved to the COLR [ 3/4 9-16 Figure 3.9-1 Minimum Required Fuel Assembly Bumup This reflects the fuel design cha ige to as a Function of Initial Enrichment VANTAGE SH with IFM. 3/4 10-2 4/4.10.2 Group Height, insenion, and Power Changes associated with the COLR Distribution Limits 3/4 10-4 3/4.10.4 Reactor Coolant Loops Changes associated with ti.e COP A B 3/4 1-2 3.4.1.1.3 Moderator Te:5rature Coefficient Limit moved to the COLR surveillance limit 8 3/0 2-1 3/4.2 Basis Deleted definition of Fxy This change is made due to the change from Fxyto F o B 3/4 2-4 3/4 2.2 Basis Revised discussion on DNBR limits This cha ige is made Jue to the 3/4.2.3 Basis and revised discussion on VANTAGE SH with IFM fueldesign surveillance from xyF to F o and Fo surveillance. 1-12 _ __~
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J Pace Section Description ,)f Chance Justirication B 3/4 24 3/42.5 DNB limit moved to tlk COLR Re-analysis performed and specired in the COLR B 3/4 34 3/4.3.32 . Movable incore Detectors Change assooated with Fo rnorutoring Fo monitoring 5-6 5.3 Maxir,A Fuel Enrichment Reflects the allowab:e ennchment for the fuet 5 as a Function of initial Enrichment design change to VANTAGE SH with IFM. 5-8 Figure SE1 Minimum Required Fuel Assembly Bumup This reflects the fuel design change to as a Func e of initial Enrichment VANTAGE SH mth IFM. 6-21 6.9.1.9 Pavised discussr o on the Radia! This change reflects the change to Peaking Factor Limit Report a COLR 1-13
1.4 Safety Evaluation 1.4.1 Ba$ ground This sofoty evaluation addressos the up0rado to the VANTAGE SH with IFM fuel design and the additional changes discussed in section 1.0 of this report. 1.4.2 Dasis The changes discussed in Section 1.0 represent changes to the plant as described in the Technical Specifications and the USAR. These changes have boon ovaluated to datormino if an unroviewed safety question exists. The conclusion that there is no unroviewed safety question involved is based on the consideration of the following criteria:
- a. Fuel damago during Condition 1 (normal operation) or Condition ll (Incidents of Moderate Frequency) ovents.
- b. The reactor can be brought to a safo state following a Condition lll (Infrequent faults) ovent with only a small fraction of fuel rods damaged, although sufficient fuel damage might occur to preclude resumption of operation without considerable outage time.
- c. The reactor can be brought to a safe ctato and the core can be kept suberitical with acceptable heat transfer geometry following transients arising from Condition IV (l.imiting faults) ovents.
To satisfy the above requirements, the following criteria have boon estab!ished.
- 1. Departure from Nucloato Boiling (DNB) Design Basis There will be at least 95% probability that DNB will not occur on the limiting fuel rods during normal operation and operational transients and any transient conditions arising from faults of moderato frequency (Condition I or 11 ovents), at a 95% confidence level.
- 2. Fuel Temperature Design Basis During modos of operation associated with Condition I and Condition ll ovents, there is at least 95% probability that the peak kw/ft fuel rods will not exceed the UO2molting temperature at the 95% confidence level.
1-14
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- 3. Reactor Coolant System Pressure Peak RCS pressure is not to exceed 110% of the design pressure duing Condition 11, Ill, and IV events.
- 4. Loss of Coolant Design Bases (10CFR50.46)
These criteria are discussed in Section 6.0 The compliance to these requirements is demonstrated in this safety evaluation. ; NON LOQA Evaluation Non LOCA safety evaluations were performed to assess the affect of the VANTAGE SH with IFM upgrade and additional changes specified in Section 1.0 on the affected transients presented in Chaoter 15 of the USAR. Each affected transient was re-analyzed. The assumptions included in the evaluation are discussed in Sections 5.0. LOCA Evaluations This evaluation is discussed in Section 6.0 Radioloalcal Conseauences of Accidents The impact on the radiological consequences of accidents has been evaluated taking into account the changes discussed in Section 1.0 above. A discussion of this evaluation is presented in Section 5.7 of this report. 1.4.3 Evaluation The information presented in the report following this section forms the basis for this safety evaluation
- 1. The probability of occurrence of an accident previously evaluated in the USAR is not increased.
The features of VANTAGE SH with IFM which are different from Standard fuel and VANTAGE SH fuel are included as modifications to the analysis performed. The analyses performed, see Sections 3.0,4.0,5.0, and 6,0, have also included the items discussed in Section 1.0 of this report. The probability of occurrence of sn accident is not increased. The changes and additional assumptions ir,corporated into the safety analysis discussed in Section 1.0 are not initiators of any accident and therefore do not affect the probability of occurrence. 1-15
- 2. The proposed changes do not increase the consequences of an occioent previously evaluated in the USAR. Consequences of an accident or malfunction of equipment important to safety previously evaluated in the safety analysis report are not increased.
The evaluation of the impact of the proposed changes specified in Section 1.0 is discussed in Sections 2.0 through 6.0 of this report. The evaluation addressed a full core of VANTAGE SH with IFMs, as well as transition cores consisting of VANTAGE SH and Standard fuel. An evaluation of the radiological consequences of the changes was included as part of the evaluation and the results are discussed in Section 5.7. These proposed changes have had the radiological consequences evaluated for the following events:
. Main Steamline Failure outside Containment
. Loss of Non-Emergency AC Power Accident
. Locked Rotor Accident
. RCCA Ejection Accident
. Failure of Letdown Line Outside Containment
. Steam Generator Tube Rupture
. Lon Of Coolant Accident
. Fuel Handling Accident
. Waste Gas Tank Rupture
. . Liquid Radwaste Tank Rupture The results of the analysis demonstrate that the consequences are well within 10 CFR 100 guideline values.
3, The probability of occurrence of a malfunction of equipment important to safety previously evaluated in the USAR is not created. Mechanical evaluations have been performed on the fuel assemblies. The mechanical design changes associated with VANTAGE SH with IFM fuel, (as discussed in Section 2.0), result in the capability for relaxation of analytical input parameters such that increased margin can be generated without violation of . any acceptance criteria. This margin can then be applied towards relaxation of operational limits. In each case, the appropriate design and acceptance criteria are met. No new performance requirements are being imposed on any system or component in order to tupport the revised analysis assumptions. Subsequently, overall plant integrity remains consistent with that established by the original licensing basis. 1-16
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- 4. There is no increase in the consequences of a malfunction of equipment
, important to safety previously evaluated in the USAR.
The impact of the changes on the safety analyses has been evaluated and the results indicate that safety limits continue to be met as detailed in Section 5 and 6.
- 5. The proposed changes do not create the possibility of an accident of a different type than previously evaluated in the USAR.
The changes discussed in Section 1.0 do not create any new accident scenarios. The impact of the changes on the accidents have been discussed in Sections 5.0 and 6.0.
- 6. The proposed changes de not create the possibility of a different type of malfunction of equipment important to safety than previously evaluated in the USAR.
Continued performance of equipment important to safety will be achieved. The changes do not alter the manner in which equipment responds once required to actuate. The changes will not degrade the performance of any safety system assumed to function in the accident analyses. Therefore the possibility of malfunctions is not increased.
- 7. The margin of safety as defined in the basis for any technical specification is not reduced.
The margin of safety in the plant licensing basis which is affected by the upgrade to VANTAGE SH with IFM fuel and associated changes discussed in Section 1.0 is defined in the Bases of the Technical Specifications. These Bases and the supporting technical specification values are defined by the accident analyses which have bee performed to conservatively bound the operating basis defined by the technical specifications and to demonstrate meeting the regulatory acceptance limits, I The changes discussed in Section 1.0 were evaluated against the applicable acceptance criteria, discussed above. The performance of analyses for the upgrade to VANTAGE SH with IFMs and other changes discussed in Section 1.0 have confirmed that the operating envelope def:ned by the Technical Specifications continues to be bounded by the analyses, which in no case exceeds the weeptance limits. Therefore, the margin of safety provided by the 1-17
, . - ~ . . . . . _ _ _ . _ _ _ _ _ _ _ . _ . . . _ . _ . . .- ___ _ _
Fi i C-analyses in accordance with these acceptanco limits is maintained and not reduced.
-s 1.5 Conclusions ,
it is concludod from this evaluation that the changes addressed in Section 1.0 does not constitute an unreviewed safety question., The results of evaluation / analysis described herein lead to the following conclusions:
- 1. The fuel assemblies ccntaining VANTAGE SH with IFM fuel design are mechanically compatible with 1e current fuel assemblies, control rods, and '
reactor internals interfaces. The changes in the nuclear characteristics due to the transition to upgraded fuel will be within the range normally seen from cycle to cyclo due to fuel management. The roload upgraded fuel assemblies are hydraulically compatible with the fuel assemblies from previous reload cores.
- 2. The changes in the design full power FaH/Fo limits (with appropriate treatment of uncertainties)is supported by design basis analyses summarized in this evaluation.
- 3. The change to a Positive Moderator Temperature Coefficient of +6 pom/0F has been incorporated into the accident analysis and been found to be acceptable.
- 4. The OTATIOPAT-setpoint changes, the decrease in allowed RCS thermal design flowrato, and the chango in the main steam safoty valve setpoint tolerance are acceptable based on the results of the analysis presented in this report.
- 5. The shutaown margin increase in mode 5 has been incorporated into the analysis and found to be acceptable.
- 6. The changes associated with the Core Operating Limit Report are in response to the NRC Generic Letter 88-16 which encourages licensees to preparo changes to Technical Specifications related to cycle-specific parameters, and are therefore acceptable.
1-18
7 The analyses have shown that all or any combination and pattern of thimble plugs may be removed from the Cycle 7 core and subsequent cores, and the tube plugging in any one steam generator may be increased to 10%. 1-19
1.6 References
- 1. Davidson, S. L., ed., et al.,' VANTAGE 5 Fuel Assembly Reference Core Report,"WCAP 10444 P A, September 1985.
- 2. Davidson, S. L., ed., et al., ' VANTAGE SH Fuel Assembly," WCAP 10444 P-A, Addendum 2A, February 1989.
- 3. " Qualification of Steady State Core Physics Methodology for Wolf Creek Design and Ana!ysis", TR 91-0018 WO1, E. W. Jackson, et. al., December 1991,
- 4. "Roload Safety Evaluation Methodology for the Wolf Creek Generating Station",
NSAG-007 Rev. O, W. S. Kennamore, et. al., January 1992.
- 5. " Core Thermal-Hydraulic Analysis Methodology for the Wolf Creek Generating Station", TR 90-0025 WO1, W. S. Kennamore, et al., July 1990.
- 6. " Transient Analysis Methodology for the Wolf Creek Generating Station", NSAG-006 Rev. O, W. D. Wagner, et. al., January 1991.
- 7. Letter from D. G. Pickett (NRC) to Wolf Creek Nuclear Operating Corporation,
" Summary of Meeting Held on April 10,1991, with the Wolf Creek Nuclear Operating Lorporation (WCNOC) to Discuss their Proposed 4.5 Percent Power Rerating Program", April 19,1991, 1 20
. ,,m.m_ . _ . . ~ .. . _ y_ _ .. --.m. , , , , .y , . . . . .., , -
2.0 Design Features 2.1 Intermediato Flow Mixer (IFM) Grid Wolf Creek Cycle 7 and subsequent reloads will contain fuel assemblies that incorporato IFM grids. The IFM grid feature is currently part of the licensing basis in other plants and meets all fuel assembly and fuel rod design criteria (Referenco 1). The IFM grid in the VANTAGE SH fuel assembly is an adaptation of the existing VANTAGE 5 IFM grid design to a 0.374 inch OD standard fuel rod. As shown in Figuro 2-1, IFMs are located in the throo uppermost spans between the VANTAGE SH Zircatoy structural grids to promoto flow mixing in the hottest fuel assembly spans. The IFM grids are fabricated of Zircoloy in the same manner as the VANTAGE SH Zircatoy grids but are not intended to be structural members. The IFM grid envelope is slightly smaller than the VANTAGE SH Zircaloy grid. Each IFM grid cell providos four (4) point fuel rod support. The simplified cell arrangement allows the IFM grid to accomplish its flow mixing objective with minimal pressure drop. Tablo 2-1 provides a comparison of the Standard, VANTAGE SH and VANTAGE SH fuel assembly with IFM grids mechanical design parameters. 2.2 Fuel Rod Performanco Tho 0.374 inch OD fuel rod used in the VANTAGE SH fuel assembly with IFM grids is the same as that used in the Wolf Creek Standard and VANTAGE SH fuel assemblies. Fuel performance ovaluations are completed for each fuel region to demonstrate that the design critoria will be_ satisfied for all fuel regions under the planned operating conditions for each reload core. Fuel rod design evaluations are performed using the NRC-approved models in References 2 and 3 and the NRC-approved extended burnup design methods in Referenco 4 to show that all fuel rod design bases are satisfied. 2.3 Seismic /LOCA Impact on Fuel Assemblies A structural integrity evaluation considering the lateral effects of LOCA and seismic loadings has boon performed for the VANTAGE SH fuel assembly with IFM grids. The VANTAGE SH fuel assembly with IFM grids is structurally equivalent to the Standard and VANTAGE SH fuel designs. The main differences between those designs are five low pressure drop (LPD) Zircaloy-4 structural grids for VANTAGE SH and three additional intormediate flow mixer (IFM) grids. The load bearing capability for the LPD and IFM grids under the faulted condition loadings has been analyzed The results indicated that VANTAGE SH grid loads are below the grid strengths. 2-1 ___________m__._-_______m.__. ..__-.-_um__.---..m%__-_a _mm_
Based on the grid load results, the VANTAGE SH grids are capable of maintaining the core coolable geometry under the design safe shutdown earthquake or asymmetric
^
. pipe rupture transient with either all VANTAGE SH fuel assemblies with IFM grids or transition core operations. The VANTAGE SH fuel assembly with IFM grids is structurally acceptable. This is also true for a transition core compesed of both VANTAGE SH and Standard fuel assemblies core configurations. The (, rids of either fuel types will not buckle due to combined impact loads of seismic and LOCA events.
There is no flow channel reduction during a LOCA transient; thus, the coolable geometry requirement is met. The stresses in the fuel assembly components resulting from seismic and LOCA induced deflections are well within acceptable limits. 2-2
1 l Figure 2-1 Fuel Assembly Comparison ! 'r ..
,,,,3 4 3 . ,, ,,,,,
,,,, 1 l
4 17X17 VANTAGE $H FUEL ASSEMBLY WTTH [FM GRIDS i l l j . .. s n i
- 1 sts 192 20 -- - 23 3
~
11 ! 71 1
\] == g : = ; = === g _ -- l. g g 3
17X17 VANTAGE SH FUEL ASSEMBLY l iss on 3 473- -; 131 to --+-- 2 0 8 3 i r.
-=j .-
17X17 STANDARD FUEL ASSEMBLY 2-3 l _ - _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ ~ _ _ _
Table 21 Comparison of Fuel Assembly Mechanical Design Parameters VANTAGE SH Standard VANTAGE SH with IFM Grids Fuel Assembly Overall Length, inch 159.975 159.975 159.975 Fuel Rod Overall Length, inch 152.20 152.20 152.20 Assembly Envelope, inches 8.426 8.426 8.426 Fuel Rod Pitch, inch 0.496 0.496 0.496 Number of Fuel Rods / Assembly 264 264 264 Number of Guide Thimbles / 24 24 24 Assembly Number of instrumentation 1 1 1 Tube / Assembly Fuel Tube Material Zirc-4 Zirc-4 Zirc-4 Fuel Tube Clad OD, inch 0.374 0.374 0.374 Fuel Rod Clad Thickness, inch 0.0225 0.0225 0.0225 Fuel Clad Gap, mit 6.5 6.5 6.5 Fuel Pellet Diameter, inch 0.3225 0.3225 0.3225 Fuel Rod End Plugs Standard Tapered Tapered Fuel Pellet Length 0.387 0.387 0.387 2-4
T Table 21 (Cont) Comparison of Fuel Assembly Mechanical Design Parameters VANTAGE SH Standard VANTAGE SH with IFM Grids Relative Clad Thickness / 1.0 1.0 1.0 Diameter Ratio Relative Moderator / Fuel 1.0 1.0 1.0 Ratio for Assembly Relative UO 2/ Rod 1.0 1.0 1.0 Guide Thimble Material Zirc-4 Zirc-4 Zirc-4 Guide Thimble OD, inch 0.482 0.474 0.474 Guide Thimble Wall 0.016 0.016 0.016 Thickness, inch Grid Material: Inner Mid Grid (6) inconal Zirc-4 Zirc-4 Edges Modified No Yes Yes Grid Material: End Grids (2) Inconel inconel inconel IFM Grids (3) No No Zirc-4 Grid Types Utilized: IFM Grica No No Yes incor'el Mid Grids Yes No No Zircaloy Mid Grids No Yes Yes-Inconel Top & Bottom Grids Yes Yes Yes Inner Spring (Mid Grids) Vertical Non Vertical Non-Vertical l l 2-5
Table 21 (Cont) Comparison of Fuel Assembly Mechanical Design Parameters VANTAGE SH Standard VANTAGE SH with IFM Grids Grid Fabrication Brazed joining Brazed joining Brazed joining loconel Grids of interlocking of interlocking of interlocking stamped straps stamped straps stamped straps Zircaloy Mid Grlo None Laser weld Laser weld joining of joining of interlocking interlocking stamped straps stamped straps Grid / Guide Thimble Attach: Inconel Grids Thimbles bulges Thimbles bulged Thimbles bulged together with together with together with sleeves prebrazed sleeves probrazed sleeves prebrazed Zircaloy Mid Grids None Thimbles bulged Thimbles bulged together with together with sleeves laser sleeves laser prewelded to prewelded to straps - straps IFM Grids None None Thimbles bulged together with sleeves laser prewelded to straps Top Nozzle Removable Removable Removable stainless steel stain less steel stainless steel reduced height reduced height reduced height Compatible with Fuel Yes Yes Yes Handling Equipment 2-6
- , - , _ -- ,-y- ,- -- , , - - , , , --- . _ _ - _ _
._m - -..
2.4 References
- 1. Slagle, W. H. and Ryan, T. L., "Operationa! Experience with Westinghouse Cores,"
(through December 31,1990), WCAP 8183, Revision 19, January 1992.
- 2. Weiner, R. A., et al., " Improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations,"WCAP-10851 P A, August 1988. ,
- 3. Miller, J. V. (ed), " Improved Analytic Model Used in Westinghouse Fuel Rod .
Design Computations," WCAP-8785, October 1975.
- 4. Davidson, S. L., ed., et al.," Extended Burnup Evaluation of Westinghouse Fuel,"
WCAP-10125 P-A, December 1985. L i i 2-7 l 1
- . . . - . . . - - - - - - - - . . . . . . ...--,.-.-.-.--,w.
3.0 Nuclear Analysis . 3.1 Introduction and Summary The effects of increasing the allowable thermal power and increasing the core peaking factor limits on the nuclear design bases and methodologies for Wolf Creek Generating Station are evaluated in this section. The increased allowable thermal power is 4.5% more than the currently licensed power level. The effects of the allowed thermal power and associated fuel and rnoderator temperature changes on core physics characteris'Jes are small and are explicitly modeled in the neutronice models. The specific values of core safety parameters, e.g., power distributions, peaking factors, rod worths, reactivity coefficients, are primarily loadng pattorn dependent. The variations in the loading pattern dependent safety parameters are expected to be typical of the normal cycle to cycle variations for the standard fuel reloads. The increase in peaking factor limits allows more flexibility when developing the fuel management scheme. Specific items concerning the evaluation are given in Sections 3.2 and 3.3. In summary, the increase in allowed thermal power from the current level will not reduce the margin of safety in the current WCGS USAR nuclear design bases. However, the design bases will be modified due to the increases to the peaking factor limits and allowed thermal power. Nuclear design methodology is not affected by the thermal power change, increased peaking factors, or any other of the changes discussed in section 1.0. 3.2 Increase in FaH Limit The limit on the nuclear enthalpy rise hot channel factor, FaH will be: FaH = 1.65(1 + 0.3(1 P)) where, P = THERMAL POWER / RATED THERMAL POWER The increase in the nuclear enthalpy rise hot channel factor limit will allow additional flexibility for fuel management and for determining core loading patterns. The new limit is applicable to standard and VANTAGE SH with IFM fuel. 3-1
i 3.3 increase in Fo Limit The limit on the heat flux hot channel factor, Fo, will take the following form in the l Technical Specifications: l Four Loop Operation Fo(z) < (2,50/P) x (K(z)) fer P > 0.5, and Fo(z) < (5.00) x (K(z)) for P < 0.5 , where, P = THERMAL POWER / RATED THERMAL POWER, and K(z) = The function obtained from the WCGS Core Operating Limit Report for a given coro height location. The increased heat flux hot channel factor limit, Fo, will allow additional flexibility in fuel management and core operation as well as accommodato the increased nuclear enthalpy rise hot channel factor limit. With thn longer cycles and the necessary higher enrichment h that will be used in the future, the radial peaking factor (FAH, discussed in Section 3.2) will increase This increase will result in higher total peaking factors in the WCGS core. 3.4 Methodology No changes to the nuclear design philosophy or methods are necessary becauso of the increased allowable thermal power or the use of increased peaking factors. The reload design philosophy includos the evaluation of the reload core key safety parameters which comprise the nuclear design dependent input to the USAR safety evaluation for . each reload cycle These key safety parameters will be evaluated for each reload cycle, if one or more of the parameters fall outside the bounds assumed in the safety analysis, the affected transients will be re-evaluated and the results documented in the Reload Safety Evaluation (RSE) for that cycle. Analyses of WCGS reload cores are performed in accordance with standard roload methodology, detailed in References 1 and 2, to ensure compliance with the new peaking factor limits. p 32 1
3,5 References 1, " Qualification of Steady State Core Physics Methodology for Wolf Creek Design and Analysis", TR 91-0018 WO1, E. W. Jackson, et. al., December 1991,
- 2. " Reload Safety Evaluation Methodology for the Wolf Creek Generating Station",
NSAG-007 Rev. O, W. S. Kennamore, et. al., January 1992. 3-3
. - . - -.- .- - - - ~ . - - - - - - - . . - _ -
4.0 Thermal and Hydraulic Design 4.1 Introduction and Summary This section describes the calcul6tional methods used for the thermal hydraulic . analysis, evaluation of the departure from nucleate boiling (DNB) performance, and the hydraulic compatibility during the transition from mixed fuel cores to an all VANTAGE SH with intermediate flow mixers (IFM) core. Based on minimal hardware design differences and prototype hydraulic tet ting of the fuel assemblies, it is concluded that the standard (STD), VANTAGE SH, and VANTAGE SH with IFM fuel assembly designs are hydraulically compatible [1]. Table 4-1 provides a summary of the thermal. hydraulic design parameters for the WCOS that were used in this analysis. The thermal hydraulic design for the upgraded fuel product was analyzed for an increase in i the design limit value for the nuclear enthalpy rise hot channel factor (F39) from 1.55 to 1.65. This increase is achieved by removing unnecessary conservatism in the design through the uss of an improved critical heat flux correlation and improved analysis methodologies as described in the following sections. The thermal-hydraulic design criteria and methods remain the same as those presented in the WCGS USAR with the exceptions noted in the following sections. All of the current USAR thermal-hydraulle design criteria are satisfied. 4.2 Methodology The existing thermal-hydraulic analysis of the 17x17 STD and VANTAGE SH fuel used in the WCGS is based on the standard thermal and hydraulic methods and the WRB-1 critical heat flux correlation as described in the Unit USAR and subsequent approved licensing submittals [2]. The DNB analysis of the core containing the 17x17 STD, VANTAGE SH, and VANTAGE SH with IFM fuel assemblies has been modified to incorporat the WRB 2 critical heat flux correlation [3] and the Statistical Core Design (SCD) analysis methodology [4,5). The W-3 correlation and standard methods are still used when conditions are outside the range of the WRB-2 correlation and of the SCD. 4.2.1 Critical Heat Flux Correlations The WRB 2 critical heat flux correlation is based entirely on rod bundle data and takes credit for significant improvement in the accuracy of the critical heat flux predictions over previous DNB correlations. As documented in the VANTAGE 5 Fuel Assembly-l Reference Core Report [3), the WRB 2 correlation was found to predict the CHF test l data witn a mean measured / predicted ratio of 1.0051 with a standard deviation of l 0.0847 over 684 test points.- The approved 95/95 limit for DNBR for standard, VANTAGE SH and VANTAGE SH with IFM fuel assemblies is 1.17. 4-1 l L
The NRC safety evaluation report for VIPRE-01, stated that "use of a CHF correlation which has been previously approved for application in connection with anoiner thermal ydraulic codo other than VIPRE-01 will require an analysis showing that, given the correlation data boce, VIPRE-01 gives the same or a conservative safety limit."[13). This section presen ~ a discussion of the qualification effort for the Westinghouse WRB 2 critical hea! Nx correlation 95/05 design limit in the VIPRE-01 core thermal-hydraulic analysis code. Qualification of a CHF corrolation for uso in a thermal-hydraulic analyfis code other than the tode usnd to develop the correlstion requires that analyses be performed which dompnstrates that given the modeling philosoph and correlation set applied, the now codo will yield conservative results. This was er /mplished for the WRG 2 critical heat flux correlation and the V! PRE-01 code by an aos9in of the entire data set used by Westirighouse in the original development of WRB 2 with the THINC code The VIPRE specific resu!:s were then used to calculate the WRB 2 design limit which will insure protecnon of at least 95 percent of the fuel at a 95 percent confide,1ce level.
'The calculation of the WRB-2 design limit in VIPRE which provides 95/r 5 protection is a function of the mean of the measured / predicted critical heat flux ratics, the standard deviation of the measured /prodicted ratios, and Owen's one-sided tolerance factor for the data set. The expression for the correlation 95/95 design limit is given by; DNHRa =
MiP-K w ,,*a,, Where, DNBRd = CHF correlation 95/95 design limit. M/P = Mean Measured / Predicted Ratio Ky ,y = Owen's one sided tolerance factor. The factor is a function of the number of members in the correlation data set, the confidence level required, and the percentito protection required. cr,,, = Standard deviation of the measure / predicted ratios Results from the VIPRE-01, Mod 1 analysis, summarized in Table 4 2, indicate a mean measured / predicted ratio of 1.022377 and a measure / predicted standard deviation of 0.079859. After adjusting for the degrees of freedom, the standard deviation of the measured / predicted ratios from the VIPRE-01 code increases to 0.080629. The Owon's one sided tolerance factor for 95% protection at a 95% confidence level is 1.75 for a data set population with 684 members. Thus, the results of the VIPRE-01, Mod 1 analysis of the WRB-2 correlation data set indicate a 95/95 design limit of; DNBR". = 1.022377 -1 75(0.080629) 4-2
l DN#Ra = 1.14 However. Westinghouse repor1ed the design limit for WRB 2 at 1.17 for use with the THINC code. The wording of the Safety Evaluation Report on the VIPRE-01 code issued by the NRC requires the t.sers of CHF correlations developed for codes other than VIPRE-01 use the same or a more conservative 95/95 design limit. Therefore the correlation design limit to be used for WRB 2 in the VIPRE-01 code is 1.17. Examination of the WRB 2 qualification in VIPRE-01 results reveals two significant points. First, the mean measured / predicted ratio from the VIPRE-01 data of
- appror.imately 1.022 indicates that the WRB 2 correlation, as implemented in VIPRE-01, .1!ightiv under predicts the experimental critical heat flux as compared to the results originelly repoAd. This is a conservative result. Secondly, the adjusted standard deviation of 0.08W20 frem the VIPRE-01 qualification results indicates that the under predicting of experimental critical heat flux is consistent across the entire CHF database. Therefore, use of an correlation 95/95 design limit of 1.17 in thermal-hydraulle analyses for the Wolf Creek Generating Station will yield conservative core thermal-hydraulic designs.
The W 3 correlation is used below the first mixing vano grid with a 95/95 limit DNBR of 1.30. The W-3 correlation is also used for the steamline break analyses with a 95/95 limit DNBR of 1.45 in the pressure range cf 500 to 1000 psia [6). 4.2.2 Statistical Core Design Statistical Core Design (SCD) is a thermal hydraulic analysis technique that provides an increase in core thermal (DNB) margin by tre:: ting core state and bundle uncertainties statistically. WCNOC has applied the Babcock & Wilcox Fuel Company SCD method to WCGS te determine DNB limitations in reactor core designs [4,5), Energy produced in the fuel of a nuclear reactor is removed from the surface of the fuel cladding by the coolart flow. Under nominal conditions, the heat transfer mechanism is highly efficient nucleate boiling. Heat transfer coefficients under these conditions are typically around 50,000 Blu/Hr-ft 2.op, Conditions adverse to heat transfer from the cladding (i.e. increased heat flux, decreasing pressure, decreasing flow, etc.) degrade the ability of the coolant to accept heat from the clad surface. Eventually, the cooiant flow past the fuel may reach conditions which result in the formation of a continuous layer of steam around the fuel rod. This phenomena is accompanied by a dramatic reduction in the heat transfer coefficient to around 500 Btu /Hr-ft 2 *F. The heat transfer coefficient falls because under these conditions, heat transfer is principally accomplished by conduction through the insulating layer of steam. This phenomena is defined as film boiling heat transfer, Figure 41 defines the various modes of heat transfer which can occur in a reactor core. As indicated, transition to the film boiling region of heat transfer must be . 4-3
~ . _ _ _ - - _ -. .. _
l J accompanied by a dramatic increase in the clad temperature to maintain the same heat transfer rate. This temperature increcse ca i be so severe as to cause fuel damage. The heat flux et which steam film starts to form around the fuel rod is known as the critical heat flux (CHF) or the point of departure from nucleate boiling (DNB). The departure from nucleate boiling ratio (DNBR), which is defined as the ratio of predicted CHF to the actual heat flux, provides a measure of the thermal margin to film boiling in the core. The greater the DNBR is above 1.0, the greater the thermal margin. Figure 41 Definition of the Bolling Curve s y . x ... ' 5.v .i . , l t,.-.+.. l ni , r . . .. c.~.w.o l %. . i . somn, l e.mo, esi.ng 6
$.$s$.'# l i
s r CHF ,
. i
- . l l ; l
$ , l l .
l u l l . l , S l . : ! l l . l l
+
l l l
/ ! ! ; !
! l l l T
OND CHF MFB Log Wou Svoetheat ( T w-TSat ) Due to the complexity of the transition from nucleate boiling to film boiling phenomena,- the critical heat tiux must be calculated from empirical correlations. These CHF correlations correlate the critical heat flux as a function of the local thermodynamic conditions and the bundle geometry. The CHF correlations are developed from experimental data obtained with electrically heated test sections which incorporate geometries and conditions typirst of those found in a reactor core. Thus, application of a CHF correlation is limited by both the range of thermal hydraulic conditions and the test bundle geometries used to develop the correlation. A CHF correlation is essentially a least squares fit to'the data obtained in the critical l-heat flux tests. As such, each correlation has an associated uncertainty. The correiation uncertainty is quantified for use in design analyses by defining a correlation 4-4 L l' L _
l l I design limit. This design limit is defined as the departure from nucleate boiling ratio at which there remains at least a 95% probability at the 95% confidence level that a departure from nucleate boiling will not occur on a specific fuel pin. Core DNBR calculations are performed utilizing the VIPRE 01 thermal-hydraulic analysis code. This code calculates DNBR in the core for a given set of core state variables (i.e. power, flow, pressure, etc.). The DND analysis performed with the thermal-hydraulic code is complicated by both the uncertainty associated with the DNB correlation and the uncertainties on the values of the core state variables. The traditional design philosophy for protection from DNB in nuclear reactor cores has folluwod an extremely conservative approach. Essentially, core thermal hydraulle design analyses were performed with all core state point variablos (i.e. power, flow, pressure, inlet temperature, radial peaking factors, axial peaking factors, axial peak elevation, etc.) assumed to be at their most adverse values with respect to DNB. Thus, design analyses were performed with the core power assumed to be at the upper bound of the associated uncertainty, the flow was assumed to be at the lower bound of the measurement uncertainty, and pressure was assumed to be at the lower bound of the measurement uncertainty. This philosophy was applied to each of the core state variables and the resulting DNB ratio was then compared to the 95/95 desigrilimit for the critical heat flux correlation. The compounding of uncertainties on the CHF correlation and the core state variables ' is unnecessarily restrictive. The simultaneous occurrence of all core state variables at their most detrimentallimit is not realistic. A mora realistic approach can be developed in which the occurrence of the core state points at both their detrimental and beneficial limits are considered. This is accomplished using a statistical method termed
" Statistical Core Design"(SCD).
In the SCD methodology, the important variables, their uncertainties, and their distributions are identified and propagated through a model to establish the overall uncertainty on the calculated DNBR. The uncertainty on the DNBR is then used to establish a stativ.ical design limit (SDL) which replaces the original CHF correlation 95/95 design limit. The SDL is by definition higher than the CHF correlation design limit since it includes uncertainties on the correlation as well as the uncertainties on the core state variables treated. The benefit of this approach is that the thermal-hydraulic analyses are then run using the nominal values for the core state variables treated in the development of the statistical design limit Results from these nominal condition analyses are then compared to the statistical design limit, rather than the correlation design limit, to determine the thermal marg)n for a given core state. The statistical core design methodology consists of four steps. These steps are;
- 1) Selection of core state variables and their associated uncertainties.
- 2) Development of a response surface model(RSM).
.45
- 3) Propagation of uncertainties through the RSM to establish the overall uncertainty on DNBR.
- 4) Application of the SCD to establish a statistical design limit.
The critical heat flux, and consequently the departure from nucleate boiling ratio, is not a directly observable parameter in the reactor core. DNBR must be inferred from . observable core state variables and then calculated with a critical heat flux correlation. l These core state variables may be classified as either global (i.e. core wide) or local (i.e. bundle or pin specific) variables. A total of five core state variables were treated in the WCGS Statistical Core Design for cycle 7. These are the core power, core flow, system pressure (i.e. the core outlet pressure), tha core inlet temperature, and the radial peaking factor. The first four of these core state variables are considered global variables and the remaining factor classified as a local variable. Each of these five core state variables may be either directly measured or inferred during operation of the reactor. In addition, these five variables completely define the thermal hydraulic environment when they are used to define the core state for the VIPRE-01. The directly mecsurable parameters, such as the system pressure or core inlet temperature, are associated with well defined uncertainties. These uncertainties are generally the result of instrumentation errors. The uncertainties on the inferred parameters, such as the radial peaking factor, has uncertainties both on the secondary instrumentation used to make the inference and on the calculational methods employed. Following selection of the core state variables und identifying their associated uncertainties, the overall uncertainty on the DNBR is established. This is accomplished by maki,sg a large number of DNBR calculations in which the core state variables are allowed to randomly vary (within their associated uncertainty ranges) about their nominal values. Thus, one DNBR calculation may allow the power to be at the upper limit of the power uncertainty while the remaining four core state variables are at their nominal conditions. A subsequent calculation may allow the power to be at the upper uncertainty bound while the pressure and temperature are at their respective lower bounds. This method is known as a Monte Carlo propagation of uncertainties. If a sufficient number of these DNBR calculations are performed, the overall distribution of DNBR about a nom!nal, or expected, value may be determined. This DNBR distribution is then used to define the DNDR uncertainty based upon the uncertainties of the individual coLs state variables. The determination of the DN.8R distribution requires a large number of DNBR calculations. The methodology employed in this analysis requires 3000 propagations to estet;lish the DNBR distribution. While the necessary data could be obtained by making 3000 VIPRE-01 runs, it is more efficient to establish a response surface model ' for use in performing the required DNBR calculations. The RSM is essentially an equation, derived from a least squares fit of VIPRE-01 results, which predicts DNBR as a function of the core state variables. The RSM developed for the WCGS is a fit of the 46
of the core state variables. The 3000 Monte Carlo uncertainties through the response surface model equation can be accomphsh
- quickly and with thermal-hydraulic analysis less expense than attempting to perform the propagations w code.
The RSM is developed using statistical methods in which " experimental" data obtained by running VIPRE-01 test cases representing a test matrix. The state variables covered in the test is chosen based on an experimental desi technique termed " central composite design" The coefficients of the respo model equation are then optimized to the DNB ratios produced by the VIPR important to realize that the primary purpose of the RSM equation is the accu propagation of uncertainty; not the absolute calculation of DNBR. The RSM is not tool for predicting the impact of changes in the core the RSM must be developed so that it provides a reasonable prediction of excellent propagation of the core state variable uncertainties. Establishing the overall uncertainty on DNBR allows a statistical calculatio critical heat flux correlation 95/95 confidence limit. This limit, termed the statist design limit (SDL), replaces the original correlation limit and incorporates the correlation uncertaintles. Necessarily, the SDL is highe correlation design limit. Subsequent thermal-hydraulic analyses with the VIPRE-01 code are then r core state variables at their nominal values. The resulting DNBR predictions are pin in a core is calculated to be at the SDL, the actual D probability at a 95% cnnfidence level of being above 1.0 (i.e. onset of film Less limiting pins in the core will have correspondingly higher protection ag boiling. An additional benefit of the Statistical Core Design methodology is that it all the number of pins in the core in danger of entering DNB for nny given pe distribution to be quantified. The SDL is required to provided hot pin protectio 95/95 confidence level and further that 99.9 percent core wide protection is The uncertainties on the axial peak and the location of the axial peak were in the determination of the cycle 7 statistical design limit due to difficult with intermadiate ilow mixers (IFM). This type of fuel, w upper part uf the bundle, exhibits a two step DNB response. THs is an effect of the increased turbulent crossflow in the upper region of the fuel bundle caused grids. Thus, the uncertainties on the axial peak and the location of the axial p 4-7
, _ _ ------"--____,____.___m-------------^~ -
fiv9 core state variables and contains linear, cross, and quadrat
- arms involving each of the core state variables. The 3000 Monte Carlo propagations c: core state variable uncer inties through the response surface model equation can be accomplished quickly and with less expense than attempting to perform the propagations with a therma.aydraulic analysis code.
The RSM is developed using statistical methods in which " experimental" data is obtained by running VIPRE-01 test cases representing a test matrix. The range of core state variables covered in the test is chosen based on an experimental design technique termed " central composite design". The coefficients of the response surface model equation are then optimized to the DNB ratios produced by the VIPRE runs. It is important to realize that the primary purpose of the RSM equation is the accurate propagation of uncertainty; not the absolute calculation of L'NBR. The RSM is not intended to be a replacement for the thermal hydraulic code, though it can be a useful tool for predicting the impact of changes in the core state variables on DNBR. Thus, the RSM must be developed so that it provides a reasonable prediction of DNBR and excellent propagation of the core state variable uncertainties. Establishing the overail uncertainty on DNBR allows a statistical calculation of a new critical heat flux correlation 95/95 confidence limit. This limit, termed the statistical design limit (SDL), replaces the original correlation limit and incorporates the uncertainties on all of the core state variables included in the analysis as well as the correlation uncertainties. Necessarily, the SDL is higher than the original CHF correlation design limit. Subsequent thermal-hydraulic analyses with the VIPRE-01 code are then run with the core state variables at their nominal values. The resulting DNBR predictions are then compared to the SDL to determine the margin to film boiling. Thus, when the limiting pin in a core is calculated to be at the SDL, the actual DNBR will have a 95% probability at a 95% confidence level of being above 1.0 (i.e. onset of film boiling). Less limiting pins in the core will have correspondingly higher protection against film boiling. An additional benefit of the Statistical Core Design methodology is that it allows the of the number of pins in the core in danger of entering DNB for any given peaking distribution to be quantified. The SDL is required to provided hot pin protection at the 95/95 confidence leval and further that 99.9 percent core wide protection is provided. The uncertainties on the axial peak and the location of the axial peak were not included in the determination of the cycle 7 statistical design limit due to difficulties in obtaining an acceptable fit for the response surface model introduced by the use of a fuel design with intermediate flow mixers (IFM). This type of fuel, with the IFM grids located in the upper part of the bundle, exhibits a two step DNB response. This is an effect of the increased turbulent crossflow in the upper region of the fuel bundle caused by the IFM grids. Thus, the uncertainties on the axial peak and the location of the axial peak were 47
treated in a traditional manner. This approach war, provided for in the orginial duelopment of the SCD methodology in that core stato variables not treated, or treatad at an inferior level, could still be incorporate ' in the thermal-hydraulic analysis by corrgunding the associated uncertainity in the traditional manner. The axial peak uncertainty, assessed at 2.5%, was applied to the DNBR calculations obtained with the VIPRE-01 code. The uncertainly of the location of the axial peak, resulting from uncertainties in the noding used in the physics analysis codes, was applied by the cora design group during the maneuvering analysis. . The determination of the statistical design limit for the rerate of the WCGS, cycle 7, included measurement and calculational uncertainties on five core state variables; power, flow, pressure, temperature, and radial peaking. The core state variables, the associated uncertainties, and the distributions assumed in the determination of the statistical design limit for cycle 7 are summarized in Table 4-3. The resulting SDL for the WRB 2 critical heat flux correlation for the cycle 7 design was established as 1.31. Generic design margin, which is used to account for various design penalties, is then added to the statistical design limit. This new DNBR limit, called the thermal design limit (TDL)'"en defines the acceptance criteria for all Di48R evaluations. For cycle 7, the generic margins allocated were; Transition Core Penalty 12.0 % Lower Plenum Flow Anomaly 3.0% Rod Bow Penalty 1.5% Axial Peaking Uncertainly Penalty 2.5% Design Mergin 8 22 % Total Generic Margin 27.22 % The thermal design limit for the cycle 7 design is then given by the expression; 7DL = 1.0-%M argin /100 {56-(27.22/100)] 7DL = 1.80 Therefore, analysis results which yield a minimum departure from nucleate boiling ratio-greater than or equal to 1.80 will meet the requirement of 95% protection from DNB at the 95% confidence level. 4-8
The thermal bydraulic design criteria is that the probability that DNB will not occur on the most limiting fuel rod is at least 95%, at the 95% confidence level, for any Condition I or Condition ll evont. ConJervative uncertainty values on the treated core state variables are used in SCD to establish overall unceriainty on DNB [S). SCD analyses use a new flow parameter, termed minimum measured flow (MMF), which is equal to the thermai desiCn flow (TDF) plus the flow uncertainty. Analyses by standard methods continue to use the thermal design flow. 4.3 Hydraulic Compatibility Fuel assembly lift forces are defined as the net upward force acting on the assembly due to interaction with coolant flow, excluding fuel assembly weight and buoyancy. Fuel assembly lift forces are used in the design of fuel assembly hold down springs and reactor vessel internals. Lift forces are calculated at hot full power, cold startup and hot pump overspeed. Designing to these conditions ensures that the hold down spring design criterion is met. The Wolf Creek Generating Station will transition to VANTAGE SH fuel assemblies with IFM grids from a core consist 'g of VANTAGE SH and Standard fuel assemblies Consequently, lift forces for all thres fuel types were evaluated. Based on this _ evaluation if was concluded that the hydraulic load on the hold-dewn spring and the core interr'als for Standard, VANTAGE SH and the VANTAGE 5 fuel assemblies with iFM grids is acceptable. 4.4 Core Bypass Flow Core bypass flow is defined as the total amount of reamtor coolant flow which bypasses ' the core region. This flow is not considered effective in the core heat transfer process. The piincipal bypass flow paths are:
- a. Baffle.' Barrel Reo!on The Daffle/ barrel region consists of vertical baffie plates that follow the perhery of the core. These are joir.ed te the core barrel by horizontal former plates spaced along the elevation of the baffle -
plates. The fraction of total flow that passes upward through this region between the core barrel and baffle plates is considered core bypass flow. This bypass flow fraction is large enough to maintain closure head fluid temperature equal to the reactor vessel inlet temperature. l 4-9 -
- b. Vessel Head Coolino Sorav Nozzles Thnse nozzles are flow paths between the reactor vessel and core barrel annulus and the fluid volume in the vessel closure head-region above the upper support plate. A fraction of the flow that anters the vesse' inlet nozzles and into - the vessel / barrel downcomer passes through these nozzles and into the vessel closure head region.
- c. Core Barrel- Reactor Vessel Outlet Nozzle Gao Some of the flow that enters the vessel / barrel downcomer will leak through the gaps between the cora barrel outlet nozzles and the reactor vessel outlet nozzles and merge with the vessel outlet nozzle flow.
- d. Fuel Assembly - Baffle Plate Cavity Gao This is the core bypass flow path between the peripheral fuel assemblies and the core baffle plates.
- e. Fuel Assembly Thimble Tubes These tubes are physically part of each fuel assembly skeleton and flow within them is partially effective in removing core heat.
However, such flow is analytically not corsidered to be effsetive in heat removal, and is consequently treated as core bypass flow. Thimble plugging devices reduce this component of the overall bypass flow. The design value of overall core bypass flow used in existing safety analyses for the Wolf Creek Generating Station is 5.8%. The change in core bypass flow resulting from the VANTAGE SH fuel assembly with IFM grids was calculated. The result is an increase to the core bypass limit to 6.4% for thin %Ie plugs in place or d.4% with thimble plugs removed. The best estimate bypass flow for the case with thimble plugs removed was determined to be 6.61%. 4.5 - Effects of Fuel Rod Bow on DNBR The phenomenon of fuel rod bowing must be accounted for in the DNBR safety analysis of Condition I and Condition il events. In the IFM region of a VANTAGE SH with IFM fuel assembly, the grid-to-grid spacing is approximately 10 inches compared to approximately 20 inches in the current fuel assemblies in the Cycle 6 core. Using approved methodology [7), the predicted channel closure in the 10 inch spans in the VANTAGE SH with IFM assemblies will be less than 50%. Thus, no rod bow penalty is required in this region. In the spans b3 low the IFM region of the VANTAGE 5H with IFM assemblies and for the resident fuel, rod bow is accounted for in available DNBR margin as summarized in s< aon 4.2.2. 4-10
The maximum rod bow penalties accounted for in the design safety analyses are cased on an assembly average burnup of 24,000 MWD /MTU, as approved by the Commission
-[8]. At burnups greater than 24,000 MWD /MTU, credit is taken for the effect of F3 s-burndown, due to the decrease in fissionable isotopes and buildup of fission product - ,
inventory. No additional rod bow penalty is required. 4.6 Fuel Temperature Analysis The 0.374 inch O.D. fuel rod used in the VANTAGE SH fuel assembly with IFM grids is the same as that used in the Standard and VANTAGE SH fuel assemblies resident in the core. Fuel performance evaluations are completed for each fuel region to demonstrate that the design criteria will be satisfied for all fuel regions under the planned operating conditions for each reload core. Fuel rod design evaluations are performed using approved models [9,10,11]. There is no change in the fuel temperature design criteria used in the safety analysis calculations between the fuel types resident for cycle 6 and the VANTAGE 5H with IFM grids to be loaded for cycle 7. 4.7 Transition Core Effects The fuel to be loaded for cycle 7 has IFM grids located in spans cetween mixing vane grids in the upper region of the fuel assembly. The resident fuel, both the Standard and VANTAGE SH assemblies, does not feature these intermediate grids. The additional grids introduce localized flow redistribution from the VANTAGE SH with IFM assembly into the Standard and VANTAGE SH assemblies at axial zones near the IFM grid positions in a transition core. Between the IFM grids, flow returns to the VANTAGE SH with IFM assemblies due to the tendency for velocity equalization in parallel open channels.. This localized flow redistribution actually benefits the Standard and VANTAGE SH assemblies. This benefit more than offsets the slight mass flow bias due to velocity equalization at non-gridded locations. Thus, the analysis for a full core of these fuel assembly types remains appropriate for that fuel in a transition core. Transition corcs are analyzed as if they were a full core of one assembly type, VANTAGE SH with IFM grids in this application. A transition core penalty is then applied to the thermal-hydraulic design analyse. to account the impact of the flow redistribution. For VANTAGE SH with IFM grids, the transition core penalty is a function of the number of VANTAGE SH with IFM grid assemblies present in the core and is determined using approved methodologies (12). The transition core penalty for Cycle 7 operation has been established at 12.0%. This penalty is included in the safety analysis limit DNBR such that sufficient margin over the design limit DNBR exists to accommodate the transition core penalty along with other appropriate DNBR penalties. 4-11
4.8 Conclusion The thermal-hydraulic evaluation of the fuel upgrade and peaking factor increase for the Wolf Creek Generating Station has shown that 17x17 Standard, VANTAGE SH, and VANTAGE SH with IFM grids are hydraulically compatible and that the DNB margin gained through the use of the Statistical Core Design methodology and the WRB-2 critical heat flux correlation is sufficient to allow in increase in the design F3g from 1.55 to 1.65. More that sufficient DNBR margin exists in the safety limit DNBR to cover the rod box and transition core penalties. All thermal-hydraulic design criteria are satisfied. 4-12
Table 41 Thermal and Hvdraulic Design Parameters -Wolf Creek Generatir.g Stathn Thermal and Hydraulic Design Parameters Parameter Values Reactor Core Heat Output (MWm) 3411 Reactor Core Heat Output (10s BTU /HR) 11,639 Heat Generated in Fuel (%) 97.4 Pressurizer Pressure, Nominal (psia) 2250.0 Design Radial Power Distribution
- 1.65[1.0 + 0.3(1.0 - P))
DNB Correlations" WRB-2 W-3 HFP Nominal Coolant Conditions Vessel Thermal Design Flow Rate, includi.79 bypass, 374,400 (GPM) Vessel Minimum Measured Flow Rate, including 384,000 bypass, based on 2.5% flow uncertainty (GPM) Core Flow Rato, excluding bypass, based on Thermal 342,950 Design Flow (GPM)*" Core Flow Rate, excluding bypass. based on MMF and 358,618 best estimate bypass flow (GPM)*" Core Flow Area (ft2) 51.28 Core inlet Mass Flux, based on TDF (106 lbm/hr-ft2) 2.484 Core Inlet Mass Flux, based on MMF (106 lbm/hr-ft2 ) 2.598 Nominal Vessel / Core inlet Temperatura (oF) 558.8 Vessel Average Temperature ( F) 588,5 Core Average Temperature ( F) 591.8 Vessel Outlet Temperature ( F) 618.2 Core Outlet Temperature ( F) 621.4 Average Temperature Rise in Vessel ( F) 59.4 Average Temperature Rise in Core ( F) 62.6
~
1.65 represent 1.59 plus 4% measurement uncertainity. W-3 is used for conditions outside the range of applicability of the WRB-2 correlation or the Statistical Core Design. Design bypass flow, including uncertainty is 8.4% with thimble plugs removed. Best estimate bypass flow with thimble plugs remove is 6.61%.- 4-13
Table 41 Continued Thermal and Hydraulic Desgin Parameters Heat Transfer Active Heat Transfer Surface (ft2) 59,742 Average Linear Power (kW/ft) 5.44 Peak Linear Power for Normal Operation (kW/ft)* 12.6 Temperature at Peak Linear Power for Prevention of 4700 Centerline Fuel Melt ( F) Based on maximum Foof 2.32. s 4-14 I I
1 Table 4 2 WRB 2 Qualification Results Cateaorv # of Points Sample Mean M/P Standard Deviation Total CHF Points 684 1.022377 0.079859 Axial Heat Flux Uniform 357 1.0327 0.0821 Non-Uniform 327- 1.0111 0.0757 Bundle Arrays 5x5, 0.374" O.D. 501 1.0247 0.0804 5x5, 0.360" O. D. 183 1.0161 0.0781 Subchannel Types Typica! Cell 508 1.02s3 0.0786 Thimble Cell 176 1.0167 0.0831 Heated Lengths 96" (8 ft.) 213 1.0266 0.0835 168" (14 ft.) 471 1.0205 0.0781 Mixing Vane Grid Spacing 10" 82 1.0038 0.0725 20" 101 1.0261 0.0810 22" 282 1.0182 0.0701 20" 219 1.0329 0.0912
.4-15
Table 4 3 Summary of Uncertainties included in SDL Variable Description Uncertainty Dist. O Heat Balance 2.0% Normal W RCS Flow 2.5% Uniform W Bypass Flow 1.79% Uniform P Pressurizer Pressure 30.0 psia Uniform T Temperature Control 4.85 *F Uniform R Radial Peaking - Measurement 5.0% Normal R Hot Channel Factors 3.0% Normal R Initial Bundle Spacing 1.5% Uniform-D WRB-2 Correlation 0.1479 DNBR Normal - D VIPRE-01 Code 5.0% Normal D RSM to VIPRE-01 Fit 4.5% Normal i 4-16
. - - . - . - . - - -.- _ .- . ._ . - . ~ . . - .. _.
4.9 References
- 1. " Vantage SH Fuel Assembly", WCAP-10444 P-A, Addendum 2, Davidson, S. L.
ed et al., April 1988 and Letter from W. J. Johnson (Westinghouse) to M. W. Hodges (NRC), " Supplemental information for WCAP-10444-P-A Addendum 2,
' VANTAGE SH Fuel Assembly"', NS-NRC-88-3363, dated July 29,1988.
- 2. Letter from Reckley, W. D. (NRC) to Withers, B. D. (WCNOC), " Wolf Creek Generating Station - Amendment No. 51 to Facility Operating License No., NPF-42 (TAC No. 79923), dated November 6.-1991.
- 3. " Vantage 5 Fuel Assembly Reference Core Report", WCAP-10444-P A, Davidson, S. L. ed. et al., September 1985.
- 4. " Statistical Core Design for Mixing Vane Cores", BAW-10170 P-A, Farnsworth, D. A. and Meyer, G. A., Babcock & Wilcox Co., August 1987.
- 5. " Core Thtrmal-Hydraulic Analysis Methodology for the Wolf Creek Generating Station", TR-90-0025 WO1, Kennamore, W. S. et al., Wolf Creek Nuclear Operating Corporation, July 1990.
- 6. Letter from Thadeni, A. C. (NRC) to Johnson, W. J. (Westinghouse),
" Acceptance for Referencing of Licensing Topical Report, WCAP-9226-P/WCAP-9227-NP, Reactor Core Response to Excessive Secondary Steam Releases", dated Jr.nuary 31,1989.
- 7. " Fuel Rod Bow Evaluation", Skaritka, J. (Ed.), WCAP-8691, Revision 1, July, 1979.
- 8. Letter from Berlinger, C. (NRC) to Rahe, E. P. Jr. (Westinghouse), " Request for Reduction in Fuel Assembly Burnup Limit for Calculation of Maximum Rod Box Penalty", dated June 18,1986.
- 9. " improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations", Weiner, R. A. et al., WCAP-10851-P-A, August 1988.
- 10. " Improved Analytic Model Used in Westingnouse Fuel Rod Design Computations", Miller, J. V. (Ed.), WCAP-8785, October 1975.
- 11. " Extended Burnup Evaluation of Westinghouse Fuel", Davidson, S. L. (Ed.) et al., WCAP-10125-P-A, December 1985.
- 12. " Extension of Methodology for Calculating Transition Core DNBR Penalties",
Schueren, P ' nd McAtee K. R., WCAP-11837-P-A, January,1990. 4-17
- 13. Letter from Rossi, C. E. (NRC) to Blaisdell, J A. (UGRA), " Acceptance fro Referencing of Licensing Topical Report, EPRI NP-2511-CCM, 'VIPRE-01: A Thermal-Hydraulic Analysis Code for Reactor Cores', Volumes 1,2,3,4", dated May 1,1986.
4
'l 4-18
5.0 Non-LOCA Accident Analysis The impact of the proposed Technical Specification changes and the other assumptions discussed in section 1.0 on the non-LOCA accident analysis is discussed - in the following section. The American Nuclear Society (ANS) classification of plant conditions continues to be used, as in the USAR Chapter 15 (1) which divides plant conditions into four categories in accordance with anticipated frequency of occurrence and potential radiological consequences to the public. The four categories are as follows (2):
- 1. Condition 1: Normal Operation and Operational Transients
- 2. Condition ll: Faults of Moderate Frequency
- 3. Condition lil: Infrequent Faults
- 4. Condition IV: Limiting Faults For this evaluation, as in Chapter 15, the events of Conditions ll, Ill, and IV have been considered here.
Radiological consequences are presented in Section 5.7 for all events that result in a radioactive rel ease to the enviroment. 5.0.1 Analysis Assumptions and Initial Conditions The initial power assumed for all transients which initiate from 100% power is the uprated value of core power equivalent to 3565 MWth. Although the Technical Specification change for the allowance to increase power is not being submitted at this time, the analy. sis has been performed in a conservative, bounding manner. The methodology as described in References 3,4, and 5 have been assumed in this evaluation. The nominal values of initial conditions are assumed in most accidents which are DNB limited. The allowances on power, pressure, temperature, and flow are determined on a statistical basis and are included in the limit DNBR, as described in Reference 4. This procedure is known as " Statistical Core Design"(SCD) and uses the
-WRB-2 correlations (Section 4.0). The initial conditions for other key parameters are selected in such a manner to maximize the impact on DNBR.
For accidents which do not use the SCD method, the WRB-2 correlation is used when coolani conditions are within the range of the correlation, otherwise the W-3 DNB correlation is used. For the case where the SCD method is not employed, the initial conditions are obtained by adding the maximum steady state errors to rated values in such a manner as to maximize the impact on the limiting parameter. The following conservative steady state errors were assumed in the analysis of these accidents: 5-1
. Core Power i 2 percent allowance for calorimetric error
. Average Reactor 4.85 0F allowance for controller deadband Coolant System and measurement error plus 2,5 0F allowa'nce Allowance for steam generctor tube fouling
. Pressurizer Pressure 130 psi allowance for steady state fluctuations and measurement error
. Reactor Coolant Flow Thermal Design Flow 5.0.1.1 Proposed Technical Specification Changes Assumed The proposed Technical Specification changes which are assumed in the non-LOCA accident analysis include the following:
Increase in FaH/Fo
. Positive Moderator Temperature Coefficient
. OTAT/OPAT Setpoint Changes
. Decrease in Allowed RCS Thermal Design Flowrate
. Man Steam Safety Valve Setpoint Tolerance Increase
. Shutdown Margin increase in Mode 5 increase in FAH-The increase in the design enthalpy rise peaking factor (FDH) assumption has been increased from the current value of 1.55 to 1.65.
Positive Moderator Temperature Coefficient A positive moderator temperature coefficient (PMTC) of +6 pcm/0F from 0% power to 70% power and then decreasing linearly to O pcm/0F at 100% power has been assumed in the analysis. The resulting boron concentration increase due to the PMTC has been included in the re-analysis. QTAT/OPaT Setooint Chanaes The non-LOCA accident analysis has been re-analyzed with increased OTAT/OPAT setpoints. The resulting Technical Specification Changes associated with the change will allow additional operating margin for the Wolf Creek plant. 2
Decrease in Allowed RCS Thermal Desian Flowrate The current reactor coolant system (RCS) thermal design flowrate is 375,500 gpm. The analysis has been performed with a revised thermal design flowrate of 374,400 gpm. Main Steam Safetv Valve Setooint Tolerance increase The analyses have been performed with an increased tolerance on the main steam safety valves (MSSVs). The current MSSV tolerance is 1%. The proposed MSSV tolerance is 13%, at which all analyses have been performed. , Shutdown Marain increase in Mode 5 The shutdown margin requirement for Mode 5 has been increased in order to regain margin for the Boron Dilution Event discussed on Section 5.4.6. 5.0.1.2 Additional Assumptions Additional assumptions which are not directly related to Technical Specification changes were included in the accident analysis. These include:
. Increased Power Assumption
. Analysis Over a Range of Temperatures
. 10% Steam Generator Tube Plugging
. Thimble Plug Removal increased Power Assumption All transients which were analyzed at 100% power correspond to the upratad core power level of 3565 MWth. This is a conservative assumption for the accident analysis.
Analysis Over a Ranae of Temoeratures An analysis over a range of hot leg temperatures was performed in the accident analysis, and extended over a range of 16.8 0F. The accident analysis presented in sections 5.0 and 6.0 are to support the Technical specification changes discussed in section 1.0, but the temperature range analysis is being presented as it was an analysis assumption for incorporation later in the Power Uprate Report. 5-3
Increased Steam Generator Tube Pluaaina.. The accident analysis has been performed assuming 10% plant total steam generator tube plugging - not to exceed 10% in any single steam generator. This assumption has - been made for those events in which the 10% tube plugging is conservative. Those events in which the assumption of 10% tube plugging is not conservative, the increase in heat removal events, have continued to assume 0% steam generator tube plugging. Thimble Pluo Removal Thimble plug removal affects the core pressure drop and increases the core bypass flow. These assumptions have been incorporated into the accident analysis and the Thermal-Hydraulic analysis discussed in section 4.0. 5.0.1.3 Determination of Design Operating Conditions The Power Rerating program required the definition of revised operating conditions. A summary of these design operating conditions have been included in Table 5.0-1. In-Table 5.0-1, the first column lists the current design parameters, while the second and third columns list the high and low range of temperatures which have been analyzed in the accident analysis. 5-4
I Table 5.0-1 Wolf Creek Generating Station Design Parameters 0 0F 150F Present - Thot - That Parameter , ower Reduction Reduction NSSS Power, MWt 3425 3579 3579 Reactor Power, MWt 3411 3565 3565 , Thermal Design Flow Per loop, gpm 95700 93600 93600-Total flow, gpm 382800 374400- 374400 Reactor Flow. Total 142.1 139.4- 142.9-6 Total (10 lbm/hr) Reactor Coolant Pressure 2250 2250 2250' Pressere, psia Core Bypass, % 5.8 8.4 8.4 Fuel Design 17x17 17x17- '17x17~ Std. V5H w/lFMs V5H w/lFMs Reactor Coolant Temperature. OF Core Outlet 621.4 625.0 608.5 Vessel Outlet 618.2 620.0 603.2 l Core Average 591.8 -593.0 575.1 Vessel Average 588.5 538.4 570.7 Vessel / Core Inlet 558.8 556.8 538.2 Steam Generator Outlet 558.6 556.6 538.0-Zero Load Temperature 557 557 557 i Steam Generator Steam Temperature, OF 544.6 538.4 519.4 L Steam Pressure, psia 1000 950 807 l Steam Flow,106 lbin/hr total 15.14' 15.92 15.83 Feedwater Temperature, OF 440 446 446 Steam Generator 0 10 10 Tube Plugging, % 5-5
t 5.1 increase in Heat Removal by the Secondary System 5.1.1 Feedwater System Malfunctions That Result in A Decrease In Feedwater Tamperature 5.1.1.1 introduction Reductions in feedwater temperature will cause an increase in core power by decreasing the reactor coolant temperature. Such transients are attenuated by the thermal capacity of the secondary plant and the RCS. The overpower /overtemperature protection (neutron overpower, overtemperature, and overpower AT trips) prevents any power increase which could lead to failure to meet the departure from nucleate boiling (DNB) acceptance criterion. A reduction in feedwater temperature may be caused by a spurious heater drain pump trip. - In the event of a heater drain pump trip, there is a sudden reduction in feedwater inlet temperature to the steam generators. At power, this increased subcooling will create a greater load demand on the RCS, With the plant at no-load conditions, the addition of cold feedwater may cause a decrease in RCS temperature, and thus a reactivity insertion due to the effects of the negative moderator coefficient of reactivity. However, the rate of energy change is reduced as load and feedwater flow decrease, so the no-load transient is less severe than the full power case. The net effect on the RCS due to a reduction in feedwater temperature is similar to the effect of increasing secondary steam flow, i.e., the _ reactor will reach a new equilibrium - condition at a power level corresponding to the new steam generator AT. A decrease in normal feedwater temperature is classified as an ANS Condition 11 event, fault of moderate frequency.
' 5.1.1.2 Methodology This transient is analyzed by computing an enthalpy that is equivalent to the conditions in the flow increase (FWMAL) accident analyzed in Section 5.1.2. This equivalent :
enthalpy is compared to the enthalpy corresponding to me maximum feedwater-temperature decrease. The following assumptions are made:
- a. Plant initial power level corresponding to guaranteed nuclear steam supply system thermal output I
5-6
g
- b. Heater drain pumps trip, resulting in a reduction in feec' water temperaturec 5.1.1.3 Results A trip of the heater drain pumps causes a reduction in feedv ater temperature that -
increases the thermalload on the primary system. The calculated reduction in feedwater temperature is less than a temperature decrease equivalent to the conditions analyzed in Section 5.1.2. The increased thermal load due to a spurious heater drain pump trip would result in a transient very similar (but of reduced magnitude) to that-presented in Section 5.1.3 for an excessive increase in secondary steam flow, which evaluates the consequences of a 10-percent step load inc'3ase. -Therefore, the transient results of this analysis are not presented. The plant is expected to reach stabilized conditions at a power level slightly higher than the initial power level. 5.1.1.4 Conclusions The decrease in the feedwater temperature transient is less severe than the increase in the feedwater flow event (see Section 5.1.2) and the increase in the secondary steam flow event (see Section 5.1.3). Based on results presented in Sections 5.1.2 and 5.1.3, the appScable acceptance criteria for the decrease in feedwater temperature event have been met. l i l' 5-7 u
i 5.1.2 Feedwater System Malfunctions That Result In An increase in Feedwater Flow 5.1.2.1 Introduction Addition of excessive feedwater will cause an increase in core power by decreasing the reactor coolant temperature. Such transients are attenuated by the thermal capacity of the secondary plant and of the RCS. The overpower /overtemperature protection (neutron overpower, overtemperature, and overpower AT trips) prevents any power increase which could lead to failure to meet the departure from nucleate boiling (DNB) acceptance criterion. An example of excessive feedwater flow would be a full opening of a feedwater control valve due to a feedwater control system malfunction or an operator error. At power, this excess flow causes a greater load demand on the RCS due to increased subcooling in the steam generator. With the plant at no-load conditions, the addition of cold feedwater may cause a decrease in the RCS temperature and thus a reactivity insertion due to the effects of the negative moderator coefficient of reactivity. Continuous addition of excessive feedwater is prevented by tne steam generator high-high level trip, which closes the feedwater valves and feedwater pump discharge valves and trips the turbine and main feedwater pumps. 5.1.2.2 Methodology The system is analyzed to demonstrate plant behavior in the event that excessive feedwater addition occurs due to a control system m 31 function or operator error that allows a feedwater control valve to open fully. Four cases are analyzed, assuming a conservotively large negative moderator temperature coefficient:
- a. Accidental opening of one feedwater control valve with the reactor just critical at zero load conditions and in automatic control
- b. Accidental opening of one feedwater control valve with the reactor just critical at zero load conditions and in manual control
- c. Accidental opening of one feedwater control valve with the reactor in automatic control at full power
- d. Accidental opening of one feedwater control valve with the reactor in manual control at full power The reactivity insertion rate following a feedwater system malfunction is calculated with the following assumptions:
5-8
- . ~ - - .. .. --
a, For the feedwater control valve accident at full power, one feedwater control valva is assumed to malfunction, resulting in a step increase to 200 percent of nominal feedwater flow to one steam generator.
- b. For the feedwater control valve accident at zero load condition, a feedwater control valve malfunction occurs, which results in an increase in flow to one steam generator from zero to 250 percent of the nominal full load value for one steam generator.
- c. For the zero load condition, feedwater temperature is at a conservatively low value of 32 0F.
- d. No credit is taken for the heat capacity of the RCS and steam generator thick metal structure in attenuating the resulting plant cooldown.
- e. The feedwater flow resulting from a fully open control valve is terminated by a steam generator high-high level trio signal, which closes all feedwater control and isolation valves, trips the main feedwater pumps, and trips the turbine.
- f. No credit is taken for steam generator tube plugging. Since the decreased heat transfer rate associated with tube plugging would reduce the severity of the transient, assuming maximum cooling is conservative.
5.1.2.3 Results The full power case (maximum reactivity feedback coefficients, with rod control) results in the greatest power increase. When the steam generator water level in the faulted loop reaches the high-high level setpoint, all feedwater control valves and feedwater pump discharge valves are automatically closed, and the main feedwater pumps are tripped. This prevents continuous addition of the feedwater, in addition, a turbine trip is initiated. Following turbine trip, the reactor is automatically tripped, either directly due to turbine trip or due to an overtemperature AT signal, if the system is in manual control. If the reactor were in the automatic control mode, the control rods would be insertad at the maximum rate following turbine trip, and the ensuing transient would then be Jimilar to a loss of load (turbine trip event), as analyzed in Section 5.2.3. i Transient results presented in Figures 5.1.1-1 and 5.1.1-2 for the full power case with automatic rod control show the increase in nuclear power and AT associated with the increased thermalload on the reactor. The DNB does not exceed the acceptance criterion. Since the power level rises during the excessive feedwater flow incident, the fuel temperatures will also rise until after reactor trip occurs. The core heat flux lags behind i the neutron flux response due to the fuel rod thermal time constant, hence the peak value of heat flux does not exceed 118 percent of its nominal value (i.e., the assumed l-l g 5-9 L
high neutron flux trip setpoint). The peak fuel temperature will thus remain well below the fuel melting temperature. The transient results show that DNB does not occur at any time during the excessive feedwater flow incident. Thus, the ability of the reactor coolant to remove heat from the fuel rod is not reduced. The fuel cladding temperature, therefore, does nc! tise significantly above its initial value during the transient. The calculated sequence of events for this accident is presented in Table 5.1.21. 5.1.2.4 Conclusions The results of the analysis show that the DNB ccnditions encountered for an excessive feedwater addition at power do not exceed the limiting value. Additionally, it has been t determined that the reactivity inst on rate which occurs at no-load conditions following excessive feedwater addition is less than the maximum value considered in the analysis of the rod withdrawal from subcritical condition analysis, presented in Section 5.4.1, t 5-10
T:b!e 'i 1.2-1 TIME SEQUENCE C EVENTS Time Accident Ele.n_1 (seci Feedwater system malfunctions that result in an increcse in feedwater flow at full power One main feedwater control 0.0 valve fails fully open OPAT setpoint in two 13.2 loops reached __ Turbine trip /feedwater isolation 25.6 on high-high steam generator water level Maximum core heat flux 29.5 reached Reactor trip on turbine trip 29.6 Rod motion occurs 29.3 Feedwater isolation 32.6 4 5-11
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E 5.1.3 Excessive increase in Secondary Steam Flow 5.1.3.1 Introduction
' An excessive load increase is a rapid increase in steam flow that causes a power mismatch between the reactor core power and the steam generator load demand. This accident could result from either an administrative violation such as excessive loading by the operator or an equipment malfunction in the steam dump control or turbine speed control.
5.1.3.2 Methodology Four cases are analyzed to demonstrate the plant behavior following a 10 percent step load increase from rated load. These cases are as follows:
- 1. Manual rod control with minimum reactivity feedback
- 2. Manual rod control with maximum reactivity feedback
- 3. Automatic rod control with minimum reactivity feedback
- 4. Automatic rod control with maximum reactivity feedback For the minimum moderator feedback cases, a zero moderator temperature coefficient of reactivity was assumed which results in the least inherent transient response capability. For this cooldown event, the zero moderator temperature coefficient produces bounding results relative to those with a positive moderator temperature coefficient. The maximum moderator feedback cases assume conservatively the most negative moderator temperature coefficient of reactivity which produces the largest amount of reactivity feedback from changes in coolant temperature.
A 10 percent step increase in steam demand is assumed, and all cases are analyzed without credit being taken for pressurizer heaters. The cases which assume automatic rod control are analyzed to ensure that the worst case is presented, although the - automatic function is not required. The primary consideration in the analysis of the excessive load increase event is to demonstrate that the DNB design basis is satisfied for the rerate conditions of 3579 MWt, a +6.0 pcm/*F MTC and 10% steam generator tube plugging. For tha analysis of this event, the initial reactor power, RCS pressure and RCS temperature are determined using Statistical Core Design methodology. 5.1.3.3 Results Figures 5.1.3-1 through 5.1.3-8 illustrate the transient results with the reactor in the. manual control mode. As expected, for the minimum moderator feedback case there is l 5-14
a slight power increase and a large average core temperature decrease. This results in a DNBR which increases above its initial value. For the maximum moderator feedback, manually controlled case there is a large increase in reactor power due to the moderator feedback. A reduction in DNBR is - experienced but DNBR remains above the limit value. Figures 5.1.3-9 through 5.1.3-16 illustrate the transient results assuming the reactor is - in the automatic control mode. Beth the minimum and maximum moderator feedback cases show that core power increases, thereby reducing the rate of decrease in coolant average temperature and pressurizer pressure. For both of these cases, the minimum DNBR remains above the limit value. For all cases, the plant rapidly reaches a stabilized condition at a h!gher power level corresponding to the increase in steam flow. Neither the primary nor secondary system pressures approach their respective overpressure limits. Reactor trip is not predicted to occur for any of the cases analyzed. The calculated seque:1ce of events for the excessive load increase incident are shown in Table 5.1.3-1. 5.1.3.4 Conclusions The analysis results presented show that four loop operatian with a ten percent step load increase, tne DNBR remains above the limit value, thereby ensuring that the DNB design basis is inet. Also, neither the primary nor the secondary system pressures exceed 110% of the design values. The results of the analysis also demonstrate that for the modeled transient, the plant reaches a stabilized condition following tha load increase without a reactor trip occurring. I l 1 5-15 i
1 Table 5.1.3-1
' TIME SEQUENCE OF EVENTS -
EXCESSIVE INCREASE IN SECONDARY STEAM FLOW Event Time (sec) Manual reactor 10 percent step load 0.0 (minimum moderator increase feedback) . Equilibrium conditions 140. reached (approximate time only) Manual reactor 10 percent step load 0.0 (maximum moderator increase feedback) Equilibrium conditions 50 reached (approximate time only)- Automatic reactor 10-percent step load 0.0 control (minimum increase moderator feedback) Equilibrium conditions 110 reached (approximate time only) Automatic reactor 10-percent step load 0.0 control (maximum increase moderator feedback) Equilibrium conditions 50. reached (approximate time only) l 5-16
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5.1.4 Inadvertent Opening of a Steam Generator Relief or Safety Valve
... 5.1.4.1 Introduction The most severe core conditions resulting from on accidente,l depressurization of the main steam system are associated with an inadvertent opening of a single turbino bypass, relief, or safety valvo. This ovent is bounded by the rupture of a main steam line, as described in Section 5.1.5.
The steam release, as a consequence of this accident, results in an initialincrease in steam flow followed by a decrease in steam flow during the rest of the accident as the steam generator pressure decreases, The increased steam flow increases the energy removal from the RCS, producing a reduction of coolant temperature and pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivity. This causes a return to power during which DNB may becomo a concern. The return to power is limited by a combination of fuel and moderator reactivity feedback effects, and by boron injected via the safety injection system. The analysis is performed to demonstrate that the following critorion is satisfied: Assuming a stuck RCCA, with offsite power available, and assuming a single failure in the engineered safety features system, there will be no consequential damage to the cc.e er reactor coolant system after reactor trip for a steam release equivalent to the spurious openirig, with failure to close, of the largest of any single stearn dump, relief, or safety valve. 5.1.4.2 Methodology The following conditions are assumed to exint at the time of a secondary steam system release:
- 1. End-of life shutdown margin at no-load, equilibrium xenon conditions, and with the most reactivo RCCA stuck in its fully withdrawn position.
Operation of RCCA banks during core burnup is restricted by the insertion - limits so that addition of positive reactivity in a secondary system steam release accident will not lead to a more adverse condition than the case analyzed. 2,. A negative moderator coefficient corresponding to the end-of-life rodded core with the most reactive RCCA in tha fully withdrawn posit'on. The variation of the coefficient with temperature and pressure is included. l 5-33 h -- -
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- 3. Minimum capability for injection of 2000 ppm boron solution corresponding to the most restrictive single failure in the safety injection system is assumed. This corresponds to the flow delivered by one centrifugal charging pump delivering flow to the cold leg header and taking suction from the RWST. Reactor coolant seal injection flow is not included in the total core delivery. The volume downstream of the RWST must be swept prior to delivery of boric acid (2000 ppm) to the reactor coolant loops.
- 4. The caso studied is a steam flow of 268 pounds per second at 1200 psia, with offsite power available. This is the maximum capacity of any single turbine bypass, relief, or safety valve. Initial hot standby conditions at time zero are assumed, since this represents the most conservative initial' condition.
Should the reactor be just critical or operating at power at the time of a steam release, the reactor will be tripped by the normal overpower protection when power level reaches a trip point. Following a trip at power, the RCS contains more stered energy than at no load, the average coolant temperature is higher than at no-load, and there is appreciable energy stored in the fuel. Thus, the additional stored energy is removed via the cooldown caused by the steam release before the no load conditions of RCS temperature and shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity insertions proceed in the same manner as in the analysis which assumes no-load conditions at time zero. However, since thi initial steam generator water inventory is greatest at no-load, the magnitude and duration of the RCS cocidown are less for - steamline release occurring at power.
- 5. In computing the steam flow, the Moody Curve for fUD = 0 is used.
G. Perfect moisture sepaistion in the steam generator is assumed.
- 7. Offsite power is assumed, since this would maximize the cooldown.
- 8. Maximum cold auxiliary feedwater flow is assumed.
- 9. Four reactor coolant pumps are operating.
5.1.4.3 Results The results presented are a conservative indication of the events which would occur, ' assuming a secondary system steam release, since it is postulated that all of the 5 34 - .- -- . -- _=
conditions described above occur simultaneously. The results are prosented in the Figures following. The corresponding event .sequerice is shown in Table 5.1.41. l The valve fails at time zero, which causes the seccndary system depressurization. As the pressure decreases, flow through the valve also decreases. The secondary system _ depressurization causes the RCS temperature to drop; the reactivity climbs, eventually becomes positive, and a return to power occurs. The return to power is limited primarily by Doppler reactivity feedback; however, the power does not decrease significantly untilinjected boron reaches the core. From this point onward the power decreases. . 5.1.4.4 Conclusions i i Results of the analysis show that, for an accidental secondary side depressurization, all ; safety criteria are met. Since the DNB remains above the allowable limit, the core is not adversely affected. Note also that this event is limited by the steamline rupture caso described in Section 5.1.5. i Table 5.1.41 4 Inadvertent Opening of a Steam Generator Relief or Safety Valve Event Sequence Event Time (cec) SG SV Failure 0.001 Aux. Feed Actuation 0.001 Low Pressure Trip 185 0 . Peak Power 372.0 Boron Arr;ves in Core - 400.0 , c 5-35 i
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b 5.1.5 STEAM SYSTEM PIPING FAILURE 5.1.5.1 Introduction The steam release arising from a rupture of a main steam line would result in an initial increase in stearn flow that decreases during the accident as the steam generator pressure decreates. The energy removal from the RCS causes a reductiori of coolant temperature and pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivity. If the most reactive RCCA is assumed stuck in its fully withdrawn position after reactor trip, there is possibility that the core will become critical and return to power. The core is ultimately shut down by the boric acid solution delivered by the safety injection system. The analysis of a main steam line rupture is performed to demonstrate that the core cooling capability is maintained. Although DNB and possible clad perforation following a steam pipe rupture are not necessarily unacceptable, the following analysis, in fact, snows that the DNB design basis is not exceeded for any rupture, assuming the most reactive control rod assembly stuck in its fully withdrawn position. Effects of minor secondary system pipe breaks are bounded by the analysis presented in this section. The major rupture of a steam line is the most limiting cooldown transient, and is analyzed at zero power with no decay heat. Decay heat would retard the cooldown, thereby reducing the likelihood that the reactor will return to power. A detailed analysis of this transient with the most limiting break size, a double ended rupture, is presented here. Fast-acting isolation valves are provided in each steam line; these valves will fully close within 5 seconds of actuation, following a large break in the steam lina. For breaks downstream of the isolation valves, closure of all valves would completely terminate the blowdown. For any break in any location, no more than one steam generator would experience an uncontrolled blowdown, even if one of the isolation valves fails to close. Flow restrictors are installed in the steam generator outlet nozzic, a integral part of the steam generator. The effe5 ve throat area of the nozzles is 1.4 squcre feet, which is considerably less than the main steam pipe area; thus, the nozzles also servo to limit the maximum steam flow for a break at any location. 5.1.5.2 Methodology 5-40
The analysis of the steam pipe rupture has been performed to determine:
- a. The core heat flux and RCS temperature and pressure resulting from the i cooldown following the steam line break.
- b. The thermal and hydraulic behavior of the core (DNBR) following a steam line break.
The following conditions are assumed to exist at the time of a main stetsmline Dreak accident:
- a. Initial plant conditions are presented in Section 5.0.
- b. End-of life shutdown margin at no-load, equilibrium xenon conditions, and the most reactive RCCA stuck in its fully withdrawn position. Operation of the control rod banks during each operating cycle is restricted by the ,
insertion limits so that addition of posillve reactivity in a steam line break accident will not lead to a more adverse condition than the case analyzed. ,
- c. A negative moderator coefficient corresponding to the end-of life rodded core with the most reactive RCCA in the fully withdrawn position. The variation of the coefficient with temperature and pressure has been included. The keff versus temperature at 1,000 psia, corresponding to the negative moderator temperature coefficient used, as shown in WCGS USAR [1] Figure 15.1-11.
The effect of power generation in the core on overall reactivity _is shown in WCGS USAR (1) Figure 15,1 14. . r The core properties associated with the sector nearest the affected steam generator and those associated with the remaining sector were conservatively combined to obtain average core properties for reactivity feedback calculations. Further, it was conservatively assumed that the core power distribution was uniformi These two conditions cause l underprediction of the reactivity feedback in the high power region near i 'the stuck rod. The stuck rod is assumed in the region of the core of j lowest temperature. To verify the conservatism of this method l the reactivity, as well as the l- power distribution, was checked for the limiting conditions during the transient for the cases analyzed. This core analysis considered the ~ Doppler reactivity from the high fuel temperature near the stuck RCCA, l moderator feedback from the high water enthalpy near the stuck RCCA, power redistribution, and non uniform core inlet temperature effects. For 5-41 L _
l i cases in which surface boiling occurs in the .igh heat flux regions of the core, the effect of void formation was also included, it was determined . 1 that the reactivity employed in the kinetics analysis was always larger I than the reactivity calculated, including the above local effects for the l limiting conditions during the transient. These results verify conservatism, l.e., underprediction of negative reactivity fsedbeck from power generation. , I
- d. Minimum capability for injection of boron solution corresponding to the j most restrictive single failure in the safety injection system. The emergency core cooling system consists of three systems: 1) the passive l accumulators,2) the residual heat removal system, and 3) the safety ;
injection system. Only the safety injection system is modeled for the steam line break accident analysis. The, fiow corresponds to full flow (less sealinjection flow) of one charging pump delivering to the RCS via ' the cold leg header. No credit has been taken for the borated water that must be swept from the lines downstream of the RWST prior to the delivery of the boron solution from the RWST to the reactor coolant loops. For the cases where offsite power is assumed, the sequence of events in the safety injection system is as follows. After the generation of the' safety injection signal (appropriate delays for instrumentation, logic, and signal transport included), the appropriate valves begin to operate, and the high head safety injection pump starts. In 27 seconds, the valves are assumed to be in their final position, and the pump is assumeJ to be at full speed . - The volume downstream of the RWST must be swept prior to delivery of boric acid to the reactor coolant loops. This deiay is included in the ; calculations. In cases where offsite power is not available, an additional 12 second , delay is assumed to start the diesels and to load the necessary safety injection equipment onto them, e
- e. Since the steam generators are provided with integral flow restrictors with a 1,4-square foot throat area, any rupture with a break area greater than '
1.4 square feet, regardless of location, would have the same effect on the - NSSS as the 1,4 square-foot break. The following cases have been , considered in determining the core power and RCS transients:
- 1. Complete severance oi ; pipe, with th'e plant initially at no-load conditions; full reactor coolant flow with offsite power available. ;
- 2. Case (1) with loss c. Offsite power simultaneous with the steam line-break, which causes initiation of the safety injection sigrmi. Loss of - ,
offsite power results in reactor coolant pump coastdown. 5-42 - ' - - .--%,~- - . . - , - --.c.,.,%-g,
- f. Power peaking factors corresponding to one stuck RCCA a.id non-uniform core inlet coolant temperatures are determined at end of core life.
The coldest core inlet temperatures are assumed to occur in the sector with the stuck rod. The power peaking factors account for the effect of the local veld in the region of the stuck RCCA during the return to power phase following the steam line break. This void in conjunction with the large negative moderator coefficient partially offsets the effect of the stuck assembly. The power peaking factors depend upon the core power, temperature, pressure, and flow, and, thus, are different for each ( ese studied. The core parameters used for each of the two cases correspond to va~ues , determined from the respective transient analysis. ; I Both cases above assume initial hot standby conditions (557 F average coolant temperature) at time zero, since this represents the most i pessimistic initial condition. Should the reactor be just critical or operating at power at the time of a steam line break, the reactor will be tripped by the normal overpower protection system when power level i reaches a trip point. Following a trip at power, the RCS contains more stored energy than at no load, the average coolant temperature is higher - than at no load, and there is appreciable energy stored in the fuel. Thus, t the additional stored energy is removed via the cooldown caused by the , steam line break before the no load conditions of RCS temperature ed shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity , insertions proceed in the same manner as in the Lnalysis, which assumes no-load condition at time zero.
- g. In computing the steam flow during a steam line break, the Moody Curve for fL/D = 0 is used.
5.1.5.3 Results The calculated sequence of events for both cases analyzed is shown in Table 5.1.5-1. The results presented are a conservative indication of the events which would occur, assuming a steam line rupture, since it is postulated that all of the conditions described above occur simultaneously. Core Power and Reactor Coolant System Transient 5-43
-1
Figures 5.1.5-1 through 5.1.5-6 show the RCS transient and core heat flux following a main steam line rupture (complete severance of a pipe) at initial no load condition for the case with offsite power (case 1). As shown in Figures 5.1.5-4 and 10, the core attains criticality with the RCCAs inserted (with the design shutdown assuming one stuck RCCA) before the RWST boron solution at 2000 ppm enters the RCS. A peak core power significantly lower than the nominal full power value is attained. The calculation assumes that the boric acid is mixed with, and diluted by, the water flowing in the RCf pnor to entering the reactor core. The concentration after mixing depends upon the relative flow rates in the RCS and in the safety injection system. The variation of rnass flow rate in the RCS due to water density changen is included in the calculation, as is the variation of flow rate in the safety injection system due to changes in the RCS pressure. The safety injection system flow calculation includes the line losses in the system as well as the pump head curve. r ,res 5.1.5-7 through 5.1.5-12 show the salient parameters for case 2, which corresponds to the case discussed above with the additionalloss of offsite power at the time the safety injection signal is generated. The safety injection syrDm delay time includes 12 seconds to start the diesel in addition to 27 seconds to siart the safety injection pump and open the valves. Criticality is achieved later, and the core power increase is slower than in the similar case with offsite power available. The ability of the emptying steam generator to extract heat from the RCS is reduced by the decreased flow in the RCS. The peak power remains well below the nominal full power value, it shou!d be noted that following a steam line break only one steam generator blows down completely. Thus, the remaining steam generators are still available for the dissipation of decay heat after the initial transient is over. In the case of loss of offsite power, this heat is removed to the atmosphere via the steam line relief or safety valves. Marain to Critical Heat Flux A DNB analysis was performed for both of these cases. It was found that both cases had a minimum DNBR greater than the acceptable. 5.1.5.4 Conclusions The analysis has shown that the criteria stated earlier in Section 5.1.5.1 are satisfied. Although DNB an_d possible clad perforation following a steam pipe rupture are not necessarily unacceptable and not precluded by the criteria, the above analysis shows i that the DNB design bt sis is met for any rrpture, assuming the most reactive RCCA stuck in its fully withdrawn position. I i S-44 4 -
.- _ _ . - _ ~ . _ _ _ _ . _ . _ _ - .-_.._. _ _._ . . - _ _ _ . _ .- . _ _ _ _ _ ._
Table 5.1.5-1 - Main Steam Line Rupture Sequence of Events Time Accident Event {sqg) ; Steam system p; ung failure
- 1. Case 1 (offsito power available)
Steam line ruptures 0.0 Si signal on low pressurizer pressure 14.8 Criticality attained 24.0 Pressurizer empties 30.0 Si actuaction and delivery of boron to the core 66.8 ,
- 2. Case 2 (concurrent loss of offsite power)
Steam line ruptures 0.0 Si signal on low pressurizer pressure 17.4 Criticality attained 28.0 Pressurizer empties 40.0 Si actuaction and delivery of boron to the core 81.4 . I u 5-45
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VESSEL-INLET AND CORE AVERAGE TEMPERATURES FOR MAIN STEAM LINE RUPTURE WITHOUT OFFSITE POWER AVAILABLE L L S-54 o i l
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- CORE REACTIVITY AND CORE-BORON CONCENTRATION FOR-MAIN STEAM LINE RUPTURE -
WITHOUT OFFSITE POWER AVAILABLE 5-55 J +
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q 5.2 Decrease in Heat Removal from the Secondary System-5.2.1 Steam Pressure Regulator Malfunction or Failure That Results in Decreasing Steam Flow This event is not applicable to the Wolf Creek Generating Station. 5.2.2 Loss of External Electrical Load 5.2.2.1 Introduction A major load loss on the plant can result from a Inss of external electrical load due to some electrical system disturbance. Offsite AC power remains available to operate plant components such as the reactor coolant pumps; as a result, the onsite emergency diesel generators are not required to function for this event. Following the loss of generator loed, an immediate fast closure of the turbine control valves will occur. The automatic turbine bypass system would accommodate the excess steam generation. Reactor coolant temperatures and pressures do not significantly increase if the turbine bypass system and pressurizer pressure control system are functioning properly, if the condenser were not available, the excess steam generation would be relieved to the atmosphere. Additionally, main feedwater flow would be lost if the condenser were not available. For this situation, feedwater flow would be maintained by the auxiliary feedwater system. For a loss of external electrical load without subsequent turbine trip, no direct reactor trip signal would be generated. The plant would trip from the reactor protection system - if a safety limit were approached. A continued steam load of approximately 5 percent would exist after totalloss of external electricalload because of the steam demand of plant auxiliaries, in the event that the turbine bypass valves fail to open following a large loss of load, the steam generator safety valves may lift, and the reactor may be tripped by the high pressurizer pressure signal, the high pressurizer water level signal, or the overtemperature AT signal. The steam generator shell side pressure and reactor coolant temperature will increase rapidly. The pressurizer safety valves and steam generator safety valves are, however, sized to protect the RCS and steam generator I against overpressure for all load losses without assuming the operation of the turbine
- bypass system, pressurizer spray, pressurizer power-operated relief valves, automatic rod cluster control assembly (RCCA) control, or direct reactor trip on it'6ine trip.
L L A loss of externalload event results in a nuclear stean supply system transient that is bounded by the turbine trip event analyzed in Sectio" 5.2.3. Therefore, a detailed-transient analysis is not presented for the loss of external load event. 5-58
The primary side transient is caused by a decrease in heat transfer capability from primary to secondary, due to a rapid termination of steam flow to the turbine, , accompanied by an automatic reduction of feedwater flow (should feedwater flow not be reduced, a larger heat sink would be available and the transient would be less severe). Termination of steam flow to the turbine following a loss of external load occurs due to automatic fast closure of the turbine control valves, approximately 0.3 seconds. Following a turbine trip event, terrnination of steam flow occurs via turbine stop valve closure, which occurs in approximately 0.1 seconds. Therefore, the transient in primary pressure, temperature, and water volume will be less severe for the loss of externalload than for the turbine trip, due to a slightly slower loss of heat transfer capability. 5.2.2.2 Methodology Refer to Section 5.2.3.2 for the method used to analyze the limiting transient (turbine trip)in this grouping of events. The results of the turbine trip event analysis bound those expected for the loss of external load. 5.2.2.3- Results The Loss of load is bounded by the Turbine Trip event due to a more rapid turbine stop valve closure. - The results of the Turbine Trip event are presented in Section 5.2.3.3. 5.2.2.4 Conclusions Based on results obtained for the turbine trip event (see Section 5.2.3.3) and considerations described in Section 5.2.2.1, the applicable acceptance criteria for a loss of externalload event are met. 5-59 L
5.2.3 Turbine Trip - 5.2.3.1 Intioduction For a turbine trip event, the reactor would be tripped directly (direct reactor trip on turbine trip is blocked below 50% power by the P-9 Interlock) from a signal derived from the turbine stop emergency trip fluid pressure and turbine stop valves. The turbine stop valves close rapidly (typically 0.1 second) on loss of trip fluid pressure actuated by one of a number of possible turbine trip signals. Upon initiation of stop valve closure, steam flow to the turbine stops abruptly. Sensors on the stoo valves detect the turbine trip and initiate turbine bypass, and, if above 50-percent power, a reactor trip. The loss of steam flow results in an almost immediate rise in secondary system temperature and pressure, with a resultant primary system transient as described in Section 5.2.2.1 for the loss of externalload event. A slightly more severe transient occurs for the turbine trip event, due to the more rapid loss of steam flow caused by the more rapid valve closure hence a more rapid loss of primary to secondary heat transfer. The automatic turbine bypass system would accommodate up to 40 percent of rated steam flow. Reactor coolant temperatures and pressure do not increase significantly if the turbine bypass system and pressurizer pressure control system are functioning properly. If the condenser were not available, the excess steam generation would be relieved to the atmosphere, and main feedwater flow would be lost. For this situation, i feedwater flow would be maintained by the auxiliary feedwater system to ensure adequate residual and decay heat removal capability. Should the turbine bypass system fail to operate, the steam generator safety valves may lift to provide pressure control. A turbine trip is more limiting than loss of external load, loss of condenser vacuum, and other events which result in a turbine trip as a result of the more rapid loss of steam flow. As such, this event has been analyzed in detail. Results and discussion of the analysis are presented in Section 5.2.3.3. 5.2.3.2 Methodology in this analysis, the behavior of the unit is evaluated for a complete loss of steam load from 102 percent of full uprated power (3636 MWt), without direct reactor trip, primarily to show the adequacy of the pressure relieving devices, and also to demonstrate core protection margins. That is, the turbine is assumed to trip without actuating all the sensors for reactor trip on the turbine stop valves. This assumption delays reactor trip until conditions in the RCS result in a trip due to other signals. Thus, the analysis assumes a worst case transient. In addition, no credit is taken for the turbine bypass 5-60 l
l system. - Main feedwater flow is terminated at the time of furbine trip, with no credit taken for auxiliary feudwater to mitigate the consequences of the transient. The turbine trip event is analyzed for RCS overpressurization and pressurizer overfilling assuming nominal initial conditions including allowances for measurement errors. DNBR _is evaluated using SCD methodology which assumes nominalinitial conditions. The major assumptions used in the analysis are summarized below:
- a. Initial operating conditions (Overpressure / Overfill)
Initial reactor power and RCS temperatures are assumed at their-maximum values consistent with steady state full power operation, including allowances for calibration and instrument errors. The initial-F 7S pressure is assumed at a minimum value consistent with steady state full power operation, including allowances for calibration and instrument errors. This results in the maximum power difference for the load loss and the minimum margin to core protection limits at the initiation = of the accident,
- b. Initial operating conditions (DNB)
The initial pressure, reactor power and RCS temperatures are assumed at their nominal values consistent with steady state full power operationc Allowances for calibration and instrument errors are treated statistically by the DNBR evaluation code,
- c. Reactivity coefficients (two cases are analyzed):
- 1. Minimum reactivity feedback A most posi_tive moderator temperature coefficient and a least negative Doppler-only power coefficient are assumed.
- 2. Maximum reactivity feedback A conservatively large negative moderator temperature coefficient and a most negative Doppler-only power coefficient are assumed,
- d. Reactor control From the standpoint of the maximum pressures attained, it is conservative to assume that the reactor is in manual control. If the reactor were in automatic control, the control rod banks would move prior to trip and reduce the severity of the transient.
- e. Steam release No credit is taken for the operation of the turbine bypass system or steam generator power-operated relief valves. The steam generator pressure 5-61
rises to the safety valve setpoint where steam relee.se through safety valves limits secondary steam pressure at the setpoint value.
- f. Pressurizer spray and power-operated relief valves Cases for DNBR, Overfill and Overpressure are analyzed for both the minimum and maximum reactivity feedback conditions:
- 1. DNBR/ Overfill Full credit is taken for the effect of pressurizer spray and power-operated relief valves in reducing or limiting the coolant pressure.
Safety valves are also available.
- 2. Overpressure No credit is taken for the effect of pressurizer spray and power-operated relief valves in reducing or limiting the coolant pressure.
Safety valves are operable.
- g. Feedwater flow Main feedwater flow to the steam generators is assumed to be lost at the time of turbine trip. No credit is taken for auxiliary feedwater flow, since a stabilized plant condition will be reached before auxiliary feedwater initiation is normally assumed to occur; however, the auxiliary feedwater pumps would be expected to start on a trip of the main feedwater pumps.-
The auxiliary feedwater flow v/ould remove core decay heat following plant stabilization.
- h. Reactor trip is actuated by the first reactor protection system trip setpoint reached, with no credit taken for the direct reactor trip on the turbine trip. . Trip signals are expected due to high pressurizer pressure, overtemperature AT, and low low steam generator water level.
5.2.3.3 Results The transient responses for a turbine trip from 102 percent of full power operation are shown for six cases in Figures 5.2.3-1 through 5.2.3-12. Minimum and maximum reactivity feedback cases are presented without pressure control (i.e., PORVs, pressurizer spray) o ensure 110% of the design pressure (2750 psla)is not exceeded. Cases with pressure control are analyzed to show that the pressurizer will not overfill and therefore this event does not lead to a more serious plant condition. Cases using nominal initial plant conditions and pressure control are analyzed to provide boundary conditions for DNB analysis. Minimum and maximum reactivity feedback is analyzed for each case to ensure the worst case is analyzed.- 5-62
The calculated sequence of events for the accident is shown in Tables 5.2.3-1 through 5.2.3-6. Figures 5.2.3-1 and 5 2.3-2 show the transient responses for the total loss of steam load with minimum reactivity feedback, assuming full credit for the pressurizer spray and pressurizer power-operated relief valves. No credit is taken for the steam bypass. Due to the positive moderator temperature coefficient, the power increases to approximately 110 percent of nominal (3565 MWt) before the reactor is tripped by the overtemperature AT trip channel. The steam generator safety valves limit the secondary steam conditions to saturation at the safety valve setpoint. Figures 5.2.3-3 and 5.2.3-4 show the responses for the totalloss of steam load with maximum reactivity feedback. All other plant parameters are the same as the above. Again the reactor is tripped by the overtemperature AT trip channel but reactor power remains essentially constant until the trip occurs. The steam generator safety valves limit the secondary steam conditions to saturation at the safety valve setpointJ The turbine trip accident was also studied assuming no credit taken for the pressurizer spray, pressurizer power-operated relief valves, or the turbine bypass system. Tho reactor is tripped on the high pressurizer pressure signal. Figures 5.2.3-5 and 5.2.3-6 show the transients with minimum reactivity feedback The neutron flux increases slightly above 102 percent of full power until the reactor is tripped, due to the positive moderator temperature coefficient. In this case, the pressurizer safety valves are actuated, and maintain RCS pressure be!aw 110 percent of the oesign value. Figures 5.2.3 7 and 5.2.3-8 show the transients with maximum reactivity feedback, with the other assumptions being the same as in the preceding case. Again, the pressurizer safety valvos are actuated to limit primary pressure. The DNB evaluation for the turbine trip event was analyzed assuming full pressure control and nominal initial conditions. The errors in initit. conditions are treated statistically. Figures 5.2.3-9 and 5.2.3-10 show the transient responses for the total loss of steam load with minimum reactivity feedback for the DNB evaluation. No credit is taken for the steam bypass. Due to the poshive moderator temperature coefficient, the power increases to approximately 110 percent of nominal (3565 MWt) before the reactor is tripped by the overtemperature AT trip channel. Steam is released through the pressurizer safety valves for this case and the maximum RCS pressure is maintained below 110 percent of the design pressure. The steam generator safety ! valves limit the secondary steam conditions to saturation at the safety valve setpoint. Minimum DNBR remains within the acceptable limits. Figures 5.2.3-11 and 5.2.3-12 show the responses for the total loss of steam load with maximum reactivity feedback. All other plant parameters are the same as the above. I The reactor is tripped by the high pressurizer pressure trip channel but the pressurizer safety valves are not actuated for this case. Reactor power remains essentially 5-63
constant until the trip occurs. The steam generator safety valves limit the secondary steam conditions to saturation at the safety valve setpoint. Minimum DNBR remains within the acceptable limits. 5,2.3.4 Conclusions Results of the analyses show that the plant design is such that a turbine trip without a direct or immediate reactor trip presents no hazard to the integrity of the RCS or the main steam system. Pressure-relieving devices incorporated in the two systems are adequate to limit the maximum pressures to within the design limits. The analyses show that the DNBR remains within the acceptable limit through the transient. Thus, the DNB design basis is met. 5-64
Table 5.2.3-1 Turbine Trip with Pressure Control, Minimum Feedback Time Sequence of Events Event Time (sec) Turbine trip; Loss of 0.025 main feedwater flow initiation of steam 4.75 l release from pressurizer relief valve Initiation of steam 5.36 release from steam generater safety valves OTAT reactor trip 6.98 setpoint reached Rods begin to drop 8.98 Peak pressurizer 10.95 pressure occurs 5-65 t-- _
Table 5.2.3-2 Turbine Trip with Pressure Control, Maximum Feedback Time Sequence of Events
= Event Time (sec)
Turbine trip; Loss of 0.025 main feedwater flow Initiation of steam 4.65 release from pressurizer relief valve initiation of steam 5.36 release from steam generator safety valves OTAT reactor trip 7.09 setpoint reached Rods begin to drop 9.09 Peak pressurizer 1C.15 pressure occurs
.i l
l 5-66 E _- j
9 Table 5.2.3-3
. Turbine Trip without Pressure Control, Minimum Feedback Time Sequenco of Events y
Event Time (sec) _ Turbine trip; Loss of 0.025 main feedwater flow Initiation of steam NONE release from pressurizer relief valve High pressurizer 5.40 pressure reactor trip setpoint reached Initiation of steam 5.35 release from steam generat ,r safety valves Rods begin to drop 7.40 Initiation of steam 7.61 release from pressurizer safety valve Peak pressurizer 7.95 pressure occurs l i 5-67 L c
Table 5.2.3-4 . Turbine Trip without Pressure Control Maximum Feedback-Time Sequence of Events Event Time (sec) _ Turb,r;e trip; Loss of 0.025 main feedwater flow initiation of steam NONE release from pressurizer relief valve High pressuriner 5.34 pressure reactor trip setpoint reached initiation of steam 5.35 release from steam generator safety valves Rods begin to drop 7.34 Initiation of steam 7.55 release from pressurizer safety valve Peak pressurizer 7.95 pressure occurs 5-68
Table 5.2.3-5. Turbine Trip DNB Evaluation Minimum Feedback
. Time Sequence of Events Event Time (sec)
Turbine trip; Loss of 0.0.25 main feedwater flow initiation of steam 3.85 release from pressurizer relief valve initiation of steam 7.00 release from steam generator safety valves
- OTAT_ reactor trip 9.03 setpoint reached Rods begin to drop 11.03 Initiation of steam 12.19 release from pressurizer safety valve Peak system pressure occurs 12.30 5-69
Table 5.2.3-6. Turbine Tr'pi DNB Evaluation, fAaximum Feedback Time Sequerice of Events
" Event Time (sec) ,
Turbine trip; Loss of 0.025 main feedwater flow initiation of steam 3.85 release from pressurizer relief valve initiation of steam 6.99 release from steam generator safety valves High pressurizer pressure 9.62 reactor trip setpoint reached Rods begin to drop 11.62 Peak system pressure occurs 12.65 i, 5-70 i 4
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- 1. 1200.0-O ,
e j 1000.0 - ;
-} 800.0 -
+] .
600.0-400.0 . . . . . . . . 0 20 40 60 . 80 .100 120 Time (seconds) RCS Temperature 650.0 640.0-C630.0- l ,, m., 3,, c , ., j .. e
= 6 20.0 - -i.es. ,v can 4 l b,610.0 e -
o,
.-------j4, em v . can - l e l 600,0- ,
{, 590.0- ,, , E 580.0 I E ;/ 0 70.0-5 / '
.: N... '" .
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550.0 . . . . . . . , . . 0 20 - AO - 60 80- 100- 120 Time (seconds)
- WOLF CREEK- -1.
L FIGURE 5.2.3-2 . TURBINE TRIP EVENT WITH l PRESSURE CONTROL MINIMUM REACTIVITY FEEDBACK - L 5-72 u L
.= . . . .
-- 5 i Nuc! ear Power (Freciton Of Uproied)
, 1.2 c na COUT +2
- 1.0 -
c.o
\ -
y 0.8'- -! 2, t 0.6 - . E i 0.4 - E o E b 0.2 - 0.0 . . . . . . . . . . . 0 20 40 60 80 100 120 Time (seconds) Primery Pressure System
' 2800.0 l "' I 2700.0-2600.0-2500.0- '
fi
. [ - s,.. * - wsas [
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-- -l n uta n u.,
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i
'2000.0- ,.
1900.0- ,. 1800.0- ' ' ' " "-- --------------- . 1700.0 O 20 40 60 80~ 100 120:
- Time (seconds) -
WOLF CREEK FIGURE 5.2.3-3 TURBINE TRIP EVENT WITH PRESSURE CONTROL. MAXIMUM REACTIVITY FEEDBACK = 5-73 qw - +- m>m- , m ,e w- r, my ,- d*t
g-7 . Pressurizer Liquid Volume
. 1600_
Como PE2 LCV 21 1400-2 5
!.1200-e
- E 1000 -
3u- 800-
*3 600-400 . . . . . . . . . . .
0- 20 40 60 80 100 120 Time (seconds) RCS Tempercture 650.0 4 M21 640.0-C 630.0- .[.. t., ,, .. l 620,0-E 610.0 - o ---.--4 l4, a t r.,, can .: l 1
, 600.0-
-590.0-It 580.0 - / .s -
g_ p . H 570.0 - j .. 560.0- ' - .r. . ;; . 7;;.g. - ; . 550.0 . . . . . > > - . . 0 20 40 .60 80 100 120 Time (seconds) WOLF CREEK FIGURE 5.2.3-4 TURBINE TRIP EVENT WITH PRESSURE CONTROL. MAXIMUM REACTIVITY FEEDBACK 1 5-74
= < n, , - , , - x
l Nuclear Power (Freciton'Of Uproied)
-1.2 ca.ns 1
ecut +2
- 1.0 -
n
>1 N
y 0.8 - E t 0.6 - 5 E i 0.4 -
'E E
8 0. ? - (w L z 0.0 . . . . . . . . . . C 20 40 60 80 100 120 Time (seconds) Primary Pressure System 2800.0 - e k= ns l 2700.0-2600.0- ; 2500.0-y, 2 400.0 -
~~~""'~'**I
/
S 2300.0- ' I '~ * '~' * " I e a 2200.0 - , N s-
-2 2100.0- '
n. 2000.0- .' , 1900.0- ', 1800.0- "~"~"'-'""- 1700.0 . . . . . . . . . . 0 20 40 60 80 100 120 Time (seconds) i WOLF CREEK FIGURE 5.2.3-5 i TURBINE TRIP EVENT WITHOUT - PRESSURE CONTROL. _ MINIMUM REACTIVITY FEEDBACK l (- 1 5-75 l l l
'I,
- .. - ~-. .- - . . - . - - -. .- -.. -_
l Pressurlier Llquid Volume 1600 ca. nY- l
._le.lL.
1400-E l1200-
=.
1000-2 800-O a 600-400 i . . . . . . . . . - 0 20 40 60 80 100 120 Time (seconds) RCS Temperature 650.0 - H M3 I 640.0-g 6 30,0 - _ l ,, ,,,, t , , ,,, can ., l l 6 20.0 - ,,,,,,,,,,,,- th ge 610.0 - .---- l c.w6 i m. can.: l
, 500.0-fw 590.0 - ...
f 580.0 - / 0 570.0- ,/ 's - 560.0-
~~N 550.0 . . . . . . . . . .
0 20 40 60 80 '100 120 ilme (seconds) WOLF CREEK FIGURE 5.2.3-6 TURBINE TRIP EVENT WITHOUT PRESSURE CONTROL. MINIMUM REACTIVITY FEEDBACK 5 76
Nuclear Power (froction of Uproted) Come M4
* +2 3.-
m 1.0 R y 0.8 - a b 0.6 - s 10.4-H 0.2 - 0.0 , . , e-- , , , , , 0 20 40 60 80 100 120 ilme (seconds) Primory Pressure System 2700.0-k* M4 l 7' 2600.0- lj 2500.0-d ~ '""* " **" I
$ 2 400.0 -
S 2300.0- ' I * '"" "'" ' "' " I e 5 2200.0-
- 2100.0-a .
2000.0- ' 1900.0- ..,,,
' ~ " ' ' ' ' " " " " - -
1800.0-1700.0 , , , , , v , , , , 0 20 40 60 80 100 120 Time (seconds) r i WOLF CREEK FinURE 5 S.3-7 TURBINE TRIP EVENT WITHOUT PRESSURE CONTROL. JAXIMUM REACTIVITY FEEDBACK 5-77
Pressurizer L1 quid Volume 1600- - Ceu M4
._WLt.L 1400-E
!_1200-C f1000-S }800-a 600- . 400 , 1 4 - - - r 0 20 40 60 80 100 120 Time (seconds) RCS Tempercture 650.0 - ha is] 640.0-g 6 30.0 - l ., , t., im can . i l { 6' t.0 - g ,, ,, , ,, ; 610.0 - , , , , _ ,,, , ,,
, 600.0-4g 5 9 0.0 -
.. s
, 580.0-g ,
0 5 70.0- !
~
560.0- - .....___ ,__,,,;,,,,, 550.0 - , , . , - r - - - 0 20 40 60 80 100 120 Time (seconds) I
~
WOLF CREEK FIGURE 5.2.3-8 TURBINE TRIP EVENT WITHOUT PRESSURE CONTROL. MAXIMUM REACTIVITY REEDBACK 5-78
Nuclear Power (frection Of Upreted) 1.2 Cue s001 Caff+2
$ 1.0 -
n n N
$ 0.8 -
L b 0.6 - a 20.4-H j 0.2 - 0.0 . . , . . . . . . 0 20 40 60 80 100 120 Time (seconds) Primary Pressure System 28')0.0
"' " I 2700.0-7,i 2600.0- /i iS00.0 - / '
3 2400.0- /
'" ~ ""' " " d S 2300.0- * I e
B 2200.0 - C g 2100.0 - a 2000.0-1900.0- .
'"~'
1800.0- -------------...-- 1700.0 . . . . . . . . . . 0 20 40 60 80 100 120 ilme (seconds) WOLF CREEK l FIGURE 5.2.3 9 l TURBINE TRIP EVENT FOR DNB EVALUATION. MINIMUM REACTIVITY FEEDBACK 5 79 i
- . . _ _ _ . - . -.. . - _ _ - . - . . - = . - - - ~ - - . . . -- Pr essurlsor Ltquid Volume 1600 LcAetti scoi Jl 1400-E I.1200-c. 1000- : 80n= a 600-400 0 2b db Bb Bb 100 120 Time (seconos) NC$ Temperature 0.0
,,, ,t,n 1 640.0-c 6 30.0 - Q, , ,, . . . . l E 6 2 0.0 ,_..,. ,,,,,,,
pto o- .___.-i.m...._ i
, 600.0-
-f $ 9 0.0 - ,,
S80.0- ' s
+- 5 7 0,0 - /
560.0- - ' ' ' ' ' ' - - - . . - . . . . , , , , , , _ _ _ ,, 0 2'O d 'O 6 'O 8 'O ido 120 Time (seconos) WOLF CREEK FIGURE 5.2.3-10 TURBINE TRIP EVENT FOR DNB EVALUATION, MINIMUM REACTIVITY FEEDBACK 5-80 g - - ty r-- ,* s ., w-%da y ey-+ *._m.- w.Wav f.m
Nuclear Po*er (frection of Uproted) 1.2 ca. sco ca'i 43
$ 1.0 -7 '
4 N
"0 18< -
b 0.6 - T 0.4 - l 0.2 - Q 0.0 . - ,-- - - - -- 60 80 > 0 20 40 100 120 Time (seeands) Primary Pressure System 2800,0 2700.0-2600.0- , 2500.0-
' "" "'" ' ~ ' " " I
$ 2 400.0 -/'
S 2300.0- '~"'~' { e g 2200.0 - U 2100.0-2000.0- - 1900.0- ~ 1800.0- '""'~'~'~'"-" ----- 1700.0 . . . . . . . . . . . 0 20 40 60 80 100 120 Time (seconds) WOLF CREEK FIGURE 5.2.3-11 l l TURBINE TRIP EVENT FOR DNB EVALUATION. MAXIMUM REACTIVITY FEEDBACK 5-81 l
- . . . . _ , , . - . < ,
- w
1 i l l Pressuriser Llawid Volwme 1600 11 1400-V n m 1200-1000" 800-600-w 400 - - - -- . 0 20 40 60 80 100 120 Time (seconds) RCS Temperature 650.0 640.0-I'** C 630.0- [,,,,,,,,,,,,., ; [ 620.0 ~ 7,,,,,,,,,,,,, I5'o o- ..
..i._,.,_m
, 600.0-f 59 0.0 -
580.0- ,
*- 5 7 0.0 - /
560.0 - ......,_,;,__,, 550.0 . . . . . 0 20 40 60 80 100 120 Time (seconds) WOLF CREEK FIGURE 5.2.3-12 TURBINE TRIP EVENT FOR DNB EVALUATION, MAXIMUM REACTIVITY FEEDBACK 5 82 l
- m- - , ,-e-.r-- e- y - --
- 1
5.2.4 Inadvertent Closure of Main Steam isolation Valves inadvertent closure of the main steam isolation valves would result in a turbino trip with no credit taken for the turbine bypass system. Turbine trips are discussed in Section 5.2.3. 5.2.5 Loss of Condenser Vacuum and Other Events Resulting in Turbine Trip Loss of condenser vacuum is one of the events that can cause a turbine trip. Turbine trips are described in Section 5.2.3. A loss of condenser vacuum would preclude the uso of steam dump to the condenser; however, since steam dump is assumed to be unavailable in the turbine trip analysis, no additional adverse effects would result if the turbine trip were caused by loss of condenser vacuum. Therefore, the analyais results and conclusions contained in Section 5.2.3 apply to the loss of the condenser vacuum. In addition, analyses for the other possible causes of a turbine trip are covered by Section 5.2.3. Possible overfrequency effects due to a turbine overspeed condition are discussed in Section 5.2.2.1, and are not a concern for this type of event. 5-83
_ . _ _ _ _ _ _ _ _ _ . _ . ~ . _ _ _ _ . - _ 5.2,6 Loss of Non Emergency AC Power to the Station Auxiliaries 5.2.6.1 Introduction A complete loss of nonemergency AC (LNEAC) power to the station auxiliaries may be caused by a complete loss of the offsite grid accompanied by a turbine generator trip at the station, or by a loss of the onsite AC distribut;on system. This transient is more i severe than the turbine trip event because, for this case, the decrease in heat removal by the secondary system is accompanied by a flow coastdown (RCP trip) which further l reduces the capacity of the primary coolant to remove heat from the core. Similarly, due to RCP trip, this transient is more severe than the loss of normal feed event to which it is otherwise similar. For the LNEAC event, the reactor will trip on either 1) due to turbine trip, or 2) upon reaching one of the trip setpoints in the primary and secondary systems, or 3) due to the loss of power to the cc.ntrol rod drive mechanisms as a result of the loss of power to the plant. , Following a loss of AC power, a transient limited by the event sequence described below will occur.
- 1. Plant vital instruments are supplied from emergency DC power sources.
- 2. Feedwater is lost as a consequence of LNEAC (condensate pumps lose their power supply). Consequently the RCS heats up and pressurizes, and the pressurizer level swells. If the reactivity feedback coefficients are negative, the ;
reactor power can be expected to fall slightly as a consequence of RCS heatup. However, in the presence of a positive moderator temperature coefficient (PMTC), which is more limiting the power will climb slightly as a consequence of RCS heatup.
- 3. As the RCS pressurizes, the pressurizer watei level will increase. The level increase will be maximized if pressurizer PORVs and spray are credited; hence, this assumption is consarvative from the perspective of water relief from the pressurizer, it such water relief does not occur, then it will be more conservative to disallow the pressurizer PORVs and spray, since this will tend to maximize the RCS pressure unde'such conditione (NFe -when being considered from the perspective of DNB, the analysis assumes that pressure relief systems are -;
operational.)
- 4. Due to loss of feedwater, the steam generator liquid level will steadily decrease.
Eventually, scram will occur on low-low steam generator ;evel (or possibly on high pressurizer pressure, or OTAT).
- 5. When reactor trip (and turbine trip) occurs, power to the reactor coolant purnps (RCPs)is lost and the pumps consequently coast down Subsequently, natural 5-84
I circulation provides the flow necessary for core cooling and removal of decay , heat.
- l 1 6. As the steam generator pressure rises following reactor trip (and turbine trip),
[' the steam generator (SG) safety valves (SVt) lift to dissipate the sensible heat of the fuel and coolant plus the decay heat produced in the core. (Note that both - the SG power-operated relief valves (PORVs) and turbine bypass to the l condenser are assumed to be unavailable.) l
- 7. A short time after scram, auxiliary feedwater is activated to replace the fluid discharged through the steam generator SVs. If the auxiliary feedwater supply is sufficient, a smooth cooldown and depressurization of the 7CS will occur. If the auxiliary feedwater supply is insufficient, or if it arrives too late, the steam generator heat removal capacity may eventually degrade to the point where decay heat can no longer be adequately removed. Should this occur, the RCS may begin to re heat and repressurize, and the pressurizer level may again swell.
- 8. Eventually the decay heat level will fall below the auxiliary feedwater heat - ,
removal capacity. From this point onward, the steam generator safety valves i (power-operated relief valves assumed unavailable) are used to dissipate the residual decay heat and to maintain the plant at the hot shutdown condition. For the limiting scenario, the auxiliary feedwater system is started automatically following a steam generator low-low level trip. The turbine-driven auxiliary feedwater pump utilizes steam from the secondary system and exhausts to the atmosphere. There are two motor driven auxiliary feedwater pumps with electrical power supplied by the diesel generators. Both types of pump are designed to supply rated flow within one minute of the initiating signal even if a loss of all non-emergency AC power occurs simultaneously with loss of normal feedwater. The auxiliary feedwater pumps normally take suction directly from the condensate stccage tank (CST). If the CST is unavailable, they take suction from the essential service water system. The analysis presented here shows that natural circulation in the RCS is sufficient to remove residual heat from the core, that the RCS does not over-pressurize, and that the pressurizer does not over fill. 5.2.6.2 Methodology Certain assumptions made in the analysis are dependent on the perspective from which the event is considered. The three perspectives include.DNB, RCS over. 5-85
m._._____.___.- _____ _ _ _ _ __. -.- . _ . . _ _ _ _ _ . _.__ pressurization, and pressurizer over-fill. The analysis perspective-dependent assumptions are summarized as follows. For the DNB perspective, the plant initial conditions are selected consistent with the WCNOC SCD methodology. Pressurizer pressure relief systems are assumed operational. ) For the RCS over pressurization perspective, the initial reactor power is 102% of , the uprated reactor power. A range of initial RCS ternperatures and pressures i were investigated to assure a limiting analysis. Pressurizer pressure relief systems are not credited. l For the pressurizer over. fill perspective, the initial reactor power is 102% of the - ! uprated reactor power. A range of initial RCS temperatures were investigated to assure a limiting analysis Pressurizer pressure relief systems are assumed operational to maximize the pressunzer level swell. , Certain assumptions are not dependent on the analysis perspective. These include the following:
- 1. Conservative core decay heat following reactor trip.
- 2. Heat transfer coefficients compatible with the flow conditions present (either forced flow or natural circulation following RCP trip).
- 3. Reactor trip occurs based on normal functioning of the reactor protection system (i.e., low low steam generator level, OTAT, high pressurizer -
pressure). t
- 4. The auxiliary feedwater system is actuated by the low low steam generator water level signal. One motor-driven pump is assumed
- inoperative. The remaining motor-driven pump and the turbine driven pump are assumed to deliver a total of 1000 gpm divided equally between the 4 steam generators.
- 5. The steam generator power-operated relief valves are assumed inoperative; secondary system pressure relief is provided by safety valves.
The assumptions used in the LNEAC analysis are essentially identical to the loss of ' normal feedwater flow analysis except that power is assumed to be lost to the reactor coolant pumps at the time of rector trip. 5-86 J e yy-W &W4gey--wne---y r'-M '--+1 w-g--r- V e ng,y--ggges.,-m.-Mq -.gm-r mmsgW-egg 1Wgy,i-p, 4g wgr em.,:y+-g-qq-6 y-y--e.9t- mewe-e::u- - W7g-p se.* sac:r=y,b-- >1.-g g--ta.* w r-rW w w"f- w
l l l l 5.2.6.3 Results . The transient response following a loss of non emergency AC power is shown in figures following. The corresponding sequence of events is listed in Table 5.2.61. As stated above, several analysis perspectives were considered; the results shown here are representative of the results from all three analysis perspectives. The analysis begins with a loss of normal feedwater that is a product of the loss of non. emergency AC power. This causes the SG level to fall, the RCS temperature gradually increases, and the pressurizer pressure and liquid level both rise. Due to PMTC, the reactor power also begins to rise. When the reactor trips on low low level, the reactor power drops toward decay heat levels, and the RCPs subsequently trip. Due to the reactor power reduction following reactor trip, the RCS temperature falls, as do the pressurizer pressure and level. Following reactor trip, the focus of the transient analysis becomes the falling steam generator inventory. While the inventory is high, there will be sufficient heat transfer to remove the core decay heat. However, ;f auxiliary feedwater arrives too late or at an insufficient rate, then the steam generator heat removal may fall below the reactor decay heat generation rate. The results show that this was not the case. Instead, auxiliary feedwater arrived at the steam generators at 341 seconds, and at a rate sufficient to reverse the steam generator inventory reduction. Results of the analysis show that, for the loss of non emergency AC power to plant auxiliaries event, all safety criteria are met. Since the DNBR remains above the design limit, the core is not adversely affected. Auxiliary feedwater capacity is sufficient to prevent water relief through the presaud:9r relief and safety valves; this assures that the RCS is not overpressurized. 5.2.6.4 Conclusion . Analysis of the natural circulation capability of the RCS demonstrates that sufficient long term heat removal capability exists following reactor coolant pump coastdown to prevent fuel or clad damage. 5-87
Table 5 2.6-1 Loss of Non-emergency AC Event Sequence Event Tirne (se_c] Feedwater Lost 0.1 Pressurizer PORVs Begin Cycling 41.0 Low-Low SG Level Setpoint Reached 43.8 Rods Begin to Drop 45.8 RCPs Trip 47.8 Pressurizer PORVs Complete Cycling 49.4 1st SG SV Bank Opens 51.8 2nd SG SV Bank Opens 52.8 Auxiliary Feedwater Reaches SGs 341. 5-8B
l m B 1.2 0 Power - 2,600 V / C 1 - ... Core Flow
. n 0 0 Pressure 0 0.8 -
2.400" a 3 - i O _ i g CL i e 3
" 0.6 2,c00 g o .
0 u 1 E 0.4 - 2,000 m O 2 ~ O _ Z E 0.2 1,800 u ~ ~ ~ - o , i i LL 0 100 200 300 400 Time (sec) WOLF CREEK FIGURE 5.2.61 RCS PRESSURE, CORE FLOW, AND POWER RESPONSE, LOSS OF NON. EMERGENCY AC 5-89 4. e
I 50 640 - - O O 40 g LL
's >
. 620 0
J
] ..-
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._____ .- 30 y v a w 600 - -
2
> J U -
20 ' I_ e
$580 - -
e m 10 m o 560 - Temperature Level O O 100 200 300 400 Time (sec)
- WOLF CREEK FIGURE 5.2.6 2 -
RCS TEMPERATUkE AND PRESSURIZER LEVEL LOSS OF NON EMERGENCY AC 5-90
-- . -. . - - - . . . = _ - . - _ ._ . . - . - . .
i i l l M 0,000 1,400 - l 120,000 m ' o E O
'Di 1,200 '-
- 100,000 O 3 i e 1 80,000 g h 1,000 -
m 60,000 } 0 - O. o
~ H 800 -
~ ~ " -
40,000 g _ w
- 20,000 600 - Pressure Macs
~
I i l 0 0 100 200 300 400 Time (sec) WOLF CREEK FIGURE 5.2.6 3 STEAf. GENERATOR PRESSURE AND MASS LOSS OF NON EMERGENCY AC 5-91
. . . . . . . . ., . . _ . _ __ _ ._ ~.. - - _ .
5.2.7 Loss of Normal Feedwater Flow 5.2.7.1 Introduction l A loss of normal feedwater (LONF) from pump failures valve malfunctions, or loss of offsite AC power results in a reduction in capability of tne secondary system to remove the heet generated in the reactor core. If an alternate supply of feedwater were not supplied to the plant, core residual heat following reactor trip would heat the primary system water to the point where water relief from the pressurizer would occur, resulting in a substantialloss of water frem the Reactor Coolant System (RCS). Since the plant is tripped well before the steam generator heat transfer capability is reduced, the i primary system variables never approach a DNB condition. -l For the LONF event the reactor will trip on either 1) due to turbine trip, or 2) upon reaching one of the trip setpoints in the primary and secondary systems (i.o., low low SG level, high pressur:zer pressure, OTAT). Following the loss of normal feedwater, a transient limited by the event sequence described beluw will occur.
- 1. Due to the loss of feedwatcr, the RCS heats up and pressurizes, and the pressurizer level swells. If the reactivity feedback coefficients are negative, the reactor powei can be expected to fall slightly as a consequence of RCS heatup.
However, in the presence of a positive rnoderator temperature coefficient (PMTC), which is more limiting, the power will climb slightly as a consequence of RCS heatup.
- 2. As the RCS pressurizes, the pressurizer water level will increase. Ncte that from an analysis perspective, the level increase will be maximizer';f pressurizer PORVs and spray are credited, hence, this assumption is conservative from the perspective of water relief from the pressurizer if such water relief does not occur, then it will be more conservative to disallow the pressurizer PORVs and spray, since this will tend to maximize the RCS pressure under such conditions.
-(Note - when being considered from the perspective of DNB, the analysis assumes that pressure relief systems are operational,)
- 3. Due to loss of feedwater, the steam generator liquid level will steadily decrease.
Eventually scram will occur on low-low steam generator level (or possibly on high pressurizer pressure, or OTAT).
- 4. As the steam generator prescure rises following reactor trip (and turbine trip),
the steam generator (SG) safety valves (SVs) lift to dissipate the sensible heat of the fuel and coolant plus the decay heat produced in the core. (Note that both the SG power operated relief valves (PORVs) and turbine bypass to the condenser are assumed to be unavailable.) 5-92
l 1 i S. A short time after scram, auxiliary feedwater is activated to replace the fluid
. discharged through the steam generator SVs. If the auxiliary feedwater supply is sufficient, a smooth cooldown and depressurization of the RCS will occur. If the auxillary feodwater supply is insufficient, or if it arrives too late, the steam generator heat removal capacity may eventually degrade to the point where decay heat can no longer be adequately removed. Should this occur, the RCS :
may begin to re heat and repressurize, and the pressurizer level may again ; swell. ,
- 6. Eventually the decay heat level will fall below the auxiliary feedwater heat removal capacity. From this point onward, the steam generator safety valves -
(power operated relief valves assumed unavailable) are used to dissipate the . residual decay heat and to maintain the plant at the hot shutdown condition. For the limiting scenario, the auxiliary feedwater system is started automatically following a steam generator low low level trip. The turbine-driven auxiliary feedwater pump utilizes steam from the secondary system and exhausts to the atmosphere. , There are two motor-driven auxiliary feedwater pumps with electrical power supplied by ) the diesel generators. Both types of pump are designed to supply rated flow within one minute of the initiating signal. The auxiliary feedwater pumps normally take suction directly from the condensate storage tank (CST). If the CST is unavailable, they take suction from the essential service water system. The analysis presented here shows that following a loss of normal feedwater, the RCS does not over. pressurize, that the pressurizer does not over-fill, that DNB is not a concern, and in particular that the auxiliary feedwater system is capable of removing the stored and residual heat. 5.2.7.2 Methodology Certain assumphons made in the analysis are dependent on the perspective from which the event is considered. The three perspectives include DNB, RCS over. pressurization, and pressurizer over. fill. The analysis perspective-dependent assumptions are summarized as follows. For the DNB perspective; the plant initial conditions are selected consistent with , the WCNOC SCD methodology. Pressurizer pressure relief systems are assumed operational. For the RCS over-pressurization perspective, the initial reactor power is 102% of , the uprated reactor power. - A range of initial RCS temperatures and pressures - 5-93
'" *
- y -- ., moe' ym ., %, ,,.,
were investigated to assure a limiting analysis. Pressurizer pressure relief systems are not credited. For the pressurizer over-fill perspective, the initial reactor power is 102% of the i uprated reactor power. A range of initial RCS temperatures were investigated to assure a limiting analysis. Pressurizer pressure relief systems are assumed l operational to maximize the pressurizer level swell. Certain assurT.ptions are not dependent on the analysis perspective. These include the following:
- 1. Conservative core decay heat following reactor trip.
- 2. Heat transfer coefficients compatible with the flow conditions present.
- 3. Reactor trip occurs based on normal functioning of the reactor protection system (i.e., low-low steam generator level, OTAT, high pressurizer pressure).
- 4. The auxiliary feedwater system is actuated by the low low steam generator water level signal. One motor driven pump is assumed inoperative. The remaining motor-driven pump and the turbine-driven pump are assumed to deliver a total of 1000 gpm divided equally between the 4 steam generators.
- 5. The steam generator _ power-operated relief valves are assumed -
inoperative; secondary system pressure relief is provided by safety valves. 5.2.7.3 Results The translent response following a loss of normal feedwater is shown in figures following. The corresponding sequence of events is listed in Table 5.2.7-1. As stated above, several analysis perspectives were considered; the results shown here are representative of the results from all three analysis perspectives. The analysis begins with a loss of nor.nal feedwater, This causes the SG level to fall, the RCS temperature gradually increases, and the pressurizer pressure and liquid level both rise. Due to PMTC, the reactor power also begins to rise. When the reactor trips on low low SG level, the reactor power drops toward decay heat levels Due to the reactor power reduction, tho RCS temperature falls, as do the pressurizer pressure and level. 5-94 p
Following reactor trip, the focus of the transient analysis becomes the falling steam generator inventory. While the inventory is high, there will be sufficient heat transfer to remove the core decay heat. However, if auxiliary feedwater arrives to late or at an insufficient rate, then the steam generator heat removal may fall below the reactor decay heat generation rate. The results show that this was not the case. Instead, auxiliary feedwater arrived at the steam generators at 336 seconds, and at a rate sufficient to reverse the steam generator inventory reduction. 5.2.7.4 Conclusions Results of the analysis show that, for the loss of normal feedwater event, all safety critoria are met. Since the DNBR remains above tbr 'isign limit, the core is not adversely affected. Auxillary feedwate'r capacity is sufficient to prevent water relief through the pressurizer relief and safety valves; this assures that the RCS is not overpressurized. I i I I i 5 95
i Table 5.2.7-1 Loss of Norrnal Feedwater Event Sequence Event Time (sec) Feedwater Lost 0.1 Low Low SG Level Trip Setpoint Reached 39.4 Pressurizer PORV Cycling Begins 39.49 Rods Begin To Drop 41.4 Pressurizer PORV Cycling Ends 44.95 1st SG SV Bank Opens 48, 2nd SG SV Bank Opens 49. Auxiliary Feedwater Reaches SGs 336. 5 96
I i l l 1.2 2,800 l Power . O #
~
E @.gssq,e 2,600 m .
~ '5 _
S l
@0.G O I 2,400 e c -
t O . - - 3
'O O.6 -
u) v (n O C 2,200 ') u - D 0.4 - Q. u , W o O _ '~
---. . ............._.... ____ - 2,ooo M 0 1 0 100 200 300 400,800 Time (sec)
WOLF CREEK FIGtIRE 5.2.7-1 RCS TEMPERATURE, CORE FLOW, AND POWER RESPONSE-LOSS OF NORMAL FEEDWATER l l 5-97 L i
50
^tu 600 - Temperature .
n CD Q) Level 40 "
/ \ Q)
V , N > v 580 -
/ s .
c)- s _J G '
..- 'y 30 .g E
c) H 560 - s...'"~ ' - - --....__ ..,____ . ...___,,_ _,- E 2 . cn - 20 $i r' O u .N-y540 -
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< m .
10 ~ c) M m u Z 520 - - b O i 0 100 200- 300- 400 Time (sec) . WOLF CREEK FIGURE 5.2.7-2 RCS TEMPERATURE AND PRESSURIZER LEVEL LOSS OF NORMAL FEEDWATER - r
+
5 98 j ,,~.; ,-na--.--,-,,,-,,--,h.n,,,,.n,.- --r, ...,-.,.n-. a r. . n e w m. w. n .s , ,, : .,w,,--..--., --,.,-,.,n.,.:w. .e r. .. e ,n p c-,,+,_,;,rrr. ,,,,w,-.,,,-, er-n.+ r..,,
1,300 120,b00 ^E Pressure - f2
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v s Vans 100,000 cn m 1 ,200 '. o
$ \ 'E u -
80,000 Q. - ',
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40,000 $ 0 c 1000 -
.._,,,,"'" ~ "~ " '; O E o 20,000 E
(n . o 3 900 O (A 0 100 200 300 400 Time (sec) WOLF CREEK FIGURE 5.2.7 3 STEAM GENERATOR PRESSURE AND MASS LOSS OF NORMAL FEEDWATER 5-99
l 5.2.8 FEEDWATER SYSTEM PIPE BREAK S.2.8.1 Introduction A major feedwater line rupture is defined as a break in a feedwater line large enough to pr6 vent the addition of sufficient feedwater to the steam generators to maintain shell side fluid inventory in the steam generators. If the break is postulated in a feedwater line between the check valve and the steam generator, fluid from the steam generator may "so be discharged through the break. Further, a break in this location could preciude the subsequent addition of auxiliary feedwater to the affected stearn generator. (A break upstream of the feedwater line check valve would affect the NSSS only as a loss of feedwater. This case is covered by the evaluation in Sections 5.2.6 , and 5.2.7.) Depending upon the size of the break and the plant operating conditions at the time of the break, the break could cause (.ither an RCS cooldown (by excessive energy discharge through the break) or an RCS heatup. Potential RCS cooldown resulting from a secondary pipe rupture is evaluated in Section 5.1.5. Therefore, only the RCS heatup effects are evaluated for a feedwater line rupture. A feedwater line rupture reduces the ability to remove heat generated by the core from the RCS for the following reasons:
- a. Feedwater flow to the steam generators is reduced. Since feedwater is subcooled, its loss may cause reactor coolant temperatures to increaso prior to reactor trip.
- b. Fluid in the steam generator may be discharged through the break, and would then not be available for decay heat removal after trip.
- c. The break may be large enough to prevent the addition of any main feedwater after trip.
An auxiliary feedwater system functions to ensure the availability of adequate feedwater so that:
- a. No substantial overpressurizaton of the RCS occurs (less than 110 percent of design presrures);_and
- b. Sufficient liquid in the RCS (3 malrteined so that the core remains in place and geometricatiy ;rtt with no loss of core cooling capability.
The severity of the feedwater line rupturd tr&.isient depends on a number of system-parameters, including break size, initial roactor power, and credit taken for the l 5100
_ . . _ _. -_- ..-. -_ . - . . . _ _ _ _ - _ . . . _ _ . . .~ functioning of various control and safety systems. The major assumptions pertinent to this analysis are defined below. The main feedwater control system is assut.ed to fail due to an adverse environment. The water levels in all steam generators are assumed to decrease equally until the low-low steam generator level reactor trip setpoint is reached. After reactor trip, a double-ended rupture of the largest feedwater line is assumed. These assamptions conservatively bound the most limiting feedwater line rupture that can occur. Analyses have been performed at full power, with and without loss of offsite power, and with no credit taken for the pressurizer power-operated relief valves. For the case without offsite power available, the power is assumed to be lost at the time of reactor trip. This is more conservative than the case where power is lost at the initiation of the event. These cases are analyzed below, i 5.2.8.2 Methodology The feedwater line rupture event is analyzed assuming initial conditions including maximum allowances for calibration and measurement error. Because no bulk boiling occurs in the hotleg, the core is assummed to remain intact and in a coolable geometry with no subsequent fuel damage. Major assumptions used in the analysis are as follows:
- a. Plant charac.teristics and initial conditions are discussed in Section 5.0,
- b. No credit is taken for the pressurizer power-operated relief valves or pressurizer spray.
- c. Initial pressurizer level is at the nominal progremmed value plus 5 percent (error); initial steam generator water level is at the nominal value.
- d. No credit is taken for the high pressurizer pressure reactor trip.
- e. Main feedwater to all steam generators is assumed to stop at the time the break occurs (all main feedwater spills out through the break).
- f. The worst possible break area is assumed. This maximizes the blowdown discharge rate following the time of trip, which maximizes the resultant heatup of the reactor coolant.
- g. A bounding feedwater line break discharge quality is assumed.
- h. Reactor trip is assumed to be initiated when the low low steam generator level reaches O percent of narrow range span in the ruptured steam generator.
5 101 v-rs - A s-- g- w -
- i. The auxiliary feedwater system is actuated by the low-low steam generator water level signal. The auxiliary feedwater system is assumed to supply a total of 1000 gpm to thres unaffected steam generators, including allowance for possible spillage through the main feedwater line break. A 60 second delay was assumed following the low-low level signal to allow time for startup of the standby diesel generators and the auxiliary feedwater pumps, An additional 374 seconds was assumed before the feedwater lines were purged and the relatively cold (120 F) auxiliary feedwater entered the unaffected steam generators.
J. No credit is taken for heat energy deposited in RCS metal during the RCS heatup.
- k. No credit is taken for charging or letdown.
- l. Steam generator heat transfer area is assumed to decrease as the shell side liquid inventory decreases,
- m. Conservative core residual heat generation is assumed based upon long-term operation at the initial power level preceding the trip.
- n. No credit is taken for the following potential protection logic signals to mitigate the consequences of the accident:
- 1. High pressurizer pressure
- 2. Overtemperature AT
- 3. High pressurizer level
- 4. High containment pressure Receipt of a low low steam generator water level signal in at least one steam generator starts the motor-driven auxiliary feedwater pumps, which in turn initiate auxiliary feedwater flow to the steam generators. The turbine-driven auxiliary feedwater pump is initiated if the low-low steam generator _ water level signal is reached in at least two steam generators Similarly, receipt of a low steam line pressure signal in at least one steam line initiates a steam line isolation signal which closes ali ....in steam line isolation valves. This signal also gives a safety injection signal which initiates flow of cold borated water into the RCS. The amount of safety injection flow is a function of RCS pressure.
Emergency operating procedures folicwing a fet: ..ater system pipe rupture require the following actions to be taken by the reactor operator:
- a. Isolate feedwater flow spilling from the ruptured feed-water line and align the system so that the level in the intact steam generators is recovered, 5-102 . .
,wn - 6- y y. - ,,- , - - - --m --w- - , - - wg-
i
- b. High head safety injection should be terminated in accordance with the emergency operating procedures. ,
Subsequent to terminating high head safety injection, plant operating procedures are followed in cooling the plant to a safe shutdown condition. No reactor control systems are assumed to function. The reactor protection system is required to function following a feedwater line rupture as analyzed here. No single act!ve failure will prevent operation of this system. i The engineered safety systems assumed to function are the auxiliary feedwater system and the safety injection system. For the auxiliary feedwater systern, the worst case , configuration has been used, i.e., only three intact steam generators receive auxiliary foodwater following the break and one motor driven auxiliary feedwater pump is assumed to fail. A discharge flow control device located on the auxiliary feedwater line to each steam generator regulates the flow from the motor-driven auxillary feedwater pump to the affected steam generator. This ensures a minimum flow of 250 gpm is delivered to each intact steam generator. Thus at least 750 gpm is delivered to the three intact steam generators. For the case without offsite power, there is a flow coastdown until flow in the loops reaches the natural circulation value. The natural circulation capability of the RCS has been shown (in Section 5.2.6) to be sufficient to remove core decay heat following reactor trip, for the loss of ac power tra sient. Pump coastdown characteristics are demonstrated in Sections 5.3.1 and 5.3.2 for single and multiple reactor coolant pump trips, respectively. 5.2.8.3 Results Calculated plant parameters following a major feedwater line rupture are shown in Figures 5.2.8-1.hrough 5.2.8-10. Results for the case with offsite power available are presented in Figures 5.2.8-1 through 5.2.8-5. Results for the case where offsite power is lost are prer ^md in Figures 5.2.8-6 through 5.2.810. The calculated sequence of events for by mes analyzed are listed in Table 5.2.81. The system response following the feedwater line rupture is similar for both cases analyzed. Results presented in . Figures 5.2.8-2 and 5.2.8-3 (with offsite power available) and Figures 5.2.8-7 and 5.2.8-8 (without offsite power) show that pressures in the RCS and main steam system remain below 110 percent of the respective design pressures. Pressurizer pressure decreases after reactor trip on low-low steam generator level. Figures 5.2.81 and 5.2.8-6 show that following reactor trip the plant remains suberitical. 5-103
i Figure 5.2.8-2 shows that the pressurizer does not empty throughout the transient so - g that the core remains covered at all times and that no boiling occurs in the r Jactor l coolant loops. . The major difference between the two cases analyzed can be seen in the plots of hot and cold leg temperatures, Figure 5.2.8-4 (with offsite power available) and Figure 5.2.8-9 (without offsite power). It is apparent that for the initial transient, the case without offsite power results in higher temperatures in the hot leg. For longer times, however, the case with offsite power results in a more severe rise in temperature until the auxiliary feodwater system is realigned. AFWS realignment is not assumed, however, in these cases. The pressurizer fills more rapidly for the case with power-due to the increased coolant expansion resulting from the pump heat addition. 5.2.8.4 Conclusions Results of the analyses show that for the postulated feedwater line rupture, the assumed auxiliary feedwater system capacity is adequate to remove decay heat, to , prevent overpresser"' v ^e RCS, and to prevent uncovering the reactor core. Radioactivity doses froi,, . postulated feedwater lines rupture are less than those previously presented for the postulated steam line break. All an.o'icable acceptance criteria are therefore met. i 5-104
1 Table 5.2.8-1
.. Feedwater Line Break Time Sequence of Events '
. Time Accident Event tsag)
Feedwater system pipe break
- 1. Case 1 (offsite power available) ,
Feedwater control system 10.0 fails Low-low steam generator 52.5 lavel reactor trip set-point reached in all steam generators Rods begin to drop an_d 54.5 feedwater line rupture occurs Steam generator safety . 55.7 valve setpoint reached in intact steam generators Low steam line pressure 83.9 setpoint reached in rup-tureed steam generators All mtin steamline iso- 90.9 . lation valves close Auxiliary feedwater to - 426.2 intact steam generators is initiated-Pressurizer safety valve 661.8 setpoint reached following feedwater line rupture Core decay heat de- . 1690.0 creases to auxiliary 4 feedwater heat removal capacity 5-105
Table 5.2.8-1'(sheet 2) Feedwater Line Break Time Sequence of Events
- 2. Case 2 (concurrent loss of offsite power)
Feedwater control system 10.0 fails Low-low steam generator 52.3 level reactor trip set- ' point reached in all steam generators Steam generator safety 53.9 valve setpoint reached in intact steam generators Rods begin to drop and o4.3 feedwater line rupture occurs Low steam line pressure 80.9 setpoint reached in rup-tureed steam generators-All main steamline iso- 87.9 lation valves ciose Auxiliary feedwater to 426.2-intact steam generatc.rs is initiated Core decay heat de- 664.0 creases to auxiliary feedwater heat removal capacity-5-106
14 m o 1,2 - f _
- o A 0.8 -
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, 0.6 0 0,4 0.2 1 3 10 30 100 300 1,000 3,000 10,000 Tim e (s e c) ,.
0.01 0 1 ( (0.01) - a u (0.0 2 ) -
~
r (0.0 3 ) - (0.04)1 3 to 30 100 300 1.000 3,000 Tim e (s e c) WOLF CREEK FIGURE 5.2.81 NUCLEAR POWER, CORE HEAT FLUX AND TOTAL CORE REACTIVITY (PCM) FOR MAIN FEEDWATER LINE RUPTURE WITH OFFSITE POWER AVAILABLE 5-107
. -_ _ __ - . ._ _ - . . ._ -._ ..._- _..~-_.._ . _. . . . . _ _ . . _ _ . -
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; 1,600 E
j_ 1,400 51 ,200 - h 5 1,000 - e
$- 800 -
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' 600 400 1- 3 -10 -30 100 300 1,000 3,000 10,000 Tlme (sec)
WOLF CREEK - < FIGURE 5.2.8 PRESSURIZER AND MAXIMUM SYSTEM PRESSURE AND PRESSURIZER WATER VOLUME FOR-FEEDWATER LINE RUPTURE WITH OFFSITE POWER AVAILABLE-5-108 e -- -><-e--- .= ;... .y , y -e-
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. Time (sec)
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j. L WOLF CREEK - FIGURE 5.2.8-3 , m -i l REACTOR COOLANT MASS FLOW RATE AND FEEDWATER LINE BREAK , FLOW RATE FOR MAIN FEEDWATER 1 ,. LINE RUPTURE WITH OFFSITE ' L POWER AVAILABLE I i 5-109 i , e
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j HOT LEG j a E 600 - O y COLD LEG $550 - 8 Y \ j 500 2 450 1 3 10 30 100 300 1,000 3,000 10,000 Time (sec) 700 C 680 - 660 SATURATION T 640 -
, HOTLEG
[620 600 - 8 f580 3 COI O LEG S 560 - 540 1 3 10 30 100 300 1,000 3,000 10,000 Time (sec) WOLF CREEK FIGURE 5.2.8-4 REACTOR COOLANT TEMPERATURE (FAULTED LOOP) AND REACTOR COOLANT TEMPERATURE (INTACT LOOP) FOR MAIN FEEDWATER LINE RUPTURE WITH OFFSITE POWER AVAILABLE G-110
1,400
- 3. 1,200 -
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$ 1,000 -
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)E 600 -
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f 1,4 s 5
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4 0.8 - 2 0.6 - 2 0.4 - 0.2 - 0 10,000-1 3 10 30 100 300 1.000 3,000 Tim e (soc) 1.4 e 2 1.2 - 5 2 1 u.
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.=?.
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q
~ 2,800
.5 e 2,600 -
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El 2,200 . Pressurizer 5 3 2,000 - b 1,800 - 8 1 3 10 30 100 300 1,000 3, BOO 10,000 , Time (sec) - 2,000 E 1,800 - d - g 1,600 m 3 1,400 - 1 Y \ 5 1,200 - b 1,000 - 8 E 800 - 600 1 3 10 30- 100 300- 1,000 3,000 10,000 Time.(sec) WOLF CREEK FIGURE 5.2.8-7 PRESSURIZER AND MAXIMUM
-SYSTEM PRESSURE AND ~
PRESSURIZER WATER VOLUME FOR
-FEEDWATER LINE RUPTURE WITHOUT OFFSITE POWER :
AVAILABLE 5-113
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WOLF CREEK - FIGURE 5.2.8 REACTOR COOLANT MASS FLOW RATE AND FEEDWATER LINE BREAK FLOW RATE FOR MAIN FEEDWATER ' LINE RUPTURE WITHOUT OFFSITE~ POWER AVAILABLE 5-114 t -~ am'r
. ~ . _ _ . . .~.. _ . - . . - . _ . _ . . _ . -. _ _ _ . . .
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SATURATION f HOT LEG
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, SATURATION 640 -
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1 3' 10 30 100 3^') 10,000 ~ Time (sec) . i WOLF CREEK FIGURE 5.2.8 9 REACTOR COOLANTTEMPERATURE (FAULTED LOOP) AND REACTOR COOLANT TEMPERATURE (INTACT- 1
- LOOP) FOR MAIN FEEDWATER LINE RUPTURE WITHOUT OFFSITE POWER AVAILABLE 5-115
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? $ 1,200 -
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$ 1,000 -
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) 600 -
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- ,.- - -- . - ~ - .
5.3 Decrease in Reactor Coolant System Flowrate 5.3.1 Partial Loss of Forced Reactor Coolant Flow 5.3.1.1 Introduction A partial loss of coolant flow accident can result from a mechanical or electrical failure in a reactor coolant pump, or from a fault in the power supply to the pump or pumps supplied by a reactor coolant pump bus, if the reactor is at power at the time of the accident, the immediate e fect of loss of coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor is not tripped promptly. 5.3.1.2 Methodoingy The loss of two reactor coolant pumps with four loops in operation has been analyzed. The accident is analyzed using the Statistical Core Design Methodology. The analysis was performed using rerate conditions of 3579 MWt, +6.0 pcm/*F MTC, and 10% steam generator tube plugging. Initial Conditions Consistent with the statistical core design approach, nominal values are assumed for the initial reactor power, pressure and RCS temperature along with a minimum measured RCS flow rate. Reactivity Coefficients A conservatively large absolute of the Doppler-only power coefficient is used. A positive moderator temperature of +6.0 pcm/*F is assumed. This results in the maximum core power during the initial part of the transient, when the minimum DNBR is
. reached.
5.3.1.3 Results Figures 5.3.1-1 through 5.3.1-5 show the transient response for the loss of two reactor coolant pumps with four loops in operation Since DNB does not occur, the ability of the primary coolant to remove heat from the . fuel rod is not greatly reduced. Thus, the average fuel and clad temperatures Jo not l incrense significantly above their respective initial values. 1 . ! The calculated sequence of events for the case analyzed is shown in Table 5.3.1-1. The ah~ected reactor coolant pumps will continue to. coast down, and the core flow will. reach a new equilibrium value corresponding to the pumps still in operation. [ l i 5-117 m w- g --- e e- -
a
)
5.3.1.4 Conclusions the analysis shows that the DNBR will not decrease below the safety analysis limit values at any time during the transient. Thus no fuel or clad damage is predicted l and all applicable acceptance criteria are met.
-Table 5.3-1 Time Sequence of Events Loss of two pumps Coastdown begins 0.0 with four loops in operation Low flow reactor trip 2.58 Rods begin to drop 3.58 Minimum DNBR occurs 4.0 f-5-118-
.. . .- . . . . .. = _ . _ . ~ - . .
2.360 2.340 - 2,320 - a w 2,300 - E 2,280 - 2 - ei ,260 d 2,240 - E 2,220 - 2,200 ' ' ' ' O 2 4 6 8 10 itME (SECONDS'l WOLF CREEK Figure S.3.1- 1 Pressurizer Preessure Transients for four loops in operation,two pumps coasting down 5-119
, . , . . . - - - ... . . . . . - , . ~ . . - . . . _ , . . .
1.05
- 1 N
5 g 0.95 - 6 2 0.9 - Q b h 0.85 8 d 0.8 - w u 0.75 -
, i t '
O.7 0 2 4 6 8 10 Tlut (SECONDS) WOLF CREEK figure 5.3.1-2 Totol Core flow Transient for four loops in operation,two pumps Coosting down r 4 4 5-120 -
1.1 i
@ 0.9 b
0,8 - a 0.7 u 8 d 0.6 - w 0.5 - 0.4 O 2 4 6 8 to TiuE (SCCONOS) WOLF CNEEK
' ' ~ ~
Figure $.3.1-3 f aulted Loop Flow Transient for four loops in operation,two pumps coosting down l 1 121
- . . _ ~ . - . - - - - - - . . . - - - . - . . - . . -
. . .,. ---- -. .- n_,..-....- -_
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6 i t a: -i W l { 0.4 er 6 ! d 0.2 - ax < 0 l 0 2 4 6 8 10 r l TIME (SECON0b) WOLF CREEK , i Ficure 5.3.1-4 Nuclear Power Transient i for four loops in operation,two i pumps Coonting down e i h i 1 i F s o a F
- i.
.I i
5-122
.6 a
, - - - m mm mm--
. . _ . . . - - - . ~ . . . . . ~ . - - . _ . . . . ~ - - . _ ~ . - . . . - - . - - . . . . - - - . - - - - . - - . - - . - . - - . - - . . . . - - - - - -
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< 0.4 -
i w ! ' r , u 0.3 i ! jo 0 2 4 6 8 TNE (SECONOS) WOLJ CREEK , figure 5.3,1-5 Core Heat Flux Transient for four loops in operation,two pumps toosting down 4 i 4 5-123
5 3.2 COMPLETE LOSS OF FORCED REACTOR COOLANT FLOW 5.3.
2.1 INTRODUCTION
A complete loss of flow accident may result from a simultaneous loss of electrical supplies to all reactor coolant pumps. If the reactor is at power at the time of the accident, the immediate effect of loss-of-coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor were not tripped promptly. Normal pcwer for the reactor coolant pumps is supplied through busses from a transformer connected to the generator. When a generator trip occurs, the busses are automatically transferred to a transformer supplied from external power lines; the pumps continue to supply coolant flow to the core. Following any turbine trip where there are no electrical faults which require the generator to trip from the network, the generator remains connected to the network before any transfer is made, thus the pumps remain connected to the generator. This ensurs full flow for 30 seconds after the reactor trip. Protection against this event is provided by the low RCS flow trip and by reactor coolant pump power supply undervoltage and underfrequency trips. Reactor trip on reactor coolant pump undervoltage is provided to protect against conditions which can cause a loss of voltage to all reactor coolant pumps, i.e., station blackout. This function is blocked below approximately 10-percent power. The reactor trip on reactor coolant pump underfrequency is provided to trip the reactor for an underfrequency condition, resulting from frequency disturbances on the power grid. A complete loss of forced reactor coolant flow is classified as an ANS Condition 111 event, an infrequent fault. The accident is not analyzed for operation with fewer than four pumps running since the WCGS Technical Specifications prohibit continuous operation under such conditions. 5.3.2.2 METHODOLOGY The complete loss of forced reactor coolant flow accident is analyzed to verify that the following two acceptance criteria are met for uprated conditions:
- a. the accident shall not result in a pressure boundary breach, and
- b. the accident shall not result in damage to the fuel.
The method of analysis and the assumptions made regarding initial operating conditions and reactivity coefficients are identical to those used in the Partial Loss of Forced Reactor Coolant Flow, see Section 5.3.1, except J,at, following the loss of power supply to all pumps at power, a reactor trip is actuated by either reactor coolant pump power supply undervoltage or underfrequency. 5-124 I I o_ . _ _ _ _ .___ _ _ _ _ _ _ _ . _ _ ____m___ ____. _ _ _ _ _ _
5.3.2.3 RESULTS Figures 5.3.2-1 and 5.3.2-2 show the transient response for the loss of power to all reactor coolant pumps with four loops in operation. The reactor is assumed to be tripped on an undervoltage signal Figure 5.3.2-2 shows that the DNB criterion is met. For the case analyzed, the plant is tripped by the undervoltage trip sufficiently fast to ensure that the ability of the primary coolant to remove heat from the fuel rod is not greatly reduced. Thus, the average fuel and clad temperatures do not increase significantly above their respective initial values. The calculated sequence of events for the event analyzed is shown on Table 5.3.2-1. Following reactor trip, the reactor coolant pumps continue to coast down, eventually establishing natural circulation and stable plant conditions. Normal plant shutdown may then proceed. ! 5.3.2.4 Conclusions The an ' lysis performed has demonstrated that, for the complete lors of forced reactor coolant flow, the overpressurization and DNB critera are met at all times during the transient; all applicable acceptance criteria are met. l. 5-125
Table 5.3.2-1 TIME SEQUENCE OF EVENTS Time Apcident . Event . (sec) . Complete loss of forced - reactor coolant flow
- 1. Loss of four pu,mps with four loops in -
operation All operating pumps 0.0 lose power and begin coasting down Reactor coolant pump 0.0 undervoltage trip point reached Rods oegin to drop 1.5 Minimum DNBR occurs 3.5 l' { l 5-126
i.1 1.0 0.9 - t i'
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.[ 2.3E+3 S
2 y 2.3E+3 - E n. w
$ 2.2E+3 5
[ 2.2E+3 2,1 E+ 3 0 2 4 6 8 10 Time (sec.) WOLF CREEK FIGURE 5.3.2-1 CORE FLOW AND. PRESSURIZER PRESSURE I 5-127
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~=
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g 0.8 - 8 5 0.6 - E t. 6 j
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$ 2.4 8
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~
WOLF CREEK-- < FIGURE 5.3.2-1: i
' CORE POWER AND.
MINIMUM DNBR .. 1
- l. 5-128 ,
t L
- i
~ . - - ... - . 5.3.3 Reactor Coolant Pump Shaft Seizure (Locked Rotor) 5.3.3.1 Introduction The accident postulated is an instantaneous seizure of a reactor coolant pump rotor. Flow through the affected reactor coolant loop is rapidly reduced, leading to an initiation of a reactor trip on a low flow signal. Following initiation of the reactor trip, heat stored in the fuel rods continue to be transferred to the coolant causing the coolant to expand. At the same time, heat transfer to the shell side of the steam generator is reduced, first because the reduced flow results in a decreased tube side film coefficient and then because the reactor coolant in the tubes cools down while the shell side temperature increases (turbine steam flow is reduced to zero upon plant trip). The rapid expansion of the coolant in the reactor core, combined with the 6 educed heat transfer in the steam generator _ causes an insurge into the pressurizer and a pressure increase throughout the reactor coolant system. The insurge into the pressurizer causes a pressure increase which in turn actuates the automatic spray system, opens the power-operated relief valves and opens the prescurizer safety valves in a sequence dependent on the rate of insurge and pressure increase. The power-operated relief valves are designed for reliable operation and would be expected to function properly during the accident; however, for conservatism, their pressure-reducing effect as well as the pressure-reducing effect of the spray are nnt included in this analysis. The general acceptance criteria applicable to this event is that, at no time during the transient can the RCS pressure exceed that which corresponds to the faulted condition stress limit.- Also, the peak clad surface temperature calculated for the hot spot must remain below 2700*F. Finally, the analysis for this event includes the calculation of the percerbge of fuel rods postulated to experience DNB. this value is used as input to the consideration of the radiolopcal consequences associated with this event. 5.3.3.2 Methodology Explicit analysis is performed for four loop operating conditions, without offsite power available. The result for the cases without offsite power bound those with offsite power available. This analysis assumes a upgraded power level, and a +6.0 pcm/*F MTC.
- Statistical Design methodology was used in the DNB analysis.
5.3.3.3 Results 5-129
- The transient results for the four loop locked rotor accle etc are shown in Figures 5.3.3-1 through 5.3.3 5. The calculated sequence of events for tnis case is shown in Table 5.3.3-1. - After pump seizure, the neutron flux is rapidly reduced by control rod insertion.-
Rod motic,n is ascumed to begin one second after the flow in the affected loop reaches 87 percent of nominal flow. The peak RCS pressure reached during the transient is - less than that which would cause stresses to exceed the faulted condition stress limits of the ASME Code Section Ill. Also, the peak clad temperature is considerably less than the 2700*F. These results represent the most limiting conditions with respect to the locked rotor event or the pump shaft break. 5.3.3.4 Conclusions Since the peak RCS pressure reached during the transient is less than that which would cause stresses to exceed the faulted condition stress limits., the integrity of the prirr.ary coolant system is nc' endangered. Since the peak clad surface temperature calculated for the hot spot during the worst transient remains considerably less than 2700'F the core will remain in place and intact with no loss of core cooling capability. L L l I 5-130
Table 5.3.3-1 Time Sequence of Events One locked rotor with Rotor on one pump 0.0 sec four loops in operation locks W/ offsite power Low flow trip point 1.042 sec reached Rod begin to drop 2.042 sec Maximum RCS pressure 3.7 see occurs One locked rotor with Rotor on one pump 0.0 sec four loops ;n operation locks W/O offsite power Low flow trip point 1.042 sec reached Rods begin to drop 2.042 see Power lost to remaining 3.042 sec pumps Maximum RCS pressure 3.7 sec occurs pumb 5-131
2.000 , l 2,5$0 - n h 2,500 - w ; W ' 3 2,450 - M O .e0 tt 2,350 - ' 2,300 - w/ Or# 511UPOWFR W/0 Orr,5fi,[ POWER . , q
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WC'J CREEK figure 6,3.5- 1 Pressuriser l'reess6pe Tronsients Lothed Rotc# with ohn without of f site pow er 6-132.
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W/ Off$iTC POWER l 1 g 1 . w/0 OrFSy,E POWER l 9 b 0.8 - 9 t~ cr 0.6 - 6 tr 0.4 - cr b 0 .2 z O O 2 4 6 8 10 TIME (SECONDS) WOLF CREEK figure S.3.3-4 Normalised power transient Locked Rotor with and without offsite power i l I i i l 5-135 i l
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t 5.3.4 Reactor Coolant Pump Shaft Break 5.3.4.1 Introduction i The accident is postulated as an instantaneous failure of a reactor coolant pump shaft. I Flow through the affected reactor coolant loop is rapidly reduced, though the initial rate s~ of reduction of coolant flow is greater for the reactor coolant pump rotor seizure event. Reactor trip is initiated on a low flow signal in the affected loop. Following initiation of the reactor trip, heat stored in the fuel rods continues to be ' transferred to the coolant, causing the coolant to expand. At the same time, heat transfer to the shell side of the steam generators is reduced. This is because the reduced flow results in a decreased tube side film coefficient and the reactor coolant in the tubes cools down while the shell side temperature increases (turbine steam flow is reduced to zero upon plant trip). The rapid expansion of the coolant in the reactor ' core, combined with reduced heat transfer in the steam generators, causes an insurge into the pressurizer and a pressure increase throughout the RCS. The insurge into the pressurizer compresses the steam volume, actuates the automatic spray system, opens the power-operated relief valves, and opens the pressurizer safety valves, in that sequence. The power-operated relief valves are designed for reliable operation and would be expected to function properly during the accident. However, for conservatism, their pressure reducing effect, as well as the pressure reducing effect of the spray, is not included in the analysis. 5.3,4.2 Conclusions
- The consequences of a reactor coolant pump shaft break are no worse than those calculated for the locked rotor incident. With a failed shaft, the impeller could conceivably be free to spin in a reverse direction as opposed to being fixed in position as assumed in the locked rotor analysis. However, the not effect on core flow is negligible, resi.iting in only a slight decrease in the end point (steady state) core flow.
For both the shaft break and locked rotor incidents, reactor trip occurs very early in the transient. l l l l 5-137 l
5,4 Reactivity and Power Distribution Anomalies 5.4.1 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal from a Subcritical e or Low Power Startup Condition 1 5.4.1.1 Introduction A rod cluster control assembly (RCCA) withdrawal accident is defined as an uncontrolled addition of reactivity to the reactor core caused by withdrawal of RCCAs ; resulting in a power excursion. Such a transient could be caused by a malfunction of the rod control system. This could occur with the reactor either subcritical, at hot zero power or at power The "at power" case is discussed in Section 5.4.2. r Should a continuous RCCA withdrawal accident occur, the automatic features of the reactor protection system available to terminate the transient are as follows;
- 1. Source range high_ neutron flux reactor trip - actuated when either of two independent source range channels indicate a neutron flux level above a preselected manually adjustable setpoint. Although the trip function may be manually bypassed after a specified flux level is reached in the intermediate range flux channel, it is automatically reinstated when both intermediate range channels indicate a flux below a specified level.
- 2. Intermediate range high neutron reactor flux trip - actuated when two out of the four power range channels indicate a power level above approximately 25 % of full power. Though this trip function may be manually bypassed when the power channels are above 10% of full power, it is automatically reinstated only after _
three of the four channels indicate a power level below approximately below 10% of full power,
- 3. Power Range High Neutron Flux Reactor Trip (Low Setting)- actuated when two out of the four power range channels indicate a power level above approximately 25 % of full power, Though this trip may be manually bypassed when the power range channels indicate a power level above 10% of full power, it is .
automatically reinstated when the power range channel indicates a power level below this value.
- 3. Power range high neutron flux reactor trip (low setting) actuated when power range channels indicate a power level above approximately 25% of full power.
This trip function may be manually bypassed above 10% of full power, it is automatically reinstated below approximately 10% power level.
- 4. Power range high neutron flux reactor trip (high setting) - actuated when the ~
power range channels indicate a power level above a preset setpoint. This trip function is always active. 5-138 n- , ~
- 5. High nuclear flux rate reactor trip - actuated when the positive rate of change of neutron flux indicates a rate above a preset setpoint. This trip function is always active.
in addition, control rod stops on high intermediate range flux and high power range flux serve t) discontinue rod withdrawal and may prevent the need to actuate the intermediate range flux trip and the power range flux trip, respectively. 5.4.1.2 Methodology The uncontrolled RCCA bann Jthdrawal from subcritical accident for the VANTAGE SH transition is analyzed using the rerate conditions of full power of 3579 MWt and a positive moderator temperature coefficient of +6.0 perr/*F under BOL conditions. The analysis is performed in three stages: first, an average core nuclear power transient calculation, then an average core heat transfer. calculation, and finally, the DNBR calculation. In orcer to give conservative results for the accident, the following assumptions are made:
- 1. Since the magnitude of the power peak reached during the initial part of the transient for any given rate of reactivity insertion is strongly dependent on the Doppler coefficient, conservatively low values are used. The Doppler defect used as an initial condition is 900 pcm.
- 2. The analysis employs a moderator coefficient which was46.0 pcm/*F at the zero power nominal temperature. Contribution of the moderator reactivity coefficient is negligible during the initial part of the transient because the heat transfer time between the fuel and the moderator is much longer than the neutron flux response time. However, after the initial neutron flux peak, the succeeding rate of power increase is affected by the moderator reactivity coefficient. A conservative value is used in the analysis to yield the maximum peak heat flux.
- 3. 557'F). This The reactor assumption is assumed is more conservative tothan be that at hot of a zero lower power in (Tavg =itial system temp The higher initial system temperature yields a large fuel. water heat transfer coefficient, larger specific heats and a less negative Doppler coefficient, all of which tend to reduce Doppler feedback effect thereby increasing the nuclear flux peak.
- 4. Reactor trip is assumed to be initiated by power range high neutron flux (low setting). The most adverse combination of instrument and setpoint errors result in a setpoint of 35%. In addition, the reactor trip insertion characteristic is based 5139 m d
on the assumption that the highest worth rod cluster control assembly is stuck in its fully withdrawn position.
- 5. The maximum positive reactivity insertion rate assumed is greater than that for the simultaneous withdrawal of the combination of the two sequential control banks having the greatest combined worth at maximum speed (45 inches / minute).
- 6. The most limiting axial and radial power shapes, associated with having the two highest combined worth banks in the high worth position, are assumed in the DNB analysis.
- 7. The initial power level was assumed to be below the power level expected for any shutdown condition (10-9 of nominal power). This combination of highest reactivity insertion rate and lowest initial power produces the highest peak heat flux.
- 8. Two reactor coolant pumps are assumed to be in operation. This is conservative with respect to DNB.
5.4.1.3 Results Figures 5.4.1-1 through 5.4.1-4 show the transient behavior for the uncontrolled RCCA bank withdrawalincident, with the accident terminated by reactor trip at 35% of nominal . power. The reactivity insertion rate used is greater than that calculated for the two highest worth sequential control banks, both assumed to be in their highest incremental worth region. Figure 5.4.1-1 shows the average nuclear power transient. The energy release and fuel temperature increases are relatively small. The thermal flux response, of consideration for DNB considerations, is shown on Figure 5.4.1-2. The beneficial effect of the inherent thermal lag in the fuel is shown by a peak heat flux much less than the full power nominal value. There is margin to DNB during the transient since the rod surface heat flux remains below the design value, and there is a high degree of subcooling at all times in the core. Figures 5.4.1-3 and 5.4.1-4 shows the response of the hoi spot average fuel and cladding temperature. The hot spot average fuel temperature increases to a value lower than the nominal full power value. The minimum DNBR at all times remains above the safety analysis limit values. The calculated sequence of events for this accident is shown in Table 5.4,4-1. 5-140 1
6 5.4.1.4 Conclusions in the event of a RCCA withdrawal accident from the subcitical condition, the core and the RCS are not adversely affected, since the combination of thermal power and the coolant temperature result in a DNBR greater than the safety limit values.' Thus, no fuel or clad damage is predicted as a result of DNB. TABLE 5.4.1-1 Time Sequence of Events For incidents Which Result in lleactivity and Power Distribution Anomatios Event Time ! (sec) Initiation of uncontrolled 0.0 rod withdrawal from 10 9 of nominal power Power range high neutron 10.37 flux low setpoint reached Peak nuclear power occurs 10.52 Rods begins to fall into 10.87 core Minimum DNBR occurs 12.6 Peak average clad temperature' 12.84 occurs Peak heat flux occurs 12.6 Peak average fuel temperature 13.19 occurs - f
'5-141 .r.-- r -e s - + -- e-r-- e4 .- m - --e .c g-- -+ v ve -e i
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i
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5-142
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_. ___ __ m ___ _ ._ ___._..._.__~_ _ _ _ _ _ . . . . _ _ _ _ _ . _ _ _ . . . - _ _ _ _ _ - _ _ . _ . _ . . m . ..m__. _ _ _ _ _ . ._ ._ I n 2,500 w ; 8 a 2,000 - i i s 1,500 - W e W 1,000 - d
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6 s I i i i I O O 5 10 15 20 25 30 TIME (SECONDS) - UPDATLD eMNAbbs REPORT Figure 5.4.1-3 Fuel Temperature Trensients for Uncontrolled RCCA bank Withdrowol from o Subtritical Coneition 1 5-144
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5.4.2 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal at Power SA.2.1 Introduction Uncontrolled rod cluster control assembly (RCCA) bank withdrawal at power results in an increase in the core heat flux. Since the heat extraction from the stearn generator lags behind the core power generation until the steam generator pressure reaches the relief or safety valve setpoint, there is a net increase in the reactor coolant temperature. Unless terminated by manual or automatic action, the power mismatch and resultant coolant temperature rise could eventually result in DNB. Therefore, in order to avert damage to the fuel clad, the Reactor Protection Syctem is designed to terminate any such transient before the DNBR falls below the safety analysis limit values. The automatic features of the Reactor Protection System which prevents core damage following the postulated accident include the following:
- 1. Power range neutron flux instrumentation actuates a reactor trip if two-of-four channels exceed an overpower setpoint.
- 2. Reactori h is actuated if any two-of four AT channels exceed an Overtemperature AT setpoint. This setpoint is automatically varied with axial power imbalance, coolant temperature and pressure to protect against DNB.
- 3. Reactor trip is actuated if any two out-of four AT channels exceed an Overpower AT setpoint. This setpoint is automatically varied with axial power imbalance and coolant temperature to ensure that the allowable heat generation rate (kW/ft) is not exceeded.
- 4. A high pressurizer pressure reactor trip actuated from any two-out-of-four pressure channels which is set at a fixed point. This set pressure is less than the set pressure for the pressurizer safety valves. ,
- 5. A high pressurizer water level reactor trip actuated from any two-out-of three level channels when the reactor power is above approximately 10 percent (Permissive P-7).
i l In addition to the above listed reactor trips, there are the following RCCA withdrawal l- blocks: I- 1. High neutron flux (one-out-of-four power range)
- 2. Overpower AT (two-out-of-four)
, 3. Overtemperature I T (two-out-of-four) 5-146
5.4.2.2 Methodology The rerate conditions of uprated power, and a positive moderator temperature coefficient of +6.0 pcm/'F were used in this analysis. This accident is analyzed with the Statistical Core Design methodology. To obtain conservative results, the following assumptions are made:
- 1. Plant characteristics and initial conditions are discussed in Section 5.0.
- 2. Reactivity Coefficients Two cases are analyzed:
- a. Minimum Reactivity Feedback (BOL). A positive moderator temperature coefficient of +6.0 pcm/'F is assumed. A variable Doppler power coefficient with core power is used in the analysis. A conservatively small (in absolute magnitude) value is assumed.
- b. Maximum Reactivity Feedback (EOL). A conservatively large positive moderator density coefficient and a large negative Doppler power coefficient are assumed.
- 3. The reactor trip on high neutron flux is assumed to be actuated at a conservative value of 118% of nominal full power. The aT trips include all adverse instrumentation and setpoint errors; the delays for trip actuation are assumed to be the maximum values.
- 4. T.1e RCCA trip insertion daracteristic is based on the assumption that the highest worth assembly is stuck in its fully withdrawn position.
- 5. A range of reactivity insertion rates is examined. The maximum positive reactivity insertion rate is greater than that for the simultaneous withdrawal of the two control banks having the maximum combined worth at maximum speed.
5.4.2.3 Results Only the most DNB limiting case is presented here. This corresponds to a slow RCCA withdrawal rate (2 pcm/sec) at full power conditions with minimum (BOL) reactivity feedback conditions. Figures 5.4.2-1 through 5.4.2-6 show the transient response for a l slow RCCA withdrawal incident starting from full power. Reactor trip on high neutron flux occurs shortly after the start of the accident. Reactor trip on Overtemperature AT l occurs after 45.5 seconds and the rise in temperature and pressure is consequently - 1 larger than for rapid RCCA withdrawal. The minimum DNBP 's greater than the safety analysis limit values throughout the transient. l 5-147
,-y** - m -- -'
was- w
Figure 5.4.2 7 shows the minimum DNBR as a function of reactivity insertion rate from initial full power operation for minimum and maximum reactivity feedback. It can be coen that two reactor trip channels provide protection over the whole range of reactivity insertion rates. These are the high neutron flux and Overtemperature AT channels. The minimum DNBR is never less than the safety analysis limit values. Figure 5.4.2-7 and 5.4.2 8 show the minimum DNBR as a function of reactivity insertion rate for RCCA withdrawalincidents starting at 60 and 10 percent power, respectively, for minimum and maximum reactivity feedback. The results are similar to the 100% power case, except as the initial power is decreased, the range over which the Overtemperature AT is effective is increased. In neither caso does the DNBR fall below the safety ana!ysis limit values. The shape of the curves of minimum DNB ratio versus reactivity insertion rate in the reference figures la due both to reactor core and coolant system transient response and to protection system action in initiating a reactor trip. Referring to Figure 5.4.2-7, for example, it is noted that:
- 1. For reactivity insertion rates above 2 pcm/sec, reactor trip is initiated by the high neutron flux trip for the minimum reactivity foodbach cases. The neutron flux in the core rises rapidly for these insertion rates while core heat flux and coolant system temperature lag behind due to the thermal capacity of the fuel and coolant system fluid. Thus the reactor is tripped prior to significant increase in heat flux or water temperature with resultant high minimum DNB ratios during the transient. As reactivity insertion rate decreases, core heat flux and coolant temperatures can remain more nearly in equilibrium with the neutron flux.
Minimum DNBR during the transient thus decreases with decreasing insertion rates.
- 2. The Overtemperature aT reactor trip circuit initiates a reactor trip when measured coolant loop AT exceeds a setpoint based on measured Reactor l Coolant system average temperature and pressure. It is important to note that the average temperature contribution to the circuit is lead-lag compensated in l
order to decrease the effect of the thermal capacity of the Reactor Coolant System in response to power increases. - For reactivity insertion rate below 2 pcm/sec, the Overtemperature AT trip terminates the transient. l Figures 5.4.2-7,5.4 2-8 and 5.4.2 9 illustrate minimum DNBRs calculated for BOL and EOL reactivity feedbacks. Since the RCCA withdrawal at power incident is an overpower transient, the fuel temperatures rise during tif 'ransient until after reactor trip occurs. For high reactivity insertion rates, the overpower transient is fast with respect to the fuel rod thermal time - 5-148
i i i constant, and the core heat flux lags behind the neutron flux response. Due to this lag, the peak core heat flux doe not exceed 118 percent of its nominal vrlue (i.e., the high neutron flux trip setpoint assumed in the analysis). Taking into account the effect of the RCCA withdrawal on the axial core power distribution, the peak fuel centerline , . temperature will still romain below the fuel rnetting temperature. i For slow reactivity insertion ratos, the core heat flux remains more nearly in equilibrium with the neutron flux. The overpower transient is terminated by the Overtemperature AT reactor trip before a DNB condition is reached. The peak heat flux again is maintained below 118 percent of its nominal value. Taking into account the effect of the RCCA withdrawal on the axial core power distribution, the peak fuel centerline < 4 temperature will remain below the fuel molting temperature. Since DNB does not occur at any time during the RCCA withdrawal at power transient, , the ability of the primary coolant to remove heat from the fuel rod is not reduced. Thu3, ' the fuel cladding temperature does not rise significantly above its initial value during the transient. The calculated sequence of events for this accident is shown in Table 5.4.2-1. 5.4.2.4 Conclusions , The high noutron flux and Overtemperaturo AT trip channels provide adequato protection over the entile range of possible reactivity insertion ratos, i.e., the minimum value of DNBR is always larger than the safety analysis limit values. Table 5.4.21 Time Sequence of Events Initiation of uncontrolled RCCA 0.0 see withdrawal at a small reactivity insertion rate (2 pcm/sec) Overtemperature AT reactor trip 45.57 sec signalinitiated Maximum Pressurizer Pressure 88.00 see Maximum Pressurizer Lovel 48.5 sec - 5149 4
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0 O 10 20 30 40 $0 60 TlVE (SECONDS) WOLF CREEK Figure $.4.2-2 Uncontrolled RCCA bonk Withorowel from full power with BOL minimum (2 ricm(/see) withdromol rate) reactivity feedbo 5-151
i i 2,400 ! 2 g ,350 l R l 1 2,300 2,250 l 2.200 - r 2,150 - i
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2,100 g 0 20 30 40 $0 60 i tut (stCON05) WOLF CREEK figure 5.4.2-3 - Uncontrolled RCCA bonk Withdrawol from full power with ; BOL minimum (2 Dem (/see) withdrowol rote) reactivity feedba 5152 1
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Wolf Creek - figure 5,4,2-6 Uncentrolled RCCA Donk withdrawal From Full Power with Beginning-of-Life ReactMty Feedback (2 pcm/sec Withdrowol Rote) t 5-155
- m. + w ,, -- i.%,_ _- --
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gg f I i. I i i l i 1 2 J 6 10 20 30 SO .Jo RoottWity inset'e n note (pien/sec) Wolf Creek Figure $.4.2-8 u Minimum DNGR Versus Recrtivity insertion Rote, Rod Withdrowol }' from 60 Percent Power - 5-157
4 s.,,aa m - w . (?tH!* High Neutron 3s - Flux Trip s v 'N High Pressure Trip 13 -
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1 2 3 b 10 20 30 50 100 Reactkity lesert,on Rote (ptm/s.c) Wolf Creek figure 5.4.2-9 Minimum DNBR Versus Reactivity insertion Rote, Rod Withdrowol from 10 Percent Power i i 5-158
5.4.3 Rcd Cluster Control Assembly Misoperation 5.4.3.1 Introduction Rod cluster control assembly (RCCA) misoperation accidents include:
- a. One or more dropped RCCAs within the same group
- b. A dropped RCCA bank
- c. Statically misaligned RCCA
- d. Withdrawal of a single RCCA Each RCCA has a position indicator chcnnel which displays the position of the assembly. The displays of assen.Lly positions aie grouped for the operator's convenience. Fully inserted assemblies are further indicated by a rod at bottom signal, which actuates a local alarm and a control room annunciator. Group demand position is also indicated.
RCCAs are always moved in preselected banks, and the banks are always mov6d in the sa'io preselected sequence. Each bank of RCCAs is divided into groups of four or five. The RCCAs comprising a group operate in parallel through multiplexing thyristors. The groups in a bank move sequentially such that the first group is always within one step of the second group in the bank. A definite schedule of actuation (or deactuation of the stationary gripper, movable gripper, and lift coils of a mechanism) is required to withdraw the PsCCA attached to the mechanism. Since the stationary gripper, movable gripper. and lift coils associated with the four RCCAs of a rod group are driven in parallel, any single failure which would cause rod withdrawal would affect a minimum of one group. Mechanical failures are in the direction of insertion or immobility. No single electrical or mechanical failure in the rod control system could cause the accidental withdrawal of a single RCCA from the inserted bank at full power operation. The operator could withdraw a single RCCA in the control twank since this feature is necessary in order to retrieve an assembly should one be accidentally dropped. The event analyzed must result from multiple wiring failures or multiple significant operator errors and subsequent and repeated operator disregard of event inoication. The probability of such a combination of cond;tions is considered low. The limiting consequences of such errors or failures may include slight fuel damage. A dropped RCCA or RCCA bank is catected by:
- a. Sudden drop in the core power level as seen by the nuclear instrumentation system 5-159
)
I .m.__._-_.__.-__.. _ _ , _ _ ..___m.-___.m_____ _w__________._____m_ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ - - - _ _ -
- b. Asymmetric power distribution as seen on out-of-core neutron detectors or core exit thermocouples
- c. Rod at bo; tom signal
- d. Rod deviation alarm (control rods only)
- e. . position inc; cation Misaligned RCCAs are detected by;
- a. Asymmetric power distribution as scen on cut-of-core neutron detectors or core exit thermocouples
- b. Rod deviation alarm (control rods only)
- c. Rod position indication The resolution of the rod position indicator channel is 112 steps ( 7.5 inches).
Deviation of any RCCA from its group by twice this distance (124 steps or 115.0 inches) will not c:Use power distributions worse than the desig, limits. The deviation alarm (control rods enly) alerts the operator to rod deviation with respect to the group position in excess of 12 steps. If the rod deviation alarm is not operable, the operator is required to take action required by the Technical Specifications. If one or more rod position indicator channel should be out of service, detailed operating instructions sht.ll b followed to assure the alignment of the nonindicated-RCCAs. The operator is also required to take action, as required by the Technical Specifications. In the extremely unlikely event of simultaneous electrical failures which would result in - single RCCA withdrawal, rod deviation and rod control urgent failure would both be - displayed on the plant annunciators, and the rod position indicators would indicate the relative positions of the assemblies in the bank.- The urgent failure alarm also inhibits automatic rod motion in the group in which it occurs. Withdrawal of a single RCCA by-l operator action, whether deliberate or by a combination of errors, would result in activation of the same alsrm and the same visual indications. Withdrawal of a single RCCA results in both positive reactivity insertion tending to increase core power and an increase in local oower density in the core area associated with the RCCA. Automatic protection for this event is provided by the overtemperature DELTA-T reactor trip, L although due to the increase in local power density it is not possible in all cases to L provide assurance that the core safety limits will not be violated. I I ! 5-160
5.4.3.2 Analysis of Effects and Consequences 5 4.3.2.1 Dropped RCCAs, Dropped RCCA Bank, and Statically Misaligned RCCA Methodoloav a,b. One or more dropped RCCAs from the same group or bank: For evaluation of the dropped RCCA event, forty-seven different combinations of dropped rods (patterns) are considered. These patterns include one, two, or three rods from a single group, an entire group, or an entire bank. The initial conditions prior to rod drop are assumed to be; full power, RCCA control system on automatic, D Bank inserted to its full-power limit, and all other rods full out. Nodal physics models (Reference 5) re used (full core) to calculate rod worths and power distributions. Simulations are made for each pattern with the dropped rods full-in and all other rods full out; an all-rods-out (ARO) case is also run. The calculations are performed at beginning of cycle (BOC) and at end of cycle (EOC). Output three dimensional power distributions are used to determine full-core minimum margin to DNBR limits for each pattern. The loss of margin from the ARO condition to the pattern-in condition is determined. This margin loss is compared to the available margin to demonstrate that the DNB limits are not violated.
- c. Statically misaligned RCCA Steady-state power distributions are analyzed for the case of static misalignment of a single RCCA. The methodology and codes used are the same as for the analysis of the dropped RCCA events. Full core power distributions are utilized to determine margin loss due to the misaligned RCCA. The loss is compared to available DNBR margin to verify that the DNBR design limit is not violated.
Results a,b. One or more dropped RCCAs from the same group or bank: i ! 5-161 l l
Those patterns which cause reactor trip are determined. Patterns with worths greater than 400 pcm are assumed to cause a negative flux rat' trip. This would occur within approximately 2.5 seconds following the drop of the RCCAs. The core is not adversely affected during this period, since power is decreasing rapidly. Following reactor trip, normal shutdown procec'ures are followed. The operator may manually retrieve the RCCA by following approved operating procedures. For those dropped RCCAs which do not result in a reactor trip, power may be re-established either by reactivity feedback or control bank withdrawal. Following a dropped rod event in manual rod control, the plant will establish a new equilibrium condition. The equilibrium process without control system interaction is monotonic, thus removing power overshoot as a concern, and establishing the automatic rod control mode of operation as the limiting case. For a dropped RCCA event in the automatic rod control mode, the rod control system detects the drop in power and initiates control bank withdrawal. Power overshoot may occur due to this action by the automatic rod controller after which the control system will insert the control bank to restore nominal power. Maximum power level (MPL), which occurs after pattern has been dropped and the control bank (D Bank in the analysis) has been fully withdrawn, is calculated for each pattern. Patterns with power overshoot (MPL) greater than 118% cause an overpower trip. The remaining patterns which do not result in reactor trip are analyzed to determine that DNBR design limits are not exceeded. For all 47 patterns, at BOC and at EOC, the available margin exceeds the margin loss due to the dropped rods; thus, for all cases the minimum DNBR remains above the limit value.
- c. Statically misaligned RCCA The most severe misalignment situations with respect to DNBR at significant power levels arise from cases in which one RCCA is fully inserted or where bank D is fully inserted with one RCCA fully withdrawn.
Multiple independent alarms, including a bank insertion limit alarm, alert the opert well before the postulated conditions are approached. The bank car .nserted to its insertion limit with any one assembly fully withdrawri wunout the DNBR falling below the limit value. The insertion limits in the Technical Specifications may vary from time to time, depending on a number of limiting criteria, it is preferable, therefore, to analyze the misaligned RCCA case at full power for a position of the control bank as deeply 5-162
. - . . . . - . - - - .-.- - -. - . - . . . . -- - - ~.- -
i Ij inserted as the criteria on minimum DNBR and power peaking factor will allow. The full power insertion limits on control bank D must then by chosen to be above that position and will usually be dictated by other criteria. Detailed results will vary from cycle to cycle, depending on fuel arrangements. For this RCCA misalignment with bank D inserted to its full power insertion limit and one RCCA fully withdrawn, DNBR does not fall below the limit value. This case is analyzed assuming the initial reactor power, pressure, and RCS temperature are at. their nominal values including uncertainties but with the increased radial peaking factor associated with the misaligned RCCA. DNB calculations have not been performed specifically for RCCAs missing from other banks; however, power shape calculations have been done as required for the RCCA ejection analysis. Inspection of the oower shapes shows that the DNB anc ,. Jak' kW/ft situation is less severe than the bank D case discussed above, assuming insertion limits on the other banks equivalent to a bank D full-in insedion litrit. For RCCA misalignments with one RCCA fully inserted, the DNBR does not fall below the limit value. This case is analyzed assuming that the initial reactor power, pressure, and RCS temperatures are at their nominal values, including uncertainties, but with the increased radial peaking factor associated with the misaligned RCCA. DNB does not occur for the RCCA misalignment incident, and thus, the ability of the primary coolant to remove heat from the fuel is not reduced. The peak fuel temperature corresponds to a linear heat generation rate based on the radial peaking factor penalty . associated with the misaligned RCCA and the design axial power distribution, The resulting linear heat generation is well below that which would cause fuel melting. Following the identification of an RCCA group misalignment condition by the operator, the operator is required to take action as required by the plant Technical Specifications - and operating instructions.- 5.4.3.2.2 Single RCCA Withdrawal Methodolooy Power distributions within the core are calculated using the computer codes as described in Reference 5. ~ The peaking factor.$ are then evaluated to calculate the number of pins failing for the event. The case of the worst rod withdrawn from bank D inserted at the rod insertion limit, with the reactor at full power, was analyzed.- This incident is assumed to occur at beginning-of-life to maxi _mize the pin peaking. Additionally, it is assumed that any pin which reaches the design radial peaking limit enters DNB and fails. - l L L 5-163 i
= . ,
Results For the single rod withdrawal event, two cases have been considered as follows:
- a. If the reactor is in the manual control mode, continuous withdrawal of a single RCCA results in both an increase in core power and coolant temperature and an increase in the local hot channel factor in the area of the withdrawing RCCA. In terms of the overall system response, this caso is similar to uncontrolled RCCA bank withdrawal, however, the increased local power peaking in the area of the withdrawn RCCA results in lower minimum DNBRs than for the withdrawn bank cases, Depending on initial bank insertion and location of the withdrawn RCCA, automatic reactor trip may not occur sufficiently fast to prevent the minimum core DNBR from >
falling below the limit value. Evaluation of this case shows that an upper limit for the number of rods with a DNBR less than the limit value is 5 percent,
- b. If the reactor is in the automatic control mode, the multiple failures that result in the withdrawal of a single RCCA will result in the immobility of the other RCCAs in the controlling bank. The transient will then proceed in the same manner as Case a described above, The cases above show that a reactor trip will, in all cases, ensure that the number of rods violating the DNBR lim.it value is no more than 5 percent. Jollowing reactor trip, normal shutdown procedures are followed.
5.4.3.3 Conclusions For cases of dropped RCCAs os dropped banks for which the reactor is tripped by the power range negative neutron flux rate trip, there is no reduction in the margin to core thermal limits, and consequently, the DNB design basis is met. For all cases that do not result in reactor trip, it is shown that the DNBR remains greater thcn the limit value and, therefore, the DNB design is met. For all cases of any RCCA fully inserted, or bank D inserted to its rod insertion limits with any single RCCA in that bank fully withdrawn (static misalignment), the DNBR remains greater than the limit value. . For the case of the accidental withdrawal of a single RCCA, with the reactor in the automatic or manual control mode and initially operating at full power with bank D at the insertion limit, an upper bound of the number of fuel rods experiencing DNB is 5 percent of the total fuel rods in the core. 5-164 1 1
5.4.4 Startup of an inactive Reactor Coolant Pump at an incorrect Temperature 5.4.4.1 Introduction If the plant is operating with one pump out of service, there is reverse flow through the inactive loop due to the pressure difference across the reactor vessel. The cold leg temperature in an inactive loop is identical to the cold leg temperature of the active loops (the resctor core inlet temperature). If the reactor is operated at power, and assuming the secondary side of the steam generator in the inactive loop is not isolated, there is a temperature drop across the steam generator in the inactive loop and, with the reverse flow, the hot leg temperature of the inactive loop is lower than the reactor core inlet temperature. Administrative procedures require that the plant be brought to a load of less than 10 percent of full power prior to starting the pump in an inactive loop in order to bring the inactive loop hot leg temperature closer to the core inlet temperature. Starting of an idle reactor coolant pump without bringing the inactive loop hot leg temperature close to the core inlet temperature would result in the injection of cold water into the core, which would cause a reactivity insertion and subsequent power increase. Should the startup of an inactive coolant pump at an incorrect temperature occur, the . I transient will be terminated by a reactor trip on low coolant loop when the power range neuron flux (two out of four channels) exceeds the P-8 setpoint, which has been previously reset for three loop operation. 5.4.4.2 Methodology Rerate conditions of upgraded power were assumed in the analysis. In order to obtain conservative results for the startup of an inactive pump accident, the following assumptions are made:
- 1. Plant characteristics and initial conditions are discussed in Section 5.0.
- 2. - Following initiation of startup of the idle pump, the inactive loop flow reverses and accelerates to its nominal flow value in approximately 30 seconds. This value is faster than the expected startup time, and is conservative for this analysis.
- 3. A conservatively large moderator density coefficient.
- 4. A conservatively small (absolute value) negative Doppler only power coefficient.
- 5. The initial reactor coolar.t loop flows are at tne appropriate values for one pump out of service.
5-165
.e
.. - - - ~ - -- - -. -- - - - - . . . -
- 6. The reactor trip is assumed to occur on low coolant flow when the power range neutron flux exceeds to P-8 setpoint. The P-8 setpoint is conservatively assumed to be 84 percent of rated power which corresponds to the nominal setpoint plus 9 percent for nuclear instrumentation errors.
5.4.4.3 Results The results following the startup of an idle pump with the abo ~.e assumptions are shown in Figures 5.4.4-1 through 5.4.4-4. As shown in these curves, during the first part of the transient, the increase in core flow with cooler water results in an increase in nuclear power and a decrease on core average temperature. The minimum DNBR during the transient is considerably greater than the safety limit values. Reactivity addition for the inactive loop startup accident is due to the decrease in core water temperature. During the transient, this decrease is due both to the increase in reactor coolant flow and, as the inactive loop flow reverses, to the colder water entering the core from the hot leg side (colder temperature side prior to the start of the transient) of the steam generator in the inactive loop. Thus, the reactivity insertion rate for this transient changes with time. The resultant core nuclear power transient, computed with consideration of both moderator and Doppler reactivity feedback effects, is shown on Figure 5.4.4-1. The calculated sequer ce of events for this accident is shown in Tabie 5.4.4-1. The transient results illuetrated in figures 5.4.4-1 through 5.4.4-4 indicate that a stabilized plant condition, with the reactor tripped, is approached rapidly. 5.4.4.4 Conclusions The transient results show that the core is not adversely afiected. There is considerable margin to the safety analysis DNBR imit values; thus, ,1o fuel or clad damage is predicted. Table 5.4.4-1 Event Time (Sec) Initiation of pump startup 0.0 Power reaches high neutron flux 12.4 P-8 trip setpoint Rods begin to drop 12.9 Minimum DNBR occurs 12.9 5-16E
5 1 50 .8 - O 6 <
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I 6 6 50.4 - B c. m 0 0.2 - da z - 1 ' 0 0 10 20 30 4 TIME (SECONOS, WOLF CREEK Figure 5.4.4- 1 Nuclear power transient for Startup of on inoctive Reactor Coolont toop 5-167
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Figure 5.4.4-2 Core Flow Transient for Startup of on inactive Reactor Coolont Loop. pilu ums 5-168
2.400 i m 2.300 - v w 2,200 - 5 8 2,100 - r d_ _
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i ' 1,800 0 to 20 30 40 TIME (SECONOST WOLF CREEK Figure S.4.4 - 3 Pressurizer Pressure transient for Stortup of on (noctive Reactor Coolant Loop I i 5-169
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520 0 10 20 3G 40 TIME (SECONDS) WOLF CREEK . 1
,igura F 5.4.4-4 Core Average Temperature TrorisleM for Stortup of on inoctive Reactor Coolant Loop D
J 5-170-m
5A.5 A Malfunction or Failure of a Flow Controller in a BWR This Event is not applicable to WCGS. 5.4.6 Chemical and Volume Centrol System Malfunction that Results in a Decrease in the Boron Concentration in the Reactor Coolant The key safety analysis parameters for modes 1-5 of the Boron Dilution Event (BDE) analysis for Cycle 7, particularly the time to flux doubling and/or the time to loss of l shutdown margin, have been revised to reflect the increate of boron concentration in the RCS as a result of implementing design changes to upgrade the core power level to 3565 MWt and to allow a positive moderator temperature coefficient (PMTC).- Additional analysis margin has been allocated in order to account for the identified analysis non-conservatisms associated with the flux doubling Boron Dilution Mitigating System (BDMS). Based upon the results of calculations described in the following sections, there is reasonable assurance that there would be sufficient time for either automatic flux doubling BDMS or operator intervention to terminate the dilution event before shutdown margin is lost should an unplanned boron dilution event occur during Mod as 1-5 of reactor operation. 5.4.6.1 Identification of Causes and Accident Description One of the two principal means of positive reactivity insertion to the core is the addition of unborated, primary grade water from the demineralized Reactor Makeup Water System (RMWS)into the RCS through the reactor makeup portion of the Chemical and Volume Control System (CVCS). Boron dilution with these systems is a manually initiated operation under strict administrative controls requiring close operator surveillance with procedures limiting the rate and duration of the dilution. A boric acid blend system is available to allow the operator to match the makeup's boron concentration to that of the RCS during normal charging. The means of causing an inadvertent boron dilution are the opening of the primary water makeup control valve and failure of the blend system, either by controller or mechanical failure. The CVCS and RMWS are designed to limit, even under various postulated failure modes, the potential rate of d.. Jon to values which, with indication by alarms and instrumentation, will allow sufficient time for automatic or operator response (depending on the mode of operation) to terminate the dilution. An 5-171
inadvertent dilution from the RMWS may be terminated by closing the primary water makeup control valve All expected sources of dilution may be terminated by closing isolation valves in the CVCS, BG LCV-112B and C. The lost shutdown rnargin (SDM) may be regained by the opening of isolation valves to the RWST, BN-l.CV-112D and E, thus allowing the addition of borated water to the RCS. Generally, to dilute, the operator must perform two distinct actions:
- a. Switch control of the makeup from the automatic makeup mode to the dilute mode
- b. Depress the start button.
Failure to carry out either of the above actions prevents initiation of dilution. Also, during normal operation the operator may add borated water to the RCS by blending boric acid from the boric acid storage tanks with primary grade water. This requires the operator to determine the concantration of the addition and to set the blended flow rate and the boric acid flow rate. The makeup controller will then limit the sum of the boric acid flow rate and primary grade water flow rate to the blended flow rate, i.e., the controller determines the primary grade water flow rate after the start button is depressed. The status of the RCS makeup is continucusly available to the operator by:
- a. Indication of the boric acid and blended flow rates
- b. CVCS and RMWS pump status lights
- c. Deviation alarms, if the boric acid or blended flow rates deviate by more than 10 percent from the preset values
- d. Source range neutron flux - when reactor is suberitical
- 1. High flux at shutdown alarm
- 2. Indicated source range neutron flux count rates -
- 3. Audible source range neutron flux count rate
- 4. Source range neutron flux - doubling alarm.
- e. When the reactor is critical
- 1. Axial flux difference alarm (reactor power > 50 percent RPT)
- 2. Control rod insertion lirnit low and low-low alarms
- 3. Overtemperature Delta T alarm (at power)
- 4. Overtemperature Delta T turbine runback (at power)
- 5. Overtemperature Delta T reactor trip
- 6. Power range neutron flux - high, both high and low setpoint reactor trips.
5-172
l l i 5.4.6.2 Analysis of Effects and Consequences To cover all phases of plant operation, boron dilution during Refueling, Cold Shutdown, Hot Shutdown, Hot Standby, Srtup, and Power modes of operation is considered in this analysis. Conservative analysis assumptions for each Technical Specification defined mode of operation were used, i.e., high RCS critical boron concentrations, high boron worths, minimum shutdown margins, and lower than actual RCS volumes. These assumptions result in conservative determinations of the time available for operator or system response after detection cf a dilution transient in progress. Dilution Durino Refuelino An uncontrolled boron dilution trancient cannot occur during this mode of operation, inadvertent dilution is prevented by administrative controls which isolate the RCS from the potential source of unborated water. Valves BG-V-178 and BG-V-601 (or BG V-602) in the CVCS will be locked closed during refueling operations. These valves block all flow paths that could allow unborated makeup water to reach the RCS Any makeup which is requirod during refueling will be borated water supplied from t..a RWST by the RHR pumps. Dilution Durino Cold Shutdown The following conditionn are assumed for inadvertent boron dilution while in this cperating mode:
- 1. Dilution flow is limited by a flow orifice in the RMWS to 150 gpm of unborated-water.
- 2. An RCS water volume of 3400 ft.3. This is a conservative estimate of the minimum active volume of the RCS and corresponds to the water level drained to mid-nozzle in the vessel while on one train cf RHR.-
- 3. The initial boron concentration required to meet the Technical Specification shutdown margin limit of 1.3% ak/k is conservatively calculated to be 1905 ppm.
This corresponds to a critical boron concentration of 1759 ppm, assuming a conservative, constant boron worth'of -8.9 pcm/ ppm. In the event of an inadvertent boron dilution transient while in this mode of operation, the source range nuclear instrumentation will detect a doubling of the neutron flux by comparison of thi murrent source range flux to that of approximately 10 minutes earlier. Upon detection of the flux doubling, an alarm is sounded for the operator, and valve movement to terminate the dilution and start boration is automatically initiated. Under the conditions defined above, these actions will occur approximately 6.4 minutes after start of dilution. Valves BN-LCV-112D and E (isolation valves to the RWST) are opened to supply borated water to the suction of the charging pumps, and valves BG-5-173
LCV-1128 and C (isolation valves in the CVCS) are closed to terminate the dilution. These automatic actions are carried out to minimize the approach to criticality and ' regain the lost shutdown margin. Action taken by the operator is to terminate boration after regaining the required shutdown margin and determine and correct the cause of - the dilution transient. l Dilution Durino Hot Shutdown Tne following conditions are assumed for an inadvertent boron dilution while in this mode:
- 1. The dilution flow r;" is limited by piping system friction losses and the capacity of two makeup water pumps to supply 260 gpm of unborated water.
- 2. An RCS water volume of 4780 ft3 . This is a conservative estimate of the minimum active volume of the RCS, while on one train of RHR and with the RCS filled and vented. The minimum active volume assumes no mixing with the reactor vessel's upper head volume.
- 3. The initial boron concentration required to meet the Technical Specification shutdown margin limit of 1.3% ak/k is conservatively calculated to be 1893 ppm.
This corresponds to a critical boron concentration of 1743 ppm, assuming a cunservative, constant boron worth of - 8.7 pcm/ ppm. In the event of an inadvertent boron dilution while in this mode of opnration, the source range nuclear instrumentation will detect a doubling of the neutron flux, automatically initiate valve movement to begin boration and terminate the dilution, and sound an , alarm for the operator. Under the conditions defined above, these actions will occur approximately 5.2 minutes after start of dilution. No operator action is required to -- terminate this transient. Dilution Durino Hot Standby The follov,ing conditions are assumed for an in6dvertent boron dilution while in this mode:
- 1. The dilution flow is limited to 260 gpm of unborated water (as in the previous !
9 case). l 2. The RCS volume is 5800 ft3 . This a conservative estimate of the minimum ' active volume of the RCS with the RCS filled and vented and one RCP running.
- 3. The initial boron concentration required to meet the Technical Specification shutdown margin limit of 1.3% ok/k is conservatively calculated to be 2160 ppm.
l l l 5-174 i-t 1
This corresponds to a critical boron concentration of 1991 ppm, assuming a conservative, constant boron worth of - 7.7 pcm/ ppm. In the event of an inadvertent boron dilution transient while in ;his mode of operation, the source range nuclear instrumentation will detect a doubling of tl.a neutron flux, autornatically initiate valve movement to begin boration and terminate the dilution, and sound an alarm for the operator. Under the conditions defined above, these actions will occur approximately 4.9 minutes after start of dilution. No operator action is required to terminate this transient. Dilution Durino Start-up In this mode, the plant is being taken from one long-term mode of operation, Hot Standby, to another, Power. The plant is maintaincj in the Start-up mode only for the purpose of start-up testing at the beginning of each cycle. During this mode of operation rod control is in manual. All normal actions required to change power level, either up or down, require operator initiation. The Technical Specifications require an available trip reactivity of 1.3% Ak/k and four reactor coolant pumps operating. Other - conditions assumed are: 1 Dilution flow is the maximum capacity of two centrifugal charging pumps with the RCS at 2250 psia (approximately 245 gpm).
- 2. A minimum RCS water volume of 9965 ft3 . This is a very conservative estimate of the acGve RCS volume, minus the pressurizer volume.
For the purpose of this analysis, the critical boron concentration is assumed to be 1853 ppm after reactor trip. This value is conservative and envelops the worst-case, actual critical boron concentration corresponding to hot zero power, all rods out, and no Xenon. Also assumed is a conservative, constant boron worth of - 6.9 pcm/ ppm.
- 4. The necessary shutdown margin, as required by the Technical Specifications, is provided by the control rods after reactor trip. No credit is taken for the excess negative reactivity inserted by the control rods which is above the shutdown margin.
This mode of operation is a transitory operational mode in which the operator intentionally dilutes and withdraws control rods to take the olant critical. During this mode, the plant is in manual control with the operator required to maintain a very high awarenesF -f the plant status. For normal approach to criticality, the operator must manually n.diate a limited dilution and subsequently manually withdrew the control rods, a process that takes several hours. The Technical Specifications require that the operator determine the estimated critical position of the control rods prior to approaching criticality, thus assuring that the reactor does not go critical with the 5-175
~ . - - - .. -.
control rods below the insertion limits. Once critical, the power escalation must be sufficiently slow to allow the operator to manually block the source !ange reactor trip after receiving P-6 from the intermediate range (nominally at 10 5 cps). Too fast a power escalation (due to an unknown dilution) would result in reaching P-6 unexpectedly, leaving insufficient time to manually block the source range reactor trip. Failure to perform this manual action results in a reactor trip and immediate shutdown of the reactor. However, in the event of an unplanned approach to criticality or dilution during power escalation while in the Start-up mode, the plaat status is suc.h that minimal impact will result. The plant will slowly escalate in power to a reactor trip on the power range , neutron flux - high, low setpoint (nominally 25 percent RTP). After reactor trip, there is approximately 15 minutes for operator action prior to return to criticality. The required operator action is the opening of valves BN LCV-112D and E to initiate boration and the closing of valves BG-LCV-112B and C to terminate dilution. Dilution Durina Full Power Operation The plant may be operated at power two ways: automatic Targ/ rod control and under operator control. The Technical Specificatinns require an available trip reactivity of 1.3% Ak/k and four reactor coolant pumps operating. With the plant at power and the RCS at pressure, the dilution rate is limited by the capacity of the centrifugal charging pumps. the analysis is performed assuming two charging pumps are in operation even though normal operat;an is with one pump. Conditions assumed for this mode are:
- 1. Dilution flow from two charging pumps is at the maximum at an RCS pressure of 2250 psia (approximately 245 gpm) when the reactor is in manual control. When in automatic control, the dilution flow is the maximum letdown flow (approximately 120 gpm).
- 2. A minimum RCS water volume of 9965 ft3 . This is a very conservative estimate of the active kCS volume, minus the pressurizer volume.
- 3. For the purposc of this analysis, the critical boron concentration is assumed to be 2041 ppm. This value is conservative and envelops the worst-case, actual critical boron concentration corresponding to hot full power; all rods out, and no Xenon. Also assumed is a conservative, constant boron worth of - 6.9 pcm/ ppm.
- 4. The necessary shutdown margin, as required by the Technical Specifications, is provided by the control rods after reactor. trip. No credit is taken for th9 excess negative reactivity inserted by the control rods which is above the shutdown margin.
With the reactor in manust control and r.a operator action taken to terminate the transient, the power and temperature rise will cause the reactor to reach the 5-176 L-_
Overtemperature Delta T trip setpoint resulting in a reactor trip. After reactor trip there is approximately 21.5 minutes for operator action prior to return to criticality. The required operator action is the opening of valves BN-LCV-112E and E and the closing of valves BG-LCV-1128 and C. The boron dilution transient in this case is essentially the equivalent to an uncontrolled rod withdrawal at power. The maximum reactivity insertion rate for a boron dilution transient is conservatively estimated to be 1.1 pcm/sec and is within the range of insertion rates analyzed for uncontrolled rod withdrawal at power. It should be noted that prior to reaching the Overtemperature Delta T reactor trip, the operator will have received an alarm on Overtemperature Delta T and an Overtemperature Delta T turbine runback. With the reactor in automatic rod control, the pressurizer level controller will limit the dilution flow rate to the maximum letdown rate, app.oximately 120 gpm. If a dilution rate in excess of the letdown rate is present, the pressurizer level controller will throttle charging flow down to match the letdown rate. Thus, with the reactor in automatic rod control, a boron dilution will result in a power and temperature increase such that the rod controller will attempt to compensate by slow insertion of the control rods. This action by the controller will result in at least three alarms to the operator;
- a. Rod insertion limit - low level alarm '
- b. Rod insertion limit - low-low level alarm if insertion continued after item a
- c. Axial flux difference alarm (Delta I outside of the target band).
Given the many alarms, indications, and the inherent slow process of dilution at power, the operator has sufficient time for action. For example, the operator has at least 30 minutes from the rod insertion limit low-low alarm until 1.3% ak/k is inserted at beginning-of-life. The time would be significantly longer at end-of-life, due to the low initial baron concentration, when shutdown margin is a concern. The above results demonstrate that in all modes of operation, an inadvertent boron dilution is precluded or responded to by automatic functions, or sufficient time is available for operator action to terminate the transient. Following termination of the dilution flow and initiation of boration, the reactor is in a stable condition with the operator regaining the required shutdown margin. 5.4.6.3 Conclusions , The results presented previously show that inadvertent boron dilution events are I precluded during refueling and automatically terminated during cold shutdown, hot shutdown, and hot standby modes. Inadvertent baron dilution events during start-up or power operation, if not detected and terminated by the operators, wi!I result in reactor l 5-177 l
l trip. Following reactor trip, there is ample time available lor the operators to terminate t"t dilution prior to a return to criticality, 5-178-
5.4.7 Inadvertent Loading and Operation of a Fuel Assembly in an improper Position The conclusions reached in Section 15.4.7 of Referenco 1 remain applicable to this event for the changes discussed in Section 1.0. t I t 5-179
5.4.8 Spectrum of Rod Cluster Control Assembly Ejection Accidents 5.4.8.1 Introduction . This accident is defined as the mec hanical failure of a control rod mechanism pressure housing, resulting in the ejection of an RCA and drive shaft. The consequence of this mechanical failure is a rapid positive reactivity insertion and system depressurization together with an adverse core power distribution, poss'bly. leading to localized fuel rod damage. The acceptance criteria are:
- 1. Average fuel-pellet enthalpy at the hot spot below 225 cal /gm for unirradiated fuel and 200 callgm for irradiated fuel; 2 Average clad temperature at the hot spot below the temperature at which clad embrittlement may be expected (2700*F);
- 3. Peak reactor ecolant pressure less than that which could cause stresses to -
exceed the faulted condition stress limits;
- 4. Fuel melting will be limited to less than ten percent of the fuel volume at the hot spot even if the average fuel pellet enthalpy is below the limits of criterion 1 above.
5.4.8.2 Methodolr.,gy - The calculation of the RCCA ejection transient is performed in two stages, first, an average core channel calculation and then, a hot region calculation. The average core calculation is performeo using spatial neutron kinetics metnods to determine the average power generation with time including the various total core feedback effects, i.e., Doppler reactivity and moderator reactivity, Enth yy and temperature transients in the hot spot are then determined by multiplying the average core energy generation by-the hot channel factor and performing a fuel rod transient heat transfer calculation. The power distribution calculated without feedback is possimistically assumed to persist throughout the transient. 5-180 l l l
Calculation of input Parameters Inout parameters for the analysis are conservatively selected on the basis of values calculated for this type of core. The more important parameters are discussed below. Eiected RoJ Worths and Hot Channel Factors The values for ejected rod worths and hot channel factors are calculated using either three-dimensional sta1 methods or by a synthesis method employing one-dimensional and two-dimensional calculations. No credlis taken for the flux flattening effects of reactivity feedback. The calculation is performed for the maximum allowed bank insertion at a given power level, as determined by the rod insertiori timits. Adverse xenon distribution are considered in the calculation. Appropriate margins are added to the ejected rod worth and hot channel factors to account for any calculational uncertainties, including an allowance for nuclear power paaking due to densification. Reactivity Feedback Weichtino Factors The largest temperature rises, and hence the largest reactivity feedbacks occur in channels where the power is higher than average. Since the weight Gia region is dependent on flux, these regions have high weights. This means that the reactivity feedback is larger than that indicated by a simple channel analysis. Physics calculations have been carried out for temperature changes with a flat temperature distribution, with a !arge number of axial and radial temperature distributions. Reactivity changes were compared and effective weighting factors determined. These weighting factors take the form of multipliers which, when applied to single channel feedbacks, correct them to effective whole core feedbacks for the appropriate flux shape, in this analysis, since a one-dimensional (ax;31) spatial kinc ics method is employed. Axial weighting is not necessary if the initial condition is made to match the ejected rod configuration in addition, no weighting is applied to the moderator ' feedback. A conservative radial weighting factor is aoplied to the transient fuel temperature to obtain an effective fuel temperature as a function of time accounting for the missing spatici dimension. these weighting factors have also been shown to be conservative compared to three dimensional analysis . Moderator and Dopoler Coefficient The critical boron concentrations at the beginning-of-life and end-ef4ite are adjusted in the nuclear code in order to obtain moderator density coefficient curves which are conservative compared to actual design conditions for the plant. The Doppler reactivity defect is determined as a function of power level using a one-dimensional . steady-state computer code with a Doppler weighting of 1.0. The Doppler defect used as an initial condition is 900 pcm. 5-181
Delayed Neutron Fraction. B Calculatiori of the effective delayed neutron fraction (perf) typically yield values no less than 0.70 percent at beginning of-life and 0.5 percent at end-of-life for the first cycle. The accident is sensitive to peff of the ejected rod worth is equal to or greater than peff as in zero power transients. In order to allow for future cycles, pessimistic estimates of perf of 0.50 percent at beginning-of-cycle and 0.44 percent at end-of-cycle were used in the analysis. Trio Reactivity insertion The trip reactivity insertion assumes the effect of one stuck RCCA. This is conservative since physics calculations for this plant have shown that the effect of two stuck RCCAs (one of which is the worst ejected rod) is to reduce the shutdown margin by about an additional one percent ak/k. Protection Reactor protection for a rod ejecticn is provided by high neutron flux trip (high and low setting) and high rate of neutron flux increase trip. 5.4.4.3 Results Cases are presented for both beginning and end-of-life at zero and full power.
-1. Beginning-of-cycle, Full power Control bank D assumed to be inserted to its insertion limit the worst ejected rod-worth and hot channel factor were conservatively calculated to be 0.23 percent .
ok/k and 6 6, respectively. The peak hot spot clad average temperature was 2347*F, The peak hot spot fuel center temperature reached melting, conservatively assumed at 4900'F. However, melting was restricted to less than 10 percent of the peWet. , 2. Beginning-of-cycle, Zero Power For this condition, control bank D was assumed to be fully inserted and banks B and C were at their insertic'n limits. The worst ejected rod is located in control. bank D and has a worth. of 0.78 percentk/k and a hot channel factor of 13.0. The peak hot spot clad average temperature reached 2694*F; the fuel center temperature was 4329*F. l 5-182 i T
. _ __ ~ __ ; 3. End-of-Cycle, Full Power -
Control bank D was assumed to be inserted to its insertion limit. The ejected rod worth and hot channel factors were conservatively calculated to be 0.25 percent Ak/k and 7.1, respectively. This resulted in a peak clad average temperature of 2252*F. The peak hot spot fuel temperature reached melting, conservatively assumed at 4800'F, Homer, melting was restricted to less than 10 percent of the pellet.
- 4. End-of-cycle, Zero Power The ejected rod worth and hot channel factor for this case were obtained assuming control bank D to be fully inserted and bank C at its insertion limit.
The results were 0.86 percent Ak/k and 21,0, respectively. The peak clad average and fuel center temperatures were 2523'F and 3610*F, respectively. the Doppler weighting factor for this c . significantly higher than for the other cases due to the very large transient hot channel factor. The nuclear power and hot spot fuel and clad temperature transients for the worst cases are presented in Figures 5.4.8-1 through 5.4.8-4 (Beginning-of-life 1ull power and end-of-life zero power). The calculated sequence of events for the worst case rod ejection accidents as shown in Figures 5.4.8-1 through 5.4.8-4, is presented in Table 5.4.8-1. For all cases, reactor trip occurs very early in the transient, after which the nuclear power excursion is terminated. As discussed previously, the reactor w" ) main suberitical following reactor trip. The ejection of an RCCA constitutes a break in the Reactor Coolant System, located in the reactor pressure vessel head. The effects and consequences of loss-of-coolant-accidents are discussed in Section 6.0. 5.4.8.4 Conclusions Even on a conservative basis, the onalyses indicate that the described fuel and clad limits are not exceeded. It is concluded that there is no danger of sudden fuel dispersal - into the coolant. Since the peak pressure does not exceed that which would cause stresses to exceed the faulted condition stress limits, it is concluded that there is no danger of further consequential damage to the Reactor Coolant system. The analyses have demonstrated that the upper limit in fission product release as a result of a number of fuel rods entering DNB amounts to ten percent. 5-183
Table 5.4.8-1 Time Scquence of Events Accident Event Time (sec) BOL, Zero Power initiation of rod ejection 0.0 Power range high 0.271 neutron flux low setpoint reached Peak nuclear power 0.301 occurs Rods begin to fall into 0.749 core Peak fuel average 2.422 temperature occurs Peak clad temperature 2.23 occurs Peak heat flux occurs 2.26 BOL, Full Power Initiation of rod ejection 0.0 Power range high 0.05 neutron flux high setpoint reached Peak nuclear power 0.135 occurs Rods begin to fallinto 0.55 core Peak clad temperature 2.322 oCCut s Peak heat flux occurs 2.334 Peak fuel temperature 2.919 occurs l 5-184
Table 5.4.8-1 Time Sequence of Events (con't)- EOL, Zero Power Initiation of rod ejection . 0.0 , Power range high 0.169 neutron flux low setpoint reached Peak nuclear power 0.2005 occurs Rods begin to fall into 0.669 core Peak clad temperature 1.337 occurs Peak heat flux occurs 1.34 Peak fuel temperature 1.657 occurs EOL, Full Power Initiation of rod ejection 0.0 Power range high 0.035 neutron flux high setpoint reached Peak nuclear power 0.13 occurs Rod begin.to fallinto 0.54 core Peak clad temperature 2.388 occurs Peak heat flux occurs 2.40 Peak fuel temperature - 2.36 occurs 5-185
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. _-______________w__________________-_w ___ _____ - .__ _ w _ _ _ - __ __._ - .__w .. ___ l I j 5.5 increase in Reacter Coolant Inventory 5.5.1 Inadvertent Operation of the Emergency Core Ccoling System During i Power Operation 5.5.1.1 Introduction The spurious ECCS operation at power could be caused by operator error or a false electrical actuation signal. A spurious signal may originate from any of the wfety injection (SI) actuation channels. An SI signel normally results in a reactor trip followed by a turbine trip. However, it cannot be assumed that any single fault that actuates the SIS will also produw s l reactor trip. If the reactor protection system does not produce an immediate trip as a result of the spurious SIS, the reactor experiences a negative excursion due to the injected boron, cau:.cs a decrease in reactor power. The transient is eventually terminated by the reactor protection system low pressurizer pressure trip or by manual trip. 5.5.1.2 Methodology The spurious ECCS ovent is relatively benign and no plant limits are actually challenged. Operating conditions do not approach the core limits because of the power and temperature reduction. The results have clso been shown to be relatively independent of time to trip. Statistical core design methodology was used in this analysis. 5.5.1.3 Results Figures 5.5.1-1 through 5.5.1-1 show the transient response to inadvertent operation of the ECCS during power operation. Neutron flux starts decreasing immediately due to boron injection, but steam flow does not decrease until later in the transient when the turbine throttle valve goes wide open, the mismatch between load an.d nuciear power causes Tavg, pressurizer water level and pressurizer pressure to drop. When the pressurizer low pressure trip setpoint is reached, the reactor trips and control rods start moving into the core. The DNBR increases throughout the transient. 5.5,1.4 Conclusions Results of the analysis show that spurious ECCS operation without immediate reactor trip presents no hazard to the integrity of the reactor coolant system. 5-190 e,-c. ,-dw- m up 4g= y- > --
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- a. -
i i Table 5.5.1 1 1 Time sequence of Events ,
.I Accident Evcot Time (sec)
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5.5.2 Chemical and Volume Control System Malfunction that increases Reactor { Coolr rit e sentory 5.5.2.1 Introduction increases in reactor coolant inventory caused by the chemical and volume control system may be postulated to result from operator error or a falso electrical signal. Transients examined in this section are characterized by increasing pressurizer level, increasing pressurizer pressure and constant boron concentration. It is demonstrated that there is adequate time for the operator to take corrective action to prevent filling the pressurizer. The most limiting case would result if charging was in automatic control and the pressurizer level channel being used for charging control failed in a low direction. This would cause maximum charging flow to be delivered to the RCS and letdown flow would be isolated. 5,5.2.2 Methodology Rerste conditions of increased power level, and a +6.0 pcm/'F MTC. Four cases were analyzed; ,
- a. Minimum reactivity feedback with automatic pressurizer spray
- b. Minimum reactivity feedback without autorratic pressurizer spray
- c. Max! mum reactivity feedback with automatic pressurizer spray
- d. Maximum reactivitv feedback without automatic pressurizer spray 5.2.2.3 Results Figures 5.5.2-1 through 5.5.2-16 show the transient response due to the charging system malfunction. In all the cases analyzed, core power and RCS average temperature remsin relatively constant.
i Cases where the pressurizer spray is inoperable show the pressurizer level increases at a relatively constant rate. This is because the pressurizer pressure initially rises very quickly to the pressure at which the relief valves open and remains there. Cases where the pressurizer spray is operable show the pressurizer level increases with varying rates. Spray actuation tends to keep the pressurizer pressure lower for several minutes, which allows the charging pumps to deliver more flow. Eventually,. pressurizer pressure does increase enough to open the relief valves. times at which the operator would receive alarms are listed in Table 5.5.2-1. 5-19G
5.5.2.3 . Conclusions Results show none of the operating conditions during the transient approach core . limits. Because the pressurizer level trip has been defeated by failures, the transient l must be terminated by the operator. The sequence of events presented in Table 5.5.2 ; i show the operator has sufficient time to take corrective action. 4 t
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TABLE 5.5.2-1 Time Sequence of Events (con't) Rerate casel 2 Pressurizer 0.0 Positive MTC level channels WI spray fail low Maximum Charging 0.0 flow from one CCP Letdown isolation 0.0 Lo-lo-pressurizer 0.0 level Alarm Hi pressurizer 1490. level alarm from the one working level channel . Pressurizer relief 1712. valve setpoint reached Rerate Casel 2 Pressurizer 0.0 Positive MTC level channels fail low W/O spray Maximum Charging 0.0 flow from one CCP Letdown isolation 0.0 Lo-lo-pressurizer 0.0 level Alarm Pressurizer relief 358. valve setroint reached Hi pressurizer level 1388. alarm from the one working level channe' Pressurizer fills 1572. 5-201
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5.6 Decrease in Reactor Cmlant Inventory 5.6.1 Inadvertent Opening of a Pressurizer Safety or Relief Valve 5.6.1.1 Introduction An accidental depressurization of the Reactor Coolant System could occur as a result of an inadvertent opening of a pressurizer relief or safety valve. Since a safety valve is sized to relieve approxi:oately twice the steam flowrate of a relief valve, and will therefore allow a mucii more rapid depressurization upon opening, the most severe core conditions resulting from an accidental depressurizat;on of the Reactor Coolant System are associated with an inadvertent opening of a pressutizer safety valve. Initially the event results in a rapidly decreasing Reactor Coolant System pressure until this pressure reaches a value corresponding to the hot leg saturation pressure. At this time, the pressure decrease is slowed considerably. The pressure continues to decrease throughout the transient. The effect of the pressure decrease would be to dec: ease power via the moderator density feedback, but the reactor control system (in the automatic mode) functions to maintain the power and average coolant temperature until reactor trip occurs. Pressurizer level increases initially due to expansion caused by depressurization and then decreases following reactor trip. 5.6.1.2- Methodolog; This analysis was performed using uprated power conditions,- and a moderator temperature coefficient of +6.0 pcm/*F. A statistical core design methodology was used in the DNB analysis. In order to obtain conservative results in calculating the DNBR during the transient, the following assumptions are rnade:
- 1. Plant characteristics a nd initial conditions are discussed in Section 5.0.
- 2. A positive moderator tempersture coefficient of +6.0 pcm/*F is assumed. The spatial effect of voids due to : ni or subcooled boiling is not considered in the snaiysis with respect to reacovity feedback or core power shape. These voids would tend to flatten the core power distribution.
- 3. A large (absolute value) Dopoler coefficient of reactivity such that the resultant amount of positive feedback is vonservatively high in order to retard any power decrease due to moderator reactivity feedback.
5-218
5.6.1.3 Results The system response to an inadvertent opening cP a pressurizer safety valve is shown on figures 5.6.1-1 through 5.6.1-4. Figure 5.6.1-1 illustrates the nuclear power transient following the depressurization. Nuclear power is maintained at the initial value until OTAT reactor trip occurs. The pressure decay transient and average temperature responses are given in Figures 5.6.1-2 and 5.6.1-3. The DNBR decreases initially, but increases rapidly following the trip, as a shown in Figure 5.6.1-4. The DNBR remains above the safety analysis limit values throughout the transient. The calculated sequence of events for the inadver: ant opening of a pressurizer safety valve incident is shown on Table 5.6.1-1. 5.6.1.4 Conclusions The result of ths analysis show that the pressurizer low pressure and the Overtemperature aT Reactor Protection System provide adequate protection against the RCS depressurization event. No fuel or clad damage is predicted for this accident. Table 5.6.1-1 Time Sequence of Events Event Time (sec) SG safety Valve Opens 0.01 Reacter Trip (OTDT) 26.7 Turbine Trip 29.2 5-219
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WOLF CREEK FAure i 5.6.1- 1 Nuclear Power Trensient inr inadvertent opening -
- of c Pressurizer Safety Valve
( -- l 5-220
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$ 1,600 -
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$ 1,400 -
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1,000 0 40 - 60- 80 TNE (SECONDS) w ar CREEK Fiqure 5.6.1-2 Pressurizer Pressure Tronsient for lnodvertent openin_ g - of a Pressur;rer Sofety Volve 5-221 . .i
F
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8 y 570 - S u L t i i O 20 40 . 60 80 TIME (SECON05) _ WOLF CREEK Figure S.6.1-3 Core Average Temp Tronsient for inodvertent openin of a Pressurizer Sofety .gvolve l l-i: l 5-222
5.6.2 Break in Instrument Line or Other Lines From Ree tor Coolant Pressure Boundary that Penetrate Containment Refer to Section 5.7 for otscussion of this event. 5-223
5.6.3 Steam Generator Tube Rupture The steam generator tube rupture (SGTR) events were re-analyzed using the approved WCGS RETRAN models (9] to assess the impact of the uprated power conditions on the transient response as well as on the radiological consequences. In particular, the SGTR analysis was performed based on the reactor power uprated to 3565 MWt. Ten percent steam generator tube plugging, as part of the rerate project, is also considered in the analysis. The transient respon'a of a postulated SGTR accident is determined using conservative assumptions and boundary conditions that are consistent with the current WCGS licensing basis. Two SGTR scenarios have been identified which result in the most ilmiting radioriuclide celeases to the environment. Detailed analyses are presented for the following two scenarios:
- a. SGTR with postulated failure of the faulted steam generator Auxiliary Feedwater (AFW) flow control valve.
- b. SGTR with postulated stuck-open Atmospheric Relief Valve (ARV) for the faulted steam generator.
In the failed-controller case, auxiliary feedwater flow is maximized in order to increase . ( probability for faulted steam generator overfill and subsequent water relief from its safety valve. The radioactive releases are maximized by assuming that the safety valve is stuck-open following water relief with an effective flow area equal to 5% of the total safety valve flow area (8]. In the stuck-open ARV scenario, the discharge of contaminated secondary fluid is maximized by assuming the faulted steam generator ARV stuck-open for 20 minutes [7]. It has been determined that the most severe radiological consequences will result from the forced steam generator overfill scenario with a stuck-open safety valve [8]. Therefore, the radiological consequences are reported for that limiting case only. 5.6.3.1 Steem Generator Tube Rupture with Failure cf Faulted Steam Generator AFW Control Valve 5.6.3.1.1 Identification of Cause and Accident Description The accident examined is the cornplete severance of a single steam generator tube. The accident is assumed to take place at power with the reactor coolant contaminated with fission products corresponding to continuous operation with a limited number of defective fuel rods. The accident leads to an increase in the contamination of the 5-224 1
secondary system due to the leakage of radioactive coolant from the RCS Less of off-- site power is assumed to occur coincident with reactor trip. Discharge of activity to the atmosphere takes place via the steam generator safety and/or power-operated relief
- valves, in view of the fact that the steam generator tube material is Inconel-600 and is a highly ductile material, the assumption of a complete severance is conservativec The more probable mode of tube failure would be one or more minor leaks of undetermined origin. Activity in the steam and power conversion system is subject to continual surveillance, and an accumulation of minor leaks which exceed the limits established in the Technical Specifications is not permitted during unit operation.
Following the occurrence of the SG tube rupture, the primary to secondary leakage causes the pressur:zer level and the RCS pressure to decrease. As the RCS pressure continues to decrease, automatic reactor trip occurs on low pressurizer pressure or over-temperature delta-T (OTAT) signal. Because of the assumed loss of offsite power, the steam dump system will not be available, and the secondary side pressure increases rapidly after reactor trip until the steam generator ARVs and/or SV lift to dissipate the energy. After reactor trip, the RCS pressure continues to decrease and the safety injection is automatically initiated on low pressurizer pressure signal. Due to the assumed loss of offsite power at the reactor trip, normal feedwater flow is terminated and the AFW is initiated. The analysis assumes failure of the AFW control valve on the discharge sioe of the motor-driven AFW pump feeding the ruptured steam generator it is assumed that this valve fails in the wide-open position to maximize the flow to the ruptured steam generator. Failure of this valve coupled with the contribution from the turbine-driven AFW pump has a greater potential for overfilling the ruptured steam generator. The operator is expected to determine that a SGTR has occurred and to identify and - isolate the ruptured steam generator on a restricted time scale to minimize contamination of the secondary system and ensure termination of radioactive release to the atmosphere from the ruptured steam generator. The recovery procedure then can be carried out on a time scale to ensure that break flow to the secondary system is terminated. Consideration of the indications provided at the control board, together with the magnitude of the break flow, leads to the conclusion that the accident diagnostics and isolation procedure can be completed so that pressure equalization between primary and secondary side of the ruptured steam generator can be achieved to stop the break flow. Assuming normal operation of the various plant control systems, the following sequence of events is initiated by a design basis tube rupture: 5-225
l
- a. Pressurizer low pressure and low level alarms are activated and charging pump i flow increases in an attempt to maintain pressurizer level. On the secondary -
side, there is a steam flow /feedwater flow mismatch before trip as feedwater flow to the ruptured steam generator is reduced due to the additional break flow being supplied to that loop.
- b. Decrease in RCS pressure (Figure 5.6.3a) due to continued loss of reactor coolant inventory leads to a reactor trip signal generated by low pressurizer pressure or over-temperature Delta T. It is conservatively postulated that loss of ot' site power (LOOP) occurs when the reactor trips. The LOOP signalinitiates auxiliary feedwater addition. Resultar.t plant cooldown (Figures 5.6-3b and 5.6-
- 30) following reactor trip leads to a rapid change of pressurizer level (Figure 5.6-3f), and the safety injection signal, initiated by low pressurizer pressure, follows soon after the reactor trip. With the loss of offsite power occurring at the reactor trip, normal feedwater is terminated and the auxiliary feedwater is being supplied to the steam generators. ,
- c. The steam generator blowdown liquid monitor alarm and/or the condenser air discharge radiation monitor alarm will activate, indicating a sharp increase in radioactivity in the seconaary system. The alarms automatically cause termination of steam generator blowdown.
- d. The reactor trip automatically trips the turbine, and if offsite power is available the steam dump valves open, permitting steam dump to the condenser. In the event of a coincident loss of offsite power, as assumed in the transient presented in this section, the steam dump valves would autoraatically close to protect the condenser. The steam generator pressure (Figure 5.6-3d) would rapidly increase, resulting in steam discharge to the atmosphere through the steam generator safety / power-operated relief valves. In Figure 5.6-3g, the steam flow is presented as a function of time. The flow is constant initially until reactor trip, followed by turbine trip, which results in a large decrease in flow, but a rapid increase in steam pressure to the safety / relief valve setpoints.
- e. Following reactor trip, the continued action of auxiiiary feedwater supply and borated safety injection flow (supplied from the refueling water storage tanis) provide a heat sink which absorbs the decay heat.
- f. Safety injection flow results in increasing the pressurizer water level (Figure 5.6-3f); the rate of which depends upon the amcunt of operating auxiliary equipment.
- g. Follow;ng water relief through the ruptured steam generator safe;y valve, the ruptured steam generator pressure is uncontrollably decreasing and the operator -
is directed to enter the Emergency Operating Procadure EMG C-31, 5-226 4
1 l 5.6.3.1.2 Method of Analysis in estimating the mass transfer from the RCS through the broken tube, t..s following assumptions are made:
- a. Reactor trip occurs automatically as a result of low pressurizer or overtemperature Delta T. Loss of offsite power occurs at reactor trip.
I b. Auxiliary feedwater flow rate is allowed to vary with the Ductuation in the faulted steam generator pressure. This results in higher AFW flow when the faulted steam gerierator pressure decrqases. At 16 minutes into the accident, the operator terminates AFW flow to the faused steam generator. AFW flow to the intact steam generators is continued to maintain the narrow range 'evel between 4% and 50% as indicated in Emergency Operating Procedure EMG E-3.
- c. Cooldown of the RCS is initiated at 24 minutes. It is assumed that steam is released through the remaining two operable steam generator ARVs in the intact S
loops until RCS temperature is less than the temperature of the core exit thermal couples corresponded to the ruptured SG pressure as !!cted in EMG E-3 procedure. Technical Specification LCO 3.7.1.6 requires that at least three SG ARVs shall be operable. With one of the required ARVs inoperable due to the single failure assumption, there are at least two ARVs still available to ensure that subcooling can be achieved for the RCS.
- d. Primary depressurization is initiated 1 minute following termination of the RCS cco!cown. The depressurization is performed by opening the pressurizer PORV until .ne pressurizer pressure is below the ruptured SG pressure.
- e. Following the depressurization termination of Slis delayed to ensure enough l'.pd entering in the ruptured SG steamline to force the safety valve open and water relief. It is assumed that the Si is terminated 10 minutes after the RCS depressurization.
- f. As the pressure in the ruptured steam generator uncontrollably decreases following water relief, the operators are directed to switch to EMG C-31 procedure. The first major step in EMG C-31 is to initiate a second RCS cooldown. It is assumed that the operators will be able to perform the second cooldown 10 minutes after water relief. This cooldown process is continued to the end of the analysis using tne two intact steam generator atmospheric relief vaives.
- g. Second depressurization is initiated 6 minutes follov, ng the second RCS cooldown. The pressurizer PORV is open to release excessive pressure and ,
closed when RCS pressure and ruptured SG pressure are equalized. : r 5-227 l
The above assumptions are conservatively made to increase the probability for faultea steam generator everfill and to maximize the radioactive releases to the atmosphere. The operator action times are taken from Reference 8. Results in Table 5.6.3-1, the sequence of events are presented. These events include postulated operator response times and normal plant responst , to the normal plant seth ' - As deph ~.i in Figure 5.6-3j, the steam generator overfilling occurs at approximately 1800 seconas. ~v' vater relief through the ruptured steam generator safety valve occurs at approximately 2830 seconds li should be noted that a static analysis of the steam lines has been performed for all steam lines completely filled with water. 5.6.3.2 Steam Generator Tube Rupture with Postulated Stuck open Atmospheric Relief Valve 5,6.3.2.1 Identification of Causes and Accident Description The accident description for this SGTR is similar tc that discussed in previous section with the assumption that the ARV for the faulted steam generators remains open 'or 20 minutes following initial secendary pressure relief. In this SGTR scenario, the aperator is expected to determitie that a SGTR has occurred, to recognize the failure of the ARV, to dispatch personnel to manually isolate the ARV, and then to isolate the ruptured steam generator. In this accident the fai. lure of the ARV allows a longer period for radionuclide release than otherwise expected. , The accident involves the complete severance of a single steam generator tube with loss of offsite power coincident with reactor trip. Fcr a discussion of normal operation of plant control systems to a design basis SGTR refer to Section 5.6.3.1.1. 5.6.3.2.2 Analysis of Effects and Consequences Methodology in estimating the mass transfec from the RCS through the broken tube, the following assumptions are made:
- a. Reactor trip occurs automatically as a result of low pressurizer pressure or overtemperature Delta T. Loss of offsite power occurs at roactor trip.
5-228
- b. As pressures rise on the secondary side, the steam generator atmospheric relief valves (ARVs) open to release excess secondary pressure. Although the ARVs in the intact steam generator close within 7 minutes, the ARV for the ruptured steam generator is assumed to remain open and steam release to continue for j
20 minutes until the ARV block valve is manually closed.
- c. AFW is initially delivered at a rate of 250 gpm to each steam generator. AFW is maintained to assure narrow range level at 15% in all steam generators.
- d. Following ruptured steam generator isolation, cocidown is initiated when the -
narrow range in the ruptured steam generator is greater than 10% and its pressure exceeds 615 psig. Cooldown continues until RCS temperature is less th m the required temperature at the core exit thermal couples as indicated in EMG E-3.
- e. Reactor coolant system depressurization is initiated three minutes after completion of cooldown. This timing is consistent with observed simulator exercises.
- f. After primary side depressurization is completed and Si termination criteria are met, a three minute time delay is assumed prior to Si termination.
- g. Followirg Si termination, the operators equalize pressure in the RCS and faulted SG in'5 minutes. During this time break flow in the faulted SG continues. After pressures are equalized, it is conservatively assumed that the transition to cold shutdown is made utilizing steam release to the atmosphere from the faulted SG.
Results in Table 5.6-1, the sequence of events are presented. These events include postulated operator response times and normal plant responses to the normal plant setpoints. Parameters of prir::ary and secondary systeras are plotted as a function of time in Figures 5.6-3k to 5.6-3t. 5.6.
3.3 CONCLUSION
S Analysis of the SGTR with postulated failure of the AFW control valve indicates that overfilling the steam generator occurs during the transient. Subsequent accident dose analysis for both SGTR scenarios shows that the radiological consequences resulting from a SGTR accident remain wc!! within the limiting values specified in 10CFR100 and Standard Review Plan 15.6.3.- A steam generator tube rupture will cause no 5-229 _ _ _ _ _ _ _ _ ________________________________w
subsequent damage to the RCS or the reactor core. An orderly reculery from the accident can be completed, even assuming simultaneously loss of offsite power. 5-230
TABLE 5.6.3-1 TIME SEQUENCE OF SGTR EVENTS Event Time (sec) Accident Steam generator tube rupture - failed-open auxiliary feedwater control valve and safety valve Tube Rupture Occurs 0.0 I Reactor Yrip Signal 45.6 c Reactor Trip 47.6 Auxiliary Feedwater injection 77.6 Safety injection Signal 160.1 Safety injection 175.1 Terminato Auxinary Feedwater 960.0 to the Faulted SG Initiate RCS Cooldown 1440.0 Terminate RCS Cooldown 1909.0 Initiate RCS Deoressurination 1967.0 Terminate RCS Depressurization 210'- Terminate Safety injection 2707.8 Water Relief Through SG SV 2831.0 Initiate 2nd Cooldown 3431.0 Initiate 2nd Depressurization 3791.0 Terminate 2nd Depressurization 3815.7 5-231
- u. -
TABLE 5.6.3-1 (Sheet 2) o TIME SEQUENUt OF SGTR EVENTS - Accidelli Event Time (sec) Steam generator tube rupture - stuck open atmospheric relief valve Reactor Trip Signal 140.0 Reactor Trip 142.0 Faulted SG ARV Open 144,0 Auxiliary Feedwater injection 202.0 Safety injection Signal 297.0 Safety injection 322.0 Faulted SG ARV isolated 1344.0 initiate RCS Cooldown 2216.0 Terminate RCS Cooldown 3024.0 initiate RCS Depressurization - 3204.0 Terminate RCS Depressurization 3364.0 Terminate Safety injection 3544.0 Pressure Equalization 3844.0 6-232
-e, F
FIGURE 5,6-3a < SGTR FORCED OVERrit L W/ STOCK-OPEN SG SV RC5 PRESSURE 2,500 J Terminate St initiate Cooldoen / 2,000
\ T ermino*.e Coolcown i / ! Ini late 2nd Cooldowr:
('
\ Termiaote AFW j initiate Si /
Cr 1,500 - h V) initicle Depressurization LJ - initiate 2nd + { Depr essurizatior 1,000 ~
/
Termincte Depressurizatien
/
Terminate 2nd Depressurization 300 O 1,000 2.000 3,000 4,000 5,000 TIME (SECONDS) 5-233 I k.
. ..- . . . . . . . - . - . . - . - . - . _ _ . . . ~ . . . . . .
FIGURE- 5.6-3b
- SGTR FORCED OVERFILL W/ STUCK-OPEN SG SV ..
- - i FAULTED LOOP RCS TEMPERATURES 700 Hot Leg 600 '
Cold Leg -
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FIGURE 5,6-3c SGTR FORCED OVERFILL W/ STUCK-OPEN SG SV INTACT LOOP RCS TEMPERATURES 650 Hot Leg J Cold Leg 600 -------
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FIGURE 5.6-3d SGTR FORCED OVERFILL W/ STUCK-OPEN SG SV STEAM GENERATOR PRESSURES 1,400
- Reactor Trip water Rel'ef s 1,200 Foutted
\
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laolots Foulted SG
... __ . -",/
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400 -
/ ',',,
Terminate Cooldown 200 = 0 1.000 2.000 3.000 4.000 5.000 TIME (SECONDS) 5-236
. FIGURE 5.6-3e SGTR FORCED OVERFILL W/ STUCK-OPEN SG SV -
STEAM CENERATOR TEMPERATURES 600
,/
Rootter Tfl0 intoct { i ! i Foutted
$50 3,
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pinitiate Cooldown
,/ initiate 2nd Cooldo*n Q500 cr' W
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& routted SG 450 -
Terminate Cooldown l 400 l l 0 1,000 2,000 3.000 4,000 5.000 l TIME (SECONDS) l i l i (; l I i 1 I l l 5-237
FIGURE 5.6-3f SGTR FORCED OVERFILL W/ STUCK-OPEN SG SV PRESSURIZER WATER LEVEL 35 Terminate si --* 30 - in;tiate 2nd Cooldown S 25 - [ 20 -
~ ; Initiate Cooldoen /
Initiate 2nd 15 -
\ o,p,essurization
/
Terminate 2nd 10 - Depressurization k Terminote : Terminate Cooidown N Depressurization - Initiate SI i , 1 i 1 3 0 1,000 2,000 3,000 4,000 5,000 TIME (SECONDS) 5-238
FIGURE 5.6-39 SGTR FORCED OVERFtLL W/ STUCK-OPEN SG SV FAULTED SG STEAM FLOW 1,200 1,000 O g 800 1 600 8 400 7 i b 200 G
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FIGURE 5.6-3h SGTR rORCED OVERFILL W/ STUCK-OPEN SG SV FAULTED SG + AFW. FLOW & NARROW RANGE LEVEL G v 160 hrow Range y gg ,
,s'
,, p isolote foulted SG z u AFW Flow g - ,' l 12 0 ~ ,'
y , O ' e i tr < i
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FIGURE 5,6-3i SGTR FORCED OVERFILL W/ STUCK-OPEN SG SV FAULTED SG - BREAK FLOW 50 initiate Cooldown Terminate $l 5 g / Initiote 40 -
/ Depress. in;tiate 2nd g
y Terminate ## to Foulted SG Cooldown
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. Terminate 2nd
- Depressurizction DepressMion
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0 1,000 2,000 J,000 4,000 5.000 TIME (SECONDS) 5-241 4
FIGURE 5.6-3J SGTR FORCED OVERFILL W/ STUCK-OPEN SG SV FAULTED SG - MlXTURE VOLUME 7.000
-6.000 -
S v$5.000 - W llE D d 4.000 - W tt J.000 5 - 2.000 - f 1.000 - 1.000 2,000 J000 4'000 5.000 TIME (SECONDS) 5-242
FIGURE 5.6-3k SGTR W/ STUCK-OPEN ARV RCS PRESSURE 2,500 Reactor Trip in;tiate Cooldown Initiate Depressurization 2.000 - AlW isolated \ W b Terminate g f Terminate Cooldown 51 3 _ initote Sofety injection ki M b O'
\..
1 Terminate Depressurization 1,000 - N
, i i , i i i i i . i 0 400 800 1.200 1.600 2.000 2,400 2,800 3,200 3.600 4,000 TIME (SECONDS)
FICURE 5.6-31 SGTR W/ LTilCK-OPEN ARV FAULTED LOOP RCS TEMPERATURES 700 ' Hot Leg Cold Leg .----- 600 -
-~s, N
- ~'~~~~~~~'
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y - i W H 200 - i ( ~.-
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'v l i i , , 1 i , i i -
L i - ' ' O 400 800 1.200 1.600 2,000 2,400 2,800 3.200 3,600 4,000 TIME (SECONDS) 5-243
l 5.' FIGURE 5,6-3m SGTR W/ STUCK-OPEN ARV INTACT LOOP RCS TEMPERATURES 700 Hot Leg Cold Leg 650 - - 6 m 600 - 6 w - {550 ,_ ( LOO - ' - --. .. _ Q-d Q ~
'~ ehD -
-. .. l Il 400 -
t t i 1 a I i ii 1 33g . 0 400 800 1,20n 1,600 2,000 2,400 2.800 3,200 3,600 4,000 TIME (SECONDS) 9 FIGURE 5.5-3n SGTR W/ STUCK-OPEN ARV STEAM GENERATOR PRESSURES 1,500 " Reactor Trip Intact f 3Ulted t I'000 -
's initiote Cooldown E ',
N g \ . ---._,..-_ .-- __.. ..._ tr s - - 3 ' M s W s LA) ct: 500 - CL l . ARV K0!cted [ i 5 Terminate Cooldown f-0 O 400 B00 1,200 ' .600 1 2,000 2,400 2,800 3.200 3,600 4,000 TIME (SECONDS) 5244
1 FIGURE 5.6-3o SCTR W/ STUCK-OPEN ARV STt.AM CENERATOR TEMPERATURES
$80 t intact Faulted----
560 -
\..
540 M\ C ' i Initlate Cooldoen g 520
\, ,. - -~..___. .._____.. $ \ '
Q 500 i ,', Z g \ /
$ 480 -
y t 460 - 'N **
- U ""
/
440 - P ARV isolated 420 O 400 800 1,200 1.600 2,000 2,400 2.800 3,200
~
' ,600 3 4.000 TIME (SECONDS)
FIGURE 5,6-3p FOTR W/ STUCK-OPEN ARV PRESSUR'ZER WATER LEVEL 35 Terminate Reactor Trip Sofety injection 30 -k' \ p 25 - 6 - d 20 - Po J Initiate Cooldown e 15 LJ
~
N N 10 Initate safety injection 5 -
- W lsolated Terminate Coolaown 0
l C 400 800 1,200 - 1.600 2.000 2.400 2,800 3,200 3,V 4.000 j TIME (SECONDS)' l l' 5245 l I
FIGURE 5.6-3q SGTR W/ STUCK-OPEM ARV FAULTED SG - STEAM FLON 1.200 1,000 - n U 800 - q v) N Si (D 600 - - d h AW Open g 400 -
/
I 200
- AW Isolated 0
0 '400 800 1.200 1,600 2.000 2,400 2,800 3200 3,600 4,000 TIME (SECONDS) 5-246-
- .c
[ ._ l_
. 2, _ - - . . . . . . , - . . . . , _ , _ _ - , - , , , _ , . , ., . . , , - , _ _ . _ - - . _ , - - , , . . , - - _ _ , _ , . _ _ ,, - - .
- . . ~ . - - . - . - _ . . - . - - _ . . . - - - . - - - - . . . - . .
i, i i FIGURE 5,6-3r SGTR W/ STUCK-OPEN ARV
-FAULTED SG - AFW FL OW AND NARROW RANGE LEVEL !
100 100
/ i AFW FLOW SG N.R. / !
--.... ,- 7 I
B0 - ' - 80 ^ - m .' h w U / , t r D 60 - ,/ - 3 ' f i 60 _ Q
/ y
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k d h a' 9 6 U1 20 ~j, 20 ,1 i
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0 0 - 0 400 800 1.200 1.600 ' ,000 2 2.400 2.800 3.200 3,600 4.000 TIME (SECONDS) i E f t l
+
E 5-247- c l.
- l. -,i.. , _ . . .L,.~. . .. . . _ . . . . . . - , .... , _..___ - . ,.. _.,....., _ ....,. ,_, ,.,,_
l flGURE 5.6-3s SGTR W/ STUCK-OPEN ARV FAULTED SG - TOTAL BREAK F LOW 60 l initiate Cooldown ) initicle Depressuritation Reactor Trip [ 50 - 1
^
s , Terminate
$0 -
ARV isolated / SI I Terrninote Cocidown ( s , m 30 - d initote Safety injection l 5 O y 20 - Termincte Depressurization 10 I 4 i i i i 1 i i 1 . I 1 . Q 0 400 800 1,200 1.600 2,000 2,400 2,800 3,200 3,600 4,000 TIME (SECONDS) FIGURE S.6-3t SGTR d/ STUCK-OPEN ARV FAULTED SG - MIXTURE VOLUME 4,500 4,000 - 3,500 - ., { 3,000 w h 2,500 - d ARV isolated 2,000 - 1,500 - t - i i . i . t i i , 3,ono 0 400 B00 1.200 1,600 2,000 2,400 2,800 3.200 3,600 4,000 TIME (SECONDS) 5-248 e m u-+-., g., g g- .#. pp ,e.m g ,,mes- _m,e
.,,r., -.,,y . e,- . . , - , , - - . , 'e-y'--%
l i 5.6.4 l . Spectrum of BWR Steam System Piping Failures Outside of Containment This event is not applicable to WCGS. 5,6.5 LOCA Sec Section 6.0 for LOCA analyses. 5.6.6 A Number of BWR Transients This event is not applicable to WCGS. .f P F
'\
4 f 249
5.7 Radioactive Release from a Subsystem o Component 5.7.1 Summary and Conc;usions Westinghouse and WCNOC have reviewed the accidents analyzed in the USAR with respect to radiological source terms and radiological consequences, , The basic radiation source terms used in the current USAR dose consequences analyses were based on a core power level of 3565 MWt and on an operating cycle of twelve months. The radiation source terms were then re-calculated when the design changes that converted the 12-month fuel cycle to 18 month was implemented. The impact of extended fuel burnup as a result of increased cycle length on the radiological consequences of accidents was found to be insignificant since the calculated radiation doses are directly related t the core inventory of radioactive isotopes which is not strongly influenced by the extension of fuel burnup. Since the source terms originally generated for dose consequences evaluations were already based on the 3565 MWt, uprating the WCGS to this power level will not affect the calculated design radiation source terms and the accident doses although the actual fission product inventory will increase by an amount equal to the increase in power level (4.5%). While the effects from the transition to an extended fuel cycle and a core thermal power of 3565 MWt on these analyses have been evaluated to be minor, WCNOC has eheted to use the revised source terms to update the radiological consequence analyses contained in the USAR. Based upon the results of the updated dore ' analyses for 'he Chapter 15 accidents, there is reasonable assurance that the potential doses of a postulated non-LOCA event will remain well within the exposure limits set forth in 10 CFR Part 100, paragraph 11. It is further confirmed that the distances to exclusion area and to the low population zone boundaries, in conjunction with the operation of the dose mitigating ESF systems, are sufficient to provide reasonable assurance that the total radiological consequences of a design-basis LOCA will be within the guidatine values o' 10 CFR 100. The generction of fission product inventories and the radiological consecuences analyses for the Chapter 15 accidents are discussed in the following sections. 5.7.2 Fission Product Inventories and Coolant Activities For the WCOS uprating the calculated fission product source terms are already based on a reactor power level of 3565 MWt, thus the calculated fission product source terms will not change as a result of the uprating. Tha activation of reactor materials which are released to the reactor coolant by corrosion 1.o other mechanisms is also directly related to power level. For WCGS these sources were alco calculated for a core thermal power of 3565 MWt, thus they will not be affected by the uprating. 5-250
., - , ,- m-*
o Fission product energy release is directly related to core power level, and to a lesser degree to operation cycle length. The effect of a longer operating cycle decreases as the cycle length is increased and does not usually have a significant effect on the total fission product energy released. Even though the cycle length does not usually have a significant offect on the fission product energy release it can cause significant changes in the amount of some isotopes in the total fission product inventory. This is because < of the differences in the half lives of the various isotopes and the effect that the longer ' cycle has on their presence. That is, cortain isotopes will be present iri increased amounts while concentrations of other isotopes will remain the same or decrease slightly. A re analysis was performed bcsed on a core thermal power of 3565 MWt and an extended fuel cycle length with higher enrichment and higher burnup to quantify the changes in the amount of fission product inventories. The calculation of the core fission product inventory is besed on a three region equilibrium cycle core which consists of 72 triple-burned assemblies in Region 1,72 twice burned assemblies in Region 2, and 49 once burned assemblies in Region 3. ! The average burnup in the regions at the end of a cycle (MWD /MTU)is 19,660,37,460 and 47,900 respectively. The calculated fission product inventory for each of the three regions as well as a combined total core inventory is given in Table 1. The maximum isotopic primary coolant activities that represent the maximum values attained over a time period ranging from 1000 seconds after startup through end of cycle is also given in Table 2. l. 1 5 251 l l-
Table 1 Region and Core Total Activity at End of Cycle (Activities in Curies / Region) Isotope 2 x Burned 1 x Burned Feed Total 1131 1.72E+ 07 3.93E+07 4.11 E+07 9.76E+07 l-132 2.48 E+07 5.72E+07 6.07E+07 1.43E+08 l-133 3.39E+07 7.98E+07 8.85E+07 2.02E+08 l134 3.61 E + 07 8.58E + 07 9.67E+07 2.19E+08 l-135 3.13E+07 7.39E+07 8.23E+07 1.88E+08 KR-83M 1.80E+06 4.54 E +06 5.71 E+C6 1.21 E+07 KR-85M 3.83E+06 9.99E+06 1.32E+07 2.70E+07 KR-85 3.07E+05 3.85E+05 2.29E+05 9.21 E+05 KR-87 6.85E+ 06 1.82E+07 2.49E+07 5.00E+07 KR-88 P79E+06 2.60E+07 3.53E+07 7.11 E+07 KR 89 1.18E+07 3.19E+07 4.46E+07 8.83E+07 XE-131M 1.09E+05 2.77E+05 2.90E+05 6.76E+05 XE-133M 4.87E+06 1.14 E+07 1.26E+07 2.89E+07 XE-133 3 41 E+07 7.98E+07 8.53E+07 1.99E4 08 XE-135M 7.07E+06 1.60E+07 1.65E+07 3.96E+07 XE-135 1.05E+07 1.87E+07 2.15E+07 5.07E+07 XE-137 2.96E+07 7.02E+07 7.88E+07 1.79E408 XE-138 2.6.' E +07 6.35E407 7.45E+07 1.64E+08 5-252
,, ..,it, p * - - - .
Table 2 Maximum Coolant Activities During 420 EFPD Cycle at 3565 MWt (Assuming 1% Failed Fuel) Isotope Coolant Activity (uCi/gm) l l-131 2.84E+00 -l 1-132 2.88E+00 l133 4.40E+00 1-134 6.40E-01 1-135 2.56E+00 KR-83M 4.80E-01 KR-85M 1.92E+00 , KR-85 7,60E+00 i KR-87 1.20E+00 ' KR-88 3.44E+00 KR-89 1.00E-01 XE-131M 2.16E+00 XE-133M 1.76E+01 XE-133 2.64E+02 XE-135M 5.60E-01 XE-135 8.80E+00 XE-137 2.00E-01 XE-138 6.80E-01 l l L i l 253-l __ _ _ . l. u .. . _ . - _ . _ - . _ _ . _ _ _ _ . _ _ - . . _ _
l i 5.7.3 Radiological Consequences for Analyzed Accidents To confirm the adequacy of the design in mitigating consequences of an accident, the radiological consequences of the postulated design basic accidents are reanalyzed using the revised source terms shown in Table 1 and/or 2. These accidents, which are the same as those previously analyzed in the USAR Chapter 15, include, Main Steamline Failure Outsids Containment Loss of Non Emergency AC Power Accident Locked Rotor Accident RCCA Ejection Accident 3 Failure of Letdown Line Outside Containment Steam Generator Tube Rupture Loss Of Coolant Accident Fuel Handling Accident Waste Gas Tank Rupture , Liquid Radwaste Tank Rupture The accident doses were re-calculated using methodology and assumptions that are consistent with the current licensing basis dose calculations and following applicable Standard Review Plan (SRP) sections and Regulatory Guides. Main Steamline Failure outside Containmen1 Two cases of iodine spiking are corsidered for the dose calculations. For the first case, it is' assumed that a reactor transient has occurred prior to the postulated main steamline break and has raised the primary coolant iodine concentration to the maximum value permitted by the Technical Specifications (i.e., a preaccident iodine spike case). For the second case, it is assumed that the reactor trip and/or primary , system depressurization associated with the MSLB creates an iodine spike in the primary system and that the iodine release rate from the fuel to the primary coolant following the accident is increased by a factor of 500 (i.e., concurrent iodine spike case). Since the engineered safeguards design power (3565 MWt) was already assumed in the original steam release for._ dose evaLation, steam releases from the secundary remain the same as that originally contained in the L:, AR: 0-2 hrs 4.04 x 105 lb 2-8 hrs 9.46 x 105 lb Utilizing thece steam releases and the revised source terms as shown in Table 2, the release of fission products from the core to the environment via the steam generator safety and relief valves were recaiculated and tabulated in Table 3. The calculated 5254 , l
l whola-body and thyroid doses at the exclusion area boundary (EAB) and the low population zone (LPZ) boundary due to release from the steamline break accident are presented in Table 4 I i 1 l 1 P il L ,
- 5 255
- 1:
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- _ . _ - . .. . _ . . _ . _ _ _ _ _ _ . - _ -.--_.-__._---._.m_.--... .._s Table 3. Activity Released to the Environment Case 1: Prr accident loe,e Spike ISOTOPE 0 2 hr (Cl) 0-8 hr (Ci) 1-131 26.33294 86.93657 l-132 21.60790 37.56772 1133 39.88802 121.10310 1-134 3.62994 4,22950 1135 22.09400 56.40312-Xe-131m 0.97729 3.86949 Xo-133m 7.87948 30.23533 Xe-133 119.07910 467.15080 Xe-135m - 0.04754 0.04777 Xe-135 3.70273 11.93058 Xe-137 0.00418 0.00418 Xe-138 0.05251 0.052G6 Kr-83m 0.15250 0.27241 Kr-85m 0.74879 1 99172 Kr-85 3.44691 -a,74573 Kr-87 0.33149 0.49257 Kr-88 1.23465 2.73822 Kr-89 0.00173 0.00173
?
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+--,e , , ~. ,,,.n- , w. -ve.,,,.w-r -,,..w.-- ~ - , . - - - - -. -..-e-
i.. I Case 2: Concurron.t lodine Spike ISOTOPE 0 2 hr (Cl) 0-8 hr (Ci) 1-131 18.49068 200.81 % 0 1-132 49.52575 252.11370 l-133 35,34147 372.42050 1-134 13.77343 27.51A85 : 1 135 28.34575 254.54000 Xe 131m - 0.97729- 3.86949
- Xe-133m - 7.87948 30.23533 Xe 133 119.07910 467.15080 Xe-135m 0.04754 -0.04777 Xe-135 3.70273 11.93058 Xe-137 0.00418- 0.00418 '
.Xe-138 0.05251 0.05266 Kr-83m 0.15250' O.27241 Kr-85m 0.74879 1.99172 Kr-85 3.44691 13.74573 Kr-87 0.33149 -0.49257 Kr-88 1.23465 2.73822 Kr-89 'O 001/3 0.00173 .
1 1 5 257' _ _- = - =,vl,,e -n e-+..sp.m-- y e p w , -e-m-w ,m-r, -e.n.e- p-.- .. w.s e-,w- ,,..,e-,-,..w---- - e m w.v - w ,.--e.,,.- e-m,e-r.. ,, , - , -g-a, oy--,,e er,,w,w-- ..e.,- .w.,w,, ,;. r,m,,
'l Table 4 Radiological Consequences of A Main Steam Line Break l Lose (rom) l i
Updated Current USAR l
'i Case 1: Pre accident lodine Spike i
Site Boundary (0-2 hr) . I Thyroid .682- 2.7 Whole Body 4.931 E-03 3.3E-3 Low-Population Zone (Duration) + Thyroid 1.151 1.2 Whole body 1.568E-03 1.1 E-3 Case . Concurrent lodine Spike ; Site Boundary (0-2 hr) : Thyroid 2.051 2.0 Whole Body 8.075E-03 4.8E 3 t Low. Population Zone (Duration) Thyroid _ 2.807 2.7 Whole body 6.663E-03 4.3E-3 [ w L
, y l- :
5-258
Loss Of Non Emeraency AC Power To The Plant Auxillarieg 3 4^ Steam releases from the secondary remain the same as that originally contained in the USAR since the uprated power (3565 MWt) was alread/ assumed in the original steam i i release for dose evaluation. 0-2 hrs 5.49 x 105 lb . 2-8 hrs 1.03 x 10G lb t Utilizing these steam teleases and the revised source terms as shown in Table 2, the I activity releases; talong with the thyroid and whole-body doses at the exclusion area : boundary and low population zone due to the releases resulting from the loss of ; nonemergency AC power to the station auxillaries are calculated and presented in Table 5 and 6. Table 5. Activity Released to the Environment .
!sotope 0-2 hour 0-8 hour .
1131 1.89 x 101 F.47 x 101 1132 1.44 x 101 2.38 x 10-1 1-133 2.84 x 101 7.64 x 101 , 1-134 2.14 x 10 2 2.49 x 10-2 1-135 1.55 x 101 3.57 x 10-1 ! Xe 131m . 9.78 x 10-1 . 3.89 Xe 133m 7.89 3 04 x 101 Xe-133 1.19 x 102 4.69 x 102 i Xe-135m 4.76 x 10-2 4.78 x 10-2 Xe-135 3.71 1.20 x 101 ' Xe-137 4.18 x 10 3 4.18 x 10-3 l Xo-138 5.25 x 10 2 5.27 x 10-2 ~ Kr-83m 1.53 x 10'1 2.73 x 101 ' Kr-85m 7.50 x 10-1 - 2.00 Kr-85 3.46 1.38 x 101 ' Kr-87 - 3.32 x 101 4.93 x 10*1 Kr-88 1.24 2.~ 5 . Kr 89 1.73 x 10*3 1.73 >.10-3 l l 5 w e + e W e W
i Table G. Radiological Consequences of Loss of Nonemergency AC Power to the l Station Auxiliaries -! Dose (rem) Updated Current USAR l Exclusion Area Boundary (0 2 hr) ! Thyroid 1.916E 2 1.93E 2 l Whole Body 3.437E-4 1.10E.4 j i Low Population Zone (Duration) i Thyroid 7.246E 3 7.30E 3- .i Whole body 1.449E-4 3.97E-5 j i I I
.I l
i 1 i o l 5-260 l 1 l
t
?
i Blaqtor Coolant Pumo Sha_ft Seizure (Lockod._, Rotor) , The steam releases are the same as those given for the loss of AC power accident ,
- analysis. Steam releases from the secondary remain the same as that originally '
l contained in the USAR since the uprated pcwer (3565 MWt) was already assumed in the original steam release for dose evaluation. 0 2 hrs 5.49 x 105 lb . 2-8 hrs 1.03 x 106 tb Utilizing these steam releases and the revised source terms as shown in Table 1 and 2, the calculated activity release from the core to the environment via the steam generator safety and relief valves were recalculated and tabulated in Table 7. The resultant whole-body and thyroid doses at the exclusion area boundary (EAB) and the low ' population zone (LPZ) boundary due to release from the locked rotor accident are presented in Table 8. I Table 7 Activity Released to the Environment Isotope 0-2 hour 0 8 hour i 1131 6.65 6.62 x 101 1-132 6.59 2.54 x 101 i 1-133 1.31 x 101 1.18 x 102 1134 5.49 9.26 l135 1.11 a 101 7.87 x 101 Xe-131m 7.81 3.10 x 101 Xe-133m 2.97 x 102 1.14 x 103 Xe-133 2.12 x 103 8.53 x 103 o Xe-135m 7.50 x 101 7.54 x 101 Xe-135 4.80 x 102 1.55 x 103 Xe-137 8.34 x 101 8.34 x 101 Xe-138 2.83 x 102 2.83 x 102 ! Kr-83m 8.60 x 101 1.54 x 102 l Kr-85m 2.36 x 102 6.29 x 102 Kr-85 3.14 x 101 1.26 x 102 l Kr-67 3.09 x 102 4.59 x 102 Kr-88 5.71 x 102 1.27 x 103 Kr-89 3.41 x 101 3.41 x 101 F 5261
.,m-+-
=- ==*ga h pm-dwFm'77-==' -F V- -+C4y .F 7 y w g-ry--g, n g *9y 7 -- 7tv---- 3 -- 9g a- tes----+p-~m9 e W Yd tygw.+t- =w*t'-5+Y.-Y*wt-t T
i Table 8. Radiological Consequences of A Locke:1 Rotor Accident Dose (rem) IJpdated Current USAR Site Boundary (0 2 hr) Thyroid 0.738 0.650 Whole Body 0.0758 0.110 Low Population Zone (Duration) Thyroid 0.938 0.81 _ Whole body 0.021 0.033 j i I I 5-262 9 44- - _,ggp$r e ,Jawng7,9.urw .g m g.gpwy-ww,-.5p ww--w..p-.,--gpcg g,.g.-w . yyg ,f p g-.gdi-,,4,-.g-g.. --.-.,g+gr y- g-p w y en. ,,---,c, w v.,.re,,,- 4 m%w--%w,-* emah d - - mwy-
Spectrum Of Rod Cluster Control A,pembly Eiection Accidents The release volume for this accident remains the same as that originally contained in the USAR. Utilizing the new source terms, the activity relaases and the radiological consequences due to a design basis RCCA ejection accident were recalculated as fo! lows: A. Activity Released to the Environment. Steam Generator Containment Releese (Cl) Release (Cl) Isotope 0-2 hr 0 2 hr 0-30 days 1-131 4.720'l+00 9.114 E+01 6.373E+03 1132 6.813E+00 1.008E+02 2.233E+02 1-133 9.695E+00 1.832E+02 - 2.218E+03 1-134 1.012E+01 1.031E+02 1.299E+02 1135 8.921 E+00 1.597E+02 8.408E+02 XE-131M 3.528E-01 1.485E+00 1.353E+02 . XE-133M 1.442E+01 6.005E+01 1483E+03 , XE-133 1.006E+02 4.219E+02 2.124E+04 XE-135M 1.857E+01 1.543E+01 1.550E+01 XE-135 2.507E+01 9.82SE+01 G.404E+02 XE-137 7.203E+01 1.716E+01 1.716E+01 XE-138 7.650E+01 5.812E+01 5.828E+01 KR-83M 5.933C+00 1.766E+ 01 3.317E+01 KR-85M 1.330E+01 4.841 E+01 1.704E+02 . KR-85 1.249E+00 5.270E+00 9.682E+02 KR-87 2.443E+01 6.343E+01 9.556E+01 , KR-88 3.496E+01 1.173E+02 3.032E+02 KR-89 3.412E+01 7.008E+00 7.008E+00 1 l
- l. 5 263
-- , . , , .~m-- . _ . , - - . , - _ ..-..-,
B. Radiological Consequences Dcse (rem) Updated Current USAR Exclusion Area i: >unciary (0-2 hr) Thyroid 1.07 E+1 9.83E40 Whole Body 511 E-2 6.77E 2 Low Population Zone (Duration) Thyroid 1.29E+ 1 1.20E+1 Whole body 2.35E 2 2.64E-2 5-264 c l 0
t Break !n Instrument Line Or Other Lines From Reactor Coolant Pressure Boundarv ! That Penetrate Containment The most severe pipo rupture with respect to radioactivity release during normal plant ; operation is rupture of the chemical and volume control system (CVCS) letdown line at , a point outside of the containment. For such a break, the reactor coolant ietdown flow would have passed sequentially from the crossover leg and through the regenerative , heat exchanger and letdown onfices. The letdown orifice reduces the letdown line pressure from 2235 psig to less than 600 psig outside containment when letdown flow is maintained at 120 gpm during normal plant operation. Increase in flow will occur due to a rupture of the letdown line downstream of the orifices. It was determined that the occurrenco of a complete severance of the letdown line would result in a loss of reactor coolant at the rate of 141 gpm. Based upon this re-calculated break flow with the revised radiation source terms, the radiological consequences were recalculated as follows: Dose (rem) Updated Current USAR A Exclusion Area Boundary (0 2 hr) Thyroid 2 49E.1 8.85E-2 Whole Bod / 2.50E-2 2.91 E-3 Low Population Zone (Duration) Thyroid 3.32E 2 1.18E 2 Whole body 3.33E 3 3.88E-4 i i f 5-265 L
Steam Generator Tube Rupture (SGTR) 1 Mass and energy balance caiculations are re-performed at the uprated power level i using the approved WCGS RETRAN models for the limiting SGTR event to determine primary to secondary mass release and to determine the amount of liquid / vapor mixture released from each of the steam generators from the occurrence of the tube rupture until RHR cut in conditions are reached. Utilizing these re-calculated steam release and the revised source terms , the activity releases and the radiological consequences resulting from a SGTR accident were recalculated for a pre-existing iodine spike and a consequential lodine spike. The results of the calculations are presented as follows; 4 A.- Activity released to the environment (Cl) Case 1. Pre-Accident lodine Spike Isotope 0-2 hour 0 8 hour 1-131 2.4645E+2 2.6902E+2 1132 1.8797E+2 1.9747E+2 l l-133 3.7106E+2 4.0307E+2 1-134 2.0760E+1 2.7295E+1 1-135 2.0153E+2 2.1658E+2 i Xe 131m 1.2525E+2 1.2769E+2 Xe-133m 1.0209E43 1.0408E+3 . Xe-133 1.5316E+4 1.5614E+4 Xe-135m 3.4126E+1 3,4771 E +1 Xe-135- 5.1152E+2 5.2158E+2 Xe-137 1.2545E+1 1.2770E+1 Xe-138 4,1510E+1 4.2291 E+1 Kr-83m 2.8150E+1 2.8715E+1 Kr-85m 1.1189E+2 1.1411 E+2 Kr-85 4.4063E+2 4.4919E+2 Kr-87 7.0699E+1 7.2117E+1 Kr-88 2.0104E+2' 2.0505E+2 Kr-89 6.2868E+0 6.3991 E+0 l l l . 5-266' l
Case 2. Concurrent lodine Spike v itope 0 2 hour 0-8 hour 1-131 7.5651 E+1 8.4804E+1 1-132 2.428 E+2 2.5993E+2 1-133 1.4796E+2 1.6485E+2 1134 8.6052E+1 8.8833E+1 1-135 1.2621 E+2 1.3881 E +2 Xe 131m 1.252SE+2 1.2769E+2 Xe-133m 1.0209E+3 1.0408E+3 Xe-133 1.5316E+4 1.5614 E+4 Xe-135m 3.4126E+1 3.4771 E+ 1 Xe-135 5.1152E+2 5.215BE+1 Xe-137 1.2545E+1 1.2770E+1 Xo-138 4.1510E+1 4.2291 E+1 Kr-83m 2.8150E+1 2.8715E+1 Kr-85m 1.1189E . 2 1.1411 E+2 Kr SS 4.4063E+2 4.4919E+2 Kr-87 7.0699E+1 7.2117E+1 Kr-88 2.0104E+2 2.0505E+2 Kr-89 6.2868E+0 6.3991 E4 0 5-267
B. Radiological Consequences ! Dose (rem) Updated Current USAR l Case 1. Pre-Accident lodine Spike l Exclusion Area Boundary (0-2 hr) Thyroid _ 25.0 23.0 Whole Body 0.088 0.042 -; Low Population Zone (Duration) , Thyroid 3.6 3.1 Whole body 0.012 0.006 P Case 2. Concurrent lodine Spike Exclut..on Area Boundary (0-2 hr) Thyroid 8.5 5.3 Whole Body 0.085 0.034 Low Population Zone (Duration) ! Thyroid 1.27 0.81 Whole body 0.012 . 0.005 r 1
~ l 5-268- -
t-- *+ -- -.c., 4 yq..e .p.e* m r_g-.e-g g 4, g - .,_my. -,w9
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i Loss-Of-Coolant Accidents Resultino From A Spectrum Of Postulated Pio'no Breaks. Within The Reactor Coolant Pressure Boundary i The current licensing basis calculations of offsite and control room doses from a hypothetical design bass LOCA include contributions from containment leakage and post-LOCA leakege from ESF systems outside coatainment. Utilizing the updated source terms as rhown in Table 1, in conjunct;on with the other assumptions listed in - the current USAR Table 15:6-6 ar.i 15A-1, the radiological consequences were recalculated. The recalculated doses are summarized as follows: . Dose (rem) Updated Current USAR Exclusion area boundary (0-2 hr.) ;
- a. Containment leakage (0 2 hr)
Thyroid 69 65 Whole Body 1.4 2.2
- b. ECCS recirci. leakage (0.47-2 hr.)
Thyroid 21.6 20 Whole Body 0.062 0.061
. Low-population zone (0-30 days)
- a. Containment leakage (0-30 days)
Thyroid 45 42 Whole Body 0.50 D.78
- b. ECCS recire. leakage (0.47 hr.-30 days)
Thyroid 48 45 Whole Body 0.047 0.045 i a 5-269
Control Room (0-30 days) Dose (rem) Updated Current USAR
- a. Containment leakage Thyroid 20.49 19.27 Whole Body 0.24 0.41 Beta Skin 4 28 7.00
- b. ECCS recire. leakage Thyroid 2.34 1.24 Whole Body 1.29E-4 6.87 E-5 Beta Skin 1.05E-3 5.64E-4 L
l
- l. 5-270 I V p w' -y up
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~ _ _ . . _ . _ . . _ _ _ _ _ _ . . _ . _ _ . _ _ . . _ _ _ . . _ . _ . _ _ . _ _ . _ _ - . _ _ _ . _ _ , . _ _ .
Radioactive Waste Gas Decav Tank Failure The release volume for this accident remains the same as that originally contained in l the USAR. Utilizing the new source terms, the radiological consequences were recalculated as follows: Dose (rem) Updated Current USAR ! l Exclusion Area Boundary (0-2 hr) l Thyroid 1.77E-2 6.7E 2 Whole Body 3.11 E-2 2.5E 2 Low Population Zone (Duration) Thyroid 2.36E-3 8.9E-3 Wnole body 4.14 E-3 3.3 E-3 t ( l T , 5-271 l. i t
,,4- -, , - . . . . , , . _ _ _ _ __ . _ , _ . . _ _ . . . _ . _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ , . _ _ _
Radiolocsca! Liould Waste System Leak O cailure The release volume for this accident rem un. the same as that originally contained in [ ;T the Ur;/e. Utilizing the new source tymt. .ne radiological consequences were ..
- recalcunited as follows:
5([. 3 ><- Dose (rern) Updated Current USAR Boron Recycle Holdup Tank: Exclusion Area Boundary (0-2 hr) Thyroid 1.79E-2 3.2 E-2 Whole Dody 5.78E-3 3 BE-3 ,, Low-Population Zor 9uration) - Thyroid . 39E-3 4.3E4 Whole body 7.71 E-4 5.1 u P,, Evaporator Bot; oms Tank: P.xclus on Atea Boundary (0-2 hr) Thyrcid 1.24E-1 2.1 E-1 Wnole Body 1.34 E-4 4.8E-5 7 Low-Populatien Zone (Duration) Thyroid 1.66E-2 2.7 E-2 r-Whale body 1.78E-5 6.4 E-6 5-272
, -- .~ . . .- - - . - . . - . . - . _ . .
Fuol Handlina Accidents
- The maximum assembly activity was recalculated for extended fuel cycle at 100 hours after shutdown at a wre power level of 3565 MWt. '!hese activity levels were adjusted by a radial peaking factor of 1.65 in accordance with the guidance of Regulatory Guide 1.25. Utilizing these values, the radiological consequences were recalcuiated as follows:
Doce (rem) Updated Current USAR in Fuel Building Exclusion Area Boundary (0-2 hr) Thyroid 8.9E+0 4.2E+0 Whole Body 1.5E-1 1.4 E-1 Low-Population Zone (Duration) Thyroid 1.2E+0 5.GE-1 Whole body 2.0E-2 1.9E-2 l In Reactor Building Exclusion Area Boundary (0-2 hr) Thyroid 1,62E+1 1.54E+1 Whole Body 5.5E-2 5 4E-2 L. Low-Population Zone (Duration) Thyroid 2.2E+9 2.0E+0 Whole body 7.3 E-3 7.0E-3 l-i. L 5-273
lt should be noted that the significant increase of the thyroid do;,es resulting from a fuel handling accident in the Fuel Building is due to the single failure assumptions associated with loss of humidity control for the Emergency Exhaust Filter-Adsorber units. Previously, loss of humidity control was not postulated beca:pe other failures (or no single failure) were assurs.;u to bound the failure of the humiday ccntrol system. A single failure of the humidity c.cntrol system would go undetected by the control rooni operator as there is no control roem annuncia or indication of a loss of charcoal adsorber humidity control or loss of the dehumiu.. er heaters. If the heater failure occurred concurrently with high ambient relative humidity a reduced efficiency in th6 removal of organic iodine would occur and doses resulting from exposure to the - processed air stream could increase. Therefore, the postuialec fuel handling accident in the Fuel Building was re-analyzed with a reduced filter eff!ciency based upon the single failure assumption that one of the
--~
Emergency Exhaust Filter-Adsorber units is operating with a failed heater or numidistet. u 5-274
5.8 - References
- 1. WCGS USAR Chapter 15
- 2. ANSI-N10.2. " Nuclear Safety Criteria for the Desion of Stationary PWR Plants,"
Section 5,1973.
- 3. "T:?nsient Anal / sis Methodology for the Wolf Creek Generating Station", NSAG-006 Rev. O, W. D. Wagner, ot.al., January 1991.
- 4. " Core Thermal-Hydraulic Analysis Methodology for the Wolf Creek Generating Station", TR-90-0025 WO1, W. S. Kennamore, et. al., July 1990.
- 5. Reload Safety Evaluation Methodology for the Wolf Creek Generating Station",
NSAG-007 Rev.0, W. S. Kennamore, et.al., January 1992.
- 6. " Qualification of Steady State Core Physics Methodology for Wolf Creek Design and Analysis", TR-91-0018 WO1, E. W. Jackson, et. al., December 1991.
- 7. Letter SLNRC 86-1, Petrick, N. A., SNUPPS to Denton, H.R. (NRC), " Steam Generator Single-Tube Rupture Analysis for SNUPPS Plants- Callaway and Wolf Creek", dated January 8,1986
- 8. Letter WM 87-0145, Withers, B. D., WCNOC to US NRC, " Response to RAI Regarding the SGTR Overfill Case Analysis", dated May 15,1987.
- 9. Letter from NRC to Withers, B. D., " Safety Evaluation Report For The Wolf Creek Gener4thg Station Steam Generator Tube Ruptt e Analysis", TAC No.-
573d3, dated May 7,19S1. I 1 l l 5-275
6,0 LOCA cnd LOCA Related Evaluations-
-- A loss-of-coolant accident (LOCA) is the result of a pipe rupture of the reactor coolant system (RCS) pressure boundary. For the analyses reported here, a small br_eak is defined as a rupture of the RCS piping with a cross-sectional area less than 1.0 ft2, in which the normally operating charging system flow is not sufficient to sustain pressurizer level and pressure. This event is considered an American Nuclear Society (ANS) Condition 111 event which are faults which may occur very infrequently during the life of a plant. A major break (large break) is defined as a rupture with a total cross-sectional area equal to or greater than - 1.0 ft2 . This event is considered an ANS Condition IV event, a lirriting fault, in that it is not expected to occur during the life of the Wolf Creek Generating Plant, but is postulated as a conservative design basis.
The Acceptance Criteria for the LOCA are described in 10 CFR 50A6 (Reference 1) as follows: A. The calculated peak fuel element clad temperature shall not exceed 2200 F. B. The amount of fuel element cladding that reacts chemically with water or steam to generate hydrogen, shall not exceed 1 percent of the total amaunt of Zircaloy_ in the fuel rod cladding. C. The clad temperature transient is terminated at a time when the core geometry is still amenable ti aling. The localized cladding oxidation limit of 17 percent is not er leded dunng or after quenching. D. The core remains amenable to cooling during and after tho break. E. The core temperature is ceduced and decay heat is removed for an extended period of time, as required by the long-lived radioactivity remaining in the core. The criteria were established to provlde a pgnificant r iargin in emergency core cooling system (ECCS) performance Tollowing a LOCA. WASH-1400 (USNRC 1975) (Reference 2) presents a study in regards to the probability o' occurrence of RCS pipe ruptures. 6-1
61- SmcIl Brcck LOCA
< The purpose of analyzing tbs small break LOCA is to demonstrate the feasibility.of maintaining conformance with the 10 CFR 50.46 requirements lis'ed in Section 6.0 for the implementation of V5H with IFMs. The analysis assumptions also supports an uprating of reactor core power to 3565 MWt and a Thot reduction.
6.1.1- Assumptions and Analysis Model The small break analysis was based upon the input assumptions and conditions listed in Table 6.1-1. For small breaks (less than 1.0 ft2) the NOTRUMP digital computar code (Referenced 3 and
- 4) is employed to calculate the transient deprossurization of the reactor coolant system (RCS) as well as to describe the mass and energy of the fluid flow through the break. The NOTRUMP computer code is a state-of-the-art one-dimensional general network code incorporating a number of advanced features. Among theses are calculation of thermal non-equilibrium in all fluid volumes, .'iow regime-dependent drift flux calculations with counter-current flooding limitatior.s, mixture level tracking logic in multiple 1 tacked fluid nodes and regime-dependent drift flux calculations with multiple-stacked fluid nodes and regime-dependent heat transfer correlations. The NOTRUMP 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 NRC concerns expressed in NUREG-0611
-(Reference 5).
The RCS model is nodalized into volumes interconnected by flowpaths. The broken loop is modeled explicitly, while the intact loops ar3 lumped into a second loop. Transient behavior of the system is determined from the governing conservation equations of mass, energy, and momentum. The multi-node capability of the program enables explicit, detailed spatial
- representation of various system compcnents which, amorig other capabilities, enables a proper calculation of the behavior of the loop seal curing a LOCA. Tne reactor core is represented as heated control vaumas with associated phase separation models to permit transient mixture height caiculations.
Clad thermal analyses are performed with a version of the LOCTA-IV code (Reference 6) using the NOTRUMP calculated core pressure, fuel rod power history, uncovered core steam flow and mixture heights as boundary conditions (Figure 6.1-1). L 6-2
= - - .- .-
1 1
- 6.1.2 Analysis j
. For small break LOCAs, the most limiting single active failure is that which results in the minimum ECCS flow delivered to the RCS. This has been determined to be tha loss of offsite power and failuie of a diesel generator resulting in the loss of an emergency power train which results in the loss of one complete train of ECCS components. This means that credit can be taken for only one high head charging pump, one safety njection pump, and one RHR (low head) pump. During the small break transient, one ECCS train is assumed to-start and deliver flow through the injection lines (one for each loop) with one branch injection line spilling to the RCS backpressure, except for the charging pump injection lines, which has one injection line spilling to containment backpressure. To minimize delivery to the core, the branch line with the least resistance is chosen to spill to the RCS backpressure. Shouw a small break LOCA occur, depressurization of the RCS causes fluid to flow into the loops from the pressurizer resulting in a pressure and level decrease in the pressurizer. The - reactor trip signal subsequently cccurs when the pressurizer low-pressure trip setpoint (1605 , psig)is reachec. Loss-of-Offsite Power (LOOP) is assumed to occur coincident with reactor trip. A safety injection signal is generated when the pressurizer low-low-pressure setpoint (1700 psig)is reached. These counter measures limit the consequences of the accident in two ways: A. Reactor trip and borated water injection supplement void formation in causing rapid reduction of nuclear power to a residual level corresponding to the delayed fission and fission product decay. No credit is taken in the LOCA analysis for the boron content of the injection water. Mowever, an average RCS/ sump mixed boron concentration is calculated to ensure that the post-LOCA cora emains subcri*ical. In addition, in the small break LOCA analysis, credit is tal .n for the insertion of Rod Cluster Control Assemblies (RCCAs) subsequent to the reactor trip signal, while assuming the most reactive RCCA is stuck in the full out position. B. Injection of borated water ensures sufficient flooding of the core to prevent excessive clad temperatures. Before the break occurs the plant is in an equilibrium condition, i.e., the heat generated in the core is being removed via the secondary system. After the small break LOCA is initiated, reactor trip occurs due to a low pressurizer signal (1805 psig). During the earlicr part of the small brcak transient, the effect of the break flow is nct strong enough to overcome the flow 6-3 i
maintained by the reactor coolant pumps through the core as the pumps coast down . following LOOP. Upward flow through the core is maintained. However, flow through the core is insufficient to prevent a partial uncovery. Subsequently, ECCS injection is required ' to recover the core. During blowdcwn, heat from fission produ:t decay, hot interne's and the vessel continues to be transferred to the RCS. The heat transfer between the RCS and the secondary system may be in either direction depending on the retrtive temperatures. In this case continued " heat addition to the secondary results in increased secondary system pressure which leads to steam relief via the main steam safety valves. Makeup to the secondary is automatically provided by the auxiliary feedwater pumps. The safety injection signal isolates normal feedwater flow by closing the main feedwater control and bypass valves. Loss-of-Offsite Power, assumed concurrent with reactor trip, initiates aux;;iary feedwater flow by starting the auxiliary feedwater pumps. The secondary flow aids in the reduction of RCS pres'sure. Also, consistent with the LOOP assumption, the reactor coolant pumps are assumed to be tripped at the time of reactor trip and effects of pump coastdown are included in the blov/down analysis. Safety injection systems consist of gas pressurized accumulator tanks and pumped injection systems. The small break LOCA analysis, assumed nominal accumulator water volume (850 ft3)with a cover gas pressure of 600 psia. Minimum ECCS availability is assumed for the analysis at the maximum RWST temperature. Assumed pumped safety injection characteristics as a function of RCS pressure are provided in Figuie 6.1 14. In order to determine the conditions that produced the most limiting case SBLOCA several sensitivity calculations were performed. These cases included the investigation of variables including RCS temperature, AFW flowrate, time in life fuel conditions, fuel rod internal pressure, and break size to ensure that the most severe postulated SBLOCA was evaluated. 6.1.3 Results The evaluations to determine the limiting reactor coolant system average temperature and the limiting auxiliary feedwater flowrate were completed before the break spectrum was performed. The evaluations were completed assuming a 4 inch equivalent diametcr break in the cold leg. This break size was selected since Wolf Creek is currently 4-inch limited-(USAR 15.6.5.3.3).' The limiting 4-inch break ccaditiors were at high RCS temperature and minimum AFW which resulted in a PCT of 1398"F. The 3-inch and 2-inch breaks which were 6-4
1 cnalyzcd at th3 limiting RCS tcmpsrature and limiting AFW flowrats with PCTs of 1510*F and 1434 F, respectively, which showed that the 3-inch break was limiting. The burnup (limiting time in life) study was completed for the limiting 3 inch break and it was determined that beginning of life fue; conditions were limiting. Results of the analyses are summarized in Tables 6.1-2 and 6.1-3. The follewing figures are included for the break spectrum at the limiting RCS temperature and AFW flowrate. The clad temperature transient of the hot rod The reactor coolant system depressurization trar: pant The core mixture level For the limiting break size the following figures are presented The core steam flow The rod film heat transfer coefficient The hot spot fluid temperature The peak clad temperature calculated for the limiting small break LOCA is 1510 F, which is less than the acceptance criteria limit of 2200 F. The maximum local metal-water reaction is 0.94 percent, which is well below the embrittlement limit of 17 percent as required by 10 CFR 50.46. The total core metal water reaction criterion of 10 CFR 50.46 of 1.0 percent was n.et for all breaks analyzed. The clad temperature transient is terminated at a time when the core geometry remains amenable to core cooling. As a result, the core temperature will continue te drop and the ability to remove decay heat generated in the fuel for an extended period of time will be provioed. l l l l i l 6-5
TABLE 6.1-1 INPUT PARAMETERS USED IN THE ECCS ANALYSES Parameter High Tave Low Tave Rer, tor core rated thermal power 1, 3565 3565 GviWt) ' Peak linear powei, (kw/fi ) 13.872 13.872 Total peaking factor (Fqt) at peak 2.39 2.39 Power shape See Figure 6.1-15 See Figure 6.1-15 FAH 1.65 1.65 Fuel assemble array 17 x 17 17 X 17 VANTAGE SH VANTAGE SH Accumulator water volume, nominal 850 850 (ft3/ accumulator) Accumulator tank volume, nominal 1350 1350 (ft3/ accumulator) Accumulator gas pressure, minimum 600 600 (psia) Pumped u aty injection flow Figure 6.1-14 6.1-14 Steam generator initial pretsure, (psia) 944.5 802.29 Steam generator tube pluggirig level (%) 10 10 Initial ficw in each loop, (Ib/sec) 9656.95 9901.1 Vessel inlet ternperature, (*F) 556.04 537.43 Vessel outlei temperature, ( F) 620.76 603.97 Reactor coolant pressure, (psia) 2300 2300 Maximum auxiliary feedwater flowrate, 34.57 34.57 (Ib/sec) Minimum auxiliary feedwater flowrate, 29.04 29.04 (Ib/sec)
\
1 Two percent is added to this power to account for calorimetric error. Reactor coolant pump heat is not modeled in the SBLOCA analyses. 6-6 I _--_ - - - - - -- - - - - - - - - - - - - - - - - - - - _ - - - - - _ - - - - _ _ - a
TASLE 6.12 SMALL BREAK LOCA ANALYSIS TIME SEQUENCE OF EVENTS Break Spectrum, Limiting Conditions (High Tave, Min. AFW) 2 inch 3 inch 4 inch Break Occurs (sec) 0.0 0.0 0.0 Reactor Trip Signal (sec) 84.07 28.48 15.9 Safety injection Signal (sec) 94.54 37.98 24.52 Safety injection Begins (sec) 133.E 4 76.98 63.F? Loop Seal Venting (sec) 1030 505 270 Top Of Core Uncovered (sec) 2074 936 623 Accumulator injection Begins (sec) N/A 3331 894 Peak Clad Temperature Occurs (sec) 3610 1446 948 Top Of Core Covered (sec) 5726 2887 1313 6-7
TABLE 6.103 SMALL BREAK LOCA ANALYSIS FUEL CLADD!NG RESULTS Break Spectrum, Limiting Conditior:s (High Tave, Min. AFW) 2 inch 3 inch 4 inch Peak Clad Temperature ( F) 1434 1510 1398 Peak Clad Tempermure Location (ft) 11.75 11.75 11.5 Peak Clad Temperaute Time (soc) 3610 1446 948 Local Zr/H2O Reaction, Maximum (%) 0.494 0.939 0.193 Local 7r/H2O Reaction Loca*. ion (ft) 11.75 11.75 11.5 Total Zr/H2O Reaction (%) < 1.0 < 1.0 < 1.0 Hot Rod Burst Time (sec) No Burst No Burst - No Burst Hot Rod Burst Lccation (ft) N/A N/A N/A ^ 3 i N L 0 0 T C s R T U cent emissunt. cent A Flow, MIXTURE t,tVEL, H AND FutL rod P powls si::'ToRv 0 < TIME < CORE COVEAf D , s FIGURE 6.1-1 ww -. o c-CODE 8NTERFACE DESCRIPT10N FOR SN1ALL BREAK AIODEL 6-9 l 1
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_ 6.2 Lcrge Brock LOCA The purpose of analyzing the large break LOCA is to demonstrate conformr.nce with the 10 CFR 50.46 requirements listed in Section 6.2 for a fuel upgrade from STD to V5H with IFMs. The analysis assumptions also support an uprating cf reactor core power to 3565 MWt and a Thot reduction. 6.2.1 Assumptions and Analysis Model The large break analysis was based upJn the input assumptions and conditions listed in Table 6.2-1. 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 witnin the containment, and the fuel and clad temperature transient of the Lottest fuel rod in the core. Based on these considerations, a system of interrelated computer codes has been developed for the analysis of the LOCA. A description of the various aspects of the LOCA analysis methodology 's given by Bordelon, Massie, and Zordan (1974) (Reference 7). This document describes the major phenomena modeled, the iriterfaces among '.he computer codes, and the features of the codes which ensure compliance with the Acceptance Criteria. The SATAN-VI, WREFLOOD, BASH and LOCBART codes, which are used in the LOCA analysis, are described in detail by Bordelon et al. (1974) (Reference 8); Kelly et al. (1974) (Reference 9); Young et al. (1987) (Reference 10); and Bordelon et al. (1974)(Reference 7). Code modifications are specified in References 11,12,13 and 14. These codes assess the core heat transfer geometry and determine if the core remains amenabli to cooling through and subsequent to the blowdown, refill, and reflood phases of the LOCA. SATAN-VI calculates the thermal-hydraulic transient, including 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 systems as a function of time during the blowdown phase of the LOCA. SATAW-VI also calculates the accumulator water mass and ir ternal pressure and also the break mass and energy flow rates that are assumed to be ver.!ed to the containment during blowdown. At the end of the blowdown, the mass and energy release rates during blowdown are transferred to the COCO code, detailed in 6-23
Rcftrcnca 15, for uso in dstcrminction of th3 contrinment pressuro responso during tho first phase of the LOCA. Additional SATAN-VI output data from the end-of-blowdown, including the core inlet flow rate and enthalpy, the core pressure, and the core power decay transient, are input to the LOCBART code. At the end of the blowdown, information from SATAN-VI on the state of the system is t'ransferred to the WREFLOOD code which calculates the time to bottom of core (BOC) recovery, RCS conditions at BOC and mass and energy relesse 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 backpre.sure, the WREFLOOD and COCO codes are interactively linked. The BOC conditions calculated by WREFLOOD and the containment pressure transient calcuiated by COCO are used as input to the BASH cc,de. Data from both SATAN-VI code and WREFLOOD code out to BOC are input to the LOCBART code which calculates core average conditions at BOC for use by the BASH code. The BASH code provides a realistic tharmal-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 6.2-1 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 detailed description of the BASH code is available in Reference 10. The BASH code provides a sophisticated treatment of steam / water flow phenomena in the reactor coolant system during core reflood. A dynamic 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. The loop model determines the loop flows and pressure drops in response to the calculated core exit flow determined by BART. The updated core inlet flow calculated by the loop model is used by BART to calculate a new entrainment rate to be fed into 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
.oupling of the BART code with a loop code produces a dynamic flooding transient, which reflects the close coupling between core thermal-hydraulics and loop be." 'ior.
The cladding heat-up transient is calculated by LOCBART which is a combination of the LOCTA code with BART, A more detailed description of the LOCBART code can be found in References 10 and 4. During reflood, the LOCBART code provides a significant improvement in the prediction of fuel rod behavior. In LOCBART the empirical FLECHT 6-24
corrclition his be:n rcpl:ced by tho BART coda, BART cmploys rigorous m:chtnistic models to generate heat transfer coefficients appropriate to the actual flow and heat transfer regimes experienced by the fuel rods. 6.2.2 Analysis A LOCA is the result of a pipe rupture of the RCS pressure boundary For the reduced THOT analysis reported here, a major pipe break (large break) is defined as a rupture with a total cross sectional area equal to or greater than 1.0 ft2 . This event is considered an American Nuclear Society (ANS) Condition IV event, a limiting fault, in that it is not expected to occur during the lifetime of the Wolf Creek Generating Station, but is postulated as a conservative design basis. For large break LOCAs, the most limiting single failure is the one which produces the lowest containment pressure. The lowest containment pressure would be obtained only if all containment spray pumps and fan coolers operated subsequent to the postulated LOCA. Therefore, for the purposes of large break LOCA analyses, the most limiting single failure would only be the loss of one RHR pump with full operation of the spray pumps and fan coolers. However, the large break LOCA analysis conservatively assumes both maximum containment safeguards (lowest containment pressure) and minimum ECCS safeguards (the loss of one complete train of ECCS components which includes one RHR pump, one safety injection pump, and one high head charging pump) which results in the minimum delivered ECCS flow available to the RCS Modeling full containment heat removal systems operation is required by Branch Technical Position CSB 6-1 and is conservative in large break LOCA. A normal FSAR large break LOCA analysis gives consideration to a full spectrum of break discharge coefficients with the above assumption. The reduced TAVG analysis was performed based on break discharge coefficient which produced the most limiting results (Co=0.6) as determined in the current licensing basis (Reference 16). Loss-of-Offsite Power (LOOP) was assumed coincident with the occurrence of the break. The reactor trip signal subsequently occurred when the pressurizer low pressure trip setpoint was reached. A safety injection signal was generated when the appropriate setpoint (high containment pressure or low pressurizer pressure) was reached. These countermeasures limit the consequences of the accident in two ways: 6-25 l 1
A. Re:ctor trip cnd boretad watcr injcetion supplemcnt vcid formation in causing rapid reduction of power to the residual level corresponding to fission product decay heat. No credit is taken in the LOCA analysis for the boron content of the injection water. However, an average RCS/ sump mixed boron concentration is calculated to ensure that the post LOCA core remains subcritical. 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. To minimize delivery to the reactor, the branch line chosen to spill is selected as the one with the minimum resistance. In addition, the high head charging pump, the safety injection pump, and the RHR pump performance curves were each degraced by 10%. Minimum safety injection performance is presented in Figure 6.2-2a and the maximum safeguards performance is presented in Figurs 6.2-2b. The time sequence of events following a large break LOCA is presented in Table 6.2-2. Before the break occurs, the unit is in an equilibrium condition; i.e., the heat generated in the core is being removed via the secondary system. During blowdown, heat from fission product decay, hot internals, and the vessel continues to transfer to the reactor coolant. At the beginning of the blowdown phase, the entire RCS cordaim subcooled liquid which transfers heat from the core by forced convection with sot
- developed nucleate bG. ling After the break develops, the time to departure from nucleate boiling is calculated, consistent with Appendix K of 10 CFR 50 (Reference 17). Thereafter, the core heat transfer is unstable, with both nucleate boiling and film boiling occurring. As the core becomes uncovered, both transition boiling and forced convection are considered as the dominant core heat transfer mechanisms. Heat transfer due to radiation is also considered.
Turbine trip on a reactor trip signal is assumed in the analysis but is not explicitly modeled. The steam generators are isolated at reactor trip and become a heat source early in the transient due to the rapid energy loss of the primary. Further, the delays associated with the AFW system prevent the secondary from having a significant effect on the transient. As such, no credit for the secondary system engineered safety features (i.e., main steam safety valves, AFW, etc.) is taken for the large break (.OCA. 6-26
Vhen the RCS d;pressuriz2s to 600 psis, the cccumulators b: gin to injact borat d water into the reactor coolant loops. The conservative assumption is made that all of the accumulator water injected during the bypass period is subtracted from the RCS after the bypass period terminates (called end-of-bypass). End-of-bypass occurs when the expulsion or entrainment mechanisms responsible for the bypassing are calculated not to be effective. This conservatism is again consistent with Appendix K of 10 CFR 50. Since LOOP is l assumed, the reactor coolant pumps are assumed to trip at the inception of the accident. The effects of pump coastdown are included in the blowdN analysis. The blowdown phase of the transient ends when the RCS pressure (initially assumed at 2300 psia) falls to a value approaching that of the containmmt atmosphere. Prior to or at the end of the blowdown, termination of bypass occurs ano *' of the reactor vessel lower plenum begins. Refill is completed when emergency core cooling water has filled the lower plenum of the reactor vessel, which is bounded by the bottom of the fuel rods (i.e., BOC recovery time). The reflood phase of the transient is defined as the time period lasting from BOC recovery until the reactor vessel has been filled with water to the extent that the core temperature rise has been terminated. From the latte'r stage of blowdown and then the beginning of reflood, the safety injection accumulator tanks rapidly discharge borated cooling water into the RCS, thus contributing to the filling of the reactor vessel downcomer. The downcomer head provides the driving force required for the reflooding of the reactor core. Tne RHR (low head), safety injection, and high head charging pumps aid in the filling of the downcomer and subsequently supply water to maintain 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 associated with dissipation of residual heat generation. After the water level of the refucling water storage tank (RWST) 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. Spilled borated water is drawn from the engineered safety features (ESF) containment sumps by the RHR (Iow head) pumps and retumed to the RCS cold legs. The containment spray pumps are manually aligned to the containment emergency sumps and continue to operate to further reduce containment pressure and temperature. Approximately 10 hours 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. 6-27
6.2.3 Results Based on the results of the LOCA sensitivity studies (Westinghouse 1974 (Reference 18); Salvatori 1974 (Reference 19); Johnson, Massie, and Thompson 1975 (Reference 20)), the limiting large break was found to be the double-ended guillotine (DECLG). Therefore, only the DECLG break was considered in the large break ECCS performance analysis. A limiting PCT of 1916*F was calculated for V5H fuel for the CD=0.4 case at reduced TAVG with minimum safeguards assumptions. Until a full core of V5H is achieved a 50'F transition core penalty is applied for a resultant PCT of 1966'F. This i', tess than the acceptanco criteria limit of 2200 F. The maximum local metal-water reaction is 3.64%, which is well below the embrittlement limit of 17 percent as required by 10 CFR 50.46. The total core metal-water reaction is less than 1.0 percent for all break.s analyzed, corresponding to less than 1.0 percent hydrogen generation, as compared with the 1.0 percent criterion of 10 CFR 50.46. The clad temperature transient is terminated at a time when the core geometry is still amenable to core 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. Figures 6.2-3 through 6.2-31 present the results for the complete break spectrum assuming reduced TAVG with minimum safeguards. A list of the parametrics given for each break size is given below.
-core pressure
-power (normalized to the nominal power) l
-core and downcomer collapsed liquid water levels during reflood
-core inlet flow velocity during reflood
-clad temperature at the hot spot on the hot rod The following additional plots are provided for the limiting break size:
-break flow rate (sum of both ends of the guillotine break)
-core flow rate during blowdown
-accumulator flow rate (intact loop) during blowdown I -combined accumulator and pumped safety injection flow rate
-break energy released to containment
-fluid quality
-fluid velocity at the hot spot l -heat transfer coefficient at the hot spot on the hot rod
-fluid temperature at the hot spot on the hot rod 6-28
- ~ , - - ,
-conteinmtnt beckprussuro ,
Additional sensitivity analyses were performed for RCS temperature, maximum safeguards ! and*once-burned STD fuel. The results of these calculations determined that reduced RCS l Temperature, minimum safeguards and V5H fuel are limiting. l Conclusions The WCGS large break LOCA has been analyzed to support operation of V5H fuel with IFMs at a uprated core power of 3565 MWt. The analysis included calculations for a range of RCS TAVG between 570.7'E and 588.4'F with 10% steam generator tube plugging. In addition sensitivity studios were performed assuming once burned standard fuel and maximir.n safeguards. The limiting PCT of 1916*F was calculated for the break with Co=0 4 at tha reduced TAVG conditions assuming minimum safeguards. Until a full core of V5H is echieved, a 50*F transition core penalty will be applied to the limiting large break PCT for a resultant PCT of 1966 F. The acceptance criteria for LOCA, as described in 10 CFR 50.46, has been met. l t i 6-29
. . - . _ . - _ . _ . . ~ _ - . . _ - __ __
.. . ym
. TABLE 6.21 LARGE BREAK LOCA ECCS ANALYSIS INPUT PARAMETERS License Core Poweri (MWth) 3565 Peak Linear Poweri (KW/ft) 14.225 Total Peaking Factor [FoT) 2.50 Axial Peaking Factor [FZ) 1.5'51 Hot Channel Enthalpy Rise Factor [FAH) 1.65 Maximum Assembly Average Power [P-BARHA) 1.469
- Power Shape Chopped Cosine Fuel Assembly Array 17X17 Accumulator Water Volume (ft 3/ accumulator) 850 Accumulator Tank Volume (ft 3/ accumulator) 13C0 Accumulator Gas Pressure, Minimum (psia) 600 Safety injection Pumped Flow Figures 6.2-2a,-2b Containment Parameters Tables 6.2-4,-5 Initial Loop Flow (gpm/ loop) 93200-Vessel Average Temperature (F)
. Normal Operation 588.4 Reduced Temperature Operation 570.7 Reactor Coolant Pressure 2 (psia)- 2300 Steam Pressure (psia) - r Normal Operation 944.5 Reduced Temperature Operation 802.3 Steam Generator Tube Plugging Level (%) 10 f
1 - Two percent is added to this power to account for calorimetric uncertainty.
= 2 This pressure includes 50 psi for measurement uncertainty.
6-30
. _ . _ _ _ . _ _ . . _ . _ - . _ . . _ _ _ . _ - _ _ - . _ . . - _ _ _ . . _ _ __ ~._
TABLE 6.2 2 (SHEET.1 OF 2) TIME SEQUENCE OF EVENTS FOR-LARGE BREAK LOCA ANALYSIS
. Accident' Event Time (sec)
Large Break LOCA Case
Description:
(TAVG=570.7'F VANTAGE 5H Fuel w/lFMs Minimum Safeguards)
- a. DECLG Co=0.4 Start - 0.0 1
, Reactor Trip Signal 0.718 ! Safety injection Signal 2.210 Accumulator injection Begins 19.756 Pumped Safety injection Begins 41.210 End-of-Bypass 42.430 End-of-Blowdown 42.430 Bottom of Core Recovery 59.363 Accumulator Empty 62.557
- b. DECLG Co=0.6 Sfart 0.0 Reactor Trip Signal _ 0.706 Safety injection Signal 1.800-Accumulator injection Begins 14.503-
-Pumped Safety injection Begins 40.800 End-of-Bypass 32.380 End-of-Blowdown 32.380 ,
Bottom of Core Recovery. 47.659 Accumulator Empty 55.300
- c. DECLG Co=0.8 Start 0.0 Reactor Trip Signal.: 0.698.
Safety injection Signal 1.580 Accumulator Injection Begins 12.005 l Pumped Safety. injection Begins 40.580 _ End-of-Bypass 27.818 End-of-Blowdown 27.818 Bottom of Core Recovery 42.362-Accumulater Empty 51.562-i l 6-31 l L - . -
TABLE 6.2 2
-(SHEET 2 OF 2)
TIME SEQUENCE OF EVENTS FOR LARGE BREAK LOCA ANALYSIS Accident Event Time (sec) - Large Break LOCA . I Case Descripton: (TAVG=570.7'F - VANTAGE SH Fuel w/lFMs Maximum Safeguards)
- d. DECLG Co=0.6 Start 0.0 Reactor Trip Signal 0.706 Safety injection Signal 1.800 Accumulator injection Begins 14.503 Pumped Safety injection Begins 40.800 End-of-Bypass 32.380 End-of-Blowdown 32.380 Bottom of Core Recovery 47.489 Accumulator Empty 55.577 9
l l l 6-32
~
TABLE 6.2-3 LARGE BREAK LOCA ANALYSIS RESULTS FUEL CLADDING DATA (TAVG=570.7*F - VANTAGE SH Fuel w/lFMs) Results DECLG Co=0.4-DECLG Co=0.6 DECLG Co=0 8 DECLG Co=0 6 MAXIMUM SI - Peak Clad Temperature ( F) 1916.0** 1828.4 1627.4 1739.4 Peak Clad Temperature Location (ft) 8.00 8.75 7.00 8.75 Peak Clad Temperature* Time (sec) 161.4 205.5 91.1 182.3 Local Zr/H2O Reaction , Maximum (%) 3.64 2.67 1.28 1.91 Local Zr/H2O Reaction Location (ft) 6.25 8.00 7.00 8.75 Total Zr/H2O Reaction (%) < 1.0 <1.0 < 1.0 <1.0 Hot Rod Burst Tim 1(sec) 71.8 47.8 49.1 47.6 Hot Rod Burst Location (ft) 6.25 6.00 6.00 6.00 Hot Assembly Channel Blockage (%) 39.78 46.24 N/A 45.89
- Total thickness at transient termination. Reaction rate decreasing as core continues to quench.
- A 50 F transition core penalty applies until a full core of V5H is attained. Penalty has NOT been added to this value.
6-33
TABLE 6.2 4 LARGE BREAK LOCA CONTAINMENT PRESSURE RESPONSE ANALYSIS INPUT PARAMETERS Initial Containment Pressure (psia) 14.7 Initial Containment Temperature ( F) 120 Relative Humidity (%) 99 Containment Net Free Volume (ft3 ) 2.7 x 106 Outside Temperature, Minimum ( F) -60.0 RWST Water Temperature, Minimum ( F) 37.0 Qontainment Mini-Purge / Exhaust Valve i Signal Delay (sec) 2.0 Closure Time (sec) 3.0 Active Heat Sink Data Containment Spray System Number of Pumps Operating (-) 2 Total Flow Rate (gpm) 7754 Time to Full Flow [ loss of offsite ;.ower (sec)) 25 Spray Temperature ( F) 37 Containment Fan Coolers Number of Fan Coolers Operating (-) 4 Actuation Time (sec) 35 Essential Service Water Temperature (oF) 33 Passive Heat Sink Data Table 6.2-5 6-33a i
TARI E 6.2 5 (SHEET 1 OF 2) LARGE BREAK LOCA CONTA!NMENT PRESSURE RESPONSE ANALYSIS PASSIVE HEAT SINK DATA 1 Wall Thickness (ft) Material Surface Area (ft2) l
- 1. 0.000167 Inorganic Paint 64,919 j 0.021 Carbon Steel ;
4.0 Concrete j
- 2. 0.000167 Inorganic Paint 34,129 0.021 Carbon Steel 3.0 Concrete
- 3. 0.000167 Inorganic Paint 13,538 1.5 Concrete 0.021 Carbon Steet 10.0 Concrete
- 4. 1.0 Concrete 8,564
- 5. 2.0 Concrete 43,497
- 6. 2.5 Concrete 17,061
- 7. 0.000167 Inorganic Paint 7,821 0.021 Carbon Steel 2.0 Concrete
- 8. 0.021 Stainless Steel 8,708 2.0 Concrete
- 9. 0.0001083 Zinc Coating 8,081 l 0.005 Carbon Steel .
2.0 Concrete
- 10. 0.0001083 Zine Coating 186,183 0.0052 Carbon Steel
- 11. 0.000167 Inorgani: Paint 17,746 0.0052 Carbon Steel l-6-33b I
r TABLE 6.2 S (SHEET 2 OF 2) LARGE BREAK LOCA CONTAINMENT PRESSURE RESPONSE ANALYSIS PASSIVE HEAT SINK DATA Wall Thickness (ft) Material Surface Area (ft2)
- 12. 0.000167 Inorganic Paint 114,205 0.0104 Carbon Steel 13, 0.000167 Inorganic Paint 49,101 0.0208 Carbon Steel
- 14. 0.000167 Inorganic Paint 31,372 0.0417 Carbon Steel 15, 0,000167 Inorganic Paint 5.631 0.0833 Carbon Steel
- 16. 0.000167 Inorganic Paint 8,355 0.1667 Carbon Steel
- 17. 0.000167 Inorganic Paint 817 0.3333 Carbon Steel
- 18. 0.0417 Carbon Steel 19,185
- 19. 0.0052 Stainless Steel 31,731
- 20. 0.0208 Stainless Steel 13,610 6-33c
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200 150 S rn S 2 100 5 E dc , 50 . i
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,= f SAFETY INJECTION FLOWRATE I MAXIMUM SAFEGUARDS - l 6-36
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FIGURE 6.2 22 O we sw c-CORE INLET FLOW VELOCITY REDUCED Tryg DECLG(Cn = 0.61. NflN SAFEGUARDS 6-56
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f
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FIGURE 6.2 30 O ww wcw CORE INLET FLOW VELOCITY REDUCED Tryg DECLB(Cn = 0.0, MAX SAFEGUARDS 6-64
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l
I
'O.3' References,
- 1. " Acceptance Criteria for Emergency Core Cooling Systems for Light Water Cooled Nuclear Power Reactors,"10 CFR 50.46 and Appendix K of 10 CFR 50, Federal Reaister 1974. Volume 39, Number 3.
- 2. U.S. Nuclear Regulatory Commission 1975, " Reactor Safety Study - An Assessment of Accident Risks in U.S. commercial Nuclear Power Plants," WASH-1400, NUREG 75/014.
3 MayN, P.E., "NOTRUMP - A Nodal Transient Small Break and General Network Code,"WCAP 1G079 P A,(Proprietary), Augus!198B.
- 4. Lee, N. et al., ' Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code " WCAP-10054-P-A, (Proprietary), Augu >t 1985. !
- 5. " Generic Evaluation of Feedwater Transients and Small Break Loss-of Coolant Accidents in Westinghouse - Designed Operating Plant," NUREG-0611, January 1980.
- 6. Bordelon, F. M. et al.,"LOCT'A-IV Program: Loss-of Coolant Transient Analysis,"
WCAP-8305, June 1974, WCAP-8301 (Proprietary), June 1974.
- 7. Bordelon, F. M.; Massie, H. W.; and Zordan, T. A. ' Westinghouse ECCS Evaluation Model -Summary," WCAP-8339, July 1974.-
- 8. Bordelon, F. M. et al., " SATAN VI Program: Comprehensive Space Time Dependent Analysis of Loss of-Coolant," WCAP-8302 (Proprietary) and WCAP-8306 (Non-Proprietaryn June 1974.
- 9. Kelly,- R. D. et al., " Calculation Model for Core Reflooding After a Loss-of-Coolant Accident (WREFLOOD Code)," WCAP 8170 (Proprietary) and WCAP-8171 (Non-Proprietary), June 1974.
- 10. Young, M. Y, et al., "The 1981 Version of the Westinghouse ECCS Evaluation-Model Using the BASH Code," WCAP 10266-P A Rev,2 (Proprietary), March 1987,
- 11. - Rahe, E. P. (Westinghouts), letter to J. R. Miller (USNRC), Letter No. NS EPRS-2679, November 1982,
- 12. Rahe, E. P., ' Westinghouse ECCS Evaluation Model,1981 Version, WCAP-9220--
P-A (Proprietary), WCAP-9221-P-A (Non-Proprietary), Revc 1, February 1982 6-66
- 13. Bordelon, F. M., et al., Westinghouse E OS Evaluation Model . Supplementary Information," WCAP-8471 (Proprietary) ted WCAP-8472 (Non-Proprietary), April 1975
- 14. Special Report NS NRC 85-3025(NP),"BART-WREFLOOD Input Revision."
- 15. Bordelon, F. M., and Murphy, E. T., " Containment Pressure Analysis Code (COCO)," WCAP-8327 (Proprittary) and WCAP 8326 (Non Proprietary), June 1974.
- 16. Wolf Creek USAR.
- 17. " Acceptance Criteria for Emergency Core Cooling Systems for Light Water Cooled Nuclear Power Reactors." 10 CFR 50.46 and Appendix K of 10 CFR 50, Federal Reaister 1974. Volume 39, Number 3.
- 18. " Westinghouse ECCS - Evaluation Model Sensitivity Studies," WCAP-8341 (Proprietary) and WCAP-8342 (Non Proprietary), July 1974,
- 19. Salvatori, R., "Westinghcuse ECCS - Plant Sensitivity Studies," WCAP-8040 (Proprietary) and WCAP-8356 (Non Proprietary), July 1974.
- 20. Johnson, W. J.; Massie, H. W.; and Thompson, C. M. " Westinghouse ECCS Four Loop Plant (17X17) Sensitivity Studies," WCAP.9565.p.A (Proprietary) and WCAP-8566-A (Non Proprietary), July 1975.
6-67
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